International Conference on Coal Science and Technology, Oviedo-Spain 2011 (ICCS&T 2011)
INSTITUTO NACIONAL DEL CARBÓN INCAR
Programme for the International Conference on Coal Science and Technology 2011, Oviedo 9-13 October Sunday 9th October 2011 017:00-19:30 Registration and documents 19:30-21:00
Ice-break party
Monday, 10th October, 2011 08:30-09:00
Registration and documents Multiusos Room
09:00-10:00 10:00-11:00 11:00-11:30
Plenary Lecture: R. Kandiyoti
11:30-13:20
Coal combustion
13:20-15:00 15:00-16:30
Coal combustion
Room 10
Room 11 Opening Ceremony
Coffee break Coal characterization and coal Coal pyrolysis and liquefaction: structure thermoplasticity of coal Lunch Coal characterization and coal Coal pyrolysis and liquefaction: structure liquids from coal Coffee break
16:30-17:00 17:00-18:00
CO2 Storage
Coal characterization and coal structure Welcome cocktail
19:30-21:30 Tuesday, 11th October, 2011 Multiusos Room Plenary Lecture: D. Harris 09:00-10:00 10:00-11:00
CO2 Transport and storage
11:00-11:30 11:30-13:20 13:20-15:00
Coal pyrolysis and liquefaction: general
CO2 Capture and storage: oxyfuel
Room 11
Room 12
Coal mineralogy and coal ash: coal Coal pyrolysis and liquefaction: ash liquefaction Coffee break Coal gasification and clean fuels Lunch
Coal pyrolysis and liquefaction: fundamentals
Room 12
Carbons from coal
Coal and the environment
15:00-16:30 16:30-17:00 17:00-18:00
Clean coal technology: Mercury emissions
Coal petrology
Coal pyrolysis and liquefaction: direct liquefaction
Coffee break & Poster Viewing Poster session: coal chemistry, coal petrology, coal mineralogy, coal upgrading, coal pyrolysis, carbons from coal Typical Asturian dinner
19:30-23:00
Wednesday, 12th October, 2011 Multiusos Room Room 11 Room 12 Plenary Lecture: R. Malhotra 09:00-10:00 Clean coal technology: Mercury Coal gasification and clean fuels: 10:00-11:00 Coal pyrolysis emissions mineral matter Coffee break 11:00-11:30 CO2 Capture and storage: Mineral matter and coal ash: Coal pyrolysis and liquefaction: 11:30-13:20 mineral matter coke microstructure chemical looping Lunch 13:20-15:00 CO2 Capture and storage: 15:00-16:30 Coal gasification and clean fuels Coal upgrading gasification Coffee break & Poster Viewing 16:30-17:00 Poster Session: Coal combustion, clean coal technologies, coal gasification, CO2 Capture and storage, 17:00-18:00 coal and the environment Conference dinner
20:00-23:00
Thursday, 13th October, 2011 Multiusos Room Room 11 Plenary Lecture: J. C. Abanades 09:00-10:00 CO2 Capture and storage 10:00-11:00 Coal gasification and clean fuels
12:50-13:30 13:30-15:00
Coal upgrading
Coffee break
11:00-11:30 11:30-12:50
Room 12
CO2 Capture and Storage
Coal gasification and clean fuels Closing ceremony Lunch
Coal pyrolysis and liquefaction: biomass
Monday, 10th October, 2011
08:30-09:00
Registration and documents Opening ceremony
09:00-10:00 10:00-11:00
Plenary Lecture: (Chair: R. Moliner ) A00 Thermal breakdown in middle rank coals R. Kandiyoti Coffee break
11:00-11:30 Session A
Session B
Session C
Session D
Coal combustion Chairs: F. Montagnaro & B. Arias
Coal characterization and coal structure Chairs: E. Suuberg & M. Sciazko
Coal pyrolysis and liquefaction: thermoplasticity of coal Chairs: S. Krzack & C. Barriocanal
Carbons from coal Chairs: G. Gryglewicz & M. Granda
11:30-12:00
B01 Visualizing the macromolecular network structure of a large-scale A01 Ash deposition (50,000 atom) Illinois No. 6 characteristics determined in pilot bituminous coal molecular plant tests burning bituminous representation in 3D and 2D lattice and sub- bituminous coals. views. M. Shimogori , N. Ooyatsu, N. Y. E. Alvarez, J. C. Katson, J.O Takarayama, T. Mine Pou, F. Castro-Marcano, J. P. Mathews
12:00-12:20
B02 Brown coal solubilisation with A02 Investigation of contributions novel ionic liquids. to unburned carbon in a 200MWe A. L Chaffee , C. Patzschke, D. power utility boiler. Russell, D. Kelley, Y. Qi, V. H. Gao , A. Majeski, A. Verheyen, M. Marshall, V. Runstedtler, M. Sybring Ranganathan, D. MacFarlane
C01 1H-NMR Study on the thermoplasticity of coking coaleffects of coal blending and additives. H. Kumagai , N. Okuyama, T. Shishido, K. Sakai, M. Hamaguchi, N. Komatsu
D01 A relationship between the structures of graphitized anthracites and isotropic graphite. M.S. Nyathi, C.E. Burgess Clifford, H.H. Schobert
C02 A systematic study of the effects of pyrolysis conditions on coal devolatilisation. M.A. Kochanek, D.G. Roberts, B. Garten, S. Russig, D.J. Harris
D02 Benzene and toluene adsorption on high surface area activated carbons obtained from an anthracene oil derivative. N.G. Asenjo , P. Álvarez, C. Blanco, R. Santamaría, M. Granda, R. Menéndez
Session A
Session B
Session C
Session D
12:20-12:40
A03 Impact of biomass on char burn-out under air and oxy-fuel conditions. T. S. Farrow, D. Zhao, C. Sun, C.E. Snape
B03 Comparison of structure and reactivity of an Australian algal coal and a Jordanian oil shale. W. R. Jackson , M. W. Amer, Y. Fei, M. Marshall, A.L. Chaffee
C03 The role of sulfur in coals plastic layer formation. L. Butuzova , R. Makovskyi, T. Budinova, S. P. Marinov
D03 High performance electric double-layer capacitor using cctivated carbon from hyper-coal. K. Sato, K. Magarisawa, T. Takarada
12:40-13:00
A04 Numerical study on the reburning of ash with high unburned carbon in pc boiler. M.-y. Hwang , G.-B. Kim, J.-h. Song, S.-Mo Kim, C.-H. Jeon
B04 Ethanol effect on the average structural parameters of IDF soot soluble organic fraction. M. Salamanca , M. Velásquez, F. Mondragon, A. Santamaria
C04 Understanding the effects of biomass addition to coking coals during carbonisation. M. Castro-Díaz , A. Dufour, N. Brosse, R. Olcese, C. Snape
D04 Degradation characteristics of SOFC by trace elements in coal gasified gas. Y. Ueki , T. Kobayashi, R. Yoshiie, I. Naruse
A05 Characteristics of hydrogen sulfide formation in pulverized coal combustion. H. Shirai , M. Ikeda, H. Aramaki
B05 An analysis of the research performed with the Argonne Premium Coals and its contribution to coal science. J. P. Mathews , Y.E. Alvarez, R.E. Winans
C05 Influence of alkali additives on the swelling behavior of a high swelling bituminous coal. C.A.Strydom, J.R. Bunt, Y. van Staden, J. Collins
D05 Coupling gasification and solid oxide fuel cells: effect of tar on anode materials. M. Millan , E. Lorente, J. Mermelstein, C. Berrueco, N.P. Brandon
13:00-13:20
Lunch
13:20-15:00 Coal combustion Chairs: E. Lester & C.-H. Jeon
15:00-15:30
A06 Numerical study of the influence of heterogeneous kinetics on the carbon consumption by oxidation of a single coal particle. P.A. Nikrityuk, M. Gräbner , M. Kestel, B. Meyer
Coal characterization and coal Coal pyrolysis and liquefaction: structure coke microstructure Chairs: A.L. Chaffee & L. Butuzova Chairs: S. Niksa & M.A. Diez B06 Porosity and gas absorption of coals studied by X-ray scattering and modeling. R.E. Winans, S. Seifert, D. Locke, P. Chupas, K. Chapman, M. R. Nariewicz, J. P. Mathews , J. M. Calo
C06 Estimation of coking pressure in coke ovens by Koppers-Incar test. R. Alvarez , C.Barriocanal, M.A. Díez
Coal and the environment Chairs: M.J. Lázaro & L. Santos
D06 Explosions in Coal Mines due to Emission of Molecular Hydrogen via Atmospheric Weathering Processes. H. Cohen
Session A
Session B
Session C
B07 Organic sulphur form alterations in consecutively chemically- and bio-treated lignites. L.Gonsalvesh , S.P.Marinov, M.Stefanova, R.Carleer, J.Yperman
C07 The potential to upgrade petroleum cokes using high temperature processing. M. Ismail , J. W Patrick, E. Lester
D07 A thermo-petrographic method to identify coals prone to self-oxidation. C. Avila, E. Lester
C08 A study of the feasibility of an anthracene oil-based pitch for isotropic carbon fibres preparation N. Díez , P. Álvarez, R. Santamaría, C. Blanco, R. Menéndez M. Granda
D08 Monitoring hot spots in bituminous coal piles stored at atmospheric conditions. H. Cohen , U. Green, F. Gildemeister, L. Metzger, M. Pesimberg, S. Wasserman
15:30-15:50
A07 Pyrolisis and combustion kinectics using the distributed activation energy model. F. Saloojee, S. Kauchali, N. Wagner
15:50-16:10
A08 Plasma supported coal ignition and combustion. V.E. Messerle, E.I. Karpenko, A.B. Ustimenko
B08 Rank dependant formation enthalpy of coal. M. Sciazko
16:10-16:30
A09 A graphite furnace atomic absorption spectrometer as an experimental platform for studying matrix effects in trace element vaporization during coal combustion. E. I. Kozliak , O. V. Klykov, A. A. Raeva, D. T. Pierce, W. S. Seam
C09 Advanced characterisation of liquid hydrocarbons from South B09 Ion beam tomography for coal African high volatile bituminous characterization. coal. A. Bhargava , P.J. Masset, N. M.H. Makgato , H.W.J.P Gordillo, C. Habchi, P. Moretto Neomagus, R.C Everson, J.H.L. Jordaan, H.H. Schobert
16:30-17:00
Coffee break
D09 Reducing the environmental impact of sponteaneous coal combustion in coal waste gobs by applying soil covers. X. Querol , X. Zhuang, J. Li, O. Font, M. Izquierdo, A. Alastuey, B.L. van Drooge, T. Moreno, J. O. Grimalt, F. Plana
Session A
Session B
Session C
CO2 Storage Chairs: M. Lupión & D. Casal
Coal characterization and coal structure Chairs : R. Menéndez & Y. Fernández
Coal pyrolysis and liquefaction: general Chairs: J.-i Hayashi & P. Álvarez
17:00-17:20
A10 Diffuse soil CO2 flux to assess the reliability of CO2 storage in the MazarrónGañuelas Tertiary Basin (Spain). J. Rodrigo-Naharro, O. Vaselli, B. Nisi, M. Lelli, R. Saldaña, C. Clemente-Jul , L. Pérez del Villar
B10 The current state of affairs of coal research in U.S. Universities J. P. Mathews , B. G. Miller, C. S. Song, H. H. Schobert, F. Botha, R.B. Finkleman
C10 Decoupling in Thermochemical Conversion: Approach and Technologies. G. Xu, J. Zhang , Y. Wang, S. Gao
17:20-17:40
A11 Carbon and storage by pH swing aqueous mineralisation using a mixture of ammonium salts. A. Sanna , M. Dri, X. Wang, M. R Hall, M. Maroto-Valer
B11 The Current state of coal research in the United Kingdom, Germany, Australia and South Africa J. P. Mathews , B.G. Miller, C. S. Song, H.H. Schobert, F.Botha, R.B. Finkleman, A. Chaffee
C11 Integrated process of coal pyrolysis with CH4/CO2 activation by dielectric barrier discharge plasma. X. He, H. Hu , L. Jin, Y. Zhao
17:40-18:00
A12 New equipment for characterization of rocks for geological CO2 storage in coal seams. P. Cienfuegos, J.Loredo
B12 Adsorption behavior and biogasification of Soma lignite. M. Baysal, S. İnan, F. Duygun , Y. Yürüm
C12 Effect of steam treatment of a sub-bituminous coal on its caking and coking properties. H. Shui , C. Shan, H. Chang, Z.Wang, Z. Lei, S. Ren, S. Kang
Tuesday, 11th October, 2011 09:00-10:00
Plenary Lecture: (Chair: C.E.Snape) A20 The role of coal science in development and deployment of high efficiency energy technologies. D.J. Harris Session A Session B Session C CO2 transport and storage. Chairs: N.R. Marcilio & M.V. Gil
Coal mineralogy and coal ash: coal ash Chairs: X. Querol & E. Kozliak
Coal pyrolysis and liquefaction: liquefaction Chairs: O. Yamada & R. García
10:00-10:20
A21 Gas adsorption capacity of coaly shales from Japan and B21 Viscosity behaviour of slags USA -Implications for CO2 from coal-petroleum coke blends. storage in coal-bearing formationA. Ilyushechkin , M. Duchesne S. Shimada , Y. Nishiiri, N. Sakimoto, K. Ohga, Y-S. Jun
C21 Interpreting coal conversion under elevated H2 pressures with FLASHCHAIN and CBK. S. Niksa
10:20-10:40
A22 Relationships between the sorption capacity of methane, carbon dioxide, nitrogen and ethane on bituminous coals. R. Sakurovs , S. Day, S. Weir
C22 Development of a ZeroEmission Coal-to-Liquids Plant. W. Atcheson, K. Myers, L. O’Sullivan, H.H. Schobert
10:40-11:00
A23 CIUDEN CO2 Transport Test Rig: Technical Description and Experimental Plan. B. Navarrete ; P. Otero; I. Llavona; M.A. Delgado
11:00-11:30
B22 Study on clinker generation control in coal combustion boiler; clinker controlling effect of Febased coal additive. N. Wakabayashi , H. Shirai B23 Properties, microstructure and leaching of coal slag with additives during high temperature gasification. Y. Ninomiya , Y. Wei, K. Honma, T. Tanosaki, H. Li, M. Kawaguchi, N. Tatarazako Coffee break
Session A
Session B
Session C
CO2 capture and storage: oxyCoal pyrolysis and liquefaction: Coal gasification and clean fuels fundamentals fuel Chairs: J. P. Mathews & E. Jorjani Chairs: R. Malhotra & J.J. Chairs: R. Davidson & A. Fernández Ustimenko
11:30-12:00
A24 Doosan Power Systems B24 What is Reactive Surface Area OxyCoal™ Technology. in Coal Chars? M.D. Maloney , B. Dhungel, D.W. E. M. Suuberg , I. Aarna, I. Külaots Sturgeon, P. Holland-Lloyd
C24 Direct CTL: innovative analyses for high quality distillates. A. Quignard , N. Caillol, N. Charon, M. Courtiade, D. Dendroulakis
12:00-12:20
A25 CIUDEN CO2 Technology Development Centre on Oxycombustion. M Lupion , V J Cortes, M Gomez, A Fernandez
B25 The properties of chars derived from inertinite-rich, high ash coals and CO2 gasification: major properties affecting reactivity. G. N. Okolo, R.C. Everson , H.W.J.P. Neomagus
C25 Direct coal-liquid hydrogenation over NiMoNX / Al2O3 catalysts. C. Qi , G. Lin, F. Jie , L. Wenying , X. Kechang
12:20-12:40
A26 Oxy-fuel coal gasification in fluidised beds. N. Spiegl, E. Lorente, N. Paterson, C. Berrueco, M. Millan
B26 Characterization and carbon dioxide gasification kinetics of high ash inertinite-rich South African coals. R. Kaitano , R. C Everson, H.W J P Neomagus
C26 Preparation of activated carbons from direct coal liquefaction residue. J. Zhang, L. Jin, B. Qiu, H. Hu
12:40-13:00
A27 Experimental and numerical investigations of oxy-coal combustion in an entrained flow reactor. L. Álvarez , M. Gharebaghi, A. Williams, M. Pourkashanian, J. Riaza, C. Pevida, J.J. Pis, F. Rubiera
B27Gasification kinetics of coal char using direct measurement of particle temperature. R. Kim, H. Lim, C. Kim, J. Song, C.H. Jeon
Session A
13:00-13:20
Session B
Session C
B28 Performance simulations for coA28 Char characterisation from gasification of coal and methane. oxyfuel combustion. S. Niksa , J.-P. Lim, D. del Rio Diaz A. Nuamah , E. Lester, T. Drage, Jara, D. Steele, D. Eckstrom, R. G. Riley Malhotra, R. B. Wilson Lunch
13:20-15:00 Clean coal technology: Mercury emissions Chairs: D.A. Spears & M. DíazSomoano
Coal petrology Chairs: I.Suárez-Ruiz & D. van Niekerk
Coal pyrolysis and liquefaction: fundamentals Chairs: W.R. Jackson & H. Hu
15:00-15:30
A29 Avoiding mercury emissions from coal combustion:postcombustion research strategies. M. R. Martínez-Tarazona , M. A. López-Antón, R. OchoaGonzález, A. Fuente-Cuesta, P. Abad- Valle, J. RodríguezPérez, F. Inguanzo-Fernández, M. Díaz-Somoano and R. García
B29 Biomarker and Petrographic Evidence for the Origin and Maturity of Oil-Prone Arctic Coal and Associated Bitumen. C. Marshall, D.J. Large, C.E. Snape , W. Meredith, B. Spiro, I Mokogwu
C29 Attempted Production of Blast Furnace Coke from Victorian Brown Coal. A.L. Chaffee , M. M. Mollah, R. S. Higgins, M. Marshall,W. R. Jackson
15:30-15:50
A30 Interpreting the re-emission of elemental mercury during wet FGD scrubbing. B. Krishnakumar, S. Niksa , N.Fujiwara
B30 Coal quality from a mega coal basin from Xianjiang, Northwest China. J. Li , X. Zhuang, X. Querol, O. Font, P.Córdoba
C30 Chlorine retention during the pyrolysis of a Western Australian lignite in a fluidised-bed reactor. J. Zhang, C. Kelly, A. Rossiter, S. Wang, C.-Z. Li
Session A
15:50-16:10
16:10-16:30
16:30-17:00
17:00-18:00
Session B B31 The effect of particle size and petrographic composition on A31 The fate of Hg at two coal combustion behaviour of selected power plants equipped with FGD. Russian coals. P Córdoba , O Font, M. V. Hudspith, A. Nuamah , A.C. Izquierdo, X. Querol Scott, T. Drage, J. Powis. G. Riley, M.E. Collinson, E.Lester A32 Aqueous chemistry of mercury in flue gas desulphurization conditions. R. Ochoa-González , M. DíazSomoano, M. R. MartínezTarazona
B32 The effect on coal flotation of segregating group macerals based on particle size. E. Jorjani , S. Esmaeili, M.T. Khorami
Session C C31 Dimethyl ether production from Victorian brown coal biomass: a comparative process modelling study. K. B. Kabir, G. Grills, J. Walter, S. Bhattacharya C32 Synchronous fluorimetric characterization of preasphaltene and asphaltene from direct liquefaction of coal. Z. Wang, C. Wei, H. Shui , Z. Wang, C. Pan, S. Ren, Z. Lei
Coffee break Poster Session: Coal Chemistry, Coal Petrology, Coal Mineralogy, Coal Upgrading, Coal Pyrolysis, Carbons from Coal P1.01 Jurassic perhydrous coals of the Lusitanian Basin, Portugal. Petrography, geochemical and textural characteristics A. Costa , C. Tomás, I. Suárez-Ruiz, P.P. Cunha, D. Flores, B. Ruiz P1.02 Rank distribution of the Cretaceous coals in the region of San Juan de Sabinas, Coahuila de Zaragoza, Mexico N. Piedad-Sánchez , I. Suárez-Ruiz, F. R. Carrillo-Pedroza, J. A. Moreno-Hirashi, G. de la RosaRodríguez, K. Flores-Castro, R. Corona-Esquivel, J. L. Cadena-Zamudio, J. O. Navarro-Lozano, B. Santiago-Carrasco, F. González-Carrillo P1.04 Correlation between fluidity and structure transformation of three coal ashes. X. Lin , J. Miyawaki, S.-H. Yoon, I. Mochida P1.05 Nanominerals and ultra-fine particles within coal ashes. L.F.O. Silva; F. Waanders; M.L. S. Oliveira ; K. da Boit P1.06 Properties of fly ash from biomass combustion. R. P. Girón , I. Suárez-Ruiz, B. Ruiz, E. Fuente, R. R. Gil P1.08 Characteristics of crush strength in coal briquette molded with polymer as a binder. S.-H. Moon , S.-J. Lee, I.-S. Ryu, Y.-W. Kim, T.-I. Ohm
P1.09 Characteristics of dried low-rank coal by hot oil immersion drying method for the upgrading T.-I. Ohm , J.-S. Chae, S.-H. Moon P1.10 Energy and exergy analysis of continuous microwave dryer. M. B. Alvarado , E. J. Muñoz, S.C. Navarro, F. Chejne, H. Velazquez P1.11 Moisture re-adsorption characteristics of coal samples of dried by a pneumatic dryer. S. Kim , Y. Rhim, S. Lee P1.12 Moisture readsorption characteristics of upgraded low rank coal. H. Choi , S. Kim, J. Yoo, D. Chun, J. Lim, Y. Lim, S. Lee P1.13 Solar drying technology of coal in the open in the Iron and Steel Research Center. A. Leyva. Mormul , A. D. Castillo, O. S. Leyva González , J.A. Trotman Gavilán, O. Figueredo Stable P1.14 A New Supercritical Solid Acid for Breaking Car-Calk Bond in Di(1-naphthyl)methane. X.-M. Yue , X.-Y. Wei, B. Sun, Y.-H. Wang, Z.-M. Zong, Z.-W. Liu 17:00-18:00
P1.15 Briquetting of carbon-containing wastes from steelmaking for metallurgical coke production. M.A. Diez , R. Alvarez, J.L.G. Cimadevilla P1.16 Co-carbonization behaviour of coal and biomass-derived products and its effect on coke structure and properties. M.A. Diez , R. Alvarez, M. Fernández P1.17 Evolution of volatile products of coal and plastic wastes during co-pyrolysis. S. Melendi, M.A. Diez , R. Alvarez P1.18 Role of selected coal- and petroleum-based additives in low- and high-temperature co-pyrolysis with coal blends. E. Rodríguez , S. Melendi, R. García, R. Alvarez, M.A. Diez P1.19 Effect of volatile matter evolution on optical properties of macerals from different rank coals A. Guerrero , M.A. Diez, A.G. Borrego P1.20 Semi-pilot scale carbonization to assess blast furnace coke quality. E. Díaz-Faes, R. Alvarez , C. Barriocanal, M.A. Díez
P1.21 Fundamental investigation of pyrolysis behavior of low rank coals. T. Harada , S. Matsuda, N. Wada, Y. Matsushita, I. Mochida P1.22 High pressure pyrolysis of different coal types – Influence of pressure on devolatilisation characteristics using TGA/MS. M. Klinger , B. Meyer P1.23 Brown coal and rape cake co-pyrolysis products in the range 5 to 40 per cent. J. Vales , J. Kusy, L. Andel, M. Safarova P1.24 Integrated coal pyrolysis with methane aromatization over Mo/HZSM-5 catalyst for improving tar yield. X. Zhou, H. Hu , L.Jin P1.25 Integrated process of coal pyrolysis and CO2 reforming of methane over Ni/Al2O3-MgO catalyst. J. Liu, H. Hu , L. Jin, S. Zhu P1.26 Kinetics of co-pyrolysis of high- and low-sulfur coal blends with additives. L. Butuzova , R. Makovskyi, V. Bondaletova, D. Dedovets, G. Butuzov 17:00-18:00
P1.27 Effect of elemental composition of various additives on the modification of coal thermoplastic properties. M.G. Montiano , C. Barriocanal, R. Alvarez P1.28 Influence of residual volatile matter in semicokes on coking pressure. E. Díaz-Faes, C. Barriocanal , R. Alvarez P1.29 Pyrolysis of wastes from tyre grinding. B. Acevedo, C. Barriocanal, R. Alvarez P1.30 Secondary reactions of HCl during coal pyrolysis: studies on reactions of HCl with model carbons. N. Tsubouchi , N. Ohtaka, A. Kawashima, Y. Ohtsuka P1.31 Solvent modulation in getting high purity of anthracene and carbazole from crude anthracene. M. Fan, C. Ye , H. Zheng, T. Wu, J. Feng , W. Li P.1.32 Study of coal, char and coke fines structures and their proportions in the off-gas blast furnace samples by X-Ray diffraction. A.S. Machado, A.S. Mexias, A.C.F. Vilela, E. Osório P1.34 Yields of the pyrolysis tests of brown coals mined in the Czech Republic. J. Kusý, L. Anděl , M. Šafářová, J. Valeš
P1.37 Wet oxidation of anthracene oil-based pitch - a way to porous carbons. H. Machnikowska, K. Torchała, G. Gryglewicz , J. Machnikowski
17:00-18:00
P1.38 Adsorption of phenol on nitrogen enriched activated carbons prepared from coal-tar pitch and polymers. E. Lorenc-Grabowska, G. Gryglewicz , M.A. Diez, C. Barriocanal P1.39 Mineral matter and heat treatment temperature effects on the development of graphitic structure in two South African anthracites as studied by Raman Spectroscopy and XRD M. Vanegas-Chamorro , K. Tamargo-Martínez, J. Xiberta, A. Martínez-Alonso, J. M. D. Tascón P1.40 Structural evolution upon thermal treatment of two South African anthracites as studied by XRD. M. Vanegas-Chamorro , K. Tamargo-Martínez, J. Xiberta, A. Martínez-Alonso, J. M. D. Tascón
Wednesday, 12th October, 2011 09:00-10:00
Plenary Lecture: (Chair: H.H. Schobert) A40 A cubic mile of oil: realities and options for averting the looming global energy crisis. R. Malhotra Session A
Session B
Clean coal technology: Coal gasification and clean fuels: Mercury emissions mineral matter Chairs:H. Cohen & N. Tsubouchi Chairs: A. Sharma & Q. Campbell
10:00-10:20
A41 Gold-impregnated carbon materials as regenerable sorbents for mercury retention. J. Rodríguez-Pérez , F. InguanzoFernández, E. Rodríguez, M. A. López-Antón, R. García, M. DíazSomoano, M. R. MartínezTarazona
B41: Entrained-flow gasification of coal under slagging conditions: properties of solid wastes and relevance of char-wall interaction phenomena. F. Montagnaro , P. Brachi, P. Salatino
Session C Coal Pyrolysis Chairs: R. Kandiyoti & N. Okuyama
C41 Characterization of the hydrocarbon components from sequential flash pyrolysis for a vitrinite-rich and inertinite-rich coal. D. van Niekerk , C. du Sautoy, J. van Heerden
Session A
Session B
10:20-10:40
A42 Speciation and fate of mercury in oxy coal combustion. O. Font , P. Córdoba P, C. Leiva, L.M. Romeo, I. Bolea, I. Guedea, N. Moreno, X. Querol, C. Fernandez-Pereira, L.I. Díez
B42 Mineral-char interaction during the gasification of high ash coals in a fluidised bed gasifier: Redistribution of mineral phases within the char matrix. B.O. Oboirien, A.D. Engelbrecht, B.C. North, R. Falcon
C42 Characterisation of coal and biomass based on kinetic parameter distributions for pyrolysis. N. Sonoyama , J.-i. Hayashi
10:40-11:00
A43 Experimental study on the SO2 emission and Calcium-based desulfurization in the coal Oxygen-enriched combustion. L. Tian , H. Chen, H. Yang, X. Wang, S. Zhang, C. Zeng
B43 Introduction of a ternary diagram for comprehensive evaluation of gasification processes for high ash coals. M. Gräbner , B. Meyer
C43 Effect of the pyrolysis conditions on the microstructure of anthracene oil-based cokes. P. Alvarez, N. Díez, R. Santamaría, C. Blanco, R. Menéndez , M. Granda
Coffee break
11:00-11:30
11:30-12:00
12:00-12:20
Session C
CO2 capture and storage: Chemical looping Chairs T. Drage & S. Bhattacharya A44 Current status of the chemical looping combustion technology. F. García-Labiano , L. F. de Diego, P. Gayán, A. Abad, J. Adánez A45 Oxygen transfer from metal oxides during chemical looping combustion of Victorian brown coal – An experimental and modelling study. C. Saha, T. X. Seng, A. Auxilio , S. Bhattacharya
Mineral matter and coal ash: Mineral matter Chairs: R.C. Everson & M.R. Martínez-Tarazona
Coal pyrolysis and liquefaction: coke microstructure Chairs: R. Sakurovs & S. Melendi
B44 The determination of trace element distributions in coal. A "new" approach. D.A. Spears
C44 Optical and scanning electron microscopy of coke: microstructure & minerals. S. Gupta, E Lester , M Ismail, G O'Brien
B45 Solid-state NMR study on mineral structure and transformation behaviors of coal ash. X. Lin, K. Ideta, J.Miyawaki, S.-H. Yoon , I.Mochida
C45 Small scale determination of metallurgical coke CSR. T. MacPhee , L. Giroux, K.W. Ng, T. Todoschuk, M. Conejeros, C. Kolijn
Session A
Session B
12:20-12:40
A46 Chemical looping combustion of coal using a residue from alumina production. T. Mendiara , G. Ferrer, P. Gayán, A. Abad, F. GarcíaLabiano, L. F. de Diego, J. Adánez
B46 The effect of minerals on the moisture adsorption and desorption properties of South African fine coal. S.M. du Preez , Q.P. Campbell
C46 3-D Structural analysis for metallurgical coke microstructure using micro X-ray CT. Y. Yamazaki , K. Hiraki, T. Kanai, X. Zhang, A. Uchida, M. Shoji, Y. Matsushita, H. Aoki, T. Miura, S. Nomura, H. Hayashizaki
12:40-13:00
A47 Chemical looping combustion of char with a Cubased carrier. A. Coppola, O. Senneca, R. Solimene, R. Chirone, L. Cortese, P. Salatino
B47 Structural changes and possible modes of interaction in bituminous coal fly ash due to treatments with neutral and acidic aqueous solutions. R.N. Lieberman , R.Nitzsche, H. Cohen
C47 Preparation of high-strength coke from hot-briquetted brown coal. A. Mori, Y. Huang, K. Norinaga, S. Kudo, T. Kanai, H. Aoki, J.-i. Hayashi
13:00-13:20
A48 Theoretical approach on the CLC performance with solid fuels: optimizing the solids inventory. A. Cuadrat, A. Abad, P. Gayán, L. F. de Diego, F. García-Labiano, J. Adánez
B48 Prediction of selective trace element emissions during oxy-CFB combustion of Victorian brown coals. B. Roy, W. L. Choo, S. Bhattacharya
C48 A mechanism of improvement in coke strength by adding a solvent-extracted coal N. Okuyama , T. Shishido, K. Sakai, M. Hamaguchi, N. Komatsu
Lunch
13:20-15:00 CO2 capture and storage: Chairs: M. Millan & M. Alonso
15:00-15:30
Session C
A49 Puertollano IGCC: Towards zero emissions power plants. F. García Peña , P. Coca Llano
Coal gasification and clean fuels Chairs: D. Harris & F. Mondragon B49 Evolution of Victorian brown coal char structure during the gasification in CO2 and steam. H.-L. Tay, S. Kajitani, C.-Z. Li
Coal Upgrading Chairs: K. Miura & T. MacPhee C49 Coal drying and dewatering for power generation – current status, research and development needs. D. Stokie, J. Yu, A. Auxilio , S. Bhattacharya
Session A
Session B
Session C
15:30-15:50
A50 Pre-combustion CO2 capture: laboratory- and benchscale studies of a sweet watergas-shift catalyst for H2 and CO2 production. J.M. Sánchez , M. Maroño, D. Cillero, L. Montenegro, E. Ruiz
15:50-16:10
A51 Regeneration of used alkali carbonates for removal of gaseous sulfur compounds in gasification process. S. Raharjo , Y.Ueki, R. Yoshiie, I. Naruse
B51 A CeO2-La2O3-based Cu catalyst for application in hightemperature water-gas shift reaction. L D. Morpeth , Y. Sun, S.S. Hla, G.J. Duffy, J.H. Edwards, D.J. Harris, D.G. Roberts
C51 Upgrading and dewatering of low rank coals realizing the suppression of self-ignition tendency through solvent treatment at around 350°C. H. Fujitsuka , R. Ashida, K Miura
16:10-16:30
A52 Step Change Adsorbents and Processes for CO2 capture “STEPCAP”. T.C. Drage , A.I Cooper, R Dawson, J Jones, C Cazorla Silva, C.E. Snape, L. Stevens, X. Guo, J. Wood, J. Wang
B52 Coal plasma gasification for clean synthesis gas production. V.E. Messerle, A.B. Ustimenko , N. Slavinskaya, O.A. Lavrichshev, E.F. Ossadchaya
C52 Upgrading of low-quality coals by thermal extraction. T. Takanohashi , N. Sakimoto, K. Koyano, Y. Harada, H. Fujimoto
16:30-17:00
17:00-18:00
B50 Development of a new synthesis gas production process from coal by catalytic gasification of HyperCoal using steam-CO2 as gasifying agent. A.Sharma , T. Takanohashi
C50 Coprocessing of low-rank coal and biomass utilizing mild solvent treatment at around 350°C. X. Li, J. Wannapeera, N. Worasuwannarak, R. Ashida , K. Miura
Coffee break Poster Session: Coal Combustion, Clean Coal Technologies, Coal Gasification, CO2 Capture and Storage, Coal and The Environment P2.01 Arsenic leachability and speciation in fly ashes from coal fired power plants. S. Kambara, M. Endo , S. Takata, K. Kumabe, H. Moritom i P2.02 Low temperature SNCR by photochemical activation of ammonia. S. Kambara , M. Kondo, N. Hishinuma, M. Masui, K. Kumabe, H. Moritomi P2.03 Optimum temperature for sulphur retention in fluidised beds working under oxy-fuel combustion conditions. A. Rufas , M. de las Obras-Loscertales, L.F. de Diego, F. García-Labiano, A. Abad, P. Gayán, J. Adáne z P2.04 Carbon based catalytic briquettes for NOx removal in flue gases. M.J. Lázaro , M.E. Gálvez, S. Ascaso, I. Suelves, R. Moliner
P2.05 Identification of operational regions in the chemical- looping with oxygen uncoupling (CLOU) process with a Cu-based oxygen- carrier. I. Adánez-Rubio , A. Abad, P. Gayán, L. F. de Diego, F. García-Labiano, J. Adánez P2.06 Influence of hydrogenation on the mercury capture by active carbons. J. Rodríguez-Pérez , M. A. López-Antón, R. García, M. Díaz-Somoano, M. R. Martínez-Tarazona P2.07 Sub-products of gasification as sorbents for mercury retention. A. Fuente-Cuesta , M. Diaz-Somoano, M.A. Lopez-Anton, M.R. Martinez-Tarazona P2.08 Biomimetic sequestration of CO2 and conversion to CaCO3 using enzyme extracted oyster. S.K. Jeong, Y.I. Yoon, S.C. Nam P2.09 Carbon dioxide sequestration by aqueous mineral carbonation of serpentine and explanation of experimental results. K. Alizadehhesari , K. Steel P2.10 Characterization of a novel flat-panel airlift photobioreactor with the internal heat exchanger L. H. Kochem, N. C. da Fré, C. Redaeli, N. R. Marcílio , R. Rech
17:00-18:00
P2.11 Clean coal technologies scenario and evaluation of present CO2 dwindling initiatives to approach zero emission power stations by coal combustion. Deployment situation and evaluation study. F. Guerrero, C. Clemente-Jul P2.12 Co-combustion of coal and biomass blends in an entrained flow reactor under oxy-fuel atmospheres. J. Riaza, L. Álvarez , M.V. Gil, C. Pevida, F. Rubiera, J.J. Pis P2.13 Effect of the activation temperature and the burn - off degree on the CO2 capture capacity of microporous activated carbons. M.V. Gil , M. Martínez, S. García, J.J. Pis, F. Rubiera, C. Pevida P2.14 Influence of light over CO2 biofixation by the microalgae Chlorella minutissima. C. Redaelli, R. Rech, N. R. Marcilio P2.15 Influence of temperature and salinity over CO2 biofixation by the microalgae Dunaliella tertiolecta. N. C. da Fré, R. Rech, N. R. Marcílio P2.16 Carbon oxides emission via the atmospheric oxidation of coals: effect of coal rank. U. Green, Z. Aizenstat, H. Cohen P2.18 Environmental pollution by migration of gas produced in the Underground Coal Gasification process. M. Ludwik-Pardał a, K. Stańczyk
P2.19 Fly ash as a potential scrubber for low activity radioactive waste. R.N. Liberman , G. Segev, E. Elish, E. J. C. Borojovich, H. Cohen P2.20 Mobility of major and minor species in fly ash-brine co-disposal systems: up-flow percolation test. O.O. Fatoba , W.M. Gitari, L.F. Petrik, E.I. Iwuoha P2.21 Removal of arsenate from water by adsorption onto lignite. Y. Yürüm , Z. Özlem Kocabas P2.22 Biomass and mineral coal in South Brazil: potential use for energy generation in bubbling fluidized bed. G.M.F.Gomes, L.Dalla Zen , A.C.F. Vilela, E. Osório P2.23 Effect of particle size in the devolatilization behaviour of coal chars of different rank. K.S. Milenkova and A.G. Borrego P2.24 The element distribution during the co-combustion of coal with wood and wood wastes. Z. Klika , L. Bartoňová P2.25 The influence of particle size and density on the combustion of Highveld coal. G.W. van der Merwe, R.C. Everson , H.W.J.P. Neomagus, J.R. Bunt 17:00-18:00
P2.26 Effect of blending waste materials with coal on minerals and reactivity of char and coke. A.M. Fernández, C. Barriocanal, S. Gupta, D. French P2.27 An investigation into the catalytic potential of coal ash constituents on the CO2 gasification rate of high ash, South African coal. B.B. Hattingh, R.C. Everson, H.W.J.P. Neomagus, J.R. Bunt P2.28 Catalyst recovery using by support in steam gasification of lignite at low temperature. Y.-K. Kim , J.-I. Park, J. Miyawaki, I. Mochida, S.-H. Yoon P2.29 Evaluation of coal gasification reaction from composition of gases produced (2). M. Kaiho, O. Yamada , H. Yasuda, S. Shimada, M. Fujioka P2.30 Effect of calcium on the interaction of CO2 with carbonaceous materials during coal gasification. J. D. González , F. Mondragón, J. F. Espinal P2.31 Interaction of calcium with carbonaceous materials: A DFT Study. J.D. González , F. Mondragón, J. F. Espinal P2.32 Use of a waste generated in the cement industry as an additive in the process of coal gasification in fluidized bed. E. Rodríguez Acevedo, F. Chejne , W. Jurado
17:00-18:00
P2.33 Modelling and simulation of a coal gasification process in the pressurized fluidized bed. F. Chejne , E. Lopera, C. A. Londoño, C. A. Gómez P2.34 Preliminary studies on ash-free coal gasification at mild condition. J. Yoo , S. Jin, H. Choi, Y. Rhim, J. Lim, D. Chun, S. Kim, S. Lee P2.37 Kinetic study on the lignite- CO2 gasification in the presence of K2CO3. V.C. Bungay , B.H. Song, S.D. Kim, J.M. Sohn, H.M. Shim, Y.J. Kim, G.T. Kim, S.R. Park, Y.I. Lim P2.38 Prediction of steam reforming of the simulated coke oven gas with a detailed chemical kinetic model. K. Norinaga , R. Sato, J.-i. Hayashi P2.39 Invention of quantitative method of char and soot to clarify soot production and reaction behavior in coal gasification. S. Umemoto , S. Kajitani, S. Hara
Thursday, 13th October, 2011 09:00-10:00
Plenary Lecture: (Chair:J. van Heerden) A60 Calcium looping technologies for CO2 capture C. Abanades Session A Session B CO2 capture and storage Chairs: F. García-Labiano & J. Chamberlain
10:00-10:20
A61 N2O and CO emissions increase during oxy-char combustion in fluidized bed. A.Sánchez, Y. Betancur, E. Eddings, F. Mondragón
Session C
Coal gasification and clean fuels Chairs: C.-Z. Li & J.M. Sánchez
Coal Upgrading C. Clemente-Jul & M.A. López Antón
B61 Chemical Looping Combustion of Coal-derived Synthesis Gas containing H2S over Supported Fe2O3 - MnO2 Oxygen Carrier. E. Ksepko , R.V. Siriwardane, H. Tian, T. Simonyi, M. Sciazko
C61 Drying behavior of brown coal under the various temperature conditions with halogen heat source and its formulation. Y. Matsushita, N. Mitsuhara, T. Harada
Session A
Session B
Session C
10:20-10:40
A62 New candle prototype for hot gas filtration industrial applications. M. Rodríguez-Galán, M. Lupión , B. Alonso-Fariñas, J. MartínezFernández
B62 Measurement of gasification rate of coal char under high pressure and high temperature using a mini directly-heated reactor. K. Miura , M. Makino, E. Sasaoka, S. Imai, R. Ashida
C62 Mobility of hazardous elements of Coal Cleaning Residues. L.F.O. Silva; F. Waanders ; M.L.S. Oliveira; K. da Boit
10:40-11:00
A63 Fluid dynamic simulation of dry filter for removal of particulates from coal and biomass gasification. C. B. da Porciúncula, N. R. Marcilio , M. Godinho, A.R.Secchi
B63 Implementation of coal gasification in a fluidized bed firing system for brick tunnel kiln. F. Chejne , C. londono, C. Gómez, J. Espinosa, F. Mondragon, J.J Fernandez, E. Arenas, L. C Cuartas Coffee break
11:00-11:30 CO2 capture and storage Chairs: C. Abanades & M.D. Maloney
11:30-11:50
A64 Capture of CO2 during low temperature biomass combustion in a fluidized bed using CaO. A new larger scale experimental facility. J.R. Chamberlain , C. Perez Ros
Coal gasification and clean fuels Chairs: F. Chejne & M. Gräbner
B64 Catalytic steam gasification of large coal particles. S. Nel , H.W.J.P Neomagus, J.R. Bunt, R.C. Everson
C63 Modelling of coal slag viscosity: Focus on the volume fraction of solid particles. A. Bronsch , P. J. Masset
Biomass Processing Chairs: Z. Klika & E. Osorio
C64 Effect of ash components on devolatilisation behaviour of coal and biomass – product yields, properties and heat requirement. D. Reichel , M. Klinger, S. Krzack, B. Meyer
Session A
11:50-12:10
A65 La Pereda CO2. A 1.7 MWt pilot to test postcombustion CO2 capture with CaO. A. Sánchez-Biezma, J Paniagua, L. Diaz, E. de Zarraga, J. López, J Alvarez, B. Arias , M. Alonso, J.C. Abanades
12:10-12:30
A66 Fluidized bed desulfurization using lime obtained after slow calcination of limestone particles. F. Scala , R. Chirone, P. Meloni, G. Carcangiu, M. Manca, G. Mulas, A. Mulas
12:30-12:50
A67 Synthetic gas bench study of CO2 capture from PCC power plants. E. Ruiz , J. M. Sánchez, M. Maroño, J. Otero
12:50-13:30 13:30-15:00
Session B
Session C
B65 The influence of particle size on the steam gasification of coal. G. H. Coetzee , H.W.J.P Neomagus, R.C. Everson
C65 How closely do low volatile bituminous coals prepared by hydrous pyrolysis of woody biomass and low-rank coals correspond to prime coking coals?. S. Kokonya, M. C. Diaz, C. Uguna, C. Snape , A.D. Carr
B66 Effect of iron and calcium catalysts on pyrolysis and steam gasification of wood. K. Murakami, M. Sato, T. Kato, K. Sugawara
C66 Biomass characterisation and link with char morphology and ashing behaviour. C. H. Pang , T. Wu, E. Lester
B67 Thermodynamic efficiency analysis of gasification of high ash coal and biomass. R. Rodrigues, N.R. Marcilio , J.O. Trierweiler, M. Godinho Closing ceremony Lunch
Authors A Aarna, I.
B24
Álvarez, P.
C08, C43, D02
Abad, A.
A44, A46, A48, P2.03, P2.05
Álvarez, R.
C06, P1.15, P1.16, P1.17, P1.18, P1.20, P1.27, P1.28, P1.29
Abad-Valle, P.
A29
Álvarez, Y.E.
B01, B05
Abanades, J.C.
A60, A65
Amer, M.W.
B03
Acevedo, B.
P1.29
Anděl, L.
P1.23, P1.34
Adánez, J.
A44, A46, A48, P2.03, P2.05
Aoki, H.
C46, C47
Adánez-Rubio, I.
P2.05
Aramaki, H.
A05
Aizenstat, Z.
P2.16
Arenas, E.
B63
Alastuey, A.
D09
Arias, B.
A65
Alizadehhesari, K.
P2.09
Ascaso, S.
P2.04
Alonso, M.
A65
Asenjo, N.G.
D02
Alonso-Fariñas, B.
A62
Ashida, R.
B62, C50, C51
Alvarado, M.B.
P1.10
Atcheson, W.
C22
Álvarez, J.
A65
Auxilio, A.
A45, C49
Álvarez, L.
A27, P2.12
Avila, C.
D07
Barriocanal, C.
C06, P1.20, P1.27, P1.28, P1.29, P1.38, P2.26
Botha, F.
B10, B11
Bartoňová, L.
P2.24
Brachi, P.
B41
Baysal, M.
B12
Brandon, N.P.
D05
Berrueco, C.
A26, D05
Bronsch, A.
C63
B
Betancur, Y.
A61
Brosse, N.
C04
Bhargava, A.
B09
Budinova, T.
C03
Bhattacharya, S.
A45, B48, C31, C49
Bungay, V.C.
P2.37
Blanco, C.
C08, C43, D02
Bunt, J.R.
B64, C05, P2.25, P2.27
Bolea, I.
A42
Burgess Clifford, C.E.
D01
Bondaletova, V.
P1.26
Butuzov, G.
P1.26
Borojovich, E.J.C.
P2.19
Butuzova, L.
C03, P1.26
Borrego, A.G.
P1.19, P2.23
C Cadena-Zamudio, J.L.
P1.02
Chun, D.
P1.12, P2.34
Caillol, N.
C24
Chupas, P.
B06
Calo, J.M.
B06
Cienfuegos, P.
A12
Campbell, Q.P.
B46
Cillero, D.
A50
Carcangiu, G.
A66
Cimadevilla, J.L.G.
P1.15
Carleer, R.
B07
Clemente-Jul, C.
A10, P2.11
Carr, A.D.
C65
Coca Llano, P.
A49
Carrillo-Pedroza, F.R.
P1.02
Coetzee, G.H.
B65
Castillo, A.D.
P1.13
Cohen, H.
B47, D06, D08, P2.16, P2.19
Castro-Díaz, M.
C04
Collins, J.
C05
Castro-Marcano, F.
B01
Collinson, M.E.
B31
Cazorla Silva, C.
A52
Conejeros, M.
C45
Chae, J.S.
P1.09
Cooper, A.I.
A52
Chaffee, A.L.
B02, B03, B11, C29
Coppola, A.
A47
Chamberlain, J.R.
A64
Córdoba, P.
A31, A42, B30
Chang, H.
C12
Corona-Esquivel, R.
P1.02
Chapman, K.
B06
Cortés, V,J.
A25
Charon, N.
C24
Cortese, L.
A47
Chejne, F.
B63, P1.10, P2.32, P2.33
Costa, A.
P1.01
Chen, H.
A43
Courtiade, M.
C24
Chirone, R.
A47, A66
Cuadrat, A.
A48
Choi, H.
P1.12, P2.34
Cuartas, L.C.
B63
Choo, W.L.
B48
Cunha, P.P.
P1.01
da Boit, K.
C62, P1.05
Diaz, L.
A65
da Fré, N.C.
P2.10, P2.15
Díaz, M.C.
C65
da Porciúncula, C.B.
A63
Díaz-Faes, E.
P1.20, P1.28
Dalla Zen , L.
P2.22
Díaz-Somoano, M.
A29, A32, A41, P2.06, P2.07
Dawson, R.
A52
Díez, L.I.
A42
Day, S.
A22
Díez, M.A.
C06, P1.15, P1.16, P1.17, P1.18, P1.19, P1.20, P1.38
de Diego, L.F.
A44, A46, A48, P2.03, P2.05
Díez, N.
C08, C43
de la Rosa-Rodríguez, G.
P1.02
Drage, T.C.
A28, A52, B31
de las Obras-Loscertales, M.
P2.03
Dri, M.
A11
de Zarraga, E.
A65
du Preez, S.M.
B46
Dedovets, D.
P1.26
du Sautoy, C.
C41
del Río Díaz Jara, D.
B28
Duchesne, M.
B21
Delgado, M.A.
A23
Duffy, G.J.
B51
Dendroulakis, D.
C24
Dufour, A.
C04
D
Dhungel, B.
A24
Duygun, F.
B12
Eckstrom, D.
B28
Engelbrecht, A.D.
B42
Eddings, E.
A61
Esmaeili, S.
B32
Edwards, L.H.
B51
Espinal, J.F.
P2.30, P2.31
Elish, E.
P2.19
Espinosa, J.
B63
Endo, M.
P2.01
Everson, R.C.
B25, B26, B64, B65, C09, P2.25, P2.27
Falcon, R.
B42
Figueredo Stable, O.
P1.13
Fan, M.
P1.31
Finkleman, R.B.
B10, B11
Farrow, T.S.
A03
Flores, D.
P1.01
Fatoba, O.O.
P2.20
Flores-Castro, K.
P1.02
Fei, Y.
B03
Font, O.
A31, A42, B30, D09
Feng, J.
P1.31
French, D.
P2.26
Fernández, A.
A25
Fuente, E.
P1.06
Fernández, A.M.
P2.26
Fuente-Cuesta, A.
A29, P2.07
Fernández, J.J.
B63
Fujimoto, H.
C52
Fernández, M.
P1.16
Fujioka, M.
P2.29
Fernandez-Pereira, C.
A42
Fujitsuka, H.
C51
Ferrer, G.
A46
Fujiwara, N.
A30
P2.04
Gómez, C.
B63
E
F
G Gálvez, M.E.
Gao, H.
A02
Gómez, C.A.
P2.33
Gao, S.
C10
Gómez, M.
A25
García Peña, F.
A49
Gonsalvesh, L.
B07
García, R.
A29, A41, P1.18, P2.06
González, J.D.
P2.30, P2.31
García, S.
P2.13
González-Carrillo, F.
P1.02
García-Labiano, F.
A44, A46, A48, P2.03, P2.05
Gordillo, N.
B09
Garten, B.
C02
Gräbner, M.
A06, B43
Gayán, P.
A44, A46, A48, P2.03, P2.05
Granda, M.
C08, C43, D02
Gharebaghi, M.
A27
Green, U.
D08, P2.16
Gil, M.V.
P2.12, P2.13
Grills, G.
C31
Gil, R.R.
P1.06
Grimalt, J.O.
D09
Gildemeister, F.
D08
Gryglewicz, G.
P1.37, P1.38
Girón, R.P.
P1.06
Guedea, I.
A42
Giroux, L.
C45
Guerrero, A.
P1.19
Gitari, W.M.
P2.20
Guerrero, F.
P2.11
Godinho, M.
A63, B67
Guo, X.
A52
Gomes, G.M.F.
P2.22
Gupta, S.
C44, P2.26
Habchi, C.
B09
Higgins, R.S.
C29
Hall, M.R.
A11
Hiraki, K.
C46
Hamaguchi, M.
C01, C48
Hishinuma, N.
P2.02
Hara, S.
P2.39
Hla, S.S.
B51
Harada, T.
C61, P1.21
Holland-Lloyd, P.
A24
Harada, Y.
C52
Honma, K.
B23
H
Harris, D.J.
A20, B51, C02
Hu, H.
C11, C26, P1.24, P1.25
Hattingh, B.B.
P2.27
Huang, Y.
C47
Hayashi, J.-i.
C42, C47, P2.38
Hudspith, V.
B31
Hayashizaki, H.
C46
Hwang, M.Y.
A04
He, X.
C11
I Ideta, K.
B45
Inguanzo-Fernández, F.
A29, A41
Ikeda, M.
A05
Iondono, C.
B63
Ilyushechkin, A.
B21
Ismail, M.
C07, C44
Imai, S.
B62
Iwuoha, E.I.
P2.20
Ínan, S.
B12
Izquierdo, M.
A31, D09
Jackson, W.R.
B03, C29
Jones, J.
A52
Jeon, C.H.
A04, B27
Jordaan, J.H.L.
C09
Jeong, S.K.
P2.08
Jorjani, E.
B32
Jie, F.
C25
Jun, Y.S.
A21
Jin, L.
C11, C26, P1.24, P1.25, P2.34
Jurado, W.
P2.32
Kabir, K.B.
C31
Kim, S.M.
A04
Kaiho, M.
P2.29
Kim, Y.J.
P2.37
Kaitano, R.
B26
Kim, Y.K.
P2.28
Kajitani, S.
B49, P2.39
Kim, Y.W.
P1.08
Kambara, S.
P2.01, P2.02
Klika, Z.
P2.24
J
K
Kanai, T.
C46, C47
Klinger, M.
C64, P1.22
Kandiyoti, R.
A00
Klykov, O.V.
A09
Kang, S.
C12
Kobayashi, T.
D04
Karpenko, E.I.
A08
Kochanek, M.A.
C02
Kato, T.
B66
Kochem, L.H.
P2.10
Kauchali, S.
A07
Kokonya, S.
C65
Kawaguchi, M.
B23
Kolijn, C.
C45
Kawashima, A.
P1.30
Komatsu, N.
C01, C48
Kechang, X.
C25
Kondo, M.
P2.02
Kelley, D.
B02
Koyano, K.
C52
Kelly, C.
C30
Kozliak, E.I.
A09
Kestel, M.
A06
Krishnakumar, B.
A30
Khorami, M.T.
B32
Krzack, S.
C64
Kim, C.
B27
Ksepko, E.
B61
Kim, G.B.
A04
Kudo, S.
C47
Kim, G.T.
P2.37
Külaots, I.
B24
Kim, R.
B27
Kumabe, K.
P2.01, P2.02
Kim, S.
P1.11, P1.12, P2.34
Kumagai, H.
C01
Kim, S.D.
P2.37
Kusý, J.
P1.23, P1.34
Large, D.J.
B29
Lim, J.P.
B28
Lavrichshev, O.A.
B52
Lim, Y.
P1.12
Lázaro, M.J.
P2.04
Lim, Y.I.
P2.37
Lee, S.
P1.11, P1.12, P2.34
Lin, G.
C25
L
Lee, S.J.
P1.08
Lin, X.
B45, P1.04
Lei, Z.
C12, C32
Liu, J.
P1.25
Leiva, C.
A42
Liu, Z.W.
P1.14
Lelli, M.
A10
Llavona, I.
A23
Lester, E.
A28, B31, C07, C44, C66, D07
Locke, D.
B06
Leyva González , O.S.
P1.13
Londoño, C.A.
P2.33
Leyva Mormul, A.
P1.13
Lopera, E.
P2.33
Li, C.Z.
B49, C30
López, J.
A65
Li, H.
B23
López-Antón, M.A.
A29, A41, P2.06, P2.07
Li, J.
B30, D09
Loredo, J.
A12
Li, W.
P1.31
Lorenc-Grabowska, E.
P1.38
Li, X.
C50
Lorente, E.
A26, D05
Lieberman, R.N.
B47, P2.19
Ludwik-Pardała, M.
P2.18
Lim, H.
B27
Lupion, M.
A25, A62
Lim, J.
P1.12, P2.34
M MacFarlane, D.
B02
Meredith, W.
B29
Machado, A.S.
P1.32
Mermelstein, J.
D05
Machnikowska, H.
P1.37
Messerle, V.E.
A08, B52
Machnikowski, J.
P1.37
Metzger, L.
D08
MacPhee, T.
C45
Mexias, A.S.
P1.32
Magarisawa, K.
D03
Meyer, B.
A06, B43, C64, P1.22
Majeski, A.
A02
Milenkova, K.S.
P2.23
Makgato, M.H.
C09
Millán, M.
A26, D05
Makino, M.
B62
Miller, B.G.
B10, B11
Makovskyi, R.
C03, P1.26
Mine, T.
A01
Malhotra, R.
A40, B28
Mitsuhara, N.
C61
Maloney, M.D.
A24
Miura, K.
B62, C50, C51
Manca, M.
A66
Miura, T.
C46
Marcilio, N.R.
A63, B67, P2.10, P2.14, P2.15
Miyawaki, J.
B45, P1.04, P2.28
Mari, A.
C47
Mochida, I.
B45, P1.04, P1.21, P2.28
Marinov, S.P.
B07, C03
Mokogwu, I.
B29
Maroño, M.
A50, A67
Moliner, R.
P2.04
Maroto-Valer, M.
A11
Mollah, M.M.
C29
Marshall, C.
B29
Mondragón, F.
A61, B04, B63, P2.30, P2.31
Marshall, M.
B02, B03, C29
Montagnaro, F.
B41
Martínez, M.
P2.13
Montiano, M.G.
P1.27
Martínez-Alonso, A.
P1.39, P1.40
Moon, S.H.
P1.08, P1.09
Martínez-Fernández, J.
A62
Moreno, N.
A42
Martínez-Tarazona, M.R.
A29, A32, A41, P2.06, P2.07
Moreno, T.
D09
Masset, P.J.
B09, C63
Moreno-Hirashi, J.A.
P1.02
Masui, M.
P2.02
Moretto, P.
B09
Mathews, J.P.
B01, B05, B06, B10, B11
Moritomi, H.
P2.01, P2.02
Matsuda, S.
P1.21
Morpeth, L.D.
B51
Matsushita, Y.
C46, C61, P1.21
Mulas, A.
A66
Melendi, S.
P1.17, P1.18
Mulas, G.
A66
Meloni, P.
A66
Muñoz, E.J.
P1.10
Mendiara, T.
A46
Murakami, K.
B66
Menéndez, R.
C08, C43, D02
Myers, K.
C22
N Nam, S.C.
P2.08
Niksa, S.
A30, B28, C21
Nariewicz, M.R.
B06
Ninomiya, Y.
B23
Naruse, I.
A51, D04
Nishiiri, Y.
A21
Navarrete, B.
A23
Nisi, B.
A10
Navarro, S.C.
P1.10
Nitzsche, R.
B47
Navarro-Lozano, J.O.
P1.02
Nomura, S.
C46
Nel, S.
B64
Norinaga, K.
C47, P2.38
Neomagus, H.W.J.P.
B25, B26, B64, B65, C09, P2.25, P2.27
North, B.C.
B42
Ng, K.W.
C45
Nuamah, A.
A28, B31
Nikrityuk, P.A.
A06
Nyathi, M.S.
D01
Oboirien, B.O.
B42
Olcese, R.
C04
O'Brien, G.
C44
Oliveira, M.L.S.
C62, P1.05
Ochoa-González, R.
A29, A32
Ooyatsu, N.
A01
Ohga, K.
A21
Osório, E.
P1.32, P2.22
Ohm, T.I.
P1.08, P1.09
Ossadchaya, E.F.
B52
Ohtaka, N.
P1.30
O'Sullivan, L.
C22
Ohtsuka, Y.
P1.30
Otero, J.
A67
Okolo, G.N.
B25
Otero, P.
A23
Okuyama, N.
C01, C48
Özlem Kocabas, Z.
P2.21
O
P Pan, C.
C32
Pesimberg, M.
D08
Pang, C.H.
C66
Petrik, L.F.
P2.20
Paniagua, J.
A65
Pevida, C.
A27, P2.12, P2.13
Park, J.I.
P2.28
Piedad-Sánchez,N.
P1.02
Park, S.R.
P2.37
Pierce, D.T.
A09
Paterson, N.
A26
Pis, J.J.
A27, P2.12, P2.13
Patrick, J.W.
C07
Plana, F.
D09
Patzschke, C.
B02
Pou, J.O.
B01
Pérez del Villar, L.
A10
Pourkashanian, M.
A27
Pérez-Ros, C.
A64
Powis, J.
B31
Qi, C.
C25
Querol, X.
A31, A42, B30, D09,
Qi, Y.
B02
Quignard, A.
C24
Qiu, B.
C26
Q
R Raeva, A.A.
A09
Rodríguez, E.
A41, P1.18
Raharjo, S.
A51
Rodríguez-Galán, M.
A62
Ranganathan, V.
B02
Rodríguez-Pérez, J.
A29, A41, P2.06
Rech, R.
P2.10, P2.14, P2.15
Romeo, L.M.
A42
Redaelli, C.
P2.10, P2.14
Rossiter, A.
C30
Reichel, D.
C64
Roy, B.
B48
Ren, S.
C12, C32
Rubiera, F.
A27, P2.12, P2.13
Rhim, Y.
P1.11, P2.34
Rufas, A.
P2.03
Riaza, J.
A27, P2.12
Ruiz, B.
P1.01, P1.06
Riley, G.
A28, B31
Ruiz, E.
A50, A67
Roberts, D.G.
B51, C02
Runstedtler, A.
A02
Rodrigo-Naharro, J.
A10
Russell, D.
B02
Rodrigues, R.
B67
Russig, S.
C02
Rodríguez Acevedo, E.
P2.32
Ryu, I.S.
P1.08
Šafářová, M.
P1.23, P1.34
Shishido, T.
C01, C48
Saha, C.
A45
Shoji, M.
C46
Sakai, K.
C01, C48
Shui, H.
C32
Sakimoto, N.
A21, C52
Shui, H.
C12
Sakurovs, R.
A22
Silva, L.F.O.
C62, P1.05
Salamanca, M.
B04
Simonyi, T.
B61
Salatino, P.
A47, B41
Siriwardane, R.V.
B61
Saldaña, R.
A10
Slavinskaya, N.
B52
Saloojee, F.
A07
Snape, C.E.
A03, A52, B29, C04, C65
Sánchez, A.
A61
Sohn, J.M.
P2.37
Sánchez, J.M.
A50, A67
Solimene, R.
A47
Sánchez-Biezma, A.
A65
Song, B.H.
P2.37
Sanna, A.
A11
Song, C.S.
B10, B11
Santamaría, A.
B04
Song, J.
B27
Santamaría, R.
C08, C43, D02
Song, J.H.
A04
Santiago-Carrasco, B.
P1.02
Sonoyama, N.
C42
S
Sasaoka, E.
B62
Spears, D.A.
B44
Sato, K.
D03
Spiegl, N.
A26
Sato, M.
B66
Spiro, B.
B29
Sato, R.
P2.38
Stańczyk, K.
P2.18
Scala, F.
A66
Steel, K.
P2.09
Schobert, H.H.
B10, B11, C09, C22, D01
Steele, D.
B28
Sciazko, M.
B08, B61
Stefanova, M.
B07
Scott, A.C.
B31
Stevens, L.
A52
Seam, W.S.
A09
Stokie, D.
C49
Secchi, A.R.
A63
Strydom, C.A.
C05
Segev, G.
P2.19
Sturgeon, D.W.
A24
Seifert, S.
B06
Suárez-Ruiz, I.
P1.01, P1.02, P1.06
Seng, T.X.
A45
Suelves, I.
P2.04
Senneca, O.
A47
Sugawara, K.
B66
Shan, C
C12
Sun, B.
P1.14
Sharma, A.
B50
Sun, C.
A03
Shim, H.M.
P2.37
Sun, Y.
B51
Shimada, S.
A21, P2.29
Suuberg, E.M.
B24
Shimogori, M.
A01
Sybring, M.
A02
Shirai, H.
A05, B22
T Takanohashi, T.
B50, C52
Tian, H.
B61
Takarada, T.
D03
Tian, L.
A43
Takarayama, N.
A01
Todoschuk, T.
C45
Takata, S.
P2.01
Tomás, C.
P1.01
Tamargo-Martínez, K.
P1.39, P1.40
Torchała, K.
P1.37
Tanosaki, T.
B23
Trierweiler, J.O.
B67
Tascón, J.M.D.
P1.39, P1.40
Trotman Gavilán, J.A.
P1.13
Tatarazako, N.
B23
Tsubouchi, N.
P1.30
Tay, H.L.
B49
U Uchida, A.
C46
Umemoto, S.
P2.39
Ueki, Y.
A51, D04
Ustimenko, A.B.
A08, B52
Uguna, C.
C65
V Valeš, J.
P1.23, P1.34
Vanegas-Chamorro, M.
P1.39, P1.40
van der Merwe, G.W.
P2.25
Vaselli, O.
A10
van Drooge, B.L.
D09
Velásquez, M.
B04
van Heerden, J.
C41
Velázquez, H.
P1.10
van Niekerk, D.
C41
Verheyen, V.
B02
van Staden, Y.
C05
Vilela, A.C.F.
P1.32, P2.22
Waanders, F.
C62, P1.05
Watson, J.C.
B01
Wada, N.
P1.21
Wei, C.
C32
Wagner, N.
A07
Wei, X.Y.
P1.14
Wakabayashi, N.
B22
Wei, Y.
B23
W
Walter, J.
C31
Weir, S.
A22
Wang, J.
A52
Wenying, L.
C25
Wang, S.
C30
Williams, A.
A27
Wang, X.
A11, A43
Wilson, R.B.
B28
Wang, Y.
C10
Winans, R.E.
B05, B06
Wang, Y.H.
P1.14
Wood, J.
A52
Wang, Z.
C12, C32
Worasuwannarak, N.
C50
Wannapeera, J.
C50
Wu, T.
C66, P1.31
Wasserman, S.
D08
P1.39, P1.40
Xu, G.
C10
Yamada, O.
P2.29
Yoon, Y.I.
P2.08
Yamazaki, Y.
C46
Yoshiie, R.
A51, D04
Yang, H.
A43
Yperman, J.
B07
Yasuda, H.
P2.29
Yu, J.
C49
Ye, C.
P1.31
Yue, X.M.
P1.14
Yoo, J.
P1.12, P2.34
Yürüm, Y.
B12, P2.21
Yoon, S.H.
B45, P1.04, P2.28
X Xiberta, J. Y
Z Zeng, C.
A43
Zheng, H.
P1.31
Zhang, J.
C10, C26, C30
Zhou, X.
P1.24
Zhang, S.
A43
Zhu, S.
P1.25
Zhang, X.
C46
Zhuang, X.
B30, D09
Zhao, D.
A03
Zong, Z.M.
P1.14
Zhao, Y.
C11
International Conference on Coal Science and Technology, Oviedo-Spain 2011 (ICCS&T 2011)
ORGANISING COMMITTEE Chair: Prof. Juan M. D. Tascón Technical Programme: Dr. Angeles G. Borrego Members Dr. Nicolás de Abajo Martínez Dr. Juan Otero Prof. Antonio Valero Prof. Vicente Cortés Mr. Francisco García-Peña Dr. Juan Carlos Ballesteros Prof. Jorge Loredo Prof. Carmen Clemente-Jul Mr. Fermín Corte Mr. Francisco Javier Alonso Ms. Yolanda Fernández Mr. Juan Ramón G. Secades Prof. Juan Adánez Dr. Rosa de Vidania Muñoz Mr. Alfonso Martínez Mr. Isaac Pola Prof. Ángel Linares-Solano
Co-chair: Dr. María A. Díez
ARCELORMITTAL CIEMAT CIRCE CIUDEN ELCOGAS ENDESA ETSIMO ETSIMinas-UPM FAEN Gas Natural Fenosa HC Energía HUNOSA ICB-CSIC IGME Industrial Química del Nalón SL Mining and Energy Office. Government of the Principality of Asturias University of Alicante
International Conference on Coal Science and Technology, Oviedo-Spain 2011 (ICCS&T 2011)
LOCAL ORGANISING COMMITTEE Dr. Mónica Alonso Carreño
Dr. Patricia Álvarez Rodríguez
Dr. Borja Arias Rozada
Dr. Carmen Barriocanal Rueda
Dr. Dolores Casal Banciella
Dr. Mercedes Díaz Somoano
Dr. Roberto García Fernández
Dr. Susana García López
Dr. Mª Victoria Gil Matellanes
Dr. Mª Antonia López Antón
Ms. Concha Prieto Alas
Ms. Juliana Sánchez Villar
International Conference on Coal Science and Technology, Oviedo-Spain 2011 (ICCS&T 2011)
INTERNATIONAL ORGANISING COMMITTEE Prof. Chun-Zhu Li
Australia
Dr. Fari Goodarzi
Canada
Dr. Tony Macphee
Canada
Prof. Wolfgang Klose
Germany
Dr. Robert Davidson
IEA CCC UK
Prof. Isao Mochida
Japan
Dr. Osamu Yamada
Japan
Prof. Jieshan Jason Qiu
P. R. China
Dr Johannes van Heerden
South Africa
Prof. Quentin Campbell
South Africa
Prof. Rosa Menéndez
Spain
Prof. Colin Snape
UK
Dr. Edward Lester
UK
Dr. Mildred Perry
USA
Dr. Jonathan Mathews
USA
International Conference on Coal Science and Technology, Oviedo-Spain 2011 (ICCS&T 2011)
TECHNICAL PROGRAMME Chairperson: Dr. Angeles G. Borrego COAL CHEMISTRY AND STRUCTURE Prof. Amelia Martínez-Alonso (INCAR-CSIC) COAL UPGRADING Prof. Carmen Clemente-Jul (ETSIMinas-UPM) COAL GEOLOGY AND ORGANIC PETROLOGY Dr. Isabel Suárez-Ruiz (INCAR-CSIC) Dr. Carlos I. Salvador (University of Oviedo) COAL COMBUSTION Dr. Luis Romeo (University of Zaragoza) Dr. Luis de Diego (ICB-CSIC) COAL GASIFICATION AND CLEAN FUELS Prof. José L.G. Fierro (ICP-CSIC) Prof. José L. Valverde (University of Castilla La Mancha) CLEAN COAL TECHNOLOGIES Dr. Francisco García-Labiano (ICB-CSIC) CO2 CAPTURE AND STORAGE Dr. Roberto Martínez (IGME) Dr. Carlos Abanades (INCAR-CSIC) COAL PYROLYSIS AND LIQUEFACTION Dr. María Antonia Diez (INCAR-CSIC) Dr. Vicente Cebolla (ICB-CSIC) COAL MINERALOGY AND COAL ASH Dr. Rosa Martínez-Tarazona (INCAR-CSIC) Dr. Diego Álvarez (INCAR-CSIC)
COAL AND THE ENVIRONMENT Dr. María Jesús Lázaro (ICB-CSIC) Prof. Xavier Querol (ICTJA-CSIC) CARBONS FROM COAL Dr. Marcos Granda (INCAR-CSIC) Prof. Rosa Menéndez (INCAR-CSIC)
International Conference on Coal Science and Technology, Oviedo-Spain 2011 (ICCS&T 2011)
INTERNATIONAL ADVISORY GROUP Dr. David Harris Dr. Richard Sakurovs Prof. Terry F. Wall Prof. Colin Ward Prof. Kechang Xie Prof. Fanor Mondragón Prof. Klaus Hein Prof. Kim Dam-Johansen Prof. Mikko Hupa Prof. Piero Salatino Prof. Kouichi Miura Prof. Takayuki Takarada Prof. Jacek Machnikowski Dr. Ibrahim Gulyurtlu Dr. Nicola Wagner Prof. Ana María Mastral Prof. Rafael Moliner Dr. Petra David Prof. Yuda Yürüm Prof. Rafael Kandiyoti Prof. John Patrick Prof. Alan Williams Prof. Ripudaman Malhotra Prof. Harold Schobert
Australia Australia Australia Australia China Colombia Germany Denmark Finnland Italy Japan Japan Poland Portugal South Africa Spain Spain The Netherlands Turkey UK UK UK USA USA
International Conference on Coal Science and Technology, Oviedo-Spain 2011 (ICCS&T 2011)
SPONSORED BY Science and Innovation Ministry
Grupo Hunosa
Science, Technology and Innovation Research Programme Asturias-Spain
HC Energía
Oviedo Council
Industrial Química del Nalón S.A. NalonChem
Ciudad de la Energía, Ciuden
International Energy Agency
Spanish National Research Council
LECO
Gas Natural Fenosa
KEYNOTE PAPER THERMAL BREAKDOWN IN MIDDLE RANK COALS* Rafael Kandiyoti Department of Chemical Engineering, Imperial College London, Prince Consort Road, London SW7 2AZ, UK
1. Introduction The past decade has seen a sharp turn away from research on coal utilization, in favour of work on the mitigation of its environmental consequences. Coal utilization has potential to pollute and activities aiming to clean up past pollution, as well as mitigating the environmental consequences of future utilization are necessary. In my opinion, however, the emphasis given to CO2 capture and sequestration needs to be critically re-evaluated. Statistics show that worldwide coal production and utilization is vast and expanding. The need to support a component of research aiming to make utilization more efficient and less polluting is self evident. And as ever, we need to know more about the structure of coals and to further develop methods for their utilization. In this report, I aim to bring together experimental information in a manner that might help improve our understanding of thermal breakdown in middle rank coals. My purpose is to sift through and collate information, rather than to innovate. The first line of approach will be to compare thermal breakdown phenomena observed during pyrolysis and liquefaction. The initial questions that need answers are: (1) When does thermal breakdown actually begin? (2) Are there similarities between reaction pathways in pyrolysis and liquefaction? (3) When and how do these reaction pathways begin to diverge? As allied questions we will explore (4) Why and when pyrolysis product distributions are affected by heating rates, and, (5) How retrogressive re-combination reactions work. These are questions relevant to improve our understanding of the fundamentals of most coal utilization activities: combustion, gasification, liquefaction and coking. 2. Electron spin resonance spectrometry of thermal breakdown Middle rank coals normally contain of the order of ~1019 free radicals per gram. The relative reactivities of free radicals depend on whether and how the host structure allows delocalization of unpaired electrons. The ESR spectrum of a coal reflects the stable free radical population embedded in the coal matrix. During and after thermal treatment, we observe additional unpaired electrons left over from completed processes. With appropriate corrections, the change in spin population as a function of the temperature may be calculated. However, observing reactive free radicals in coals by ESR is difficult, due to their short lifetimes and their low concentrations relative to much higher stable free radical concentrations [1]. As with ordinary pyrolysis experiments, acquired ESR spectra of pyrolysing samples reflect, to some degree, the configuration of the sample and the experimental design [2]. Figure 1 presents a schematic diagram of typical spin population vs. temperature curves, observed when coal samples are heated in a quartz fixed-bed reactor fitted with a gas sweep facility and placed inside the cavity of an ESR spectrometer. Spin populations (S) have been defined as free radicals per gram of initial sample. Three distinct types of thermally induced processes have been identified. *
This review was developed using material first presented at the BCURA “Coal Science Lecture” of 2006. The experimental results discussed may be found in Ref [9(a)].
Region I (T < T1): S increases to a relatively shallow maximum (near 200°C), due to the recovery of signal through desorption of gases adsorbed on sample surfaces. These gases, primarily moisture and oxygen, would have been adsorbed by previous exposure to air. Region II (T2 > T > T1): S decreases to a minimum (near 300°C). This decline is thought to be associated with recombination reactions, resulting from the thermally induced mobility of occluded material with sufficient reactivity. Region III (T > T2): Above T2, S increases monotonically with rising temperature, signalling an increase in the free-radical population, as the coal pyrolyzes within the ESR cavity. T2 is thought to mark the onset of covalent bond cleavage reactions. For the range of coals tested, T2 increased from 310ºC for a lignite, to about 340ºC for higher rank coals (Table 1; also cf. [3]).
Figure 1. Schematic diagram showing main characteristics of spin population vs. temperature diagrams [3]. (Reproduced with permission: Carbon 1989, 27, 197; Copyright 1989 Elsevier) 19
Table 1 ESR parameters of coals given as “spin populations x 10 ” in the flow cell [3].
In coal liquefaction, extract yields increase and molecular mass distributions get broader as the temperature reaches 375-400°C [4, 5]. This range is well above the 310 – 340 °C interval signalled for the onset of covalent bond cleavage, suggesting that several bonds must rupture before large molecular fragments are released from the solid matrix. Figure 2 shows that the rate of dissolution accelerates with increasing temperature from about 375 °C onwards. Point of Ayr (UK) coal shows this more graphically than many other coals.
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Figure 2 Sample weight loss from Point of Ayr and Pittsburgh No. 8 coals as a function of temperature in the flowing -1 solvent reactor. Samples were heated at 5°C s to 450°C with 400 s hold. Tetralin flow rate: 0.9 ml/s at 70 bar [6]. (Reproduced with permission: Energy & Fuels 1996, 10, 1115; Copyright 1996 Am.Chem.Soc.)
Liquefaction thus allows recovering large amounts of coal as extract at temperatures near 400 °C, provided a “good” solvent for coal derived materials is used. In pyrolysis, covalent bond cleavage patterns appear to follow similar paths; however, much of the “extractable” material that detaches from the solid matrix remains within the coal particles and, at 400 °C, sample weight loss remains low (about 5 % in Figure 3).
Figure 4 The atmospheric pressure wire-mesh reactor with the early (a) and present (b) tar trap designs. Legend: [1] Copper Current Carrier; [2] Live Electrode; [3] Brass Clamping Bar; [4] Sample Holder Support Plate; [5] Mica Strip; [6] Wire-mesh Sample Holder; [7] Electrode; [8] Stainless Steel Tubes; [9] Mica Layer; [10] Brass Pillars; [11] Sintered Pyrex Glass Disk; [12] Base Plate; [13] Pyrex Bell; [14] O-ring Seal; [15] Off-take Column; [16] O-ring; [17] Carrier Gas Entry Port; [18] Connection for Vacuum Pump. (Reproduced with permission: Fuel 68, (1989), 895; Copyright 1989 Elsevier).
Figure 3 Effect of peak temperature on pyrolysis tar and total -1 volatile yields. Heating rate: 1,000°C s . Linby coal. 30 s holding at peak temperature; sweep gas, helium at 1.2 bar -1 flowing at 0.1–0.2 m s . Particle size range: 106–152 μm [7]. (Reproduced with permission: Fuel 1989, 68, 895; Copyright 1989 Elsevier).
Initial conclusions: Up to about 375 – 400 °C, no significant divergences were observed between reaction pathways during pyrolysis and liquefaction. This is understood in terms bond rupture being primarily a
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function of the temperature. Pathways diverged during and after “solvent extractable” tar precursors broke off from the coal matrix and gradually accumulated within the coal particles. In liquefaction, increasing amounts of material, soluble in the solvent used, could be removed from the coal particles. In pyrolysis, however, little of the material released into the particles could exit into the gaseous environment surrounding the particle. The “extractables” (tar precursors) mostly remained within the coal particles. The next task was to trace the fate of “extractable” material residing within pyrolysing coal particles, as the temperature was raised. It was also relevant to explore why product distributions change when the heating rate is increased. 3. The fate of “extractables” accumulated in coal particles: pre-pyrolysis temperatures Figure 4 presents a schematic diagram of the wire-mesh reactor used in experiments described below. The apparatus is best known through seminal work by Howard and co-workers at MIT [8]. The reactor has a useful configuration for approximating single particle behaviour. Stages in the evolution of its design in various laboratories have been described elsewhere [9(b)]. The atmospheric pressure version shown in Figure 4 is capable of variable heating rates between 1°C s-1 and 10,000°C s-1 and temperatures up to 2,000°C. In early work, the reactor enabled observing changes in product distributions with increasing heating rate (Figure 5; [7]). In Figure 5, the lines tracing tar and total volatile yields very nearly rise in parallel with increasing heating rate. The weakly-coking Linby coal showed plastic behaviour only when heated rapidly; progressively more extensive plasticity was observed as the heating rate was increased [7, 10]. We will see below that tar yields and plasticity are related to the amount of extractable material (tar precursors) that accumulates within coal particles during heat-up.
Figure 5 Effect of heating rate on tar and total volatile yields. Peak temperature of 700°C. Linby coal. 30 s holding at peak -1 temperature; sweep gas, helium at 1.2 bar flowing at 0.1–0.2 m s . Particle size range: 106–152 μm [7]. (Reproduced with permission: Fuel 1989, 68, 895; Copyright 1989 Elsevier).
In the iron making industry, these and similar observations have led to an interesting pilot application. At Nippon Steel Corporation, crushed coal was rapidly pre-heated in a riser to about 400ºC. The sticky mass of particles was then slowly heated in a retort to 800-900ºC. The effect was to form a stronger coke than would have otherwise been possible, using the same coal (or blend) heated slowly from ambient temperature [11].
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The procedure was found to be effective for improving the coking properties of weakly-coking coals. For prime coking coals, the initial rapid heating step provided no significant improvement in the amount or strength of product coke. Table 2 Characteristics of three Australian coal blends used in the study.
Coal A
Coal B
Coal C
Volatile Matter, % Fixed Carbon, % Ash, % Crucible Swelling No.
35.2 52.4 9.0 3.5
24.1 65.1 9.3 6.5
17.9 72.3 8.9 8.0
C (%,db) H (%,db) S (%,db) N (%,db) O (%,db)
83.6 5.6 0.55 1.8 8.3
87.7 5.0 0.57 1.7 4.8
90.7 4.6 0.15 0.8 3.3
4. Examining pre-pyrolysis phenomena in coals The accumulation of extractable material within coal particles prior to full blown pyrolysis was used as a diagnostic tool, in order to explore how the plastic properties of weakly coking coals may be improved by faster heating [12]. Part of the answer was known from previous work. In the 1960ies, Brown and Waters had shown that coking properties correlated well with amounts of chloroform extractable material found in coals heated (slowly) to between 300 and 400ºC [13]. The experiments we conducted aimed to explore relationships between (i) “extractable” contents of heated particles, (ii) the temperature of exposure and (iii) the heating rate. Three coal blends from the pilot study by Nippon Steel were used. Weakly coking Newcastle Blend Coal was labelled as “Coal A”, strongly coking Goonyella as “Coal B” and the “very” strongly coking K-9 Blend as “Coal C”. Table 2 presents some of the properties of the three samples. The initial set of experiments followed one of two sequences: Sequence I: (1) Fast heating (1,000ºC s-1) to 400ºC, (2) 30 seconds holding at 400ºC, (3) Slow (1ºC s-1) heating to an intended temperature, between 400 and 500ºC, (4) 30 seconds holding at the intended temperature, followed by cooling and extraction with NMP (1-methyl-2-pyrrolydinone) [12]. Sequence II: Differed from Sequence I by heating “slowly” to 400ºC (at 1ºC s-1) during Step 1. Figure 6 shows the differences in the amounts of extractable material recovered from Coal A particles heated to 400ºC rapidly (1,000ºC s-1; ~66% extract), compared to slow heating (1ºC s-1; about 33 % extract) in Step 1. The extract yield from untreated Coal A was about 35 %. As the temperature was then raised at 1ºC s-1 from 400ºC, the amount of extractables within the initially fast heated particles increased slowly to about 80 % near 475ºC before declining rapidly. These results give rise to two follow-up questions. First, is there a characteristic temperature which must be reached, before differences between fast and slow heating become apparent? Second, how fast does “fast heating” have to be for differences between extractable yields begin to emerge?
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Extract Yield (mass%,daf)
90 80 70 60 50 40 30 20
rapid400+slow slow
10 0 400
450 Temperature (° C )
500
Figure 6. Comparison of NMP-extractables recovered from Coal A particles treated by one of two heating sequences: Fast (1,000ºC s-1) or slow (1ºC s-1) heating to 400ºC in Step 1 [12]. (Reproduced with permission from Energy & Fuels 2004, 18, 1140; Copyright Amer. Chem. Soc 2004).
To answer the first question, Coal A particles were heated at 1,000ºC s-1 to temperatures between 350 and 400ºC, followed by cooling and extraction with NMP. Figure 7 shows a sharp transition to greater “extractable” accumulation after 370ºC. Near 400ºC, the extractable yields rose towards the same point (~65 %) reported in Figure 6, showing satisfactory internal consistency. The faster rise in extractables above 370ºC was consistent with the temperature for the onset of massive breakdown inferred from ESR data and liquefaction experiments in Figure 2. Recalling that results from ESR spectroscopy indicated the onset of individual bond rupture in the 310-340ºC range, data in Figure 7 appear consistent with the proposition that several bonds need to break before larger molecular mass material might detach from the solid matrix.
Extract Yield (mass%,daf)
80 70 60 50 40 30 20 10 0 350
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370 380 Temperature (° C)
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400
Figure 7 Extractables accumulating in particles of Coal A during “fast” heating to between 350 and 400ºC. (Reproduced with permission from Energy & Fuels 2004, 18, 1140; Copyright Amer. Chem. Soc 2004)
The next point is perhaps more difficult to explain. The sharper increase in extract accumulation above 370ºC is not observed when slow heating (1ºC s-1) is applied. In general, bond cleavage is known to be mainly a
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function of the temperature. The lower proportion of extractables recovered after the slow heating experiment suggests that more retrogressive recombination reactions occurred during slow heating. Furthermore, data in Figure 8 show that internally released extractables, which had survived during heat-up to 400°C at either heating rate, were stable at 400°C for as long as 120 s. Thus, in contrast to the extractable loss during heat-up at slow heating rates, the retrogressive repolymerization reactions between extractables that survive to 400°C proceed through far slower recombination reactions, if at all. Figure 6 shows that the temperature needs to rise to above 450°C for recombination reactions within the residual “extractable” mass to lead to significant char formation. No clear char formation trend was observed within the particles heated initially at 1ºC s-1, which seem surprisingly inert between 400 and 500ºC.
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Figure 8. Effect of holding time at 400°C on NMP-extract yields for Coal A samples heated at 1 and 1,000°C s . (Reproduced with permission from Energy & Fuels 2004, 18, 1140;Copyright Amer. Chem. Soc 2004)
These data provide evidence that the more reactive free radicals within the “extractables” recombine during slow heat-up between 370 – 400°C and that during rapid heating, these recombination reactions are blocked. Below, we will suggest that when fast heating is applied, internally released hydrogen, native to the coal, may be entering the reaction mixture to quench the more reactive free radicals, blocking at least a part of the potential retrogressive recombination reactions. The larger pool of extractables (tar precursors) contained in rapidly heated particles goes some way towards explaining the higher tar yields from some middle-rank coals during rapid heating. However, tests on many coals show that the maximum difference in tar yields measured (between slow and fast heating) is in the range of 4 – 8 %, compared to a 30 % difference in extractable (tar precursor) accumulation. It appears therefore that char formation rather than tar ejection remains the dominant reaction route when temperatures are raised above 600ºC. The stability of extractable materials in coal particles for up to 2 minutes (and probably beyond, had the experiment been continued; see Figure 8) is consistent with data from Fong et al. [14]. These authors reported a depletion rate for pyridine extractables for the higher temperature interval of 600 – 800°C, characterized by the reaction rate constant k = 1.9 x 1010 exp (-21,200/T) (s-1). The depletion rate at 400°C of “extractable” materials calculated using this equation is almost negligible. It also appears that depletion by volatilization of the extractables is not a significant factor at 400°C. The possibly relatively small weight loss through devolatilization during the first few seconds after reaching 400°C appears to be made up by new release of “extractable” material.
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These data lead indicate that rapid recombination reactions during heat-up take place between more reactive free radicals, unless they are somehow blocked. Meanwhile, less reactive free radicals in the extractable mass surviving to 400°C (Figure 6), produce char at far slower rates at temperatures above 400°C (Figure 6).
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Figure 9. Relationship between heating rate and NMP-extract yield; heating rate: 1,000°C s . (Reproduced with permission from Energy & Fuels 2004, 18, 1140; Copyright Amer. Chem. Soc 2004)
5. The heating rate and extractables accumulation in weakly/strongly coking coal particles Experiments were carried out by heating sample particles to 400°C at increasing rates, followed by extraction with NMP. Figure 9 shows how the amounts of extractable material accumulated in heated particles increased as a function of the heating rate. The transition above 500C s-1 was sharp and repeatable. In response to the two questions raised at the bottom of page 5, we found that the sample must be heated to ≥ 375ºC before large amounts of extractable may be released and the effective “fast” heating rates lie above 500 – 1,000ºC s-1 for this coal. The enhanced formation of NMP-extractables is thus observed to be a direct function of the heating rate. Whilst we should be hard put to explain why the particular heating rate threshold value of 500 – 1,000ºC s-1 turns out to be the critical one, it seems sufficient for present purposes to note that a “high” heating rate is required for recovering larger amounts of extractables from Coal A particles. We have meanwhile arrived at a working explanation for observations at Nippon Steel. We know from previous work that coal plasticity and extractable content are linked. Work at MIT has also shown (for temperatures above 600ºC) that minimum viscosity and maximum pyridine-extractable contents correlate [14]. Prior characterization work had already shown that the temperature of maximum extractables of Coal A was near its temperature of maximum thermo-plasticity, between 400 and 430ºC [11]. Finally, the stability of extractables during at least two minutes (Figure 8) appears to allow sufficient time for the particles in plastic state to form coherent lumps within the retort. Rapid heating thus leads to improved coke strength via the related increase in thermo-plasticity of Coal A. However, we still need to explain why greater amounts of extractable material are recovered from heated particles after “fast” heating. To explore this further, let us recall that “Coal A” was a weakly coking coal. How do “good” coking coals behave in analogous experiments?
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rapid400+slow
100
slow
90
rapid350
80
untreated coal
Extract Yield (mass%, daf)
Extract Yield (mass%, daf)
100
70 60 50 40 30 20 10 0
1000°C/s 1°C/s
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70 60 50 40 30 20 10 0
0
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200 300 400 500 Temperature (°C)
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(a) (b) Figure 10. Relationship between NMP-extract yield and temperature for samples of Goonyella (Coal B) and K-9 (Coal C), -1 both coking coals, heated at 1 and 1,000°C s to 400°C, followed by heating at 1°C to a variable peak temperatures up to 600°C (30 s holding at peak temperature). Carbon contents and crucible swelling numbers are given in Table 2. (Reproduced with permission from Energy & Fuels 2004, 18, 1140; Copyright Amer. Chem. Soc 2004)
Figure 10 shows that prime coking coals, Coal B and Coal C, show marginal heating rate sensitivities with regard to the accumulation of extractables. When heated, samples from the two coals gave high extract yields and became fluid, irrespective of the heating rate. High heating rates thus appear to improve the plastic behaviour of only weakly coking coals. Higher tar yields during fast heating was previously explained by the explosive ejection of tar precursors (i.e. “extractables”). In this sense, rapid heating is seen as reducing the probability of repolymerization reactions of tar-precursors through their rapid removal [15]. What we observed in our work was that during fast heating, a larger pool of tar precursors develops within the particles of coals that have marginal coking ability. The two elements are internally consistent and provide a possible contributing factor to increased tar yields observed at higher heating rates (Figure 5). The question that remains to be answered is how the larger pool of “extractable” material came about when particles of Coal A, the marginally coking coal, were heated rapidly. 6. Effect of rapid heating: Possible mechanistic explanations It is widely known that when middle rank coals are heated, small amounts of hydrogen are released from about ~ 285 - 300ºC [e.g. cf. 16]. There is also a “consensus” view from previous work [13, 17], that before tar evaporation, pyrolysis works as an internal liquefaction process, where free radicals are quenched and stabilised by internally released hydrogen. Fast heating telescopes the sequence of events into a narrower time frame and shifts upwards the temperature scale of successive pyrolytic events. Thus, any internally released hydrogen is more likely to remain in contact with the pyrolysing mass up to higher temperatures during “fast” heating (say 1,000ºC s-1 or faster) compared with “slow” heating. The internally released hydrogen is thus more likely to react with some of the free radicals and block (quench) some recombination reactions. We have no direct proof for this; however, this hypothesis would explain data showing Coal A retain more extractable material during heat-up at 1,000ºC s-1 compared to 1ºC s-1. This explanation also lends us a language by which we can begin to explain the behaviour of coals like Goonyella (above) or indeed the behaviour of H2-rich species such as liptinites, which show little sensitivity to heating rate [11(b), 18, 19]. The latter samples present relatively high elemental H2 and relatively low
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elemental O2-contents, evidently providing a combination capable of swamping the “internal liquefaction process” with sufficient hydrogen to block some of the rapid recombination reactions, irrespective of the heating rate. By comparison, marginally hydrogen deficient vitrinites show more pronounced heating rate sensitivity during pyrolysis, with respect to both plastic behaviour and tar yields [11(b), 18, 19]. 6.1 Tar yields and (alkyl+hydroaromatic) structures in coals: Figure 11 presents data for a rank ordered (and otherwise randomly selected) set of Northern Hemisphere coals, showing reasonably smooth trends of increasing tar yields with increasing “alkyl” content. Both sets of data lump together signal from alkyl and hydroaromatic (alicyclic) structures.
Figure 11. (a) Pyrolysis tar and total volatile yields as function of FT-ir-derived aliphatic:aromatic hydrogen ratios. Rank-1 ordered series of Northern Hemisphere coals: + tar, Δ total volatile. Pyrolysis in atm. pressure helium at 1,000 ºC s to 700°C with 30 s holding time. Non-melting coals: Taff Merthyr, Emil Mayrisch, Tilmanstone; melting coals: Heinrich Robert, Santa Barbara,Longannet, Candin, Bentinck, Thoresby, Gelding, Linby, Illinois No.6. (b) Pyrolysis tar yields as a function 13 of C-nmr-derived 15–37 nm aliphatic band intensity for a rank-ordered series of Northern Hemisphere coals. Coals and pyrolysis conditions as in (a). (Reproduced with permission: Fuel 1994, 73, 851; Copyright 1994 Elsevier).
It is widely thought that H-donor ability during the “internal liquefaction process” resides in the hydroaromatic component of coals [13, 17]. However, we do not have a reliable, rapid method for determining hydroaromatic contents in coals. It thus seems reasonable to conclude, but difficult to prove, that increased tar and total volatile yields in Figure 11 are due to progressively improving internal hydrogen transfer from the gradually increasing hydroaromatic component, with increasing coal rank. What we can show, on the other hand, is that the straight chain alkane molecule, hexadecane, contributes little to the coal liquefaction process, either in terms of solvent power or of H-donor activity. 6.2 Differentiating between the effects of alkyl and hydroaromatic structures: Figure 12 presents a schematic diagram of a “flowing solvent” coal liquefaction reactor, which allows products released from the fixed bed of coal to be swept away by solvent pumped through the system [6]. This reactor configuration limits the residence times of coal derived products in the reaction zone. Table 3 presents data from experiments performed with samples of Point of Ayr coal (UK), liquefied in (1) hydrogen donor solvent tetralin, (2) a mixture of non-H-donor compounds quinoline and phenanthrene, which are known as strong solvents for coal derived materials, and (3) hexadecane, a straight chain alkane, which is neither a good solvent for coal derived materials, nor an H-donor solvent. Results were compared with pyrolysis experiments under comparable conditions.
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PRESSURE REGULATOR PURGE SOLVENT RESERVOIR
PRESSURE RELIEF VALVE (AUTO) (MANUAL)
N2 STEPPER MOTOR C/W OUT
COAL/SAND BED FLOW METER
C/W IN SURGE CHECK VALVE
COOLING WATER HEAT EXCHANGER
DRAIN VALVE
FILTER DRAIN VALVE
REACTOR PRESSURE GAUGE
SOLVENT RESERVOIR
Figure 12. Schematic diagram of the flowing solvent reactor system. Solvent is forced through the fixed bed of coal and sweeps dissolved product away from the reaction zone and into the heat exchanger. The letdown valve is attached to a computer controlled stepper motor and serves to control the flow rate [6]. (Reproduced with permission: Energy & Fuels 1996, 10, 1115; Copyright 1996 Am.Chem.Soc.). -1
Table 3 Comparing conversions in different solvents [20]: Liquefaction experiments with solvent flow rate of 0.9 mL s at 70 bar (g). All data are given on % w/w dry ash free basis. (Reproduced with permission: Copyright 2006 Elsevier).
The positive contribution made by tetralin in the dissolution of the samples showed what we already knew, that the hydroaromatic component performs an H-donor function during coal thermal breakdown. Furthermore, when using n-C16H34 (hexadecane) as the liquefaction medium, the absence of both H-donor activity and solvent power was evident. The data show that coal conversion in hexadecane was comparable to the conversion observed in helium. 7. More/less reactive free radicals: fast/slow recombination reactions Table 3 shows that up to 350 °C, more coal derived material could be removed by quinoline or the mixture of quinoline and phenanthrene, compared to tetralin. Coal sample weight loss at up to 350°C is likely to represent (mostly) the extraction of already soluble material native to the original coal, rather than material made soluble through thermal breakdown. These data confirm that quinoline and phenanthrene are stronger solvents for coal derived materials compared to tetralin.
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7.1 Rapid retrogressive reactions in liquefaction: Table 3 also shows that, when the temperature was raised from 350 to 450°C, the proportion of coal sample dissolved and removed from the reaction zone increased significantly in the presence of both tetralin and the non-donor solvents. In closed (batch) reactors, conversions in non-H-donor materials decline at longer residence times. In this instance, the high conversions in the non-donor solvents can be explained in terms of the short residence times of dissolved products in the reaction zone (less than 10 s), followed by rapid cooling of products and the high product dilution ratio: about 150 mg coal-derived material in more than a litre of solvent. However, the systematically higher conversion to soluble material at 450 °C in the presence of tetralin, compared to the non-donor “strong” solvents requires explanation. The reversal of the trend, compared with results at 350°C, is consistent with the H-donor ability of tetralin being better able to block rapid retrogressive reactions during heat-up, compared to the non-donor solvents that are unable to provide this function. In other words, data in Table 3 suggest that, in the absence of H-donor ability, rapid retrogressive reactions occur with greater frequency during heat-up. The analogy with the “internal liquefaction process” posited for coal pyrolysis is clear. 7.2 Slow retrogressive reactions in liquefaction: Figure 13 compares conversions of Point of Ayr (UK) coal in two different reactor configurations: the “flowing-solvent” reactor and a small “mini-bomb” batch reactor. During initial experiments, using a tetralin-to-coal ratio of about 4:1 in the batch reactor, only minor differences were observed with conversions observed in the “flowing-solvent” reactor [21]. However, it took longer, about 1-hour, in the batch reactor to reach the level of conversion, which the “flowing-solvent” reactor achieved in several minutes. The difference appears due to slower diffusion of extracts out of coal particles in the batch reactor, caused by the gradually increasing extract concentration in the fluid surrounding the particles.
Figure 13 Flowing Solvent Reactor & Mini-Bomb reactor comparison using 1-methylnaphthalene as the liquefaction medium. Solvent/coal ratio in mini-bomb reactor: 4/1 by weight. Flowing-Solvent Reactor: Heating at 5°C s-1; solvent flow -1 rate: 0.9 mL s at 70 bars [21]. (Reproduced with permission: Fuel 1991, 70, 380; Copyright 1991 Elsevier).
In subsequent experiments, a non-donor solvent, 1-methylnaphthalene, was used as the liquid medium. Comparing conversions from the two reactors at 400 and 450°C showed that, in the “flowing-solvent” reactor, conversions increased with contact time at both 400 and 450°C. The differences between the two conversion lines for 400 ºC reflect conversion loss due to rapid recombination reactions in the batch reactor, where products cannot be moved away from the reaction zone before the termination of the experiment [cf. Ref 21].
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However, at 450°C, Figure 13 shows clear evidence for retrogressive char forming reactions in the mini-bomb reactor. At contact times longer than 100 s, conversions diminished and residual solids increased. In fact, there is nothing unusual about liquefaction with non-donor solvents in batch reactors giving yields that trace a maximum and eventually decline. This aspect of the data is consistent with trends observed during earlier work [e.g. cf. 22]. The relevance of these results in the present context is the demonstration of slow, indeed very slow, char forming reactions. Figure 13 showed a loss of 12 – 13% soluble material in about 25 minutes (1500 s). These slow char forming reactions contrast sharply with inferred rapid rates of retrogressive reactions during heat-up in non-donor solvents. The latter are only apparent when conversions are compared (i) with liquefaction in a donor solvent and (ii) when compared with short contact time yields in the flowing solvent reactor. 8. Conclusions: Overview of coal thermal breakdown Initial stages: Similar effects are observed during the initial stages of coal pyrolysis and liquefaction. This may be understood in terms of bond rupture being a function of the temperature alone. The onset of covalent bond cleavage, observed by ESR, was near 310 °C for a lignite and rose to 340 °C with increasing coal rank. Massive thermal breakdown in middle rank coals was observed near 375 ºC. This temperature is consistent with several bonds rupturing before larger molecular mass materials detach from the solid matrix and are released within the coal particles. Reaction pathways in pyrolysis and liquefaction begin to diverge during and after solvent extractable materials accumulate within coal particles at 375 - 400 ºC. In liquefaction, “extractables” may be removed from coal particles using a solvent. During the early stages of pyrolysis, on the other hand, accumulated “extractables” mostly remain inside the coal particles, as the temperature rises. Larger contents of extractable materials tend to improve the plasticity of coal particles. Greater tar yields correlate well with larger pre-pyrolysis extractable (tar precursor) contents and the greater plasticity of coal particles. Likely mechanisms: As coal particles are heated, internally released hydrogen is thought to quench reactive free radicals and effectively block some of the rapid recombination reactions. Fast heating tends to telescope the sequence of events into a narrower time frame and to shift the temperature scale of pyrolytic events to higher temperatures. Hydrogen released from pyrolysing solids from about ~ 285 - 300ºC is thought to remain in contact within the pyrolysing mass, as higher temperatures are reached more rapidly. Donatable hydrogen in coals is thought to reside in the hydroaromatic component. Internally released donatable hydrogen appears to play an analogous role during pyrolysis to donor solvents (e.g. tetralin) during liquefaction. In both cases, the availability of donatable hydrogen appears to block potentially rapid free radical recombination reactions. Weakly coking coals are thought to be marginally deficient in hydrogen. More “extractables” accumulate and greater plasticity is observed during the fast heating of these coals. The effect is observed at heating rates greater than 500ºC s-1. Meanwhile, “good” coking coals have sufficiently high donatable hydrogen (and correspondingly low hydrogen-scavenging oxygen) contents to show plastic behaviour irrespective of the heating rate.
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Fast and slow recombination reactions: The evidence presented suggests that, as thermally induced bondcleavage accelerates between 375 and 400⁰C, the more reactive of the free radicals show a tendency to recombine. The outcomes of these reactions are determined by the relative abundance of locally available “internally released” hydrogen. During slow heating (~1⁰C s-1), internally released hydrogen may escape before effective reaction temperatures are reached. Thus rapid recombination reactions taking place during slow heating produce more char and less “extractable” material compared to rapid heating; the effect is particularly observable in coals that are marginally deficient in hydrogen. Similarly, during liquefaction in the flowing solvent reactor, the use of a “strong” but non-H-donor solvent allows more repolymerization reactions to take place during heat-up, reducing the extract yield by about 5 – 8 %, compared to liquefaction in tetralin. In pyrolysis, the extractable mass that survives the heat-up stage to 400ºC and accumulates within coal particles was observed to be relatively stable and to convert to char only slowly above 450⁰C. Similarly, during liquefaction in non-H-donor solvents in batch reactors, already dissolved extracts were observed to form solids at 450⁰C, at a very slow rate (12–13% in 25 minutes.). These reactions appear distinct from, and slower than, rapid recombination reactions observed during heat-up. These observations provide evidence for wide ranges of free radical reactivities and wide ranges of rates for radical recombination reactions. Acknowledgements The work reported in this paper was supported by UK SRC (and its successor organizations SERC, EPSRC), British Coal, the European Coal and Steel Community, the European Union and the Nippon Steel Corporation. Of the many colleagues who contributed to this study, I am only able to thank a few by name in the space available: Tim Fowler, Keith Bartle, Jon Gibbins, Li Chunzhu, Alec Gaines, Alan Herod, Geoff Kimber, Şebnem Madralı, Xu Bin and Koichi Fukuda. Many others contributed with discussions and criticism. Trevor Morgan read a late draft, but I take responsability for my errors. References [1] [2] [3] [4] [5] [6] [7] [8] [9] [10] [11] [12] [13] [14] [15] [16] [17] [18] [19] [20] [21] [22]
Fowler, T.G., Bartle, K.D. & Kandiyoti, R., Energy and Fuels 1989;3; 515-522. Fowler, T.G., Bartle, K.D. and Kandiyoti, R., Carbon 1987; 25; 709-715. Fowler, T.G., Kandiyoti, R., Bartle, K.D. and Snape, C.E., Carbon 1989;27;197-208. Xu, B., Madrali, E.S., Wu, F., Li, C-Z., Herod, A.A. and Kandiyoti, R., Energy & Fuels 1994; 8; 1360-1369. Li, C-Z., Wu, F., Xu, B. and Kandiyoti, R., Fuel 1995; 74; 37-45. Xu, B. and Kandiyoti, R., Energy and Fuels 1996;10;1115-1127. Gibbins, J.R. and Kandiyoti, R., Fuel 1989;68;895-903. Howard, J. B. (1981) Chemistry of Coal Utilization Second Supplementary Volume (ed. Elliott, M. M.), Wiley, NY, Chapter 12 Kandiyoti, R., Herod, A.A. and Bartle, K.D., “Solid Fuels and Heavy Hydrocarbon Liquids: Thermal Characterization and Analysis” Elsevier Science Pub. (2006), Amsterdam Oxford London New York. (a) p.199; (b) p. 43. Gibbins-Matham, J.R. and Kandiyoti, R., Energy & Fuels 1988;2;505-511 Aramaki, T.; Arima, T.; Yamashita, Y.; Inaba, A.; Tetsu to Hagane 1996; 82; 5-34 Fukuda, K., Dugwell, D.R., Herod, A.A. and Kandiyoti, R., Energy & Fuels;2004;18;1140-1148. Brown & Waters: Fuel 1966; 45; 17; Brown, H. R., Waters, P. L. Fuel 1966; 45; 41 Fong, W. S., Khalil Y. F., Peters W. A., Howard J. B. Fuel 1986; 65;195 Gray V.R., Fuel 1988; 67;1298 Neuburg, H.J., Kandiyoti, R., O'Brien, R.J., Fowler, T.G. and Bartle, K.D., Fuel 1987;66;486-492. Neavel, R. C. Coal Science, Academic Press, 1981, 1-19 Li, C-Z., Madrali, E.S., Wu, F., Xu, B., Cai, H-Y., Güell, A.J. and Kandiyoti, R., Fuel 1994;73;851-865. Li, C-Z., Bartle, K.D. and Kandiyoti, R., Fuel 1993; 72;3-11. Xu, B., Madrali, E.S., Wu, F., Li, C-Z., Herod, A.A. and Kandiyoti, R, Energy & Fuels 1994;8; 1360-1369. Gibbins, J.R., Kimber, G., Gaines, A.F., and Kandiyoti, R., Fuel 1991;70;380-385. Clarke, J.W., Kimber, G.M., Rantell, T.D. and Shipley, D.E. (1980) ACS Symp. Ser. No. 139, D.D. Whitehurst, Editor, 111
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Oviedo ICCS&T 2011. Extended Abstract
Ash deposition characteristics determined in pilot plant tests burning bituminous and sub-bituminous coals
Miki Shimogori, Noriyuki Ooyatsu, Noboru Takarayama, Toshihiko Mine Kure Research Laboratory, Babcock-Hitachi K.K., 5-3 Takaramachi, Kure-shi, Hiroshima-ken 737-0029, Japan e-mail:
[email protected] The original paper regarding the contents presented here was submitted to "FUEL" in April 2011 and the paper is currently under review.
Abstract The purpose of this study was to obtain practical knowledge in selecting suitable coals for boiler operations without ash related problems. For this purpose, the effects of fine ash particles and alkali metals on heat transfer characteristics at the early stages of ash deposition have been evaluated in 1.5 MWth pilot plant tests. Ash deposition characteristics of four bituminous and three sub-bituminous coals were studied using a slagging probe set in the high temperature zone simulating the water wall in actual boilers. To determine ash deposition characteristics quantitatively, heat flux through the slagging probe was calculated and the decreasing rate of heat flux per unit weight of ash(dq/dt･ash) was compared for each coal as slagging potential. Deposits analysis results showed that particle fraction with a diameter less than 10 µm (R10under) of bituminous coals varied from 5 to 45 wt %, while it was almost the same, around 20 30 wt %, for sub-bituminous coals. In terms of alkali components, deposits of bituminous coals enriched Fe and Ca. On the other hand, mainly Na and K were condensed in deposits of sub-bituminous coals. To evaluate the influence of these factors on slagging potential (dq/dt･ash) quantitatively, a multiple regression analysis of heat flux data was performed using the following parameters: R10under in deposit
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1
Oviedo ICCS&T 2011. Extended Abstract
samples and sum of Na2O and K2O values in each parent coal. Calculation using the regression equation agreed well with experimental data. 1. Introduction Ash deposition on heat transfer surfaces during coal combustion is a common concern of all coal firing boilers. In tackling ash-related problems, a number of studies have been made on characteristics of coal-ash deposition. These studies [1-5] show that the initial stage of coal ash deposition has a significant effect on heat transfer characteristics during the whole ash deposition process. According to such studies, the inner layer formed at the initial stage of ash deposition consists of fine particles smaller than 10 µm and condensed alkali vapor. However, these effects have not been evaluated quantitatively, specifically the impact of these parameters on decreasing heat flux at the initial stage of deposition growth. The difficulty in selecting suitable coals for boilers without ash related problems stems from this lack of quantitative knowledge. In this study, to obtain practical knowledge in selecting coals, the effects of fine ash particles and alkali metals on heat transfer characteristics were evaluated quantitatively by conducting slagging tests at a pilot scale test plant. 2. Experimental Slagging characteristics of seven types of bituminous and sub-bituminous coals have been studied in a 1.5 MWth pilot plant with a slagging probe simulating the water wall of actual boilers (Fig.1). The furnace is a cylindrical down flow furnace, 6 m in height with a diameter of 1.35 m. The burner is at the top and three After Air Ports (AAP) are located 2, 3 and 4m down from the burner exit for staged combustion. During tests, exhaust gas passing through the furnace is cooled by an air heater then introduced to the flue-gas treatment system. Table 1 lists coal identification and operation times of tested coals. Fuel ratio of tested coals varies from 1.1 to 1.5 and ash content changes from 1.9 to 11.9 wt %. The pilot plant was operated at a minimum of 4hours and at a maximum of 100 hours. A decrease in heat flux due to growing ash deposits on the test probe is evaluated in each test run. After each run, ash deposits on the probe surface are sampled and analyzed for particle sizes and their chemical components.
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2
Oviedo ICCS&T 2011. Extended Abstract Flue-gas treatment system
Burner
Combustion Air
Air heater Vertical probe （Water-cooled）
Furnace Horizontal probe （Water-cooled）
AAP （After Air Port）
Probe temperature control（873K ） Horizontal probe （Air-cooled） Water seal
Fig.1 Schematic diagram of 1.5 MWth pilot plant.
Table 1 Coal identification and operation times of tested coals. Coal classification Coal ID Fuel ratio (-) Ash (wt%,db) Operation time (h)
Bituminous coals MP 1.2 5.2 4
AN SW 1.5 1.1 9.2 11.1 4 100
VC 1.5 11.9 69
Sub-bituminous coals DA PS RB 1.1 1.3 1.1 1.9 4.9 4.5 75 10 8
3. Results and Discussion For all test cases, heat flux through the test probe decreased dramatically at the beginning of the test. We analyzed the initial decrease of heat flux by a linear regression method and compared the decreasing rate of heat flux per unit weight of ash (dq/dt･ash) to discuss slagging potential of each coal. Slagging potential (dq/dt･ash) of tested coals was greatest for MP, followed by PS, AN, DA, RB, VC and SW. All deposits sampled from the test probe after each test were classified into three types: deposits with molten particles, powdered and agglomerated deposits. Deposits consisted of multiple layers, and in every case, the innermost layer was formed by numerous fine particles. Deposits analysis results showed that particle fraction with a diameter less than 10 µm (R10under) varied from 5 to 45 wt % for bituminous coals, while it was almost the same for subbituminous coals: around 20 - 30 wt %,. In terms of enriched components, deposits of bituminous coals enriched Fe and Ca, whereas those of sub-bituminous coals contained
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3
Oviedo ICCS&T 2011. Extended Abstract
higher ratio of Na, K and Si. From the correlation of analysis results and slagging potential (dq/dt･ash) of each coal, it was found that slagging potential of bituminous coals was high when R10under in the deposit was high and that the potential of subbituminous coals was high when the coal enriched Na2O and K2O. Given these results, a multiple regression analysis of slagging potential was performed using the following parameters: R10under in sampled deposits, and the sum of Na2O and K2O values in each parent coal. Calculation using a regression equation agreed well with experimental data. 4. Conclusions To obtain practical knowledge in selecting coals for boiler operations without ash related problems, slagging potential (dq/dt ･ ash) of seven types of coals including bituminous and sub-bituminous coals have been studied in slagging tests at a 1.5 MWth pilot plant. The effects of fine particles and alkalis on slagging potential were evaluated quantitatively. Calculation using regression equation having parameters of R10under, Na2O and K2O agreed well with experimental data and will provide useful knowledge for selecting suitable coals for actual boilers. References [1] Raask E. Mineral impurities in coal combustion, behavior, problems, and remedial measures. USA: Hemisphere Publication Corporation; 1985. [2] Benson SA and Sondreal, Impact of Mineral Impurities in Solid Fuel Combustion 1999; 121. [3] Laursen K, Frandsen FJ and Larsen OH, Impact of Mineral Impurities in Solid Fuel Combustion, 1999; 357-366. [4] Bryers RW, Fireside slagging fouling , and high-temperature corrosion of heat-transfer surcface due to impurities in steam-raising fuels. Prog Energy Combust. Sci 1996; 22(1):29120. [5] Courch G., Understanding slagging fouling in PF combustion. Research, I.C. Editor. London, UK; 1994: 118.
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4
Investigation of Contributions to Unburned Carbon in a 200 MWe Utility Boiler H. Gao, A. Majeski, A. Runstedtler, M. Sybring CanmetENERGY, Natural Resources Canada, 1 Haanel Drive, Ottawa, Ontario, Canada K1A 1M1 Tel No: 001 613 9965194, Fax: 001 613 9929335, E-mail:
[email protected] Keywords: Utility boiler, CIA, Residence time, CFD Abstract This study presents a novel examination of coal behavior in the computational fluid dynamics (CFD) model results for a 200 MWe tangentially-fired utility boiler. Data for thousands of particles were extracted and summarized for characteristics such as injection location, size, residence time, and unburned carbon content in fly ash (carbon-in-ash, or CIA). The analysis revealed that residence times for coal particles vary widely, the availability of oxygen along the trajectories of the coal particles at top levels plays an important role in the burnout process. Contributions to unburned carbon from different particle size classes, burner and burner levels were also investigated. Additionally, it was found that a noticeable amount of burning may still be occurring in the platen super heater region beyond the furnace outlet. The findings from this study can provide useful information for boiler design and retrofit, particularly in clean coal technology areas. Suggestions for possible design and operating improvements are put forward. 1. Introduction Unburned carbon content in fly ash (CIA) reduces both power utility boiler efficiency (since it is unburned fuel) and fly ash salability (since it reduces concrete strength). Minimizing unburned carbon in fly ash is therefore in the interests of power utility companies. Moreover, the wide application of low NOx technologies in utility boilers has caused an increase in CIA in many boilers. Currently, the only measureable data for unburned carbon in utility boilers is from fly ash sample analysis. The fly ash used for these analyses typically contains a mixture of coal particles from all burners in the unit and all of the different particle size classes. This obscures contributions to CIA from individual burners and particle size classes. Particle residence time is an important criteria for boiler design, and it is usually taken to be the travel time for the particles from burner outlet to the furnace outlet (shown in Figure 1). It is typically assumed that the majority of the coal particles complete combustion within that time. However, residence time in a utility boiler is traditionally estimated using a plug flow assumption, which may not be totally accurate given that a utility boiler usually has multiple coal burners
arranged at different locations. Moreover, the trajectories of coal particles from even a single burner are varied and complicated. With their rapid development in recent decades, computational fluid dynamics (CFD) modeling technologies can now provide a reasonable prediction of coal combustion in utility boilers. As a result, the application of CFD to utility boilers [1, 2, 3, 4, 5] is growing rapidly. CFD modeling technology solves the aerodynamics equations in a boiler simulation, as well as the heat transfer and combustion related equations, to predict boiler performance. In the mean time, thousands of coal particles with different properties, representing different size classes, from different coal burners are tracked in their travels throughout the boiler while undergoing both devolatilisation and char oxidation processes [6]. Particle properties including diameter, char and volatile fractions, particle travel time and coordinates are all recorded in a data file for further processing. These data are used to investigate the coal particle burnout behaviour in the boiler and it is the focus of this work to process and present this data in a novel way, in order to gain a level of understanding beyond the usual analysis of CFD results. 2. The Boiler Figure 1A shows the 200 MWe tangentially-fired utility boiler and its two groups of platen super heaters in the upper furnace. The coal burners (labeled by levels A, B, C and D), the air nozzles
Figure 1: The 200MWe boiler, its air nozzles, coal burners and injection angles. (labeled by levels AA, AB, BC, CD and DD) and oil guns (OA, OB), identical for each corner, are shown in Figure 1B. The injection directions of the air nozzles, coal burners and oil guns are the
same in each corner and are shown in Figure 1C. These injections angles are designed to generate a rotating flow in the boiler. The boiler is firing a bituminous coal; the coal size distribution listed in Table 1 is calculated according to the sieve data sampled at the power plant. Table 1 Coal particle size distributionc Particle diameter, micron Mass fraction, %
Class 1 29 43.8
Class 2 67 21.7
Class 3 105 13.6
Class 4 143 8.3
Class 5 181 12.6
3. Results and discussion A CFD simulation was performed for this boiler at full load operating conditions. The air nozzles, coal burners and the two platen super heater groups in Figure 1A are all included in the simulation domain. The predicted temperature distributions appear reasonable compared with field observations and the predicted CIA (3.27%) is close to the field measurement (3.0% to 6.2%). 3.1 Unburned Carbon in Fly Ash CIA measured at power plants results from the mixture of coal particles from all burners and all size classes. The contribution of each burner or size class to the total unburned carbon can not be identified. In a CFD simulation however, the individual particle information for thousands of coal particles along their trajectories, including the char and volatile fractions, travelling times, diameters can all be recorded. To process the data file, a script was written to extract the recorded particle information as particles crossed a designated plane. By analyzing this data, we could determine the CIA of the coal particles at this designed plane. Further, it was possible to identify the CIA for a particular burner by sorting the data by burner. Table 2 summaries the results at the furnace plane I (refer to Figure 1). CIA varies widely among burners in this table: it is higher than 8% for some burners and lower than 1% for other burners. Even on the same level, the CIA is quiet different among burners. From bottom level (Level A) to top level (Level D), CIA increases from less than 1% to 4.26%. The predicted CIA of all tracked coal particles taken together is 3.27%, which is close to the fly ash sample analysis results in the field as mentioned earlier. Table 2 CIA (%) at the furnace plane I Level D Level C Level B Level A Average
Corner 1 8.02 3.73 1.95 1.05
Corner 2 7.52 6.64 2.90 1.47
Corner 3 0.22 1.33 4.37 0.52 3.27
Corner 4 0.71 5.15 1.59 0.75
Average 4.26 4.25 2.77 0.91
Further processing of the data enables the contribution of each burner on total unburned carbon to
be calculated, Table 3 shows these results. More than 73% of the unburned carbon in fly ash is from Level D and Level C. Six burners (Corner 1 and Corner 2 of level D, Corner 1, Corner 2 and Corner 4 of Level C, Corner 3 of Level B) contribute about 80% of the unburned carbon among the 24 coal burners. This type of information can guide boiler operation, because fewer than half of burners need to be targeted to reduce CIA rather than all of them. Table 3 Contribution (%) of burners on total unburned carbon in ash Level D Level C Level B Level A
Corner 1 17.99 7.98 3.51 1.73
Corner 2 16.74 14.65 5.78 1.49
Corner 3 0.46 2.77 9.40 0.75
Corner 4 1.47 11.17 2.99 1.11
Total 36.67 36.57 21.68 5.08
The relationship between particle sizes and unburned carbon in fly ash is presented in Table 4. Despite of the fact that the mass fraction of the largest coal particles (Class 5, 181 micron) entering the boiler is only 12.6% (Table 1), almost 50% of the unburned carbon in fly ash is found in this class. The smallest particles (Class 1 and 2) contribute only 11% of the unburned carbon in ash despite their accounting for over 65% (Table 1) of the coal entering the boiler. This data provides guidance for operators to adjust fineness for pulverisers at different levels. Coal particles at the top levels need to be finer, while the coarser coal particles at bottom may not necessarily increase CIA. Moreover, the finding suggests using different pulverizers or classifiers for different levels right in the boiler design stage. Table 4 Contribution (%) of particle size classes on unburned carbon in ash Class 1 Class 2 Class 3 Class 4 Class 5
Level A 0.23 0.64 1.08 1.12 1.90
Level B 0.29 1.35 4.06 5.11 10.43
Level C 0.39 2.88 6.59 7.88 18.11
Level D 0.84 4.39 7.46 7.90 17.33
Total 1.75 9.26 19.20 22.01 47.77
3.2 Residence time Table 5 summaries the residence times extracted from particle data.
As elevation increases from
level A to level D, particle residence time reduces from 5.96s to 1.60s. On any given level, the averages among the corners tend fall within 2s. For these particles with a relative short residence time (level D and C), a relative longer residence time doesn’t necessary create a better burnout, this conclusion can be clearly seen on Figure 2 which plotted the CIA versus the residence time for the particles from each coal burner. To find the reason behind this, the average oxygen volume fraction along particle trajectories of level D was calculated. The average oxygen volume fraction for particles of Corner 1 and Corner 2 is 3.2% to 4.6%, while it is 5.4% to 6.5% for Corner 3 and
Corner 4. The CIA of them (Corner 1, 2 and Corner 3, 4) are quiet different as indicated in Table 2 although their residence time are similar. Maybe the oxygen availability along the particle trajectories is more important for burnout for this level. Table 4 Residence time ( s ) of each coal burner Corner 1 1.80 1.97 4.23 5.31
Level D Level C Level B Level A
Corner 2 1.39 1.88 3.37 7.78
Corner 3 1.51 1.77 2.44 5.59
Corner 4 1.68 2.05 3.72 5.16
Average 1.60 1.92 3.44 5.96
c
9
8
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C I A, %
6
5
4
3
2
1
0 1
2
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7
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Residence time,s
Figure 2:
Particle residence time versus CIA of each burner
Furnace plane I
Furnace outlet plane
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C I A, %
8
6
4
2
Level D
Level C
Level B
Corner 4
Corner 3
Corner 2
Corner 1
Corner 4
Corner 3
Corner 2
Corner 1
Corner 4
Corner 3
Corner 2
Corner 1
Corner 4
Corner 3
Corner 2
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0
Level A
Figure 3: CIA at furnace outlet plane and furnace plane I
The particles of level A and B have longer residence times as well as a good chance to pass through an oxygen rich region because most of the air nozzles are above (downstream of) the burners. This
result reveals the importance of creating a uniform oxygen distribution beyond the burner region to unsure all particles encounter with oxygen-rich regions. 3.3 Combustion beyond the furnace outlet plane The furnace outlet plane is defined in Figure 1A. Although field observations have suggested that coal particles may still be burning beyond the furnace outlet plane, there have been no measurements to prove it. Traditionally it is believed that most combustion of coal particles stops beyond this plane because of the quenching effect of the platen super heater groups. To examine if combustion is still occurring beyond the furnace outlet plane, Figure 3 compares the CIA at the furnace outlet plane and at furnace plane I (labeled in Figure 1) for all coal burners. This figure indicates that there is a noticeable reduction in CIA beyond the furnace outlet plan for all burners, average CIA reduces to 3.27% at furnace plane I from 4.1% at furnace outlet plane. This figure supports the supposition that coal particles are still burning beyond the furnace outlet, led to a reduction in CIA in the upper furnace. Continued burning in the upper furnace may create some potential to further reduce NOx emissions. Some separated over fire air (SOFA) nozzles may be arranged in this region to achieve a deeper air staging. Also a small fraction of air injection in this region may help reduce CIA. 4. Conclusions CFD modeling, combined with novel data analysis, has provided an unprecedented capacity to characterize coal particle burnout history in a utility boiler and can answer questions that have long existed in utility power industry. Predicted residence times for coal particles could vary widely among burners in a utility boiler, and it ranges from 1.6s to 6s in this boiler. The CIA of each coal burner could be quiet different and the unburned carbon in fly ash could be mainly from some larger coal particles of a few coal burners as predicted for this 200MWe utility boiler. This study also supports the notion that coal combustion may continue to occur in the upper furnace region, leading to a reduction in CIA in the upper furnace. Oxygen availability along particle trajectories appears important for burnout for the coal burners at top levels. This approach to investigating the coal particle burnout characteristics in a boiler can provide new ideas to design and operate a utility boiler as noted in the discussion. Acknowledgements This study was supported by the Federal Program on Energy Research and Development (PERD) of the Canadian Federal Government.
References
[1] Yin CG, Caillat S., Harion JL, Baudoin B., Perez E. Investigation of the flow, combustion, heat-transfer and emissions from a 609MW utility tangentially fired pulverized-coal boiler. Fuel 2002; 997:1006-81. [2] Philip j. Stopford, Recent application of CFD modeling in the power generation and combustion industries, Applied Mathematical Modelling 2002; 26:351-374 [3] Eddy H. Chui, Haining Gao, Estimation of NOx emissions from coal-fired utility boilers. Fuel 2010; 89:2977-2984. [4] Haining Gao, Eddy H. Chui, Numerical Investigation of the Impact of Air Distribution on the Performance of a 360 MW W-Fired Boiler. International Pittsburgh Coal Conference, Osaka, Japan, 2004 [5] Chui E.H., Haining Gao, Reduction of emissions from coal-based power generation, International conference on climate change 2007. Hong Kong; 2007. [6] AEA Technology Engineering Software Limited, CFX-TASCflow Theory Documentation Version 2.12, Sept. 20 2002
Oviedo ICCS&T 2011. Extended Abstract
Visualizing The Macromolecular Network Structure Of A Large-Scale (50,000 atoms) Illinois No. 6 Coal Molecular Representation In 3D And 2D Lattice Views 1
Alvarez, Y. E.; 2 Watson, J. C. K.; 3 Pou, J. O.; 1Castro-Marcano, F.; 1Mathews, J. P.
1
John and Willie Leone Family Department of Energy & Mineral Engineering and the EMS Energy Institute, 110 Hosler Building, 2The Applied Research Laboratory, The Pennsylvania State University, University Park 16802, USA
[email protected] 3 Departament d’Enginyeria Química, Escola Tècnica Superior IQS, Via Augusta 390, 08017 Barcelona, Spain Abstract The utility of large-scale structural models of coal is currently limited by their scale (>20,000 atoms) and complexity. A novel computational approach that performs systematic simplifications of complex 3D structural models into corresponding 2D lattice representations was used to generate equivalent reduced models for an Illinois no. 6 bituminous coal molecular representation of >50,000 atoms. The visualization of molecular entities (clusters, cross-links and molecules) from the original representation is enhanced by the 2D lattice model which is able to capture the network and portray the properties of the original representation under various color schemes. 1. Introduction Capturing the continuum of coal structure requires large-scale coal models that enable inclusion of molecular weight distribution and structural diversity. However, construction of large-scale representations is time-consuming, challenging, and requires considerable expertise. Recent work [1, 2] has attempted to capture aromatic alignment and stacking and the distribution of structural features for carbonaceous materials such as coals, chars and soots from HRTEM lattice fringe images directly via Fringe3D. With this approach, a large-scale model for Illinois no. 6 Argonne Premium coal was constructed based on available experimental data [3] in an effort to move towards capturing the continuum structure [4]. The model contains 50789 atoms in 728 crosslinked aromatic clusters with a continuous molecular weight distribution ranging from 100 to 2850 Da and it is the largest, most complex coal representation currently available.
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Oviedo ICCS&T 2011. Extended Abstract
The utility of large-scale structural models of coal has been challenged by their scale (>20,000 atoms) and complexity. The very feature that allows them to successfully represent coal behavior by the incorporation of a molecular weight distribution and structural diversity hinders their use in bond-altering simulations due to the associated computational cost. Hence, in view of the visualization and utility challenges involving complex structural models, an enabling computational tool for the simplification of large-scale molecular representations [5, 6] of coal was developed, consisting of two main scripts, Col2D and Molecwalk, capable of generating equivalent 3D coarse-grained and 2D lattice molecular models from the complex original molecular representation. This is accomplished by pattern recognition of hydroaromatic/aromatic clusters and cross-links, which are correspondingly reduced into lattice nodes and linkage lines. The approach enables a reduction of scale down to 3% of the original number of atoms, and while the view is simplistic, the coarse-graining process is able to retain all of the structural and chemical information of the reduced units (clusters and cross-links), facilitating the manipulation for visualization and simulation of the lattice and coarse-grained models. The various stages of simplification process by Col2D and Molecwalk are shown in Figure 1 for a 3D Wiser bituminous coal model [7]. Initially rings (stage 1) and clusters (stage 2) are identified and adjacent/connected elements (such as functional groups and sidechains) are collapsed into nodes (stage 3). Cross-links that may consist of multiple atoms are further collapsed into linkages (stage 4), thus Col2D creates a 3D coarsegrained view. The functional group information is retained along with the node and cross-links’ composition, for further consideration in reactivity calculations. A 2D lattice is generated from the simplified 3D structure (stage 5) through the Molecwalk script which performs ‘random walks’ of the skeletal molecules in a lattice space, obtaining a unique 2D lattice model of the original coal representation.
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Oviedo ICCS&T 2011. Extended Abstract
Figure 1. Simplification process by Col2D and Molecwalk for the Wiser bituminous coal model [7]. In stage 5 white nodes are empty lattice slots. Generic non-structure specific lattices have been used in the past to represent the macromolecular structure of coal and to describe thermal network decomposition through inexpensive statistical approaches [8-11]. Linking complex models to structurederived lattices could continue to provide paths for exploring the reactive behavior of coal and similar carbonaceous systems. It is expected that the lattice model will be a means of simulating, through network statistics, thermal breakdown processes related to pyrolysis and/or liquefaction; thus, this approach adds a visual feature and the incorporation of a coal-specific network model that is lacking in previous network decomposition modeling approaches [8-11]. In the current study, the Col2D and Molecwalk approach was utilized to generate corresponding 3D coarse-grained and 2D lattice representations of the Illinois no. 6 coal model to demonstrate the utility of structure-specific 2D lattice molecular models. 2. Methodology Here, the 3D coarse-grained and 2D lattice models are referred to as reduced models. During the execution of Col2D, an extra set of molecular files are generated which enable a connection between the coarse-graining results and the original atomic file. This extra set of InsightII files are known as the ‘Link’ files. These files treat the reducible molecular parts of the original atomic model (i.e. hydroaromatic clusters and cross-links) as independent entities (molecules), thus allowing the later extraction of
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Oviedo ICCS&T 2011. Extended Abstract
their chemical information based on individual atomic composition. However, a feature not yet included is that the functional groups are not transferred into the reduced molecules, and thus their presence is not accounted for select property calculations. The naming scheme assigned to the molecular parts in the reduced models (nodes and linkages) is consistent with the atomic reducible groups in the Link files (hydroaromatic clusters and cross-links), keeping reduced models and original atomic model relatable. This is illustrated in Figure 2.
Figure 2. Demonstration of molecular variations from original atomic file to 'Link' file and naming scheme A series of scripts (in Perl and Visual Basic) were developed to extract the chemical information from each molecular part of the Link files. By submitting the “Link” files to these model-analysis scripts, it is possible to index each node and linkage line to a series of composition-based chemical properties. This enables the formation of a summary table termed ‘reduced unit property table’ that contains properties such as elemental composition, molecular weight, chemical groups in cross-link, cross-link classification, naming scheme, etc. of each reduced unit (RU). A sample of this table showing a few properties for various reduced units in the Illinois no. 6 coal model is shown in Table 1. Table 1. Reduced unit property table for 3 molecules from Illinois no. 6 coal model Molecule_RU Name
C
H
O
N
S
MW
RU Type
Cross-link groups
Sketch520_C0
6
6
0
0
Sketch520_C1
10
8
0
0
0
78
cluster
n/a
0
128
cluster
n/a
Sketch520_C2
10
12
0
0
0
132
cluster
n/a
Sketch520_B0
2
6
Sketch520_B1
4
1
0
0
46
cross-link
-O-
0
0
0
58
cross-link
- CH2 CH2-
Sketch528_C0
10
8
0
0
0
128
cluster
Sketch536_C0
17
19
0
1
0
237
cluster
n/a
Sketch536_C1
10
12
0
0
0
132
cluster
n/a
Sketch536_B0
4
10
1
0
0
74
cross-link
- CH2 CH2O -
Table lines separate reduced units of individual molecules
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Oviedo ICCS&T 2011. Extended Abstract
Other properties this table may include are: NMR aromaticity parameter (fa), solubility parameter (calculated directly from the model with scripts written from previous work [12]), and atomic composition by force field type, which is basically the list of atoms in the original component under their force field type. Reduced unit property tables organize the chemical data of the model’s components making it accessible for manipulation of the lattice view based on any of the table properties. Once the desired properties are tabulated, a different series of Perl scripts executed in a modeling platform (Materials Studio 5.0) can be executed to change the visualization of the lattice, by coloring it under selected color-schemes for each property. A modelsensitive legend is also generated so that the range of property values within the molecular weight distribution is well represented. 3. Results and Discussion The large-scale Illinois no. 6 bituminous coal representation was processed by the simplification approach [5, 6], hence obtaining reduced models as illustrated in Figure 3. Given the density and scale of the original model, it is still very difficult to visually distinguish the reduced molecules in the 3D coarse-grained (Figure 3b); hence it was more desirable to portray the all molecular entities through a 2D lattice model [5, 6] (Figure 3c), which distributed the 728 molecules in an 80x80 grid space.
Figure 3. Atomic large-scale representation of Illinois no. 6 bituminous coal model transformed to 3D coarse-grained and 2D lattice equivalents by Col2D and Molecwalk approach. The reduction of scale was from 50879 atoms to 2459 pseudo atoms, comprised of 1592 nodes and 867 linkage lines. The lattice may aid visualization, to portray the properties Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
of molecules, clusters or cross-links. For example, the script “Colorby_MW.pl” analyzes the range of the molecular weight of the clusters in the model (not including attachments), and assigns each node to a color representative of molecular weight. The molecular weight of the Illinois no. 6 clusters is shown in Figure 4.
Figure 4. Color-coding by molecular weight of lattice clusters from Illinois no. 6 bituminous coal model. Linkage lines are colored grey to aid visualization.
Figure 5 shows the lattice after being colored by cross-link types. Cross-links were classified according to their chemical composition and based on a predominant functionality (methylene, carbonyl, ether oxygen, etc.); some cross-links may be composed of a mix of functionalities, hence their type is termed “plus” to indicate the possible presence of other groups; in most cases, this “mix” is composed of the namedgroup plus methylene groups, for example -CH2-S-CH2- for “sulfur-plus” cross-links. Analyzing the chemical composition of the cross-links also enables the assignment of cleavage or reaction probabilities that would aid in a statistically-based reaction model.
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Oviedo ICCS&T 2011. Extended Abstract
Figure 5. Illinois no. 6 bituminous coal lattice with linkage lines color-coded by crosslink type. Nodes are gray.
The model and hence network is mainly cross-linked with oxygen and methylene bridges [-(CH2)n-(O)m-(CH2)p-; n,p=0=3, m=0-1]. There is also a significant incidence of ‘sulfur plus’ bridges, which are any bridges that contain sulfur mixed with other groups (thioethers). This is characteristic of the high organic sulfur content in Illinois no. 6 coal [3]. Other colored-coding examples at the molecular level are included in Figure 6. These result from running scripts on the original atomic model, and not the Link files, thus extracting the full chemical properties of the original molecules. The Painter solubility parameter [13] (cal. cm-3)0.5 was calculated for each molecule and these were classified and colored under the 12 color-ranges that are a function of the minimum and maximum solubility parameters in the model, providing a visual sense of the model’s theoretical extractability by a specific solvent (Figure 6a). Similarly, distances (Å) of the molecular-centroid to the center of the model in 3D space are individually calculated and classified by a 12 division of the model’s total distance range. The 3D distance
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Oviedo ICCS&T 2011. Extended Abstract
lattice color scheme may be useful for mass transfer and extractability studies where knowledge of the molecule’s position to the surface of the model is relevant (Figure 6b).
Figure 6. Illinois no. 6 lattice colored by a) Painter solubility parameter (cal. cm-3)0.5 b) Proximity of molecular centroid to the center of the model in 3D space (Å)
It is expected as future work that the lattice may serve as the visual interface for the structural changes generated from reactive simulations performed directly on the coarsegrained 3D model. 4. Conclusions Visualization has increasingly become a valuable research tool exploited by scientists to obtain insights of phenomena and to make scientific knowledge more accessible. Currently, the visualization capabilities are significantly enhanced by the generation of 3D and 2D lattice molecular representations of specific complex large-scale representations, thus reproducing a representation that captures the network and its molecular properties analyzed under various color schemes through the lattice view; however, it is expected that this will move forward to include the visualization of the lattice post-simulation, displaying transitions in the properties of the model components and in the network structure as a whole. Acknowledgements. This project was funded by the Illinois Clean Coal Institute with funds made available through the Office of Coal Development of the Illinois Department of Commerce and Economic Opportunity. We thank Alan Chaffee for the 3D Wiser model.
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Oviedo ICCS&T 2011. Extended Abstract
References [1] V. Fernandez-Alos, J.K. Watson, R.v. Wal, J.P. Mathews, Soot and char molecular representations generated directly from HRTEM lattice fringe images using Fringe3D, Combustion and Flame, available in ASAP articles DOI: doi:10.1016. (2011). [2] V. Fernandez-Alos, Improved molecular model generation for soot, chars, and coals: high resolution transmission electron microscopy lattice fringes reproduction with Fringe3D, in: Master Thesis, Energy and Geo-Environmental Engineering,, Pennsylvania State University, University Park, PA, 2010. [3] F. Castro-Marcano, J.P. Mathews, Constitution of Illinois no. 6 Argonne Premium coal: a review, Energy Fuels, 25 (2011) 845–853. [4] F. Castro-Marcano, V. Lobodinb, R. Rodgers, A. McKenna, A. Marshall, J. Mathews, A molecular model for Illinois no. 6 Argonne Premium coal: moving toward capturing the continuum structure, Fuel, Subtmitted (2011). [5] Y.E. Alvarez, Development of a reactive coarse-graining approach for the utility enhancement of complex large-scale molecular models of coal, in: Masters Thesis, Energy and Mineral Engineering, The Pennsylvania State University, University Park, PA, 2011. [6] Y.E. Alvarez, J.K. Watson, J.P. Mathews, Improving the utility of large-scale coal molecular models by simplifying the view: 3D models to reactive lattice grids, Prepr. Pap. -Am. Chem. Soc., Div. Fuel Chem., 55 (2010). [7] W.H. Wiser, Conversion of bituminous coals to liquids and gases, NATO ASI Series C, 124 (1983). [8] D.M. Grant, R.J. Pugmire, T.H. Fletcher, A.R. Kerstein, Chemical-model of coal devolatilization using percolation lattice statistics, Energy & Fuels, 3 (1989) 175-186. [9] S. Niksa, A.R. Kerstein, On the role of macromolecular configuration in rapid coal devolatilization, Fuel, 66 (1987) 1389-1399. [10] P.R. Solomon, D.G. Hamblen, R.M. Carangelo, M.A. Serio, G.V. Deshpande, General-model of coal devolatilization, Energy & Fuels, 2 (1988) 405-422. [11] P.R. Solomon, D.G. Hamblen, Z.-Z. Yu, M.A. Serio, Network models of coal thermal decomposition, Fuel, 69 (1990) 754-763. [12] D. Van Niekerk, Structural elucidation, molecular representation and solvent interactions of vitrinite-rich and inertinite-rich South African coals, in: Energy and Geo-Environmental Engineering, Pennsylvania State University, University Park, PA, 2008. [13] P.C. Painter, J. Graf, M.M. Coleman, Coal solubility and swelling. 1. Solubility parameters for coal and the Flory x parameter, Energy & Fuels, 4 (1990) 379-384.
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9
Oviedo ICCS&T 2011. Extended Abstract
Brown Coal Solubilisation with Novel Ionic Liquids
Alan L Chaffee1, Christin Patzschke1, Douglas Russell1, Daniel Kelley1, Ying Qi1, Vincent Verheyen2, Marc Marshall1, Vijay Ranganathan1 and Douglas MacFarlane1 1
School of Chemistry,Monash University, Clayton, Victoria 3800, Australia School of Applied Science and Engineering,, Gippsland Campus, Monash University, Churchill, Victoria 3842, Australia
2
Corresponding Author: Alan L. Chaffee Email:
[email protected]
Abstract Ionic liquids (ILs) are now recognised as an important medium for solubilisation and reaction in a variety of chemical applications. For example, cellulose, lignin and other biopolymers can be selectively solubilized and separated from complex biomass mixtures. A recent study has shown that significant extraction yields can also be obtained from some bituminous coals with the use of specific ILs or IL/solvent mixtures. However, there are unresolved issues with the separation of the IL from the extract and also with the recovery / recycling the IL. We have commenced a study into the solubility of Victorian brown coal in novel ILs that are based on the association of CO2 with low molecular weight amines. Victorian brown coal is observed to disperse very well in these ILs and significant solvent-free extract yields (~20%) can be obtained. The structures of the products (extracts and residues) have been evaluated by a variety of analytical techniques, including FTIR, 13CNMR and pyrolysis-GC-MS. Structural differences are illustrated and the character of the separations with different ILs described.
1. Introduction Ionic liquids (ILs) consisting of large organic cations associated with inorganic anions have attracted considerable attention as potential alternatives to conventional organic solvents in a variety of synthetic, catalytic and electrochemical applications. Research in recent years has found ILs to be able to dissolve plant materials such as cellulose and lignin [1-3]. Painter et al [4-5] used specific ILs with or without catalyst to extract bituminous coals and achieved significantly extraction yields. Victorian brown coal has provided an abundant and cheap energy resource for the State of Victoria, Australia, for decades. However, with the increasing price of petroleum
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Oviedo ICCS&T 2011. Extended Abstract
products, there may be an opportunity to diversify its end-use into new fields of application. The objective of this study is to investigate the solubility of Victorian brown coal in novel ILs that are based on the association of CO2 (a weak acid) with low molecular weight amines and characterize the chemical structures of the products (extracts and residues). The advantage of such ILs is that they dissociate into their precursor components at relatively low temperatures and can be readily removed from the products by this means. 2. Experimental section ILs used in the study were N,N-dimethylammonium N’,N’-dimethylcarbamate (DIMCARB) synthesized in the laboratory and 1-ethyl-3-methylimidazolium chloride ([emim]Cl) obtained from Sigma-Aldrich. Whilst [emim]Cl has very low volatility, DIMCARB dissociates into dimethylamine and carbon dioxide at ~ 60°C. The coal used was run-of-mine Loy Yang coal from the Latrobe Valley, Australia, milled to less than 1 mm, and with a determined moisture content of 50 %. Raw coal with 50% moisture was mixed with the selected IL at a mass ratio of 1:10 in a beaker or a centrifuge tube using a magnetic stirrer (20-24 hr) or an ultrasonic bath (3 min). All sonicated mixtures were left over night for further solubilisation and to ensure comparable conditions. Due to the solid state of [emim]Cl at room temperature, the experiments were carried out by heating the mixture to the corresponding melting points. The solid residue was then separated by vacuum filtration or centrifugation. The residual and extracted products of DIMCARB were dried in an oven at 105 °C to remove any residual of IL and water. After separating the extract and residue of the [emim]Cl extraction, the IL was removed by washing with water. The products were then dried in an oven at 105 °C. The dried materials were weighed after being cooled to room temperature in a desiccator. The dried samples of coal and extraction products (residues and extracts) were analysed for FTIR (Fourier Transform Inferred Spectroscopy), Pyrolysis-GC-MS and Solid state 13C-NMR (Nuclear Magnetic Resonance Spectroscopy). 3. Results and Discussion 3.1. Extraction yield The yields of extracted material were mostly between 10 % and 20 % depending on the IL used and the method for solubilisation (Table 1). Mixing the IL and coal by stirring gave higher extraction yields than those observed using an ultrasonic bath. Separation of the coal residue and the extract by centrifuging or filtration seemed to have little effect on the achievable yield. Using the stirring method, DIMCARB demonstrated better extraction capabilities than [emim]Cl.
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Oviedo ICCS&T 2011. Extended Abstract
Table 1 Extraction Yield using ILs. IL (g) Coal (g, db) DIMCARB
[emim]Cl
5.16 5.21 4.19 20.98 22.18 4.14 5.45
0.28 0.28 0.22 1.11 1.10 0.11 0.17
Solubilisation/ Separation method Stirring/filtration Stirring/centrifuging Sonication/filtration Stirring/filtration Stirring/filtration Sonication/filtration Sonication/centrifuging
Extraction yield (%) 22.2 21.7 13.4 20.5 15.5 17.7 12.5
3.2. FTIR analysis FTIR spectra of the extraction products show the similarities of the major absorption bands to that of the original coal (Figure 1). However, the band in the 3200-3400 cm-1 range for the extracts and residues of both [emim]Cl (not shown) and DIMCARB is stronger and sharper than for the coal. This may be due to the N-H bond from the residues of the ionic liquids in the extraction products. The C-O stretching in the range of 1250-1300 cm-1 in the spectra of the extraction products is also weaker and very broad compared to that of the coal. The relative intensity of the band of C=O stretching (an indication of carboxyl group) at 1700 cm-1 compared to that of the C=C stretching at 1600 cm-1 is more prominent in the extract than in the residue and all weaker than in the coal. coal
CO2 artifact CH2 O-H
C=O
CH3
C-O-H C=C
C-O
DIMCARB extract
DIMCARB residue
3600
3100
2600
cm ‐1
2100
1600
1100
600
Figure 1 FTIR spectra of original Loy Yang coal, DIMCARB extract and DIMCARB residue. 3
Oviedo ICCS&T 2011. Extended Abstract
3.3. Solid state 13C-NMR Figure 2 compares the spectra of Solid state 13C-NMR of coal and DIMCARB residue and extract. It shows that the extract and residue are clearly different from each other and from the coal. The extract contains significantly higher portion of carboxyl carbons (160-180 ppm) than the residue. Higher proportions of aromatic carbons (100-150 ppm) are present in the residue than in the extract and there is an interesting difference in the distribution of aliphatic carbons between the two products. There is some evidence of entrainment of the DIMCARB in the extract (peak at ~40 ppm may be due to aminated carbon) in spite of the fact that the sample was treated at 105°C prior to analysis.
Coal
DIMCARB residue
DIMCARB extract
260
240
220
200
180
160
140
120
100
80
60
40
20
0
ppm
Figure 2 Solid state 13C-NMR Spectra of original Loy Yang coal, DIMCARB extract and DIMCARB residue. 3.4. Pyrolysis-GC-MS analysis The volatile components of each sample were thermally desorbed at 340°C and analysed by GC-MS. Following this analysis, the residue from thermal desorption was pyrolysed at 720°C and analysed by GC-MS in a similar fashion. Figure 3 compares the thermal desorption chromatograms (only) for the original Loy Yang coal and the extracts. Figure 4 compares the chromatograms of the 720°C pyrolysis products (only) for the coal and residues. Mass spectrometric characterisation of some lower molecular components in the chromatograms, such as the labelled peaks in Figures 3 and 4, indicates the presence of nitrogen containing compounds, which almost certainly point to the entrainment of some ionic liquid within the products. It can be seen that entrained DIMCARB is removed by thermal desorption at 340°C. However, entrained [emim]Cl is principally evident from the higher temperature (720°C) pyrolysis run.
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Oviedo ICCS&T 2011. Extended Abstract
Coal 340°C
[emim]Cl extract 340°C
IL residue
DIMCARB extract 340°C IL residue
Figure 3 Comparison of GC-MS chromatograms of thermal desorption products of coal and solvent extracts at 340°C. Coal 720°C
[emim]Cl residue 720°C
IL residue
DIMCARB residue 720°C
Figure 4 Comparison of pyrolysis-GC-MS chromatograms of pyrolysis products of coal and residues at 720°C. At the lower (thermal desorption) temperature, it can be seen that the extracts provide a different product distribution than the whole coal itself. At the higher temperature (not
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Oviedo ICCS&T 2011. Extended Abstract
shown), the DIMCARB extract produces more low molecular weight compounds than the [emim]Cl extract. This is in agreement with the extraction yields, suggesting DIMCARB as a better solvent than [emim]Cl. Pyrolysis of the residues shows that the composition resembles that of the original coal with the presence of hydrocarbon homologues (Figure 4). A more detailed evaluation of the mass spectroscopy results is underway. 4. Conclusions The extraction yields of raw coal without any pre-treatment by [emim]Cl and DIMCARB range between 10 and 20 % depending on the solubilisation and separation methods used. FTIR data indicate differences in the C=O and C-O regions of the spectra products between the extracts and the residues and compared to the original coal. Solid state 13C-NMR measurement indicates that the DIMCARB extract contains higher proportions of carboxyl carbons whereas the residue contains higher proportions of aromatic carbon. Pyrolysis-GC-MS indicates the entrainment of some ionic liquid within the products. It is also showed that the composition of thermal desorption products from the extracts differs considerably from the whole coal. The DIMCARB extract provides a lower molecular weight distribution of products than the [emim]Cl extract. Acknolwedgement This work has been supported via a BCIA (Brown Coal Innovation Australia) Research Leadership Fellowship. References [1]
[2]
[3]
[4]
[5]
Binder Joseph, B, Raines Ronald, T. Simple chemical transformation of lignocellulosic biomass into furans for fuels and chemicals. J Am Chem Soc 2009;131:1979-85. Fort, DA, Remsing, RC, Swatloski, RP, Moyna, P, Moyna, G, Rogers, RD. Can ionic liquids dissolve wood? Processing and analysis of lignocellulosic materials with 1-n-butyl-3-methylimidazolium chloride. Green Chem. 2007;9:63-9. Fu, D, Mazza, G, Tamaki, Y. Lignin Extraction from Straw by Ionic Liquids and Enzymatic Hydrolysis of the Cellulosic Residues. J Agri Food Chem 2010;58:2915-22. Painter, P, Cetiner, R, Pulati, N, Sobkowiak, M, Mathews, J. Dispersion of Liquefaction Catalysts in Coal Using Ionic Liquids. Energy Fuels 2010;24:308692. Painter, P, Pulati, N, Cetiner, R, Sobkowiak, M, Mitchell, G, Mathews, J. Dissolution and Dispersion of Coal in Ionic Liquids. Energy Fuels 2010;24:1848-53.
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Oviedo ICCS&T 2011. Extended Abstract 1
H-NMR Study on the Thermoplasticity of Coking Coal - Effects of coal blending and additives H. Kumagai1, N. Okuyama2, T. Shishido2, K. Sakai2, M. Hamaguchi2, N. Komatsu2 1
Graduate School of Engineering, Hokkaido University, Kita-13, Nishi-8, Kita-ku, Sapporo 060-8628, Japan. E-Mail:
[email protected] 2 Coal & Energy Technology Dept., KOBE STEEL, Ltd., 2-3-1, Shinhama, Arai-cho, Takasago 676-8670, Japan Abstract This study aim to investigate the effects of HPC addition on the thermoplasticity of coal blends. The thermolasticity of coal blends with HPC were monitored with in-situ high temperature 1H-NMR relaxation measurement. The solid echo pulse sequence was employed to generate 1H-NMR transverse relaxation signals. The echo signals obtained during heat treatment under a flow of nitrogen at a heating rate of 3K/min were deconvoluted into a set of one Gaussian and two Exponential decay components which represent the immobile, intermediate and mobile component, respectively. The changes in the fractional intensity of mobile component, fHm, calculated from the signals well corresponded to the softening and resolidification phenomena of coal blends. The fHm and molecular mobility of mobile component represented by Spin-Spin relaxation time, T2Hm, increased with HPC addition at lower temperature range. At higher temperature range, that is thermoplastic temperature range, fHm and T2Hm shows almost the same value as those for original coal blends. These results indicate that the HPC can be added as a substitution for coking coals. 1. Introduction Because of the importance of thermoplastic properties, which have definitive effects on the properties of the resultant coke, a wide range of experimental techniques have been applied to study the thermal transformation of materials. Most established dynamic measurement techniques specifically devised for the characterization of thermoplastic behavior of coking coal are concerned with rheological or dilatometric properties1). The dynamic measurement techniques, such as the Gieseler plastometer, and other conventional testing techniques, such as crucible swelling and reflectance measurement, are useful to characterize the properties of coking coal and enable the empirical evaluation of the quality of the resultant coke. However, relationships between the
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1
Oviedo ICCS&T 2011. Extended Abstract
changes in the parameters obtained from the conventional techniques and actual changes in the reactivity and structure of the coal during heating are not clear yet. Of the in-situ measurement techniques available, high temperature in-situ
1
H-NMR relaxation
measurement allows the progress of thermal transformation to be monitored by quantitatively recording the hydrogen content of residual specimens and as a result the sensitivity of the relaxation signal to the molecular properties of specimens2). In principle, 1H-NMR relaxation measurement is applicable to materials which have nuclear spins regardless of the state of the materials, that is, solid, liquid, gaseous or a mixture of these states. Furthermore, relaxation characteristics of nuclear spins are closely related to the mobility and motion of molecules. Thus, the phase transformation of materials can be detected as a variation in relaxation behavior of nuclear spins. The coal extracts named High Performance Caking additive, HPC, was produced by thermal extraction of coal using 2-ring aromatic solvent. HPC appeals an excellent thermal plasticity although the parent coal appeals no thermal plasticity. HPC softens at low temperature, keeps a highly fluidity in a wide temperature range and re-solidifies at high temperature. These thermoplastic properties exhibit that HPC can be available as a caking additive to make strong coke for the blast furnace. Presented in this study is a quantitative and qualitative evaluation of the thermoplasticity of coal blend with HPC, which was attempted using in-situ 1H-NMR relaxation measurement. 2. Experimental section 2. 1. Samples In this study, three Australian strongly caking coals (A, B, C) and one slightly caking coal (D) was used as sample. Table 1 shows the results of proximate analysis and Gieseler plastometry of the coal samples, and Table 2 shows the blending ratio of each coal sample. Base-1 consists of 75wt% of strongly caking coals, coal-A : B: C : D = 15: 26: 34: 25, and Base-2 consists of 50wt% of strongly caking coals, coal-A : B: C : D = 15: 30 : 5: 50. Coal-C was replaced with HPC for Base-1, and coal-B was replaced for Base-2. HPC was extracted from Australian steaming coal (MO) at 400ºC for 1 hour, and insoluble matter was separated using the settling and filtration system. The properties of MO coal are shown in Table 1. 2. 2. NMR measurements Spin-spin relaxation time, T2, measurements were conducted using a Techmag Apollo Submit before 31 May 2011 to
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2
Oviedo ICCS&T 2011. Extended Abstract
Pulse NMR spectrometer equipped with a CSIRO Australia produced high temperature 1
H-NMR probe.
The solid echo pulse sequence3), 90x-τ-90y, was employed to
generate 1H-NMR transverse relaxation signals. The advantage of the solid echo pulse sequence is that loss of information from the rapidly decaying free induction decay, FID, signals, which is encountered due to dead time, can be avoided. Thus, it becomes possible to observe the entire transverse relaxation signal. Theoretically, for sufficiently small τ compared with the transverse relaxation time T2, the peak amplitude of the solid echo is proportional to the total hydrogen content of the specimen and the subsequent signal decay is a full representation of the transverse relaxation4).
3. Results and Discussion 3. 1. Data analysis Change in the signal intensity, It, with decay time, t, is defined by Equation 1. It = I0 exp[-(t/T2)mi] mi=1 for exponential, mi=2 for Gaussian
1)
where I0 is the intensity of the signal at decay time t=0 corresponding to the number of protons in the specimen, and T2 is the spin-spin relaxation time reflecting the mobility of molecules in which the protons are embedded. When the specimen consists of several components, the obtained signal can be deconvoluted on the basis of the relaxation
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3
Oviedo ICCS&T 2011. Extended Abstract
characteristics and the signal can be described as a sum of each of the components. ItTotal = ∑Iti 2) where Iti is signal intensity for component i described in Equation 1 and ItTotal is the sum of the components corresponding to the observed signal. The spin-spin relaxation time, T2, and the fractional intensity, fH, of
250000 Observed Hm component Hint component Him component Fitting curve
each component was calculated according to measurements,
each
1
H-NMR relaxation component
was
deconvoluted on the basis of relaxation characteristics of protons which are closely related to the molecular mobility, so
200000
Intensity, a.u.
Equation 1. In the
150000 100000 50000
therefore the term “component” here does
0 0
not refer directly to particular molecules or groups of molecules. Figure 1 depicts a typical
solid
echo
signal
and
curve
Coal-C at 420℃
200
400
600
800
1000
Decay Time, µs Figure 1 Typical solid echo signal and curve deconvolution result of coal.
deconvolution result of coal. 3. 2. Effects of coal blending The echo signals of coal blends obtained during heat treatment were deconvoluted into three components: the rapidly decaying component, Him, described with a Gaussian function arising from the proton attached to immobile components in the coal; an intermediate component, Hint; and the most slowly decaying tail of the signal, Hm, described with an Exponential function arising from the proton attached to mobile components. It has already been ascertained that the deconvolution of the signal into three components is appropriate for evaluating the thermoplastic phenomenon of coal upon heating, and also that the thermoplastic phenomenon is closely related to the change of the Hm component, which has the highest molecular mobility5). Change in the fractional intensity of the Hm components, fHm, calculated from the signals during heat treatment are presented in Figure 2. For both of Base-1 and Base-2, the fluidity determined by the Gieseler plastometer varies in agreement with the variation in the fractional intensity of Hm, and the temperature of the maximum in the fluidity curve corresponds to the temperature at which fHm shows its maximum value. The fHm for Base-1 which consists of 75wt% of strongly caking coals shows higher value than for Base-2. The temperature at which fHm for Base-1 shows its maximum value is about Submit before 31 May 2011 to
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Oviedo ICCS&T 2011. Extended Abstract
20˚C lower than that for Base-2. The temperature dependence of Spin-Sipn relaxation time for the Hm, T2Hm, is shown in Figure 3. In the thermoplastic temperature range, the T2Hm for Base-1 maintains larger value than for Base-2. These results indicate that the necessity of an ample amount of strongly coking coal for sufficient termoplasticity of coal blend, furthermore, the thermoplasticity of coal is affected by the quantity and the quality of Hm component.
fHm, -
Base-1 Base-2
0.3
4.0 3.0
0.2
2.0
0.1
1.0
0 100
200
300
400
500
Base-1 Base-2
Log(Fluidity/ddpm)
0.4
80
5.0 Base-1 Base-2
60
T2Hm, µs
0.5
20
0 100
0.0 600
200
300
400
500
600
Temperature, ℃
Temperature, ℃ Figure 2 Correspondence between fHm and Gieseler fluidity as a function of temperature.
40
Figure 3 Temperature dependence of T2Hm for coal blends.
3.3. Effects of HPC addition As described above, compared with Base-1 which consist of 75wt% of strongly coking coal, the thermoplasticity of Base-2 is insufficient. In order to improve the thermoplastic properties of Base-2, coal-B was replaced with HPC. Temperature dependence of the fHm for Base-2 and Base-2 with HPC is plotted in Figure 4. The fHm for Base-2 increase with increase in the amount of HPC addition. The maximum temperature of fHm shifts to lower temperature with HPC addition. The fHm for B1+HPC10% shows almost the same maximum value as the fHm for Base-1 which is shown in Figure 2. Figure 5 depicts the variation of T2Hm for Base-2 and HPC added Base-2. In addition to the T2Hm peak at around 440˚C, new peak at below 400˚C appear with HPC addition. This peak might be result from thermal plasticity of HPC. The T2Hm value increases with increase in the amount of HPC addition. Thus, the thermoplastic properties of Base-2 have been improved not only quantitatively but also qualitatively with HPC addition. The strength of coke can be evaluated with drum index, DI, number. The DI15015 means the percentage of remaining +15mm square hole after the impact of 150 revolutions. The DI15015 for Base-2 is improved from 77.4 to 85.4 with 10wt% of HPC addition, Submit before 31 May 2011 to
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Oviedo ICCS&T 2011. Extended Abstract
B2+HPC10%. The DI15015 for B2+HPC10% is comparable to that for Base-1 of 84.0. The improvement of the coking phenomenon of Base-2 with HPC addition can be attributed to the quantitative and qualitative change of Hm component.
fHm, -
0.4 0.3
80
Base-2 B2+HPC5% B2+HPC10% B2+HPC15% B2+HPC20%
60
T2Hm, µs
0.5
0.2
20
0.1 0.0 100
40
200
300
400
500
0 100
600
Temperature, ℃ Figure 4 Temperature dependence of fHm for Base-2 and Base-2 with HPC.
Base-2 B2+HPC5% B2+HPC10% B2+HPC15% B2+HPC20% 200
300
400
500
600
Temperature, ℃ Figure 5 Variation of T2Hm for Base-2 and HPC added Base-2.
4. Summry The information obtained by means of 1H-NMR relaxation measurements of coal blend with HPC are listed as follows. 1. The thermoplastisity of coal can be evaluated both quantitatively and qualitatively with high temperature in-situ 1H-NMR relaxation measurement. 2. There is a needs for an ample amount of strongly coking coal for sufficient termoplasticity of coal blend, furthermore, the thermoplasticity of coal is affected by the quantity and the quality of Hm component. 3. The thermoplastic properties of coal blend are able to improve not only quantitatively but also qualitatively with HPC addition. 4. The improvement of the coking phenomenon of coal blend with HPC is attributed to the quantitative and qualitative change of Hm component. Acknowledgement This study was carried out as a part of Japanese national project called COURSE50 (CO2 Ultimate Reduction in Steelmaking process by innovative technology for cool Earth 50). We thank NEDO for the support of this study.
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Oviedo ICCS&T 2011. Extended Abstract
References 1. van D.W.Krevelen: Coal, 3rd revised ed., Elsevier, Amsterdam, (1993), Chapters 23, 24 2. L.J.Lynch, D.S.Webster, N.A.Bacon and W.A.Barton: Magnetic Resonance; Introduction, Advanced Topics and Applications to Fossil Energy, NATO Adv. Study Inst. Ser., ed. By L.Petrakis, J.P.Fraissard, Reidel Publ., Dordrecht, Netherlands, (1984), 617 3. J.G.Powles and P.Mansfield: Phys. Lett., 2(1962), 58 4. J.G.Powles and J.H.Strahge: Proc. Phys. Soc. London, 6(1963), 82 5. H. Kumagai, K. Tanabe, and K.Saito: 1H-NMR study of Relaxation Mechanisms of Coal Aggregate in Structure and Thermoplasticity of Coal, Nova Science, (2005), 35
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Oviedo ICCS&T 2011. Extended Abstract
A Systematic Study of the effects of Pyrolysis Conditions on Coal Devolatilisation M A Kochanek1, D G Roberts*1, B Garten2, S Russig2, and D J Harris1 1 2
CSIRO Energy Technology, Brisbane QLD, AUSTRALIA
Department of Energy Process Engineering and Chemical Engineering, TU Bergakademie Freiberg, GERMANY
ABSTRACT The effects of pressure and temperature on specific aspects of the devolatilisation process are generally understood; however, there remains considerable uncertainty associated with any predictions of coal pyrolysis behaviour under intense, entrained flow gasification conditions. The net effects of temperature, pressure, and heating rate on volatile yields and char structure are difficult to predict with any certainty, and how such behaviour is affected by coal properties is also unclear. This paper, as part of a wider study into high pressure, high temperature coal devolatilisation and char formation, presents some results of a systematic study into the effects of temperature and pressure on volatile yields and char properties. New data on volatile yields as a function of temperature and pressure are presented, and chars formed are studied for their physical characteristics and reactivity to CO2 and H2O. Whilst the separate effects of temperature and pressure on volatile yields are, in general, consistent with previous work, it is shown that the effects of increasing pressure and temperature are strongly influenced by coal type, and that there is some interaction between the two (such that effects of pressure are different at different temperatures). More work is required to clarify the impacts of these two important parameters on coal pyrolysis behaviour and char formation.
1
INTRODUCTION
The competing effects of pressure, temperature, and heating rate on coal devolatilisation make understanding coal-specific pyrolysis behaviour under gasification conditions difficult. It is generally understood that increasing pressure decreases volatile yields, and that increasing temperature increases volatile yields. The coal-specific nature of these effects, however, means that predicting the net outcome of these (and other) phenomena for a specific coal (under specific gasification conditions) is a difficult task.
*
Corresponding author:
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Oviedo ICCS&T 2011. Extended Abstract This issue is particularly important when models of the gasification system require such information so they can be used with some confidence for fuel assessment or process optimisation studies. Volatile matter data from proximate assays are often used as part of an assessment of a coal for use in entrained flow gasification; however, the significant differences in the conditions of the standard to those found in practical gasifiers makes this an unreliable indicator. It has also been shown that the reactivity of the chars produced can be as significant to overall gasification behaviour as the volatile yield, therefore understanding the impact of pyrolysis conditions on char properties is also important. This paper presents some new experimental measurements made as part of a wider study whereby the effects of temperature, heating rate, and pressure on volatile yields and char properties are investigated using a suite of coals ranging in rank from lignite to semi-anthracite.
2 2.1
EXPERIMENTAL Coal selection
Two separate studies were performed. In the first study, the volatile yields of six coals with ranks ranging from lignite to semi-anthracite were measured as a function of pyrolysis temperature and pressure. In the second study, the reactivity and surface area of selected chars formed from two (lignite and sub-bituminous) of the coals investigated in the first study were measured. The coals used in these studies were crushed and sieved to the particle size range -260+105 μm. The results of proximate analysis of the coals used in the volatile yield study are shown in Table 1. Rank
Lignite
Coal
TUF102 40.6 5.1 30.8 23.5 8.6 51.9 39.6
Moisture % (ar) Ash % (ar) VM % (ar) FC % (ar) Ash % (db) VM % (db) FC % (db)
Sub‐Bituminous
CRC252 8.7 10.1 40.1 41.1 11.1 43.9 45.0
CRC704 7.0 9.7 42.4 40.9 10.4 45.6 44.0
CRC701 25.0 5.3 28.3 41.4 7.1 37.7 55.2
Bituminous
Semi‐ Anthracite
CRC272 4.0 9.6 34.8 51.6 10.0 36.3 53.8
CRC703 6.6 9.5 7.4 76.5 10.2 7.9 81.9
Table 1: Proximate analysis of coals used in this study.
2.2
High Pressure Pyrolysis: Wire Mesh Reactor
In the first study the volatile yields of the six pulverised and sieved coals were performed in a Wire Mesh Reactor (WMR) using a procedure similar to that outlined previously [1, 2]. The following conditions were used to produce six volatile yield measurements for each coal: • • •
Pressure: 1bar, 10bar and 20bar Heating rate: 1000°C/s Hold time: 10s
Oviedo ICCS&T 2011. Extended Abstract • • •
Hold temperature: 900°C, 1100°C Initial dry sample mass: 20 to 30mg Sweep gas: High purity helium flow rate of 6 L/min at standard conditions to prevent secondary reactions
In the second study, chars produced in the pyrolysis experiments from coals TUF102 and CRC704 were characterised further. For each sample, the surface area and low-temperature reactivity were investigated using established techniques [2, 3]. 2.3
Char Characterisation
2.3.1
Surface Area
The CO2 adsorption isotherm for each pyrolysis char was measured non-destructively using a Tristar II 3020 surface area analyser. Adsorption isotherm analysis using the DubininRadushkevich (DR) equation was used to determine each char’s surface area. Char sample masses of 200±100mg were required for the surface area analysis. To produce the required quantity of char, the chars from multiple identical WMR experiments were combined to form a single sample. 2.3.2
Gasification Reactivity
Char reaction rates were measured at low temperatures, under experimental conditions designed to allow determination of ‘intrinsic’ reaction rates; that is, those obtained under conditions where diffusion through char porosity does not limit the reaction rate. A thermogravimetric analyser (TGA) was used to measure these rates using techniques previously described [2, 3]. The following conditions were used to measure the gasification reactivity of the chars: • • • •
Pressure: 5bar and 15bar Reactant gas: CO2 or H2O at a flow rate of 3 L/min at standard conditions with 100% concentration Temperature: 900°C for CO2 and 800°C for H2O reactant gas Initial dry sample mass: 40±20mg
The specific reaction rate is calculated as a function of conversion using the equation:
rate = −
1 dw ⋅ w dt
g g-1 s-1
where w is the instantaneous mass of reacting sample. The reaction rates used for analysis are initial reaction rates, determined on a specific (as measured) and intrinsic (normalised to surface area) basis.
Oviedo ICCS&T 2011. Extended Abstract 3 3.1
RESULTS AND DISCUSSION Volatile Yields
Figure 1 shows the effects of pyrolysis pressure and temperature on the volatile yields of the six coals used in this study. The results are presented as a ‘volatile enhancement’, which is a representation of the difference between the yield measured under the experimental conditions (volatile yield) and that determined in a proximate analysis (volatile matter). Here, volatile enhancement is calculated by:
⎛ %VY − %VM ⎞ VE = ⎜ ⎟ × 100 %VM ⎝ ⎠ Where %VY is volatile yield as determined under experimental conditions, and %VM volatile matter as determined by proximate analysis performed to Australian Standard AS1038A.
Figure 1: Volatile yield enhancement for a range of coals as a function of pressure at 900°C and 1100°C and with a heating rate of 1000°C/s.
For the lignite, the volatile enhancement is positive (i.e. more volatiles are released under these conditions than suggested by the proximate analysis) and most significant of all the samples. However this does not seem to be an effect of temperature or pressure, suggesting that perhaps heating rate is important for this coal. The absolute vales of volatile enhancement for the subbituminous and bituminous coals were less than that for the lignite; however the effects of temperature and pressure were clearer. This is also true for the semi-anthracite. At 900°C, the samples are readily differentiated in terms of the volatile enhancement, with lower rank coals having a greater VE than the higher ranked coals. These coal-specific effects are significantly reduced at 1100°C, however, with all coals (except for the lignite) having similar VE data as a function of pressure. These results suggest that significant differences between samples in terms of proximate volatile matter need not necessarily translate to real differences in volatile release under process conditions. For lignite samples, fast heating rate pyrolysis is likely to lead to a significantly
Oviedo ICCS&T 2011. Extended Abstract greater release of volatiles. For bituminous and sub-bituminous samples the effect is less certain, with interactions apparent between the effects of temperature and pressure; however the differences between different coal types appears to be significantly reduced at the higher temperature. 3.2
Surface Areas
Figure 2 shows measurements of char surface area, determined using CO2 adsorption at 0°C, for chars made from CRC704 and TUF102. Both coals produce chars with high surface areas, consistent with gasification chars studied in previous work [4-6]. For chars from TUF102, there is a clear effect of increasing temperature, with chars made at 1100°C having significantly less surface area than chars made at 900°C. For CRC704, this effect is less apparent. There is no strong effect of devolatilisation pressure on char surface area. Previous work [6] has suggested that for specific coals, increasing devolatilisation pressure can increase the surface area of the char produced from pyrolysis (the previous work was undertaken using flow reactors, suggesting that there may be a role of secondary reactions and/or partial reaction in the trends observed). A slight increase in surface area with increasing pyrolysis pressure is notable for the chars made from TUF102 at 900°C; the remaining sets of data suggest little effect, with some evidence to suggest a decrease in surface area with increasing pyrolysis pressure for CRC704 chars made at 1100°C. Ongoing work using the more extensive set of samples discussed above may lead to more clarity around this issue.
Figure 2: Char surface area (CO2) as a function of pyrolysis pressure at 900°C and 1100°C.
3.3
Char Reactivity
Figure 3-Figure 6 present specific and intrinsic reaction rate measurements of chars made from TUF102 and CRC704, to CO2 (at 900°C) and H2O (at 800°C). The surface areas discussed in the previous section were used to calculate initial intrinsic rate data from the measured specific reaction rates. Chars from TUF102 are consistently more reactive than those from CRC704; this is consistent with the expectation that lignite chars are, in general, more reactive than chars from sub-
Oviedo ICCS&T 2011. Extended Abstract bituminous coals. Figure 3 and Figure 4 show that in general, increasing devolatilisation pressure decreases the measured reaction rate to both CO2 and H2O for chars made from both coals. This effect is quite clear for the CO2 data (Figure 3), and the data for H2O suggest that this effect is less significant for the chars made at higher temperatures (Figure 4). The effect of pyrolysis temperature is clear: consistent with the significant amount of work reported in the literature, chars made at high temperatures have lower reaction rates to CO2 and H2O than chars made at lower temperatures. Surface areas can account for some of this difference; however as will be shown in the next section, temperature is also likely to significantly affect the crystalline nature of the carbon matrix, and consequently the intrinsic reactivity of the chars.
Figure 3: Specific reaction rates of chars made from CRC704 and TUF 102 at 900°C and 1100°C, to 5 and 15 bar CO2 at 900°C.
Figure 4: Specific reaction rates of chars made from CRC704 and TUF 102 at 900°C and 1100°C, to 5 and 15 bar H2O at 800°C.
Oviedo ICCS&T 2011. Extended Abstract Intrinsic reaction rates are shown in Figure 5 and Figure 6. As with the specific rate data discussed above, the lignite chars are also more reactive than the coal chars in both reactants on an intrinsic basis, reinforcing the notion that factors other than surface area alone (such as catalysts, carbon crystallinity, etc) are significant in determining the reactivity of lignite chars. For chars reacting with CO2, increasing pyrolysis pressure at 900°C leads to a decrease in intrinsic reactivity, consistent with those observed above for specific reaction rate data. This observation is far less clear for the chars produced at 1100°C, suggesting some interaction between the effects of temperature and pressure on the pyrolysis process, in particular the char formation aspects. This will be explored in more detail in ongoing work.
Figure 5: Intrinsic reaction rates of chars made from CRC704 and TUF 102 at 900°C and 1100°C, to 5 and 15 bar CO2 at 900°C.
Figure 6: Intrinsic reaction rates of chars made from CRC704 and TUF 102 at 900°C and 1100°C, to 5 and 15 bar H2O at 800°C.
Oviedo ICCS&T 2011. Extended Abstract 4
CONCLUSIONS
This work has presented some results from a wider study investigating the pyrolysis behaviour of coals under entrained flow gasification conditions. This work focussed on the volatile yields (and how they are affected by temperature and pressure) and the reactivity of the chars produced. Consistent with previous investigations, increasing pyrolysis temperature increased the volatile yields. Whilst increasing the pyrolysis pressure generally decreases the volatile yields, the effect of pressure is strongly dependent on coal type and pyrolysis temperature. There is some evidence of interactions between these variables; for example, increasing temperature from 900 to 1100°C removes a lot of the influence of coal type on volatile yields, in particular for the sub-bituminous and bituminous coals used in this study. For the lignite sample, the volatile yields were significantly greater than those expected based on proximate analysis, although the data presented here showed very little impact of temperature or pressure. This suggests that heating rate is a particularly significant variable determining the volatile yields from lignite pyrolysis. Heating rate is not a variable that was explicitly studied here; however, it is a parameter that will be studied in more detail in ongoing work. Chars made at higher temperatures were less reactive to CO2 and H2O than those made at lower temperatures, consistent with literature studies. Increasing devolatilisation pressure generally led to less reactive chars (to both CO2 and H2O) on both a specific (as-measured) and intrinsic (normalised to surface area) basis. This was more apparent for chars made at 900°C than those made at 1100°C, and suggests a pressure effect on the carbon structure (or chemical composition) of the chars.
5
REFERENCES
1.
C J Mill, D J Harris and J F Stubington. Pyrolysis of Fine Coal Particles at High Heating Rates and High Pressure, 8th Australian Coal Science Conference, Sydney (1998).
2.
M D Kelly, D J Harris, D G Roberts and C J Mill. Volatile Yield and Char Gasification Reactivity of Australian Coals at Elevated Pressures, 18th Annual International Pittsburgh Coal Conference, Newcastle, Australia (2001).
3.
D G Roberts and D J Harris. The Role of Bench-Scale Reactivity Data in the Assessment of Coals for use in Gasification, 12th International Conference on Coal Science, Cairns, Queensland, Australia (2003).
4.
L M Lu, C H Kong, V Sahajwalla and D Harris. Char structural ordering during pyrolysis and combustion and its influence on char reactivity, Fuel 81 1215- 1225 (2002).
5.
D G Roberts and D J Harris. Char Gasification in Mixtures of CO2 and H2O: Competition and Inhibition, Fuel 86(17-18) 2672-2678 (2007).
6.
D G Roberts, D J Harris and T F Wall. On the Effects of High Pressure and Heating Rate during Coal Pyrolysis on Char Gasification Reactivity, Energy and Fuels 17(4) 887-895 (2003).
Oviedo ICCS&T 2011. Extended Abstract
A Relationship Between the Structures of Graphitized Anthracites and Isotropic Graphite M.S. Nyathi, C.E. Burgess Clifford, and H.H. Schobert Department of Energy and Mineral Engineering and The EMS Energy Institute, The Pennsylvania State University, University Park, PA, USA.
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Abstract Two anthracites from Pennsylvania (USA) were evaluated as possible filler materials for the production of isotropic or near-isotropic graphites. Native and demineralised samples were graphitized at 3000°. Products were characterized by an array of analytical techniques, including X-ray diffraction, Raman spectroscopy, and temperatureprogrammed oxidation. The anthracite of higher rank produced, upon graphitization, materials with better-developed crystalline structure. Demineralization of either anthracite prior to graphitization yielded products having better degree of graphitization, density, and oxidation resistance (in temperature-programmed oxidation). The enhancement of la in graphitization of native anthracites was facilitated by formation and subsequent decomposition of silicon carbide. Graphitizing the lower-rank sample gave a product of lower anisotropy, closer to an isotropic material that would be more desirable as starting material for a nuclear graphite. Chances of obtaining a near-isotropic material are also enhanced by demineralizing the samples before graphitization.
1. Introduction We and our colleagues have been investigating the non-fuel uses of anthracitic coals for some time, with a focus on the anthracites of Pennsylvania in the United States [1-6]. Much of our previous work has been directed to graphitization of molded or extruded formulations of anthracite and coal tar pitch, for use in electrode or specialty applications. Currently we are investigating the potential of producing isotropic, or nearisotropic graphites from anthracites for, e.g., applications as nuclear graphite. A general approach to the production of finished synthetic graphite artifacts involves use of a graphtizable carbonaceous solid, such as petroleum coke, as filler, and mixing the coke with a binder of coal tar pitch [7]. This mixture is formed into the desired shape, e.g. by extrusion or molding. The shaped item is baked. If needed, the baked item is impregnated with petroleum pitch and re-baked. Several impregnation and Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
re-baking cycles may be required. When the density of the baked item is adequate, it is graphitized, and finally, machined to the desired size. The purpose of the work reported here was to investigate the factors affecting the graphitization behavior of two anthracites to assess their potential for use as fillers in the production of nuclear graphite. We did not investigate the blending of the anthracites with coal tar pitch, nor measure some of the physical properties important for showing that a particular graphite is fit for use in nuclear applications. Thus the focus of this work is on evaluation of prospective new filler materials, and not on actually producing nuclear graphite.
2. Experimental Section Extensive details on the selection, characterization, graphitization, and analysis of the samples are available elsewhere [8]. Here we present a brief overview of the principal aspects of the experimental work. Samples. Two anthracites were used in this work. Their characteristics are given in Table 1. The samples PSOC-1515 and DECS-21 were obtained from the Penn State Coal Sample Bank. Demineralization of PSOC and DECS samples was performed following the earlier work of Pappano [6]. Table 1. Characteristics of the anthracites used in this study. Proximate analysis, dry basis, % Fixed carbon Volatile matter Ash Ultimate analysis, daf basis, % Carbon Hydrogen Nitrogen Sulfur Oxygen
DECS-21
PSOC-1515
84.34 4.51 11.15
62.39 8.44 29.17
90.3 4.0 0.8 0.6 4.3
88.1 3.9 1.1 0.8 6.1
Graphitization. Graphitization of the PSOC and DECS samples was performed at GrafTech, Parma, Ohio, USA, at was 3000°C. Analyses and characterization. X-ray diffraction analysis was performed using a PANalytical X’Pert Pro powder diffractometer and Cu-Kα radiation. Raman spectroscopic measurements were obtained using a WITec CRM200 argon-laser spectrometer. Temperature-programmed oxidation experiments were run in a LECO
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Oviedo ICCS&T 2011. Extended Abstract
RC612 multiphase carbon analyzer, with 750 mL/min oxygen flow, heat-up rate of 30°/min to a maximum temperature of 900°, and hold at 900° for 6 min. BET surface area was measured with a Micromeritics ASAP20 instrument, using nitrogen as the adsorbate at 78 K and densities with a Qunatichrome Multipycnometer MVP-1 in helium.
3. Results and Discussion X-ray diffraction. The X-ray diffraction results of graphitized native and demineralized PSOC and DECS anthracites show the characteristic diffraction patterns of graphite, Figure 1. DECS21‐DM DECS21 PSOC1515
Intensity, a.u.
PSOC1515‐DM
10
20
30
40
50
60
70
80
90
2 theta, degrees
Figure 1. X-ray diffractograms of graphitized native and demineralized PSOC-1515 and DECS-21 anthracites. All diffractograms show a sharp, intense, narrow (002) peak at 2Θ = 26.4°. Also important is the (112) peak at 2Θ ≈ 83°, since peaks for which all of the Miller indices are non-zero signal three-dimensionality in the structure. The structural parameters are summarized in Table 2.
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Oviedo ICCS&T 2011. Extended Abstract
Table 2. X-ray diffraction parameters for anthracites graphitized at 3000°. Samples identified with –DM suffix were demineralized prior to graphitization. Data on native anthracites are provided for comparison. Sample
Status
PSOC1515 PSOC1515-DM DECS21 DECS21-DM PSOC1515 PSOC1515-DM DECS21 DECS21-DM
Native Native Native Native Graphitized Graphitized Graphitized Graphitized
d002, nm 0.3522 0.3521 0.3521 0.3498 0.3371 0.3365 0.3369 0.3361
lc, nm
la, nm
DOG
p
N
lc /la
1.4 2.8 1.7 3.1 46.0 48.0 53.6 58.0
60.9 59.1 62.4 61.2
0.7930 0.8162 0.8244 0.8267
0.2069 0.1837 0.1755 0.1732
136 142 159 172
0.76 0.80 0.85 0.94
Heat treatment at 3000° provides marked improvement in d002 and lc. DECS-21 shows slightly better structural development, as indicated by, e.g., d-spacing, degree of graphitization (DOG), and probability of random orientation between layer planes (p). TEM observation of the native anthracites showed that DECS-21 had a better developed microstructure and more flattened pores than PSOC-1515. This is consistent with reports of a relationship between graphitizability and preferred planar orientation [9,10], and consistent with reports of relationships with porosity or pore shape [10-13]. For these two anthracites, slightly better structural development is obtained using the demineralized samples; i.e., there is no evidence for in situ catalytic graphitization by aluminum compounds, or aluminum-rich minerals. Although it is well established that some minerals catalyze graphitization of anthracites, in this case X-ray analysis showed that the minerals in these anthracites were dominated by kaolinite. It has been established that kaolinite does not catalyze graphitization [6], likely because it transforms to a highly stable mullite structure during calcination or heat treatment [13].. Formation of mullite was confirmed by X-ray analysis of the calcined anthracites. Nonetheless, a silicon carbide phase that formed during graphitization of the native anthracites, appeared to enhance la, consistent with earlier work [6] on catalytic graphitization of Pennsylvania anthracites. The measured densities of the graphitized samples agree well with the degree of graphitization (DOG) as measured by X-ray diffraction, as shown in Table 3.
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Oviedo ICCS&T 2011. Extended Abstract
Table 3. Relationship of densities and degrees of graphitization of graphitized anthracite samples. Sample Density, g/cm3 DOG PSOC-1515 1.78 0.7930 PSOC-1515-DM 1.85 0.8162 DECS-21 1.89 0.8244 DECS-21-DM 2.01 0.8267 Raman spectroscopy confirms the results obtained by X-ray diffraction. Table 4 provides a comparison of spectroscopic results for native and graphitized anthracites. Table 4. Raman spectroscopy parameters for native and graphitized anthracites Sample
Status
PSOC1515 PSOC1515-DM DECS21 DECS21-DM PSOC1515 PSOC1515-DM DECS21 DECS21-DM
Native Native Native Native Graphitized Graphitized Graphitized Graphitized
FWHM (cm-1) D-band G-band 387.93 74.43 386.76 73.86 386.11 73.44 385.07 72.69 50.86 25.04 53.25 27.86 51.68 23.86 50.12 24.42
ID / (ID+IG) (%) 42.03 41.78 41.52 40.44 9.69 13.22 5.54 9.15
Products from DECS-21 have lower disorder parameters than those produced from PSOC-1515. In addition, demineralization leads to an increase in the demineralization parameter. The disorder parameters measured by Raman spectroscopy relate to changes in the value of la, as shown in Figure 2.
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Oviedo ICCS&T 2011. Extended Abstract
Figure 2. Relationship of la measured by X-ray diffraction to Raman disorder factors. Temperature-programmed oxidation results are shown in Figure 3.
PSOC1515 DECS21
Intensity, a.u.
PSOC1515‐DM
100
DECS21‐DM
200
300
400
500 600 Temperature, oC
700
800
900
Figure 3. Temperature-programmed oxidation of graphitized samples of native and demineralized PSOC-1515 and DECS-21. Products from DECS-21 are less reactive (as indicated by a shift of peak position to Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
higher temperatures) than products from PSOC-1515. Further, products of graphitization of demineralized anthracites are less reactive than those from native anthracites. These oxidation results can be related to the degree of structural ordering, and in particular, the value of N, the number of carbon layer planes in a stacked crystallite. This relationship is illustrated in Figure 4.
Figure 4. Relationship between the oxidation reactivity, as measured by peak position in temperature programmed oxidation, and the number of stacked layer planes.
4. Summary and Conclusions The DECS-21 sample, which is of higher rank than PSOC-1515, has a more developed lamellar structure. This is particularly noticeable in transmission electron microscopy, results of which are presented elsewhere [8]. Upon graphitization, the anthracite with the more developed structure, i.e., DECS-21, produces materials with better-developed crystalline structure. Demineralization of either anthracite prior to graphitization yields products having better degree of graphitization, density, and oxidation resistance (in temperature-programmed oxidation). However, the enhancement of la in graphitization of native anthracites appears to be facilitated by formation and subsequent decomposition of silicon carbide, consistent with earlier work from our group [6]. Graphitizing the lower-rank sample, i.e., PSOC-1515, this gives a product of Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
lower anisotropy, closer to an isotropic material that would be more desirable as starting material for a nuclear graphite. Chances of obtaining a near-isotropic material are also enhanced by demineralizing the samples before graphitization.
Acknowledgements Various aspects of this work were supported by the U.S. Air Force Office of Scientific Research, and the U.S. Department of Energy Consortium for Premium Carbon Products from Coal, and we gratefully acknowledge this support.
References [1] Gergova K, Eser S, Schobert HH. Preparation and characterization of activated carbon from anthracite. Energy Fuels 1993;7:661-8. [2] Gergova K, Eser S, Schobert HH, Klimkiewicz M, Brown PW. Environmental scanning electron microscopy of activated carbon produced from anthracite by one-step pyrolysis activation. Fuel 1995;74:1042-8. [3] Atria JV, Rusinko F, Schobert HH. Structural ordering of Pennsylvania anthracite on heat treatment to 2000-2900°C. Energy Fuels 2002;16:1343-7. [4] Andrésen JM, Burgess CE, Pappano PJ, Schobert HH. New directions for non-fuel uses of anthracite. Fuel Proc. Technol.2004;85:1373-92. [5] Pappano PJ, Rusinko F, Schobert HH, Struble DP. Dependence of physical properties of isostatically molded graphites on crystallite height. Carbon 2004;42:30079. [6] Pappano PJ, Schobert HH. Effect of natural mineral inclusions on the graphitizability of a Pennsylvania anthracite. Energy Fuels 2009;23:422-8. [7] Pierson, HO. Handbook of carbon, graphite, diamond and fullerenes. Park Ridge, NJ: Noyes; 1993. [8] Nyathi, MS. Evaluation of coal-petroleum blend coke and anthracites as precursors to isotropic or near-isotropic graphite. PhD Dissertation, University Park: The Pennsylvania State University; 2011. [9] Oberlin A, Terriere G. Graphitization studies of anthracites by high resolution electron microscopy. Carbon 1975; 13: 367-76. [10] Blanche C, Rouzaud JN, Dumas D. Proceedings of the 22nd biennial conference on carbon. American Carbon Society 1995;695. [11] Atria JV. Novel approach to the production of graphite from anthracite. MS Thesis, Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
University Park: The Pennsylvania State University; 1995. [12] Pusz S, Kwiecinska BK, Duber S. Textural transformation of thermally treated anthracites. Int. J. Coal Geol. 2003;54:115-23. [13] Gonzalez D, Montes-Moran MA, Suarez-Ruiz I; Garcia AB. Structural characterization of graphite materials prepared from anthracites of different characteristics: a comparative analysis. Energy Fuels 2004;18:365-70.
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Oviedo ICCS&T 2011. Extended Abstract
Benzene and toluene adsorption on high surface area activated carbons obtained from an anthrecene oil derivative N.G. Asenjo1, P. Álvarez, C. Blanco, R. Santamaría, M. Granda, R. Menéndez Instituto Nacional del Carbón (CSIC), c/ Francisco Pintado Fe, 26 – 33011 Oviedo (Spain) – Tel. +34 985119090; Fax. +34 985297662 1
[email protected]
Abstract A carbon obtained from a coal-derived pitch was chemically activated to produce a high surface area (∼3,200 m2/g) carbon with a porosity made up of both micropores and mesopores. Its adsorption capacities under the conditions tested in this work were found to be among the highest ever reported in literature, reaching values of 860 mg/g and 1,200 mg/g for the adsorption of benzene and toluene, respectively, and 1,200 mg/g for the combined adsorption of benzene and toluene from an industrial wastewater. Such high values are only possible if the entire pore system, including the mesopore fraction, is involved in the adsorption process. The filling of almost the entire pore system is thought to be due to the high relative concentrations of the tested solutions, resulting from the low saturation concentration values for benzene and toluene, which were obtained by fitting the adsorption data to the BET equation in liquid phase. The kinetics of adsorption in the batch experiments which were conducted in a syringe-like adsorption chamber was observed to proceed in accordance with the pseudo-second order kinetic model. The combined presence of micropores and mesopores in the material is thought to be the key to the high kinetic performance, which was outstanding in a comparison with other porous materials reported in the literature.
1. Introduction The main objective of this work is to analyze the performance of an anthracene oil-based activated carbon in the liquid phase adsorption of benzene and/or toluene from synthetic solutions and an industrial wastewater. The total adsorption capacities will be analyzed and compared with those reported in the literature for other systems tested under gas and liquid phase conditions. Special emphasis will be placed on a kinetic analysis of the adsorption process, which will serve as a basis for comparison with other adsorbent materials described in the literature. A combination of equilibrium and kinetic analyses
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allows us to conclude that, under the conditions of this study, the coal-derived activated carbon offers the best results ever reported in terms of adsorption capacity and kinetic rate
2. Experimental Section To produce the activated carbon an anthracene oil-based coke was subjected to chemical activation with anhydrous KOH at a weight ratio of 1:5, carbonized at 700 ºC in N2, and thoroughly washed with HCl and water. The textural characteristics of the samples were characterized by means of N2 adsorption at 77 K and CO2 adsorption at 273 K. For the kinetic experiments a set up comprising a purpose-designed syringe-like adsorption chamber was employed. Four different aqueous solutions were analyzed: synthetic aqueous solutions of (i) benzene (190-210 ppm), (ii) toluene (175 225 ppm), (iii) a mixture of both (175-225 ppm each) and (iv) an industrial wastewater from a local chemical company containing a mixture of benzene (~120 ppm) and toluene (~120 ppm), together with trace amounts of chloroform. The collected liquid samples were analyzed by means of High Performance Liquid Chromatography (HPLC) (Agilent 1100 series apparatus) and UV spectrometer (Shimadzu UV 1800). The amount of benzene and toluene adsorbed on the activated carbon at equilibrium, be, and the concentration of adsorbate in solution at equilibrium, Ce, were obtained from the kinetic data experiments at the longest time (always over 10 min).
3. Results and discussion The activated carbon has a very high volume of micropores, slightly over 1 cm3/g and a high mesopore volume (~0.6 cm3/g) which is a specific characteristic of the material developed here. At the other end of the pore distribution, the narrow microporosity (below 0.7 nm) displays a total volume of ~0.6 cm3/g. Figure 1 shows the benzene and toluene adsorption isotherms for the aqueous solutions. The results for the industrial wastewater (Figure 1, solid squares) indicate that benzene was adsorbed up to ~400-500 mg/g, whereas toluene was simultaneously adsorbed up to ~700 mg/g. Therefore, the maximum combined adsorption of both aromatic species was around 1,200 mg/g. Similar values were obtained for the synthetic mixed solution (Figure 1, empty squares), indicating that the presence of other substances in the industrial wastewater did not seem to significantly affect the performance of the adsorbent.
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Single component Synthetic water (210 ppm toluene)
1000
Industrial wastewater (115 ppm toluene)
800 be (mg/g)
C s=1,800 ppm (solubility)
600
Cs=223 ppm
400 200 Benzene adsorption isotherms 0 0
20
40
60
80
100
120
140
Ce (ppm) Single component Synthetic mixture (200 ppm benzene) Industrial wastewater (120 ppm benzene)
1400 1200
Cs=470 ppm (solubility)
be (mg/g)
1000
C s=69 ppm
800 600 400 200
Toluene adsorption isotherms 0 0
10
20
30
40
50
60
Ce (ppm)
Figure 1. Benzene and toluene adsorption isotherms for the activated carbon analyzed in different adsorption media. Lines through the single component points represent fittings to BET equation for Cs values equal to the solubility values of the single components in water at RT (dashed lines) and for Cs values that maximize the R2 coefficient (continuous lines)
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For the single component solutions (Figure 1, solid circles), the maximum amounts of adsorbed benzene and toluene (non competitive adsorption) reached ~870 and ~1,200 mg/g, respectively. These results evidence the high adsorption performance of this activated carbon and its ability to remove both molecules from polluted water streams with equal efficiency. Furthermore, these high capacities prove that adsorption is not restricted to the narrow microporosity of the activated carbon but involves the entire pore system. The use of the entire porosity of the adsorbent used in this study might imply high values of the relative concentrations (Ce/Cs) of benzene and toluene in the solutions employed in this work. Figure 1 shows the fitting of BET equation to the adsorption curves for the single components in two different cases; (a) where Cs=S and (b) where Cs is a variable that is modified to maximize the value of R2. Only when Cs is adjusted to maximize R2, the fittings are acceptable. By applying the Cs values obtained in this case (Figure 1), the relative concentrations for the highest values of adsorption capacity become 0.53 and 0.63 for benzene and toluene, respectively. Thus, in terms of the adsorption isotherm, the liquid phase equilibria take place at relatively high values of relative concentration, making the adsorption of aromatics in liquid phase more similar to that in saturated gas streams than to that in diluted gas streams. Consequently almost total occupation of the activated carbon pores can be expected, as our results indeed confirm. The results from all the batch kinetic experiments were found to follow a pseudo-second order kinetic model with the kinetic values obtained from the pseudo-second order kinetic model, it is possible to compare the kinetic performance of the anthracene oilbased activated carbon with that of other materials reported in the literature [1-6]. For the purpose of this comparison the best data representation is that offered by Figure 2, which shows the variation of kSE with be for toluene adsorption. Points on the right of a given isokinetic line (diagonal lines) are thought to represent faster adsorption systems than points on the left of the same isokinetic line. The effect of the presence of benzene in solution does not affect the toluene adsorption rates of our activated carbon; in fact the kinetic performance of the activated carbon is almost independent of the type of solution analyzed. In this figure, the kinetic points obtained for the different materials are grouped according to their main textural properties. From the pseudo-isokinetic lines, it can be seen that the adsorption rates of the microporous materials are lower than
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Oviedo ICCS&T 2011. Extended Abstract
Figure 2. Variation of the pseudo-second order kinetic rate constant (kSE) with the value of be calculated by means of the PSOKM for the toluene adsorption experiments carried out in this work in different adsorption media (legend), together with the kinetic data obtained from the literature values. that of the mesoporous material, and that this is in turn lower than the adsorption rates achieved by the coal derived activated carbon used in this study, whose porosity is made up of both micropores and mesopores. The results of our work suggest that mesopores act both as a transport and as a concentrating medium and thus enhance the adsorption rate in the micropore system. This could be the explanation for the outstanding performance of the coal-derived activated carbon in the adsorption of benzene and toluene from aqueous solutions.
4. Conclusions The benzene and toluene adsorption capacity of the anthracene oil-based activated carbon is among the highest ever reported in the literature (860 mg/g for benzene, 1,200 mg/g for toluene and 1,200 mg/g for a mixture of both molecules in solution). The extensive pore filling is thought to be due to the high relative concentrations of the tested solutions, resulting from the low saturation concentration values for benzene and
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toluene. The adsorption capacities of the activated carbon were similar both for the synthetic mixtures and for the industrial wastewater. The kinetics of adsorption in the batch experiments was observed to follow the pseudo-second order kinetic model. The combined presence of micropores and mesopores in the activated carbon is thought to be the key to its high kinetic performance, which can be described as outstanding when compared with other porous materials reported in the literature.
Acknowledgments The authors thank the Spanish Science and Innovation Ministry (CONSOLIDER INGENIO 2010, Ref. CSD2009-00050) for financial support and Dr. Patricia Álvarez for her Ramón y Cajal contract.
References [1] Lesage G, Sperandio M, Tiruta-Barna L. Analysis and modelling of non-equilibrium sorption of aromatic micro-pollutants on GAC with a multi-compartment dynamic model. Chem Eng J 2010:160(2):457-65. [2] Choi JW, Yang KS, Kim DJ, Lee CE. Adsorption of zinc and toluene by alginate complex impregnated with zeolite and activated carbon. Current Applied Physics 2009;9(3):694-7. [3] Lee SJ, Chung SG, Kim DJ, Lee CE, Choi JW. New method for determination of equilibrium/kinetic sorption parameters. Current Applied Physics 2009;9(6):1323-5. [4] Arora M, Snape I, Stevens GW. The effect of temperature on toluene sorption by granular activated carbon and its use in permeable reactive barriers in cold regions. Cold Regions Science and Technology 2011;66(1):12-6. [5] Su F, Lu C, Hu S. Adsorption of benzene, toluene, ethylbenzene and p-xylene by NaOCl-oxidized carbon nanotubes. Colloids and Surfaces A: Physicochemical and Engineering Aspects 2010;353(1):83-91. [6] Wang D, McLaughlin E, Pfeffer R, Lin YS. Aqueous phase adsorption of toluene in a packed and fluidized bed of hydrophobic aerogels. Chem Eng J 2011;doi: 10.1016/j.cej.2011.02.014.
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Oviedo ICCS&T 2011. Extended Abstract
Impact of biomass on char burn-out under air and oxy-fuel conditions
Timipere S. Farrow, Donglin Zhao, Chenggong Sun and Colin E. Snape
Department of Chemical & Environmental Engineering, Faculty of Engineering, University Park, Nottingham NG7 2RD, UK
[email protected] Abstract Although biomass co-firing is now well established in pulverised fuel (PF) combustion, there is little information available on how biomass affects coals during oxy-fuel firing. Therefore, this study involving thermo gravimetric analysis (TGA) and a drop tube furnace (DTF) examines the impact of co-firing biomass and coal under oxy-fuel and air fired conditions with particular emphasis on the catalytic effect of biomass-contained alkali and alkaline metals on coal char burnout. Sawdust and a South African coal, Kleinkopje have been used.
For the burn-out studies, the sawdust and coal chars
prepared in the DTF and under slow-heating conditions at different temperature were blended using a 50:50 mass ratio. The addition of the sawdust chars to the coal chars improved burnout and the effect was slightly more pronounced under oxy-fuel conditions. To confirm that the improved burnout rates arise from the catalytic effects of inorganic alkali and alkaline metals, the sawdust was extracted with 5M hydrochloric acid. The burn-out rates for the blends were reduced considerably and these were close to the predicted values.
1. Introduction For effective utilisation of the biomass and coal as pulverised fuel (PF), compatibility of the fuels during combustion is desirable and chemical interactions between the two fuels during co-firing could increase the carbon burnout of coal char, allowing the fly ash to meet up specification for other uses. Interactions between biomass and coal have been investigated during co-pyrolysis. For example, Haykiri-Acma and Yaman [1, 2] reported that synergetic interaction occurred during combustion of biomass and coal in air using TGA from ambient to 900oC. This effect has resulted in higher volatile yields [3, 4] . Presumably, the improved product yields may be dependent upon the contact time of the fuel particles, and the relative rates of pyrolysis of the different fuels, though blending
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ratios, temperature and hydrogen content in biomass could also influence interaction between the two fuels [5, 6]. While the synergetic interaction between biomass and coal in co-pyrolysis has been extensively studied, the limited literature exists on the impact of biomass char on coal char burnout during co-combustion or gasification. However, Kastanaki et al [7] assessed the combustion of biomass-coal char blend using non isothermal TGA with two coal types and different biomass fuels over a temperature range of 20-850 oC and at slow heating rate with biomass/coal chars blend ratios of 5:95; 10:90; 20:80 wt%. The results showed that the burnout times of the coal char was slightly reduced and the burnout (final) temperatures were lowered by 22-45 oC for 20 wt% biomass blends. Similar reports on biomass/coal char combustion and gasification have revealed some interaction depending on biomass type [8, 9]. However, these studies have not considered the effect of DTF biomass chars which simulates pulverised fuel combustion better than TGA with high temperatures, high heating rates and short residence times for coal char burnout being achieved. Additionally, the improved char reactivity which is likely to be the potential catalytic effect of biomass contained alkali and alkaline metals have not been fully investigated. Therefore, this paper examines how coal char burnout will be affected by the addition of biomass char prepared at high temperature and high heating rate during co-firing and the catalytic effect of biomass-contained alkali and alkaline metals. TGA reactivity assessment of the blend chars will be investigated in terms of burnout times and reaction rate constants for low and high heating rate chars. To prove the catalytic influence of mineral matter in biomass on coal char combustion, demineralised sawdust/coal char were investigated. Synergetic effects will be examined by comparing the predicted behaviour of the mixture with the experimental data available for the single fuels.
2. Experimental section 2.1 Sample Preparation Sawdust and pine wood chars were produced in a laboratory horizontal tube furnace (HTF) by imitating TGA devolatilisation conditions in Nitrogen and Carbon dioxide atmospheres respectively, while coals chars were produced in the HTF at 900 – 1100oC DTF coal char was produced from South African coal, Kleinkopje (KK) 53-75 µm particle size at 1100 - 1300oC 200 and 600 ms in a nitrogen atmosphere with 1 % oxygen to prevent tar acumulation. Detailed description of the HTF and DTF char preparation methods can be found elsewhere[10]. Properties of the samples used are presented in Submit before May 15th to
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Table 1. 2.2 Char blending methods Sawdust char produced from 125-250 µm was used to blend with coal char. The two chars were measured and manually mixed together in a sample bottle. The ratio of blending of biomass/coal was 10:90 wt%, 25:75 wt% and 50:50 wt%. 2.3 TGA analysis About 2-5 mg of char is heated at the rate of 50oC min-1 to burnout temperature under nitrogen to remove possible moisture and volatiles and then the furnace is switched to air at 500 oC for 80 minutes for combustion. From the TGA burnout profiles the reaction rate constants were calculated between 5-95 % conversion levels and burnout times up to 90 % carbon conversion were taken. 2.4 Demineralisation of sawdust. Removal of inorganic minerals in the sample was carried out by adding 1g of the sample in 50 ml of 5 M hydrochloric acid solution. The mixture was stirred at 60 oC for 12hrs and then allowed to cool. It was then filtered and the sample was washed with deionised water until the acid is neutralised. Char was produced from the demineralised sample for co-blending and TGA analysis as stated above. Also ash samples were produced in a furnace at 350 oC in air for ICP elemental analysis.
Table 1. TGA proximate analysis of the samples used Sample
Moisture
Volatiles
Fixed carbon Ash
(wt %)
(wt % daf)
(wt % daf)
(wt %)
Raw sawdust
6.4
83.9
16.1
0.9
5M HCl washed sawdust
8.6
87.3
12.7
0.04
Kleinkopje (KK)
3.1
36.6
63.4
15.5
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Figure 1: Schematic diagram of the experimental approach
3. Results and Discussion
3.1 Combustion reactivity of low heating rate sawdust and coal char blend
Figures 2 and 3 show carbon burnout profiles of sawdust and coal chars produced at low heating rate and their 50:50 wt% blends under air and oxy-fuel fired conditions respectively. The plots demonstrate how the burnout of coal char is greatly improved by the addition of sawdust chars under both conditions but slightly more pronounced under oxy-fuel firing. This is further illustrated in Table 2 with the reaction rate constants and the 90% carbon burnout times. The difference between the experimental burnout profile of the blend and the predicted burnout is an indication of a synergetic effect during cofiring.
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100
KK char N₂ 1000⁰C
Carbon burnout (wt%)
Predicted 50:50 blend 80
50: 50 wt% blend sawdust N₂ char 700⁰C
60
40
20
0 0
20
40
60
80
Time (min)
Figure 2. TGA burnout profiles of sawdust and coal chars produced at low heating rate and their blend under air fired conditions 100
Carbon burnout (w%)
KK HTF CO₂ char 1000⁰C Predicted 50:50 wt% Blend
80
Expt 50-50 wt% blend sawdust HTF CO₂ char 700⁰C
60
40
20
0 0
20
40
60
80
Time (min)
Figure 3. TGA burnout profiles of sawdust and coal chars produced at low heating rate and their blend under oxy-fuel fired conditions
Table 2. Comparison of combustion reactivity of low heating rate sawdust, coal char and their blend Samples
N2 chars and air firing
CO2 chars and oxy-fuel firing
Rate constants 90 % burnout Rate constants 90 % burnout (min-)
time (min)
(min-)
time (min)
50:50wt% blend
0.1002
22.15
0.1089
20.65
Predicted blend
0.0829
25.60
0.0720
31.60
Co-firing pine wood char with coal char also exhibited similar improved burnout
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Oviedo ICCS&T 2011. Extended Abstract
behaviour of coal char as presented in Figures 3 and 4. Here two particle size fractions were used to investigate the effect of particle during co-firing. The burnout profiles indicated that particle size had no significant effect for the particle sizes considered in this study. A possible explanation for the improved coal char burnout propensity lies in the alkali and alkaline metals contained in biomass fuel which catalysed the reactions. It therefore means that all biomass fuels containing alkali and alkaline metal will potentially give similar effect during co-firing. 100 Coal char 106-150 @1100⁰C Predicted 50:50 wt% Blend
Carbon burnout (wt%)
80
50:50 wt% Blend Pinewood 90-106 @700⁰C
60
40
20
0 0
20
40
60
80
Burnout Time (min)
Figure 4. TGA carbon burnout profiles of pine (90-106 µm) and coal chars and their blends 100
Carbon burnout (wt%)
Coal char 106-150 @1100⁰C Predicted 50:50 wt% Blend
80
50:50 wt% Blend Pinewood 125-250 @700⁰C
60
40
20
0 0
20
40
60
80
Burnout Time (min)
Figure 5. TGA carbon burnout profiles of pine (125-250µm) and coal chars and their blends
3.2 Combustion reactivity of High heating rate sawdust and coal chars In order to produce chars which correlate directly to chars produced in a pulverised
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Oviedo ICCS&T 2011. Extended Abstract
furnace, both sawdust and coal chars were produced in the DTF with high heating rate, high temperatures and at different residence times under air fired and oxy-fuel conditions (Figures 6 and 7, respectively). The burnout profiles of the blend demonstrate that sawdust char improved the burnout of coal char even at chars produced closer to pf furnace. However, the burnout of coal in oxy-fuel condition was about 2 times faster. This is because CO2 promotes both combustion and gasification reaction. Additionally, the CO2-char gasification reaction activates the char surface areas for easy accessibility of oxygen into the char active sites for combustion. 100 KK DTF N₂ char 1300⁰C 200ms
carbon burnout (wt%)
Predicted 50:50 wt% blend
80
50:50 wt% blend saw DTF N₂ char 1300⁰C 200ms
60
40
20
0 0
20
40
60
80
Time (min)
Figure 6. TGA burnout profiles of DTF devolatilised chars and blends in air condition 100
Carbon burnout (wt%)
KK DTF CO₂ char 1300⁰C 200 ms predicted 50:50 wt% blend
80
50:50 wt% blend saw DTF CO₂ char 1300⁰C 200 ms
60
40
20
0 0
20
40
60
80
Time (min)
Figure 7. TGA burnout profiles of DTF devolatilised chars and blends in air condition 3.3 Effect of biomass blending ratio on coal char reactivity In order to maximise the catalytic impact of biomass inherent minerals on coal char burnout, 50:50 wt% blend have been used in this work. However, 25:75 and 10:90 wt% biomass/coal char blends were investigated to examine the effect of smaller biomass
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blend ratios on coal char reactivity. As demonstrated in the burnout profiles in Figures 8 and 9, coal char burnout efficiency was increased with increase in biomass ratio in the blend. Also, interaction of the two fuels is minimal in the 10:90 wt% blend compared to the 25:75 wt% blends. Assessment of the reactivity parameters in Table 3 reaffirms the impact of blending ratio. 100
Carbon burnout (wt%)
KK DTF char 1100C 600ms Predicted 25:75 wt% blend
80
25:75 wt% blend sawdust DTF 1100C 600ms
60
40
20
0 0
20
40
60
80
Time (min)
Figure 8. TGA Carbon burnout profile of DTF chars for 25:75 wt% under air fired condition
Carbon burnout (wt %)
100
KK DTF char 1100⁰C 600ms Predicted 10:90 wt% blend
80
10:90 wt% blend sawdust DTF char 1100⁰C 600ms
60
40
20
0 0
20
40
60
80
Time (min)
Figure 9. TGA Carbon burnout profile of DTF chars for 10:90 wt% under oxy-fuel condition Table 3. Effect of biomass/ coal char blending ratio on co-combustion reactivity Sample
1st order rate constants -
(min ¹)
90% burnout time (min)
25:75 wt % N₂ char blend
0.0725
34.00
10:90 wt % N₂ char blend
0.0601
43.00
3.3. Effect of sawdust demineralisation on coal char burnout
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The char produced from the acid washed sawdust sample was blended with the coal char in 50:50 wt% ratios. It was observed that the removal of alkali/alkaline metals reduced the burnout efficiency of coal char under both conditions compared to the burnout observed earlier in Figures 2 and 3. This confirmed the fact the improved coal char burnout was as a result of the catalytic influence of the alkali/alkaline metals inherence in biomass. The reduction in the concentration of metals due to acid wash is presented in Table 4. However, improved coal char burnout is still observed in oxy-fuel condition due to CO2-char gasification. 100 kk HTF N₂ char 1000⁰C
Carbon burnout (wt%)
50:50 wt% blend 80
Predicted 50:50 wt% Acid washed sawdust HTF char 700⁰C
60
40
20
0 0
20
40
Time (min)
60
80
Figure 10. TGA burnout profiles of sawdust/coal char blend highlighting the impact of mineral catalysis during co-combustion under air fired condition. 100
Carbon burnout (wt%)
kk HTF CO₂ char 1000⁰C Predicted 50:50 wt% blend
80
50:50 wt% blend Acid washed sawdust HTF CO₂ char 700⁰C
60
40
20
0 0
20
40
60
80
Time (min)
Figure 11. TGA burnout profiles of sawdust/coal char blend highlighting the impact of mineral catalysis during co-combustion under oxy-fuel condition. Table 4. Concentration of alkali and alkaline metals in raw and acid washed sawdust
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Sample
Na (wt %)
Mg (wt %) K (wt %)
Ca (wt %)
Raw sawdust
0.12
0.09
0.21
0.75
5M HCl washed sawdust
0.02
0.02
0.008
0.006
4. Conclusions 1. The addition of sawdust char into coal char improved the burnout efficiency of coal char under air fired and oxy-fuel conditions for both low and high heating rate chars. 2. The improved burnout rates were as a result of the catalytic inorganic alkali and alkaline metals present in biomass. 3. Improved coal char burnout in oxy-fuel condition after demineralisation is due to CO2char gasification. Acknowledgement. The authors wish to the authors wish to acknowledge the financial support from the Petroleum Trust Development Funds (Nigeria) for this research. References [1] H. Haykiri-Acma, S. Yaman, Synergy in devolatilization characteristics of lignite and hazelnut shell during co-pyrolysis, Fuel, 86 (2007) 373-380. [2] H. Haykiri-Acma, S. Yaman, Effect of co-combustion on the burnout of lignite/biomass blends: A Turkish case study, Waste Management, 28 (2008a) 20772084. [3] D.K. Park, S.D. Kim, S.H. Lee, J.G. Lee, Co-pyrolysis characteristics of sawdust and coal blend in TGA and a fixed bed reactor, Bioresource Technology, 101 (2010) 61516156. [4] K. Sjöström, G. Chen, Q. Yu, C. Brage, C. Rosén, Promoted reactivity of char in cogasification of biomass and coal: synergies in the thermochemical process, Fuel, 78 (1999) 1189-1194. [5] H. Haykiri-Acma, S. Yaman, Interaction between biomass and different rank coals during co-pyrolysis, Renewable Energy, 35 (2010) 288-292. [6] L. Zhang, S. Xu, W. Zhao, S. Liu, Co-pyrolysis of biomass and coal in a free fall reactor, Fuel, 86 (2007) 353-359. [7] E. Kastanaki, D. Vamvuka, A comparative reactivity and kinetic study on the combustion of coal - biomass char blends, Fuel, 85 (2006) 1186-1193. [8] J. Fermoso, M.V. Gil, C. Pevida, J.J. Pis, F. Rubiera, Kinetic models comparison for non-isothermal steam gasification of coal-biomass blend chars, Chemical Engineering Journal, 161 (2010) 276-284. [9] S.G. Sahu, P. Sarkar, N. Chakraborty, A.K. Adak, Thermogravimetric assessment of combustion characteristics of blends of a coal with different biomass chars, Fuel Processing Technology, 91 (2010) 369-378. [10] K. Le Manquais, C. Snape, I. McRobbie, J. Barker, V. Pellegrini, Comparison of the Combustion Reactivity of TGA and Drop Tube Furnace Chars from a Bituminous Coal, Energy & Fuels, 23 (2009) 4269-4277.
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10
Oviedo ICCS&T 2011. Extended Abstract
Numerical simulation on the re-burning of ash with high unburned carbon in pc boiler Min-young Hwang1, Gyu-Bo Kim2, Ju-hun Song3, Seung-Mo Kim4 and Chung-Hwan Jeon*
1. School of Mechanical Engineering, Pusan National University, Pusan Clean Coal Center, Republic of Korea, Ph: 82-51-510-3035, Fax: 82-51-582-9818, Email:
[email protected] 2. School of Mechanical Engineering, Pusan National University, Pusan Clean Coal Center, Republic of Korea, Ph: 82-51-510-3035, Fax: 82-51-512-5236, Email:
[email protected] 3. School of Mechanical Engineering, Pusan National University, Pusan Clean Coal Center, Republic of Korea, Ph: 82-51-510-7330, Fax: 82-51-512-5236, Email:
[email protected] 4. School of Mechanical Engineering, Pusan National University, Pusan Clean Coal Center, Republic of Korea, Ph: 82-51-510-3035, Fax: 82-51-512-5236, Email:
[email protected] * Corresponding author, Associate Professor, School of Mechanical Engineering, Pusan National University, Pusan Clean Coal Center, Republic of Korea, Ph: 82-51510-7324, Fax: 82-51-582-9818, Email:
[email protected]
Abstract In thermal power generation companies, refined ash (LOI<6%) from a PC boiler has not only suitable to making light weight aggregate but also brought cost benefit to the companies through recycling. However, the ash which has high unburned carbon is not able to reuse and still burying in the ground. To obtain refined ash, re-burning of high LOI ash(LOI>6%) along with pulverized coal in PC boiler was considered in this study. In a real scale application, direct re-burning method seems to be a worthwhile subject to investigate. It allows economics, safety and stability without many change of the PCboiler. Hence, it is necessary to focus on obtain appropriate operating condition to acquire the refine ash by re-burning. Ash sample with 8% LOI was selected by the proximate analysis. Also, combustion kinetic parameters which were used to numerical
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1
Oviedo ICCS&T 2011. Extended Abstract
study of the ash and a coal were obtained through the results in a drop tube furnace. For numerical studies of ash re-burning, a 500MW PC-boiler geometry was mashed with 2000000 cell. As a reaction model, 2-step devolatilization model and kinetic-diffusion limited char model were adopted. A number of simulations have been performed as a function of fuel ratio and supply position. Through the results, the supplying position and the feeding ratio of the ash were decided at conditions with appropriate boiler temperature and emission
1. Introduction Loss on ignition (LOI) is defined by unburned carbon fraction in ash after combustion. Constantly increasing coal cost and decreasing overall amount of high lank coal cause using of low rank coal. Using a low rank coal on the conventional tangential firing boiler brought out high LOI contents because of the boiler was optimally designed for standard coal condition. In thermal power generation companies, refined ash (LOI<6%) from a PC boiler has not only suitable to making light weight aggregate but also brought cost benefit to the companies through recycling. However, the ash which has high unburned carbon is not able to reuse and still burying in the ground. To solve this problem, one of the suggestion is re-burning of high LOI ash(1~3). It is simple system and economical without high cost of facility. For the application of ash re-burning in the tanjential coner fireing boiler, deciding supply position and amount are impotant which were related with efficiency, combustion instability and emmision. To gain a insight about coal combustion phenomina in the interior of the boiler, computational fluid dynamics have been applied widely which is not only gives good approximated results but also reduce experimantal case and cost. In this study, we performed numerical simulation of coal firing boiler. Through the coal particle moving pathway and reaction, suggest suppling position and amout of ash for re-burning. 2. Experimental section For the more realistic coal combustion CFD simulation, fundamental properties of coal such as composition, moisture, carbon contents and reaction rate of each species are truly important. Also these parameters only determined by the experimental results. Proximate analysis was performed by SDT-Q600 (TGA) which experiment show the fraction of fixed carbon, moisture, volatile matter and ash. Ultimate analysis was Submit before 31 May 2011
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2
Oviedo ICCS&T 2011. Extended Abstract
performed by element analyzer (Tru-spec, LECO) which results indicate the fraction of C, H, O, N in sample. Through these analyses, chemical equilibrium could be calculated but which data could not explain how much reaction occurring as a function of time. The reaction rate constant; activation energy (Ea), pre-exponential factor (A) should be determined by additional experiment. Drop tube furnace being used widely to finding reaction rate constant which experiment provide similar atmospheric condition with interior of boiler. Ash (LOI 8%) kinetic was deducted by DTF and other kinetic parameters were determined by literature review. 3. Results and Discussion Mass, momentum, energy, species conversion and heat-transfer (conduction, convection, radiation) equation were selected for the simulation. The mass and energy transfer between gas-phase with coal particle was calculated by Euler-lagrangian model. Figure 1 shows coal particle trajectory in the interior of the boiler. The coal particles where injected each burner has a spiral flow because of increasing volume of the gas phase and the injecting momentum on the normal direction of each burner. Most of particles injected on the lower burner, going through the core axis of the boiler. This flow tendency creates fuel rich areas, in other words the coal particles which passing through this pathway experiencing leak of oxygen. Complete combustion is the function of partial pressure of the particle surface and residence time for reaction. Nevertheless of long particle pathway, locally leak of oxygen cause incomplete combustion. This result gives insight where the ash particles supply for the re-burning. In other to obtain complete re-burning of ash, supplying upper burner is more effective than lower burner. Figure 2 shows interior temperature distribution with supply amount of ash. Up to 4ton/h , there is no change in the flame and flow pattern on the boiler. However when we increase to 16ton/h, temperature distribution was unstable and calculation became divergence. It means the flow balance had broken by the momentum of supplied ash. 4. Conclusions Coal combustion simulation shows some insight for re-burning of ash. When we consider the particle moving pathway, injecting upper burner is more effective than lower burner. In a view of combustion instability, injecting less than 6ton/hour of ash is appropriate for re-burning.
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3
Oviedo ICCS&T 2011. Extended Abstract
Figure 1 Particle traces colored by a) particle char fraction, supply 200ton/h coal, b) oxygen fraction of particle surface, supply 200ton/h coal, c) particle char fraction, supply 198ton/h coal + 2ton/h ash, d) heat of reaction, supply 198ton/h coal + 2ton/h ash
Figure 2 interior temperature distribution coal: 200ton/h + ash: 4ton/h, 6 ton/h
Submit before 31 May 2011
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4
Oviedo ICCS&T 2011. Extended Abstract
References [1] Sen Li, et. al., "Optimization of coal re-burning in a 1MW tangentially fired furnace", Fuel, No. 86, 2007, pp. 1169-1175 [2] H. P. Wan et al., "Controlled LOI from coal re-burning in a coal-fired boiler", Fuel, No. 87, 2008, pp. 290-296. [3] H. Y. Park et al., "Re-burning of bottom ash in a coal-fired power plant and its effect on the plant management", Korea Society of Waste Management, 24, vol 5, 2007, pp. 472-481
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5
Characteristics of Hydrogen Sulfide Formation in Pulverized Coal Combustion Hiromi Shirai1, Michitaka Ikeda1 and Hiroshi Aramaki2 1
Central Research Institute of Electric Power Industry 2-6-1 Nagasaka, Yokosuka-shi, Kanagawa-ken 240-0196, Japan Email:
[email protected] 2
IHI 1 Shin-nakahara-cho, Isogo-ku, Yokohama-shi, Kanagawa-ken 235-0031, Japan Abstract Low excess air combustion and two-stage combustion are currently being attempted to reduce NOx emission at the outlet of the boiler in pulverized coal-fired power plants. However, a strong reducing atmosphere is formed in the region between burners and air ports in the case of two-stage combustion. This atmosphere promotes the formation of hydrogen sulfide (H2S), which causes the sulfidation corrosion at the boiler wall. The control of H2S formation and the prevention of corrosion are important in power plants. In this study, the characteristics of H2S formation in two-stage combustion were investigated using a test furnace for pulverized coal combustion. H2S is formed at the point where the O2 concentration decreases sharply. The H2S concentration distribution is similar to the concentration distributions of H2 and CO. The H2S concentration is affected by the fuel ratio and the sulfur content in coal. Furthermore, it was found that after the release of sulfur has finished, H2S and SO2 reach equilibrium concentrations in accordance with the equation SO2 + 3H2 = H2S + 2H2O.
1. Introduction In Japan, to comply with air pollution control laws, it is necessary to use low-NOx combustion technology and install a flue gas treatment system in coal-fired power plants. Low excess air combustion and two-stage combustion are currently being attempted as means of realizing low-NOx combustion. These combustion methods can reduce NOx emission from a boiler to reduce the capital and operating costs of De-NOx units. However, a strong reducing atmosphere is formed in the region between burners and air ports in the case of two-stage combustion. This atmosphere promotes the formation of hydrogen sulfide (H2S), which causes sulfidation corrosion at the boiler wall [1]. The control of H2S formation and the prevention of the corrosion are important
1
in power plants. To control H2S formation, one strategy is to raise the excess air ratio and another is to lower the two-stage combustion ratio. However, NOx emission increases in both cases. To determine suitable combustion conditions, it is necessary to investigate the characteristics of H2S formation in pulverized coal combustion. In this study, the characteristics of H2S formation in two-stage combustion and the effect of coal properties (sulfur concentration and fuel ratio) were investigated using a test furnace for pulverized coal combustion.
2. Experimental 2.1 Coal samples Three types of bituminous coal were used. The properties of each coal type are shown in Table 1. Coals with different sulfur contents and fuel ratios were selected, where the fuel ratio is the ratio of fixed carbon to volatile matter. The combustibility of coal increases with this ratio. Using these coals, the effect of the sulfur content and fuel ratio on the characteristics of H2S formation was investigated. Table 1 Properties of coals Coal Moisture（equilibrium） Ash（dry） Volatile matter（dry） Fixed carbon（dry） Fuel ratio C（dry） H（dry） N（dry） O（dry） Total S（dry） HHV（dry） LHV（dry）
％ ％ ％ ％ － ％ ％ ％ ％ ％ MJ/kg MJ/kg
SK 4.3 16.1 41.7 42.2 1.01 67.4 5.65 1.32 9.1 0.51 27.6 26.3
TH 10.9 8.8 45.6 45.7 1.00 68.7 4.78 1.53 15.4 0.91 28.3 27.3
WB 3.4 14.4 27.0 58.6 2.17 71.6 3.99 1.61 7.9 0.53 28.1 27.2
2.2 Experimental apparatus and methodology A schematic diagram of the pulverized coal combustion furnace is shown in Figure 1. The furnace is a water-cooled horizontal and cylindrical furnace made of steel. The furnace is 0.85 m in inner diameter and 8 m in length. Refractory materials are coated onto the inside wall of the furnace, the ports for air injection during two-stage
2
combustion are installed on the furnace side wall. Thermocouples and gas-sampling probes are inserted into the furnace through these ports. An advanced low-NOx burner [2], which is designed to have a combustion capacity of approximately 100 kg/h for bituminous coal combustion, is used. In the combustion tests, the furnace is preheated by the combustion of kerosene. After sufficiently preheating the furnace, the pulverized coal is introduced into the furnace, and the supply of kerosene is stopped. The coal feed rate is controlled through the thermal input of coal (760 kW). The median diameter of the pulverized coal particles is 37 – 50 µm in this study. The air ratio is set at 1.24, which means that the excess O2 concentration in the exhaust gas is 4.0%. The two-stage combustion ratio is 30%, which is a typical value for Japanese pulverized coal-fired power plants. The distance between the burner and the two-stage combustion air ports is 2.99 m, which is the most suitable distance for the reducing NOx concentration and unburned carbon concentration in fly ash in the case of bituminous coal combustion [3].
Secondary and tertiary air BUF
FDF
Furnace outlet
Primary air and pulverized fuel Two-stage combustion air ports
Stack
IDF Bag filter
Multicyclone
Gas cooler
Figure 1 Schematic diagram of the test furnace The measured gas components are O2, CO, CO2, H2, H2O, NOx, SO2, H2S and COS. CO and CO2 are measured with a non dispersive infrared gas analyzer, O2 with a magnetic oxygen analyzer, NOx with a chemiluminescence analyzer, H2, SO2, H2S and COS are measured by gas chromatography with a thermal conductivity detector. The
3
H2O concentration, if required, is calculated from the weight of moisture absorbed by CaCl2. Gas temperatures are measured with a Pt/Pt–Rh sheath thermocouple. Fly ash is collected from the gas cooler, the multi-cyclone and the bag filter, which are installed downstream of the furnace. The unburned carbon concentration, Uc [% by weight] is measured in this ash. Combustion efficiency Eff [%] is obtained by the equation,
(1) where Cash [%] is the ash content in coal.
3. Results and Discussion 3.1 Behavior of H2S and other sulfur compounds The concentration distributions of H2S, COS, SO2, NOx, CO and H2 between the burner and the two-stage combustion air ports at the center axis of the furnace for SK coal are shown in Figure 2. In this area, the gas temperature varies from 1200 to 1400 ◦C as shown in Figure 3. The O2 concentration decreases rapidly and disappears when the pulverized coal undergoes combustion near the burner. H2S is formed at the position where O2 decreases sharply. The concentration distribution of H2S has a peak and is similar to the concentration distributions of H2 and CO, which indicate the intensity of the reducing atmosphere. The concentration distribution of COS is similar to that of H2S.
H2S H2S
COS
SO2 SO2
NOX NOx
H2 H2
CO
O2 O2
600
25
500
20
400
15
300 10
200
5
100 0
1,600 H2,CO,O2 ［%］ Gas temperature [◦C]
H2S, COS, SO2, NOx ［ppm］
However, the COS concentration is less than one-tenth of that of H2S. On the other hand,
0 0.0
0.5 1.0 1.5 2.0 2.5 Distance from burner [m]
3.0
Figure 2 Concentration distributions of H2S, COS, SO2, NOx, CO and H2 between the burner and the two-stage combustion air ports at the center axis of the furnace for SK coal
1,200 800 400 0 0.0
0.5 1.0 1.5 2.0 2.5 Distance from burner [m]
3.0
Figure 3 Gas temperature distribution between the burner and the two-staged combustion air ports at the center axis of the furnace for SK coal 4
SO2 is formed near the burner and its concentration increases with the progress of combustion. Next, the distribution of the ratio of released sulfur contained in the coal to gas phase is shown in Figure 4. The flow rates of H2S, COS and SO2 were calculated from the radial distributions of their concentrations and the axial gas velocity distribution on the consideration that the gas velocity distribution was similar to the particle velocity distribution measured with a laser Doppler velocimeter [4]. Up to a distance of 1.0 m from the burner, the gas flow involved Therefore, flow rates could not be obtained. As shown in this figure, it was found that the total amount of released sulfur is constant after 1.3 m. This means that the release of sulfur is finished at this distance and that the behavior of sulfur compounds depends on only the gas-phase reactions.
Relesed sulfur ratio [%]
in recirculation was very complex.
H2S H2S
COS
SO2 SO2
Total S T-S
100 80 60 40 20 0 0.0
0.5 1.0 1.5 2.0 2.5 Distance from burner [m]
3.0
Figure 4 Distribution of the ratio of released sulfur contained in the coal to gas phase
3.2 Effect of coal properties The combustion efficiency and SO2 concentration at the outlet of the furnace are shown in Table 2. It was confirmed
that
the
combustion
efficiency deteriorates as the fuel ratio increase, and that the measured SO2
Table 2 Combustion efficiency and SO2 concentration at the outlet of the furnace Coal SK TH WB
Combustion efficiency 99.40% 99.50% 99.15%
SO2 378 ppm 759 ppm 334 ppm
concentration and the SO2 concentration estimated from the sulfur content in coal correspond. The concentration distributions of H2S, COS, SO2, H2 and O2 between the burner and the two-stage combustion air ports at the center axis of the furnace for the three types of coal are shown in Figure 5. The concentration distribution of H2S has a peak and is similar to the concentration distributions of H2 for TH coal and SK coal. However, for WB coal, the H2S concentration is very low. In the O2 concentration distribution, it was found that the progress of combustion of WB coal is later than those of the other coals because the O2 concentration is higher at 0.6 m from the burner. This causes the 5
formation of H2 and H2S to be delayed, whose concentrations are lower than those for the other coals. The relationship between H2 and H2S concentrations is shown in Figure 6. The H2S concentration has a strong correlation with that of H2. The H2S concentration is very low below of a H2S concentration of 2% of H2, above which it increases rapidly. In a high-fuel-ratio coal such as WB coal, the H2S concentration is very low because the
400
100
300
80
COS [ppm]
H2S [ppm]
SK (FR:1.01, Sulfur:0.51%) TH (FR:1.00, Sulfur:0.91%) WB (FR:2.17, Sulfur:0.53%)
200 100
40 20 0
0 0.0
0.5 1.0 1.5 2.0 2.5 Distance from burner [m]
3.0 8
600
6 H2 [%]
800
400
0.0
0.5 1.0 1.5 2.0 2.5 Distance from burner [m]
3.0
0.0
0.5 1.0 1.5 2.0 2.5 Distance from burner [m]
3.0
4 2
200 0
0 0.0
0.5 1.0 1.5 2.0 2.5 Distance from burner [m] 30
O2 [%]
SO2 [ppm]
60
3.0
20
10
0 0.0
0.5 1.0 1.5 2.0 2.5 Distance from burner [m]
3.0
Figure 5 Concentration distributions of H2S, COS, SO2, H2 and O2 between the burner and the two-stage combustion air ports at the center axis of the furnace for three types of coal 6
concentration of formed H2 is low. On
SK (FR:1.01, Sulfur:0.51%) TH (FR:1.00, Sulfur:0.91%) WB (FR:2.17, Sulfur:0.53%)
the other hand, it was confirmed that the H2S concentration is affected by the
400
3.3 Reaction of H2S formation The chemical equilibrium for the
H2S [ppm]
sulfur content of a coal. 300 200 100
reaction between H2S and SO2 (Reaction 0
1) and the water shift reaction (Reaction
0.0
2) was investigated.
1.0
2.0 3.0 H2 [%]
4.0
5.0
Figure 6 Relationship between H2 concentration and H2S concentration
Reaction 1: SO2 + 3H2 = H2S + 2H2O Reaction 2: CO + H2O = H2 + CO2
The equilibrium concentration CH2S [ppm] of H2S is calculated by the following equations. (2) (3) Here, PH2S, PSO2, PH2, PCO and PCO2 are partial pressures [Pa] of gas components, Ptotal is the total pressure, and Kp1 and Kp2 are chemical equilibrium constants of Reaction 1 and Reaction 2, respectively. The relationship between the measured 800
concentration and the equilibrium which no correlation between the two concentrations in the region up to 1.3 m from the burner is apparent. This result indicates that equilibrium is not reached in the region where sulfur is released.
Equilibrium H2S [ppm]
concentration is shown in Figure 7, in
1.3 - 2.6 m from burner 0 - 1.3 m from burner 600
400
200
Although a strong correlation is apparent in the region of 1.3 - 2.6 m from the burner, the measured concentration is higher than the equilibrium concentration. Therefore, the chemical equilibrium
0 0
200 400 600 Measured H2S [ppm]
800
Figure 7 Relationship between H2S equilibrium concentration and measured H2S in the combustion of SK coal
7
constants of Reaction 1 and Reaction 2 calculated from the measured gas compositions are compared with the theoretical chemical equilibrium constants. The relationship between the two equilibrium constants for each reaction is shown in Figure 8. Reaction 1 reaches equilibrium but Reaction 2 does not reach equilibrium. Consequently, it was found that the H2S concentration can be estimated from the equilibrium constant of Reaction 1 only in the region where the release of sulfur has finished.
Theoretical value Value Calculated from gas composition
1.0
8
0.8 Kp2
Kp1
6 4
0.6 0.4
2
0.2
0 1,000
1,200 1,400 1,600 Gas temperature ［◦C］
1,800
0.0 1,000
1,200 1,400 1,600 Gas temperature ［◦C］
1,800
Figure 8 Relationship between theoretical equilibrium constant and equilibrium conctant calculated from measured gas composition for each reaction
4. Conclusions The characteristics of H2S formation in two stage combustion were investigated using a test furnace for pulverized coal combustion. H2S was formed at the point where the O2 concentration decreased sharply. The H2S concentration distribution was similar to H2 and CO concentration distributions. The H2S concentration was affected by the fuel ratio and the sulfur content in coal. Sulfur was released from coal owing to its volatile, and H2S and SO2 were formed. After the release of sulfur, H2S and SO2 reached equilibrium concentration in accordance with the equation of SO2 + 3H2 = H2S + 2H2O.
References [1] Morinaga M, Najima S, Aramaki H, Wakabayashi N, Shirai H, Evaluation of Sulfide Corrosion Conditions in a Pulverized Coal Fired Thermal Power Plant Boilers, Thermal and Nuclear Power Generation Convention 2010; 78-85, in Japanese. 8
[2] Makino H, Kimoto M, Kiga T, Endo Y. Development of New Type Low NOx Burner for Pulverized Coal Combustion. Therm Nucl Power 1997; 48:64–72, in Japanese. [3] Makino H, Kimoto M. Low NOx combustion Technology in Pulverized Coal Combustion. Kagaku Kogaku Ronbunshu 1994; 20:747–57, in Japanese. [4] Hashimoto N, Wang S, Kurose R, Tsuji H, Shirai H, A Numerical Simulation of Pulverized Coal Combustion Field using a Tabulated-Devolatilization-Process Model (TDP Model).(Part 2: Application to a 100 kg-coal/h Low NOx Swirl Burner), Trans JSME, Series B 2010; 769; 1396-71, in Japanese.
9
Numerical Study of the Influence of Heterogeneous Kinetics on the Carbon Consumption by Oxidation of a Single Coal Particle P.A. Nikrityuk1 , M. Gr¨abner2 , M. Kestel1 and B. Meyer2 1
CIC Virtuhcon, Technische Universit¨ at Bergakademie Freiberg, Reiche Zeche, Fuchsm¨ uhlenweg 9, Haus 1, 09596 Freiberg, Germany
2 Department of Energy Process Engineering and Chemical Engineering, Technische Universit¨ at Bergakademie Freiberg, Reiche Zeche, Fuchsm¨ uhlenweg 9, Haus 1, 09596 Freiberg, Germany
Abstract This work is devoted to the numerical study of the influence of heterogeneous kinetics on the oxidation rates of a single carbon particle in quiescent and non-quiescent environments. The coal particle is represented by moisture and ash free nonporous carbon while the coal rank is implemented by several kinetic rate expressions. The model includes six gaseous chemical species (O2 , CO2 , CO, H2 O, H2 , N2 ). Three heterogeneous reactions (C + O2 , C + CO2 and C + H2 O) and two homogeneous semi-global reactions, namely carbon monoxide oxidation and water-gas shift reaction, are employed. Several semiglobal reaction rate expressions taken from the literature were utilized. The Navier-Stokes equations coupled with the energy and species conservation equations are used to solve the problem in pseudo-steady state approach. At the surface of the particle, the balance of mass, energy and species concentration is applied including the effect of the Stefan flow and the heat loss by radiation at the surface of the particle. The model and the code used are validated against analytic two-film model. Good agreement is observed. The numerical simulations performed reveal significant effect of heterogeneous kinetic on the carbon comsumption rates of the particle. In particular the maximal discrepancy between results is achieved in kinetic controlled regime and is proportional to the factor of 10 in respect to carbon mass flux on the particle surface. Additionally, the influence of Reynolds number, the ambient mass fraction O2 and the temperature on the regimes of combustion and Preprint submitted to Int. Conf. Coal Science & Technology (ICCS&T)
June 14, 2011
gasification is discussed. Keywords: 1. Introduction Coal has been extensively used as a primary fuel for energy production for many years throughout the world. However, due to the continuous increase in CO2 emission around the world, energy generation using coal combustion is becomming a thing of the past. Nowadays, coal is basically considered as a primary chemical feedstock for the production of gasoline, fertilizers or other chemicals using coal gasification. But also for fuel gas production for Integrated Gasification Combined Cycle (IGCC) power generation, research on coal gasification is necessary, since it is more ecological and efficient in comparison to a standard power plants operating with coal combustion, e.g. see [1]. In the design of novel combustors or gasifiers working on solid carbonaceous fuels (particles) the important issue is the prediction of the heating and burning rates of such fuels. Considering the coal in the role of a solid carbonaceous fuel, chemically reacting coal particles have been extensively studied over the last one hundred years due to the practical importance of coal in the production of energy and chemicals. Nusselt [2] proposed the one-film model and Burke and Schuman [3] developed the two-film model. Both models are typically used as a starting point in modeling a chemically reacting coal particle, cf. [4]. In the early 1980 Amundson and coworkers performed a large number of numerical studies investigating diffusion and reaction in a stagnant boundary layer around a spherical carbon particle, e.g. see works of [5, 6]. Only diffusion-limited regimes were investigated. It was shown that the two-film model is capable of adequately predicting the combustion of a coal particle at higher temperatures. For reviews of theoretical and numerical works produced before 1980 and 2008, Sundaresan and Amundson [7] and Higuera [8], respectively, provide good overviews. In spite of the numerous works about the modeling of particle combustion the studies about the influence of heterogeneous kinetics on the combustion of carbon appears to have received relatively little attention. Chelliah [9] studied numerically the role of carbon-radical reactions on the carbon removal rate utilizing detailed heterogenous mechanism of Bradley et al. [10] under quasi-steady burning conditions. However it should be noted that 2
detailed surface mechanism including carbon-radical reactions are rare or almost impossible to find in the literature for different types of coal. Basically, experimental works for different types of coal report global heterogeneous reactions. Recently, Stauch & Maas [11] studied numerically the burning process of a single non-movable carbon particle using two different mechanisms of heterogeneous reactions. In particular, the surface mechanisms of Bradley et al. [10] and of Libby and Blake [12] were utilized taking into account five (including carbon-radical reactions) and three heterogeneous reactions, respectively. No significant influence of heterogeneous mechanism on the carbon consumption rate of particles has been detected. Motivated by the needs to understand the effect of different heterogeneous rate constants known from the literature, e.g. see [13] and [4], on the partial oxidation of a moving coal particle we present numerical study of the influence of heterogeneous kinetics on the oxidation rates of a single carbon particle in quiescent and non-quiescent environments. In particular, the purpose of this paper is to extend the analysis of the effect of the kinetics of heterogeneous reactions on the oxidation rate of a coal particle taking into account the effect of the flow. 2. Problem and Model Formulation We consider a single spherical coal char particle with a diameter dp = 2 mm placed in a stationary position in a hot oxidizing environment, with the main gas flow passing around it. The inflow velocity was assumed to be uniform and was determined by means of the Reynolds number calculated as Re = ρ∞μu∞∞ dp , where ρ∞ and μ∞ are the density and molecular viscosity, respectively, corresponding to the inflow temperature Tin and the composition. Two Reynolds numbers are considered: 0 and 10. Two cases with different inflow gas composition were considered: the first case corresponds to the so called ’dry air’ atmosphere with 0.233 mass fraction (Y ) of O2 and 0.001 mass fraction of H2 O and second case refers to the so called ’gasification’ condition with YO2 = 0.11 and YH2 O = 0.074. For the fixed composition of the inflow gas the inflow gas temperature Tin was varied from 1000 K to 3000 K. The modeling configuration and the size of the domain are illustrated in Figure 1 and Tab. 1. It can be seen that this configuration allows the use of 2D cylindrical coordinated r and z to model the flow past the sphere.
3
H2O O2
Outlet
L1
Inlet
CO 2
u Tin Carbon particle
r L2
D
H2
O2
Tsurf Symmetry axis L3
z
Solid Carbon
CO O2
flame sheet
CO
CO 2
O2
H2 CO2
H2O
C O2
H 2O
Figure 1: Principle scheme of the computational domain (left) and schematic representation of coal oxidation by the presence of water vapor (right).
Case Type Re = 0 Re > 0
L1 L2 L3 Number of Nodes 75D 75D 75D 23205 40D 30D 100D 59668
Table 1: Size of the domain and grid resolution
In this work we utilize the pseudo-steady-state approach (PSS), see [7], due to the fact that the conversion time of the mm-size particle is always long compared to the convective and diffusion time scales for the gas phase (see [4]), which is typical for fluidized bed systems. The Navier-Stokes equations coupled with the energy and species conservation equations are used to solve the problem in pseudo-steady state approach. At the surface of the particle, the balance of mass, energy and species concentration is applied including the effect of the Stefan flow and the heat loss by radiation at the surface of the particle. The complete formulation of governing equations, interfacial boundary conditions and model validation against analytic two-film model can be found in [14, 15, 16]. The transport properties of the gas are calculated using kinetic theory and polynomials [17]. The gas flow was treated as incompressible ideal gas following model described in [18].
4
Figure 2: Comparison of the reaction rate kr for several carbon types reacting with O2 , H2 O and CO2 . The sources of the data are presented in tables 3 to 6
3. Selection of heterogeneous and homogeneous reactions The chemistry is modeled using semi-global homogeneous and heterogeneous reactions written as follows [4]: 1 CO + O2 + H2 O → CO2 + H2 O 2 CO + H2 O → CO2 + H2 CO2 + H2 → CO + H2 O 1 C + O2 → CO 2 C + CO2 → 2CO C + H2 O → CO + H2
(exothermic)
(1)
(exothermic) (endothermic)
(2) (3)
(exothermic)
(4)
(endothermic) (endothermic)
(5) (6)
The semi-global reaction rates of homogeneous chemical reactions are given f b in Tab. 2. The rate constants kCO , kshif t and kshif t are computed using the 5
Arrhenius expression:
Er
kr = Ar T nr exp− R T
(7)
where Ar is the pre-exponential factor, nr is the temperature exponent (in this work nr = 0 for all homogeneous reactions), Er is the activation energy and R is the universal gas constant. The rate expression for the reaction describing CO oxidation, eq. 1, was proposed by [19]. Notice that the reaction order is not related to the stoichiometry of the reaction due to the global character of this reaction, see Tab. 2. Basically, semi-global homogeneous chemical reactions are widely used by the modeling of industrial combustors or gasifiers using computational fluid dynamics software, e.g. see [20]. However it should be noted that global reactions rates are only valid in a narrow range of conditions and should be used very cautiously. In this work we employ four sets of semiglobal heterogeneous surface rate data, which have been adopted from [4], [13], [12] and [21], see tables 3, 4, 5 and 6, respectively. The data is provided in terms of extended Arrhenius expression according to eq. 7. The analysis of the tables shows that most of the data is gathered in 1970 and 1980s. We note here that in order to find origin source of kinetic data published in the literature we had to follow long chains of links. Finally we found out that lots of authors refer back to only a few works. In some publications the original data were even changed due to recalculating in new dimensions. For example Turns in his book [4] utilizes the heterogeneous kinetic data from Mon and Amundson [6], who recalculated data from Howard [22] and Dutta [23] and fitted it in order to eliminate the linear temperature dependency of the Arrhenius expression. Libby and Blake [12] refer to Field [24] and Dobner [25] corresponding to not archival literature in the form of internal reports. However good news is that most cited data comes from archival literature and was established by Mayers [26] and Field [27]. For detailed analysis of the kinetic data published in the literature we refer to [28]. As illustration of the influence of coal type on the reaction rate constants kr , the selected kinetic data sets are exhibited in Figure 2. Distinguished by the gaseous reactant, it is obvious that the C + O2 is fastest whereas the Boudouard reaction C + CO2 is the slowest. Also the comparison indicates the temperature range of the experimental investigation. A distinct difference between the several types of carbon is not observed due to many overlaps. Hence, the numerical study is essential to assess the effect of the combined 6
reactions. Reac.
i,r [kmol/m3 · s] R
Ar
1 2 3
kCO [CO] [H2O]0.5 [O2 ]0.25 f kshif t [CO] [H2 O] b kshif t [CO2 ] [H2 ]
2.24 · 1012 2.74 · 109 1.00 · 108
Er [J/kmol] nr
source
1.6736 · 108 8.368 · 107 6.28 · 107
[4], p. 211 [29] [29]
0 0 0
Table 2: Reaction rates for homogeneous reactions
4. Results and Discussions The results of simulations of a single coal particle behavior in a hot quiescent air (corresponding to YO∞2 = 0.233, YH∞2 O = 0.001) and reduced-air (YO∞2 = 0.11, YH∞2 O = 0.074) atmospheres are presented in Figs. 3 and 4. In particular, Figs. 3a and 3b show the carbon net mass flux m ˙ C as a function of the ambient temperature Tin calculated under the condition of quiescent air and reduced-air environments using four sets of heterogeneous kinetic data denoted by Turns, Smoot & Smith, Libby & Blake and Hla et al., see tables 3, 4, 5 and 6), respectively. The comparison of both figures reveals two significant effects. First, it can be seen that the use of Turns and Hla et al. data sets produces similar qualitative behavior of m ˙ C in dependence on the ambient temperature Tin for Tin > 1500 K. However for low ambi ˙C ent temperatures Tin < 1300 K a descrepancy can be observed between m predicted using both data sets, see Fig. 3. In particular, the use of Turns kinetic data produces earlier ignition in comparison to results obtained by use of Hla et al. kinetics. The comparison of the carbon net mass fluxes Reac. C + O2 C + CO2 C + H2 O
Ar [m/s] 3.007 · 105 4.02 · 108 1.21 · 109
Er [J/mol] 1.49 · 105 2.48 · 105 2.48 · 105
nr 0 0 0
source [24], [5], [4] [23], [6], [4] [6], [4]
Table 3: Rate coefficients for heterogeneous reaction mechanism called as Turns. Carbon type is bituminous coal char. The order of all reactions is unity.
7
Reac. C + O2 C + CO2 C + H2 O
Ar [m/(s K)] 1.692 25.9 1.33
Er [J/mol] 0.852 · 105 2.25 · 105 1.47 · 105
nr 1 1 1
source [30] [30] [26]
Table 4: Rate coefficients for heterogeneous reaction mechanism called as Smoot & Smith. Carbon type is an American standard bituminous coal (Pittsburgh #8). The order of all reactions is unity.
Reac. C + O2 C + CO2 C + H2 O
Ar [m/(s K nr )] 3.007 · 105 4.605 11.25
Er [J/mol] 1.4937 · 105 1.751 · 105 1.751 · 105
nr 0 1 1
source [24], [5], [4] [25] [25]
Table 5: Rate coefficients for heterogeneous reaction mechanism called as Libby & Blake. Carbon type is the average for coal char. The order of all reactions is unity.
Reac. C + O2 C + CO2 C + H2 O
Ar [m/(s K nr )] Er [J/mol] 20.2 1.57 · 105 6 1.59 · 10 2.91 · 105 75900 2.68 · 105
nr 0.8 0.3 0.4
source [21] [21] [21]
Table 6: Rate coefficients for heterogeneous reaction mechanism called as Hla et al. gained at pressurized atmospheres. Carbon type is an Australian high-volatile bituminous coal (CRC299). The reaction order of each reaction corresponds to nr .
8
predicted using Smoot & Smith and Libby & Blake data sets for ’air’ and ’gasification’ conditions show relative close behavior of m ˙ C in dependence on Tin . At the same time we found out that at high ambient temperatures Tin > 3000 K all four data sets produce comparable carbon net mass fluxes. This fact is explained by approaching the diffusionally controlled combustion. Second, it can be observed that at reduced atmosphere conditions the transition between kinetically and diffusionally controlled regimes occurs at higher ambient temperatures in comparison to the air atmosphere. For the definition of the diffusionally and kinetically controlled regimes we refer to the book [4]. In order to demonstrate the influence of different kinetic data on local characteristics near the particle surface we plot in Fig. 4 the spatial distribution of the temperature calculated using Turns’s and Libby & Blake’s data sets. The simulations performed using Turns’s kinetic data reveal the flame sheet at Tin = 2000 K. Whereas the use of Libby & Blake’s set predicts no flame sheet around the particle. The comparison of Figs. 4a and 4c shows that the flame sheet thickness increases with decrease of O2 . The effect of the motion of a reacting coal particle relative to the surrounding gas is demonstrated in Figs. 5 and 6. In particular, Fig. 5 illus trates the comparison of the surface net mass flux of carbon m ˙ C in dependence on Tin calculated using four data sets for the case of a movable particle (Re = 10). It can be seen that qualitative behavior of m ˙ C is similar to the case Re = 0. Thus, the calculations made using Turns and Hla et al. kinetic data sets produced close results. However, the oxidation rate of the particle increases with the velocity of the particle due to the flow-enhanced transfer of O2 between the gas and the surface. The same is true for the carbon net mass flux calculated using Smoot & Smith and Libby & Blake kinetic data. Additionally, the comparison of Figs. 3 and 5 demonstrates that increase of the Reynolds number leads to the prolongation of kinetically controlled regime to higher temperatures. Fig. 6 plots spatial distribution of the temperature near the particle predicted using Turns’s and Libby & Blake’s data sets for the air and reduced air atmospheres. It can be seen that the utilization of Turns’s kinetic constants for the case Re = 10 and Tin = 2000 K results in combustion regime in the form of classical flame sheet around the particle. At the same time the adaptation of Libby & Blake’s data for the same case leads to the gasification regime, where no flame exists. To sum up the obtained results converge to the conclusion that Turns (bituminous coal char) and Hla et al. (Australian high-volatile bituminous coal) data sets correspond to coal with high reactivity whereas Smoot & 9
.
.
0.015 0.01
0.015 0.01 0.005
0.005 0 1000
0.02
Hla et al. Turns Smoot and Smith Libby and Blake
2
0.02
0.025
2
m ’ [kg/m s]
0.025
0.03
Hla et al. Turns Smoot and Smith Libby and Blake m ’ [kg/m s]
0.03
1500
2000 Tin [K]
2500
0 1000
3000
a
1500
2000 Tin [K]
2500
3000
b
Figure 3: Predicted carbon net mass flux m ˙ C as a function of the ambient temperature Tin by quiescent environment corresponding to a) - air: YO∞2 = 0.233, YH∞2 O = 0.001 and b) - rediced O2 atmosphere: YO∞2 = 0.11, YH∞2 O = 0.074.
Smith (Pittsburgh #8) and Libby & Blake (average for coal char) data set characterize the coal with lower reactivity. This is to a certain extent in contrast the impression one can get while only interpreting the reaction rate constants as illustrated in Figure 2. 5. Summary In this work we performed the numerical study of the influence of heterogeneous kinetics on the oxidation rates of a single carbon particle in quiescent and non-quiescent environments. Three heterogeneous reactions (C + O2 , C + CO2 and C + H2 O) and two homogeneous semi-global reactions, namely carbon monoxide oxidation and water-gas shift reaction, were employed. Several semi-global reaction rate expressions taken from the literature were utilized. Based on the results of the numerical simulations carried out for a coal particle with a diameter of 1 mm we found out that using Turns and Hla et al. data sets for heterogeneous kinetics produces identical carbon net mass flux and others qualitative features of the carbon particle oxidation. The same is true for Libby & Blake and Smoot & Smith. For all kinetic adat sets used in this work we found out that the end of kinetically controlled regime is shifted to the higher temperatures with the decrease of O2 concentration in the ambient atmosphere and the increase of Re. The flame thickness increases with decrease of YO∞2 . 10
a b
c d Figure 4: Contour plots of the temperature calculated for the air (a,b) and reduced O2 atmosphere (c,d) using different data sets for heterogeneous kinetics: a) and c) - Turns, b and d - Libby & Blake. Here the ambient temeparture is Tin = 2000 K. Reynolds number is Re = 0.
Acknowledgment The authors appreciate the financial support of the Government of Saxony and the Federal Ministry of Education and Science of the Federal Republic of Germany as a part of the Centre of Innovation Competence VIRTUHCON. References [1] M. Gr¨abner, O. von Morstein, D. Rappold, W. G¨ unster, G. Beysel, B. Meyer, Constructability study on a German reference IGCC power plant with and without CO2-capture for hard coal and lignite, Energy Conversion and Management 51 (2010) 2179–2187. [2] W. Nusselt, Der Verbrennungsvorgang in der Kohlenstaubfeuerung., Zeitschrift des Vereins Deutsche Ingenieure 68 (1924) 124–128.
11
0.03
0.03
0.025
0.025 2
m ’ [kg/m s]
0.035
0.02 0.015
Hla et al. Turns Smoot and Smith Libby and Blake
0.02 0.015
.
.
2
m ’ [kg/m s]
0.035
0.01
Hla et al. Turns Smoot and Smith Libby and Blake
0.005 0 1000
1500
2000 Tin [K]
2500
0.01 0.005 0 1000
3000
a
1500
2000 Tin [K]
2500
3000
b
Figure 5: Predicted carbon net mass flux m ˙ C as a function of the ambient temperature Tin by non-quiescent environment (Re = 10) corresponding to a) - air: YO∞2 = 0.233, YH∞2 O = 0.001 and b) - reduced O2 atmosphere: YO∞2 = 0.11, YH∞2 O = 0.074.
a
b
c
d
Figure 6: Contour plots of the temperature calculated for the air (a,b) and reduced O2 atmosphere (c,d) using different data sets for heterogeneous kinetics: a) and c) - Turns, b and d - Libby & Blake. Here the ambient temeparture is Tin = 2000 K. Reynolds number is Re = 10.
[3] S. Burke, T. Schuman, in: Proceedings of the 3rd Int. Conf. Bituminous Coal, pp. 485–489.
12
[4] S. Turns, An Introduction to Combustion, McGraw-Hill, Boston, 2nd edition, 2000. [5] H. Caram, N. Amundson, Diffusion and Reaction in a Stagnant Boundary Layer about a Carbon Particle, Industrial & Engineering Chemistry Fundamentals 16 (1977) 171–181. [6] E. Mon, N. Amundson, Diffusion and Reaction in a Stagnant Boundary Layer about a Carbon Particle. 2. An Extension, Industrial & Engineering Chemistry Fundamentals 17 (1978) 313–321. [7] S. Sundaresan, N. Amundson, Diffusion and Reaction in a Stagnant Boundary Layer about a Carbon Particle. 6. Effect of Water Vapor on the Pseudo-Steady-State Structure, Industrial & Engineering Chemistry Fundamentals 19 (1980) 351–357. [8] F. Higuera, Combustion of a coal particle in a stream of dry gas, Combustion and Flame 152 (2008) 230–244. [9] H. Chelliah, The influence of heterogeneous kinetics and thermal radiation on the oxidation of graphite particles, Combust. Flame 104 (1996) 81–94. [10] D. Bradley, D. G., S. El-DinHabik, E. Mushi, The oxidation of graphite powder in flame reaction zones, Proc. Combust. Inst. 20 (1984) 931–940. [11] R. Stauch, U. Maas, Transient detailed numerical simulation of the combustion of carbon particles, International Journal of Heat and Mass Transfer 52 (2009) 4584–4591. [12] P. Libby, T. Blake, Burning carbon particles in the presence of water vapor., Combust. Flame 41 (1981) 123–147. [13] L. Smoot, P. Smith, Coal Combustion and Gasification (The Plenum Chemical Engineering Series), Springer, 1st edition, 1985. [14] M. Kestel, P. Nikrityuk, B. Meyer, Numerical study of partial oxidation of a single coal particle in a stream of air, in: 14th Int. Heat Transfer Conf., Washington D.C., USA, pp. CD: ISBN 978–0–7918–3879–2, 2010 by ASME.
13
[15] M. Kestel, P. Nikkrityuk, O. Henning, C. Hasse, Numerical study of the partial oxidation of a coal particle in steam and dry air atmosphere, IMA Journal of Applied Mathematics (under review 2010). [16] S. Schulze, P. Nikkrityuk, B. Meyer, Microscale modeling of co2 utilisation by carbon gasification, Energy & Environmental Science submitted (2011). [17] B. Bride, S. Gordon, M. Reno, Coefficients for Calculating Thermodynamic and Transport Properties of Individual Species, Technical Report, NASA., 1993. [18] A. Tomboulides, J. Lee, S. Orszag, Numerical simulation of low Mach number reactive flows, J. Scietific Computing 12 (1997) 139–167. [19] F. L. Dryer, I. Glassman, High temperature oxidation of CO and CH4 ., Proc. Combust. Inst. 14 (1973) 987–1003. [20] Y. Wu, P. J. Smith, J. Zhang, J. N. Thornock, G. Yue, Effects of turbulent mixing and controlling mechanisms in an entrained flow coal gasifier, Energy Fuels 24 (2010) 1170–1175. [21] S. S. Hla, D. J. Harris, D. G. Roberts, A coal conversion model for interpretation and application of gasification reactivity data, in: International Conference on Coal Sci. and Tech., Okinawa, Japan, p. 2005. [22] R. H. Essenhigh, R. Froberg, J. B. Howard, Combustion behavior of small particles, Industrial & Eng. Chem. 57 (1965) 32–43. [23] C. Dutta, C. Wen, R. Belt, Reactivity of coal and char. 1. In carbon dioxide atmosphere, Industrial & Eng. Chem. Proc. Design and Develop. 16 (1977) 20–30. [24] M. Field, R. Roberts, Measurement of the ratio of reaction of carbon particles with oxygen in the pulverized coal size range for gas temperature between 1400 K and 1800 K., Technical Report 325, BCURA Memb. Circ., England, 1967. [25] S. Dobner, Modelling of Entrained Bed Gasification: the Issues, EPRI, Palo Alto, CA, 1976.
14
[26] M. Mayers, The rate of reduction of carbon dioxide by graphite, J. American Chem. Society 56 (1934) 70–76. [27] F. M.A., Rate of combustion of size-graded fractions of char from a low-rank coal between 1200 K and 2000 K, Comb. Flame 13 (1969) 237–252. [28] M. Gr¨abner, Modeling-based Evaluation of Gasification Processes for low-grade Coals, Ph.D. thesis, Technische Universit¨at Bergakademie Freiberg, Germany, 2011. [29] W. Jones, R. Lindstedt, Global reaction schemes for hydrocarbon combustion, Combustion and Flame 73 (1988) 233–249. [30] G. Goetz, K. Nsakala, K. Patel, T. Lao, Combustion and gasification kinetics of chars from four commercially significant coals of varying rank, in: Proceedings of the 2nd Annual Conf. on Coal Gasification, EPRI Palo, Alto, CA, p. 1982.
15
Oviedo ICCS&T 2011. Extended Abstract
Comparison of structure and reactivity of an Australian algal coal and a Jordanian oil shale W. Roy. Jackson1,2* , Mohammad W. Amer1,2, Yi Fei1,2, Marc Marshall1,2, and Alan L. Chaffee1,2 1
School of Chemistry, Monash University, Clayton, Victoria 3800, Australia
2
Centre for Green Chemistry, Monash University, Victoria 3800, Australia
Corresponding author name: W. Roy. Jackson Email:
[email protected]
Abstract An Australian algal coal (torbanite) and a Jordanian oil shale (from El-Lajjun) are compared. Literature data indicate that both materials are highly aliphatic in nature. Their reactivity towards pyrolysis and hydropyrolysis, both in the presence and absence of tin catalysts, is compared and a detailed analysis of the products of these reactions is presented. Experiments in our laboratory confirm that the depolymerization of the torbanite and the oil shale occurs in a narrow temperature range. Establishment of a clearer understanding of this depolymerization process could lead to the discovery of more efficient methods for the extraction of oil from the kerogen in the Jordanian oil shale. 1.
Introduction
This paper describes experiments designed to provide information on the structural features of the organic component of a Jordanian oil shale from the El Lajjun deposit, by comparison of its reactivity with that of an algal coal, torbanite, from Muswellbrook, NSW, whose organic structure is reasonably well characterized [1, 2]. The organic matter in the oil shale is derived from marine algae [3], whereas the organic matter in torbanite is from fresh water algae [1]. The deposition conditions for the Jordanian oil shale are responsible for its significantly higher sulfur content [4]. In addition the oil shale is much younger, originating no earlier than the Maastrichtian age (Cretaceous period) [3, 4], whereas the torbanite comes from the much older Permian period, leading to its low oxygen content [1]. The influence of these differences on reactivity will be discussed. Pyrolysis reactions under N2 and hydropyrolysis both with and without added Sn catalysts have been carried out, and the yields and product compositions determined. 2.
Experimental
2.1
Oil shale and coal preparation and characterization
Oil shale from the El Lajjun deposit in Jordan was received as -2mm particles and was ground to -180 µm. The torbanite sample was obtained as a solid block which was ground to -180 µm. Ash 1
Oviedo ICCS&T 2011. Extended Abstract
yields were determined by heating at 490oC in air to constant weight. The water content of the samples was taken as the loss of weight observed when they were heated under a flow of N2 for 3h at 105oC, and samples were all dried by this procedure before reaction. The sample of shale for elemental analysis was washed with 0.5M HCl following the procedure of Redlich at al [5] to remove the carbonate minerals. Elemental analyses were carried out by HRL Technology Ltd for C, H, N, S, Cl and Fe for raw and acid-washed oil shale and for C, H, N and S for raw torbanite by the Campbell Microanalytical Laboratory, University of Otago. The values of organic C, H, N and O contents and the concentrations of the different forms of S were calculated from the elemental analyses, the ash yields and the loss of weight on acid washing. The solid state 13C NMR spectra for acid-washed oil shale and untreated torbanite were determined using a Bruker 400 (1H)/100(13C) MHz spectrometer with cross polarization-magic angle spinning (CP/MAS).To introduce the tin catalyst, the samples was mixed with an aqueous slurry of SnO2 (0.5 mol/kg db for the Jordanian oil shale, 1 mol/kg db for the torbanite) stirred for a few minutes under vacuum in order to promote good mixing, and then overnight in a nitrogen atmosphere. The liquid water was removed by blowing nitrogen through the mixture with continued stirring. 2.2
Reactions and workup
Reactions were carried out in 27 ml stainless steel autoclaves fitted with a stainless steel liner and charged with 2.1g of 105oC-dried oil shale or torbanite (unless otherwise stated) and 3MPa (cold) of the appropriate gas (10 MPa H2 for the Sn-treated torbanite reaction). The autoclave was evacuated and weighed before and after the gas was charged, so that the free space in the autoclave could be calculated. In the case of H2 reactions, the autoclave was evacuated and weighed before and after charging a known pressure of N2 to determine the free space in the autoclave, and then the autoclave was vented and evacuated before again weighing, charging with H2 and weighing. The autoclave was lowered into a preheated sand bath and came to the required temperature in 2-4 minutes. It was continuously shaken while in the sand bath. The autoclave was held at temperature for 1 hour, removed, allowed to cool and weighed. The gas was vented through an Agilent 3000 Micro Gas Chromatograph, equipped with four columns: MS 5A PLOT, 10m×0.32mm (110oC column temperature) for N2 and CH4, PLOT U, 8m×0.32mm (100oC column temperature) for CO, CO2, C2 hydrocarbons, H2S and COS. Alumina PLOT, 10m×0.32mm (140oC column temperature) for C3-C5 hydrocarbons and OV-1, 10m×0.15mm×2.0μm (90oC column temperature) for isobutane and n-hexane. The inlet and injector temperatures were 100oC. After the analysis the autoclave was vented and the total weight of the gas in the autoclave was calculated. On the assumption that the gases detected (and hydrogen or nitrogen by difference) were the only gases present, the yields of all the gases detected and the hydrogen consumption (in runs with hydrogen) could be calculated. The solid and liquid products were washed and scraped out of the autoclave with dichloromethane into a flask, and subjected to Lundin distillation to remove water. The waterfree material was ultrasonicated for 10 minutes, filtered, more dichloromethane added to the filter cake and the process repeated. The dichloromethane insolubles were dried at 105oC in flowing N2 for at least 2 hours, cooled and weighed. Most of the dichloromethane was removed from the dichloromethane-soluble material by rotary evaporation and n-hexane (20:1 by weight) added. The mixture was ultrasonicated for 3 minutes and filtered to give insoluble material, asphaltene, which was dried in a vacuum oven (0.1KPa) at 55oC and weighed. Most of the n-hexane was removed from the filtrate by rotary evaporation and the residual hexane solubles (oil) were stored at 4oC for later analysis. The oil yield was determined by difference. 2
Oviedo ICCS&T 2011. Extended Abstract
Based on weighing uncertainties and the spread of results for replicated runs, the uncertainty in CH2Cl2 solubles and gas was about ±1.5 wt% db, and in oil+water and asphaltene yields about ±2 wt% db. The uncertainty in the yields of individual gases taking into account weighing and calibration uncertainties was about 15% of the result, but the spread of results for replicated runs indicated higher uncertainties. 2.3
Product analysis
Some oil samples were analysed by gas-chromatography-mass spectrometry (GC-MS) on a HP6890 instrument in split mode. For GC, a HP 19091S-433 capillary column (HP-5MS 5% phenylmethyl siloxane), 30m long, 0.25mm diameter, 0.25μm nominal film thickness, was used. The inlet temperature was 230oC. The oven temperature was initially held at 50oC for 2 min then raised to 200oC at a rate of 4oC/min, held at 200oC for 2 min, then raised to 300oC at 8oC/min and held at 300oC for 3 min. For MS, the ionizing potential was 70 eV, the accelerating voltage 1.9 kV, the mass range scanned 45-600 m/z and the ion source temperature 200-250oC. For 1H NMR, the oils were dissolved in CDCl3 and spectra were obtained using a 400MHz instrument, with a 90o pulse flip angle (9.5μs). 3.
Results and Discussion
3.1
Characterisation of the oil shale
The Jordanian oil shale was obtained from the El Lajjun deposit in the Karak region. The moisture content was 1.3 wt%. The organic content as determined from the ash yield at 490oC was 24.2±0.5 wt% db. The acid-soluble fraction of the oil shale was 51.9±0.5 wt% db, indicating a high carbonate concentration. The total S and Fe were 3.6 and 1.9 wt% db respectively and the organic material in the shale had C, 71.9; H, 8.7; N, 1.7; S, 9.9 and O (by difference) 7.8 wt% dmmf of the total oil shale. These values are typical of those reported previously [6]. A solid state 13 C NMR CP/MAS spectrum of the acid-washed shale showed a Car to Caliph of 0.21: 0.79 indicating a high aliphatic content, as implied by the H/C atomic ratio for the total organic material of 1.44. 3.2
Characterization of torbanite
The torbanite came from the Greta coal measures at Muswellbrook (Bayswater Colliery) in the Sydney coal basin, NSW. The moisture content was 0.7 wt%. The ash yield at 490 oC was 4.3± 0.5 wt% db. The ultimate analysis gave C, 82.62; H, 10.64; N, 0.79; S, 0.10; O (by difference), 5.85 wt% dmmf of the torbanite. The solid state 13C NMR CP/MAS spectrum showed, as for the oil shale, a strongly aliphatic character with a Car to Caliph of 0.12 : 0.88. The atomic H/C ratio of 1.64 was even higher than that of oil shale. 3.3
Extraction yields
Table 1 compares the yields of product fractions from reactions at 355 oC, 390 oC and 425 oC. The Jordanian oil shale (JOS) and torbanite (TOR) samples showed very different reactivity patterns. The JOS broke down to give significant yields of asphaltene and oil even at 355 oC in either H2 or N2, whereas the TOR showed hardly any conversion. Reactions of brown coals with much higher organic O content than JOS also show very low conversions to liquid products in this temperature range [7].
3
Oviedo ICCS&T 2011. Extended Abstract
Table 1: The extraction yields (wt % dmmf) of product fractions for JOS and TOR at different temperatures. T (oC) Reactant Gas
JOS
Asphaltene
N2
70.7±6.6
4.9±1.8
62.8±6.5
3.0±3.6
1.5±1.6
H2
89.6±1.1
6.1±0.1
72.2±5.0
11.3±6.0
1.5±0.2
H2
82.6
6.2
64.0
12.4
0.9
N2
86.1±2.6
12.6±0.4
72.4±2.2
1.2±0.04
0.8±0.03
H2
86.9±2.6
13.8±0.4
70.3±2.2
2.8±0.1
1.7±0.1
c H2
71.9
13.4
55.5
3.0
1.8
N2
78.5±5.4
18.4±0.3
57.9±4.8
2.1±0.9
1.2±0.1
H2
84.1±2.5
9.2±1.6
74.7±3.9
0.3±0.2
0.2±0.2
N2
14.8±0.5
1.1±0.04
13.5±0.4
0.1±0.00
0.4±0.01
H2
14.0±0.4
1.3±0.04
12.7±0.4
0.04±0.00
0.1±0.00
N2
64.3
26.8
33.8
3.6
3.3
H2
69.7
30.0
38.3
1.5
0.7
N2
6.4
0.4
5.9
0.1
0.6
H2
3.5
0.3
3.2
0.0
0.1
c
425 TOR
JOS 390 TOR
JOS 355 TOR
b
CH2Cl2-CO2
Oil+H 2O HC+Sulfide gas
CO2
a - extraction for 1 hour, errors are standard deviations calculated from the results for duplicate reactions. b - oil + asphaltene + hydrocarbon and sulfide gas c - tin treated The H/C ratios for JOS and TOR samples are similar and the dmmf O content is only slightly higher in the JOS. The main difference in the elemental composition is the presence of a 10% organic S content on a dmmf basis in JOS. The dissociation enthalpy of C-S bonds in aliphatic systems is less than that of C-O or C-C bonds (CH3-S-CH3, 301 kJ/mol, CH3-O-CH3, 352 kJ/mol and CH3-CH3, 352 kJ/mol) [8] and this could explain the greater reactivity at lower temperatures of JOS. Increasing the reaction temperature to 390 oC resulted in only a small change in the reactivity of TOR but for JOS there was not only an increase in the total conversion, but also a major shift in product distribution from asphaltene to oil. This shift preference for oil was particularly noticeable for reactions in H2, which may be associated with the significant Fe content of the shale (1.9 wt% db) having a catalytic effect. Reactions of JOS at 425 oC showed no significant increase in the oil yield over reactions at 390 oC and the increase in the total conversion was mainly associated with increase in gases. In contrast, reactions of TOR at 425 oC in both N2 and H2 showed high conversion to oil with small amounts of asphaltene and gas. A structural formula proposed for torbanite has alkene bonds, remote from the termini of the long aliphatic chains [9]. Preferential thermal cleavage of C-C bonds near to the alkenes in this homogeneous polymeric structure could be responsible for the facile conversion to oil products. The use of Sn catalysts was examined for reactions of both JOS and TOR with H2 at 425oC. Addition of tin compounds to hydrogen reactions has been shown to have beneficial effects on oil yield from plant derived coals of all ranks [10]. However, addition of Sn to reactions of both JOS and TOR surprisingly reduced the overall conversion by reduction of the oil yield. Asphaltene and gas yields were unchanged.
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Oviedo ICCS&T 2011. Extended Abstract
3.4
Product Characterization
The oils were examined by 1H NMR and GC-MS. 3.4.1.
1
H NMR
The oils obtained from reactions of both JOS and TOR at 425oC are listed in Table 2. The hydrogen types are classified following Bartle and Jones [11]. Table 2: 1H NMR data for JOS and TOR at 425oC 1
Reactant
H NMR
gas Har
Hα
Hβ
Hγ
JOS
N2
0.04
0.10
0.56
0.30
JOS
H2
0.07
0.17
0.59
0.17
TOR
N2
0.05
0.10
0.67
0.18
TOR
H2
0.08
0.15
0.59
0.17
Under these conditions, where high oil yields were obtained (Table 1), low Har and high Hβ values were found in all cases. The oils from reactions of JOS with H2 and TOR with both H2 and N2 showed very similar spectra. The oil from reaction of JOS with N2 was obtained in lower yield (Table 1) and had a higher Hγ value, suggesting that more extensive fragmentation of the aliphatic chains had occurred. 3.4.2. GC-MS data GC-MS data for the oils from reactions of JOS and TOR at 425 oC in both N2 and H2 are shown in Figure 1.
a
b
c
d
Figure 1. Total ion chromatogram of GC-MS (a) JOS oil at 425oC under N2 (b) JOS oil at 425oC under H2 (c) TOR oil at 425oC under N2 (d) TOR oil at 425oC under H2 The chromatograms in general are consistent with 1H NMR data in that the trace for the oil from reaction of JOS with H2 is very similar to those of TOR with both H2 and N2, whereas the oil from JOS with N2 shows a lower proportion of long chain hydrocarbons.
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Oviedo ICCS&T 2011. Extended Abstract
4.
Conclusion
Although the reactivity of JOS and TOR are significantly different, the structure of the major product, the oils, from reactions with H2 at 425oC are remarkably similar. The oil from reaction of JOS at 425oC with N2 was obtained in lower yields and 1H NMR and GC-MS data showed evidence for more fragmentation of the long chain hydrocarbons. The comparison of reactions suggests that catalysis of scission of C-S bonds in Jordanian oil shale may open the path to obtaining good oil yields under even milder conditions. Addition of Sn compounds to reactions of both JOS and TOR showed no evidence of catalytic effects, in contrast to what is observed for similar reactions of plant-derived coals. Acknowledgments We thank the Sentient Group and Jordan Energy for financial support and provision of the oil shale sample. References [1] [2] [3] [4] [5] [6] [7] [8] [9] [10] [11]
Cane RF, Albion PR. Organic geochemistry of torbanite precursors. Geochimica et Cosmochimica Acta 1973;37:1543-1549. Zilm KW, Pugmire RJ, Larter SR, Allan J, Grant DM. Carbon-13 CP/MAS spectroscopy of coal macerals. Fuel 1981;60:717-722. Hamarneh Y, Al-Ali, J., Sawaqed, S., Oil Shale Resources Development in Jordan. Amman, Jordan. Natural Resources Authority; 2006 Abed AM, Arouri KR, Boreham CJ. Source rock potential of the phosphorite-bituminous chalk-marl sequence in Jordan. Marine and Petroleum Geology 2005;22:413-425. Redlich P, Jackson WR, Larkins FP. Hydrogenation of brown coal. 9. Physical characterization and liquefaction potential of Australian coals. Fuel 1985;64:1383-1390. Haddadin RA, Mizyed FA. Thermogravimetric analysis kinetics of Jordan oil shale. Industrial & Engineering Chemistry Process Design and Development 1974;13:332-336. Hulston CKJ, Redlich PJ, Jackson WR, Larkins FP, Marshall M, Sakurovs RJ. Reactivity and structure of two coals containing significant methoxy group concentrations. Fuel 1995;74:1865-1869. Lide DR. CRC Handbook Of Chemistry And Physics Version 2005. Boca Raton, Florida, U.S.A. CRC Press; 2004 Knights BA, Brown AC, Conway E, Middleditch BS. Hydrocarbons from the green form of the freshwater alga Botryococcus braunii. Phytochemistry (Elsevier) 1970;9:13171324. Redlich PJ, Jackson WR, Larkins FP. Studies related to the structure and reactivity of coals. 15. Conversion characteristics of a suite of Australian coals. Fuel 1989;68:231-237. Bartle KD, Jones DW. Nuclear magnetic resonance spectroscopy. Anal. Methods Coal Coal Prod. 1978;2:103-160.
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Oviedo ICCS&T 2011. Extended Abstract
Ethanol effect on the average structural parameters of IDF soot soluble organic fraction Maurin Salamanca, Mauricio Velásquez, Fanor Mondragon, Alexander Santamaria* Universidad de Antioquia, Medellín, Colombia *Corresponding author:
[email protected] Abstract In this study, conventional analytical methods such as infrared spectroscopy, 1H-NMR, elemental analysis and vapor pressure osmometry (VPO) were combined to characterize the soot precursor material present in the soluble organic fraction (SOF) of young soot of aliphatic and aromatic inverse diffusion flames doped with ethanol. The results of this study indicated that the aliphatic fraction of SOF decreases as the height above the burner increases due to thermal effects, and although this behavior was also observed for the ethanol-doped flames, the aliphatic character and the evidence of oxygenated species of SOF will depend on the fuel chemical nature. It could be observed that the aromatic-ethanol flames produce SOF with higher aliphatic and oxygenated species compared to the undoped flame, whereas the aliphatic-ethanol flames showed the opposite. The increase in the oxygenated functional groups on samples indicates that part of the oxygen coming from ethanol was incorporated into the SOF.
1. Introduction In the lasts decades the use of biofuels has more acceptance because it reduces fuel petroleum dependence as well as the greenhouse gases emissions from fossil fuels. The development of low-priced processes for obtaining biofuels and the increment on petroleum prices made biofuels an interesting choice [1-2]. In the last decades a variety of oxygenated compounds such as alcohols, ethers and acetals have been studied because of their capacity to change the combustion dynamics of the fuels, especially because they can introduce different kinds of radicals that could increase particulate matter precursors or change the species involved in oxidation processes [3-5]. Ethanol is one of the most studied oxygenated additives because it can be obtained from biomass at reasonable cost and reduces the petroleum dependence on fuels [6-7]. Ethanol as an additive has been studied in several combustion systems, on engines applications and different flames configurations [5]. These works have established a
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Oviedo ICCS&T 2011. Extended Abstract
reduction on soot, low molecular weight PAHs and acetylene emissions [8-10]. However, it has been found that the amount of oxygenated species, specially acetaldehyde and formaldehyde, increased when ethanol was used [11-12]. Some studies have found that the differences in particle nanostructure as well as soot oxidation rate are due to differences on chemical nature of initial fuel [13-14]. However, few studies have been focused on the influences of additives in the chemical composition of soluble organic fraction of soot, also referred as SOF. Santamaria et al, using combined information obtained by nuclear magnetic resonance (NMR), infrared spectroscopy (IR), vapor pressure osmometry (VPO) and elemental analysis, described the SOF coming from benzene and ethylene flames in terms of average structural parameters [15]. Therefore, the purpose of this study was to evaluate the chemical effect of ethanol addition on SOF coming from hexane and benzene flames in terms of average structural parameters, such as chain length and number of fused aromatic rings.
2. Experimental section 2.1. Burner and sampling In this study, benzene, hexane and blends with 20% of ethanol were used as fuels to generate an inverse diffusion flame. The burner is composed of three concentric stainless steel tubes; the central tube was used to supply the air, the annular tube to supply the fuel and the outer tube to generate a nitrogen shield to avoid the interaction of the flame with the surrounding air. The liquid fuel (benzene and hexane with and without ethanol) was delivered using an HPLC pump to a vaporizer at 150°C. A flow rate of 0.75 cm3.min-1 and 1.25 cm3.min-1 for benzene and hexane, respectively was used in the experiments. Then the fuel vapors were mixed and carried towards the burner mouth with nitrogen as carrier. Under these conditions, a flame height of 60 mm was obtained in all cases. Soot sampling was made using a filter system at the inception point and 60 mm height above the burner mouth at the lateral axis of the flame, (Figure 1).
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Oviedo ICCS&T 2011. Extended Abstract
Filter Filtro Llama Flame Venteo Vent
Bomba de Vacum vacio Pump
Trampa Tramp
N2
N2
Air Aire Fuel Combustible
Fuel Combustible
Figure 1. Experimental set up for sampling from flame 2.2. Chemical characterization of SOF Chemical analysis of the soot fraction soluble in chloroform was done. The infrared analysis was made using the KBr pellet method in a Nicolet Magna IR 560 spectrometer with a MCT/A detector operated at 77K on the wavelength range from 600 to 4000 cm1
. Each sample was taken at least three times to estimate reproducibility. For the 1H-
NMR analysis, the extracts were dried and re-dissolved in CDCl3 containing trace amounts of tetramethylsilane (TMS), which was used as an internal chemical shift reference. All spectra were taken in a Bruker AMX 300 spectrometer. Then, each spectra were baseline corrected and integrated manually at least four times and the results were averaged to reduce the uncertainty (less than 5%) generated by the manual adjustment. Elemental analysis was carried out in a CHNSO Leco Instruments. Average molecular weight data of the extractable material of soot generated in the flame was determined by vapor pressure osmometry (VPO) in a Knauer K7000 osmometer using chloroform as solvent and benzyl as calibration standard. All measurements were carried out using sample and standard solutions of 1g/kg and 0,0160 mol/kg respectively.
2.3. Structural Parameters The spectrum of each sample was separated into characteristic regions according to the Santamaria’s work [16]. Then this information was correlated with molecular weight and elemental analysis data to obtain structural parameters, such as those reported in Table 1.
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Oviedo ICCS&T 2011. Extended Abstract
Table 1. Structural parameters definition fa
Aromatic atomic carbon fraction 0.33
fal
0.4
%
2
0.5
1
2
1600
Chain length number of carbon atoms per average structural unit #
#
Ra
0.5
Average number of oxygen on carbonyl groups #
L
1
Aliphatic atomic carbon fraction 0.33
#O
0.4
#
Number of fused aromatic rings per average structural unit 1
#
#
2
Where: C* Molar fraction of carbons obtained by elemental analysis. H* Molar fraction of hydrogen obtained by elemental analysis. O* Molar fraction of oxygen obtained by elemental analysis. MW Average molecular weight of the samples Hγ (Aliphatic hydrogen in methyl group on γ position or further to an aromatic ring), Hβ1 (Alycyclic hydrogen in β position to two aromatic rings), Hβ2 (Aliphatic hydrogen in methyl or methylene groups in β position to an aromatic ring), Hα (Hydrogen of CH, CH2, CH3 on α position to aromatic rings), Hf (Hydrogen fluorene type), Ho (Oleofinic hydrogen) and Ha (Hydrogen on aromatic rings).
3. Results discussion 3.1. Soluble Organic Fraction The Figure 3 shows the mass percentage of soluble organic fraction as function of the flame height and ethanol addition. The amount of SOF decreases as a function of flame height for both reference and ethanol-doped flames. This reduction is mainly caused by competition between thermal and oxidative processes of soot particles upstream of the flame. However, upon comparing the amount of SOF coming from the ethanol-doped benzene flame with the reference benzene flame at a particular position, a 10% increase was observed. This behavior is opposite to that observed in hexane flames, which indicates that oxidation and pyrolysis processes will depend on the chemical nature of starting fuel. For instance, the increment in the amount of SOF observed for ethanol-doped benzene flame is due to an increment in the pyrolysis process through the ethyl radical coming from ethanol decomposition, leading to the production of acetylene and low molecular weight PAHs, whereas, the opposite behavior observed in hexane flame can be explained by an increase in the oxidation process caused by the OH radical, which in turn reduces the soot precursors material. Although the oxidation can also take place in ethanol benzene flames, the degree of pyrolysis overcome the oxidation.
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Oviedo ICCS&T 2011. Extended Abstract 90 75 60 45 30
B
75
SOF (%)
SOF (%)
90
A
15
60 45 30 15
0
0 Benzene
Ben:20%EtOH
Hexane
HAB (mm)
Hex:20%EtOH
HAB (mm) Incepction
60 mm
Figure 3. SOF on chloroform for the soot samples (A) Benzene (B) Hexane
3.2. Infrared spectroscopy The Figure 4 shows the infrared spectra of the SOF coming from hexane and benzene flames with and without ethanol. For hexane flames, it is observed that the C-H stretch signal of aliphatic groups decreases as a function of height from soot inception point up to 60 mm, however the intensity of C-H aromatic signal depends on ethanol addition, for instance, the relative intensities of the CHar/CHal observed for samples taken at 60 mm height of hexane flames, decreased when the ethanol was added. This behavior can be interpreted by two competitive processes, soot precursors oxidation and polycyclic aromatic condensation, which agrees with the information obtained by solubility. On the other hand, when ethanol was added to the benzene flame, a significant increase of the aliphatic and oxygenated functional groups were observed, which is opposite to result obtained for hexane flame. As stated early, the effect of ethanol addition depends on the chemical nature of the starting fuel. Therefore, the increase in the aliphatic character of benzene SOF can be attributed to an increase in the pyrolysis process over the oxidation, leading to the formation of low molecular weight PAHs and aliphatic compounds. Additionally, the presence of oxygenated groups on samples coming from ethanol-doped flames indicates that some additional oxidation steps are taking place in the flame compared to those normally occurring with just the molecular oxygen coming from air.
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Oviedo ICCS&T 2011. Extended Abstract Benzene Benzene:20EtOH
c
d
3500
3000
C=O
b
C-Hal
a
C-Har
Absorbance (a.u)
Hexane Hexane:20EtOH
1600
1200
800
3500
3000 -1
1600
1200
800
Wavelenght (cm )
Figure 4. Infrared spectra for SOF obtained from hexane, benzene, hexane 20%EtOH and benzene 20%EtOH. a, b sample taken at the inception point and c,d sample taken at 60 mm. 3.3. 1H-Nuclear Magnetic Resonance Figure 5 shows the 1H-NMR spectra of SOF of hexane and hexane-20% ethanol flames at the inception point. Ethanol addition did not affect the aromatic hydrogen fraction, which remains constant at this point, but caused a relative increases in the aliphatic hydrogen content, result that corroborates what has been published in the literature (15). At low flame positions, the temperature is high enough to cause ethanol fragmentation into aliphatic and oxygenated radicals. Although, it is though that the aliphatic fragment can be bonded to the aromatic species, this fact has not been corroborated yet. Figure 6 summarizes the hydrogen distribution of SOF coming from hexane flames. Ethanol addition caused an increase in the aromatic character (aromatic hydrogen, Ha) as a function of flame height, followed by a reduction in the aliphatic content (aliphatic hydrogen: Hα, Hβ, Hγ), except for sample taken at the inception point as was described early. On the other hand, the SOF of soot from benzene flames had opposite behavior compared to hexane flame samples (data not shown).
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Oviedo ICCS&T 2011. Extended Abstract Hexane
Hexane:20EtOH Hβ
Signal (a.u)
Ha
Hγ
10
9
8
7
2.5
2.0
1.5
1.0
0.5
0.0
Frequency (ppm)
Figure 5. 1H-NMR spectra of the SOF obtained from hexane and hexane 20%EtOH sample taken at the inception point.
Hexane
0,80 0,70
13
0,60
30
40
0,70 0,60
0,50
0,50
0,40
0,40
0,30
0,30
0,20
0,20
0,10
0,10
0,00
Hexane:20%EtOH
0,80
60.00
Ha Ho
Hf
Hα Hß2 Hß1
H γ
0,00
Ha
Ho
13
30
Hf
Hα
40
60
Hß2 Hß1 H γ
Figure 6. Hydrogen distribution of SOF coming from hexane soot as a function of flame height.
3.4. Elemental analysis and vapor pressure osmometry Table 2 shows the average molecular weight obtained by VPO. The average molecular weight increases with the flame height; this is an expected tendency for precursors and soot growing processes. For benzene flames, the average molecular weight of SOF was ~25% higher compared to samples from hexane flames. This fact is attributable to the chemical nature of starting fuel and its aromatic polymerization rate, being faster for benzene than for hexane, since the last one required much more time to generate the basic aromatic moieties. On the other hand, the ethanol addition caused about 6% reduction on the average molecular weight for both fuels. This reduction can be explained by both a promotion of the oxidation process and/or an increase in the rate of formation low molecular weight PAHs, as apparently occurs in the benzene flame.
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Oviedo ICCS&T 2011. Extended Abstract
Table 2 also shows the C/H ratio obtained by elemental analysis, although there is not an apparent effect of ethanol on this parameter, the results indicate that the aromatic condensation degree is consistent with flame height. Table 2. Average molecular weight and C/H of elemental analysis for SOF Hexane Hexane:20% EtOH Benzene Benzene:20%EtOH HAB MW* C/H MW* C/H MW* C/H MW* C/H Inception 326 1.34 296 1.33 471 2.04 446 1.97 point 60 mm 393 1.41 363 1.41 528 2.22 503 2.11 * MW Molecular weight (g/mol)
3.5. Structural parameters The main structural parameters calculated for samples are shown on Table 3. In general, all samples showed a tendency to increase the aromatic fraction, (fa), followed by a subsequent reduction on the aliphatic fraction, (fal), as flame height increases. However, upon comparing hexane and benzene results, it is clear that benzene flames generate soot precursors with a higher degree of aromatic condensation, (Ra), since the polymerization mechanism through aromatic moieties is faster. On the other hand, whereas the number of fused aromatic rings is higher for samples coming from benzene flame, the average chain length (L) is higher for samples coming from hexane, especially when ethanol is added to the flame, indicating that the pyrolysis process of the aliphatic fragment of ethanol plays an important role in the soot precursor’s formation. It is also observed, that the oxygen content found in SOF samples, related to #O, has a slight increase when ethanol is added to the benzene flame; however this tendency was not observed for hexane, which indicates that the degree of oxidation also depends on the chemical nature of fuel.
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Oviedo ICCS&T 2011. Extended Abstract
Table 3. Structural parameter calculated for SOF.
fa
Hexane Hexane:20% EtOH Benzene Benzene:20%EtOH In. 60 In. 60 In. point 60 mm In. point 60 mm point mm point mm 0.78 0.76 0.78 0.82 0.90 0.93 0.85 0.88
fal
0.18
0.17
0.17
0.11
0.07
0.05
0.09
0.07
#O
0.86
2.09
0.98
1.82
0.80
0.13
1.02
0.24
L
4.44
4.76
3.74
3.00
1.01
0.40
2.71
2.45
Ra
5.32
6.03
4.74
4.96
12.8
15.1
11.4
13.5
Structural Parameter
4. Conclusions The effect of ethanol added to the fuels evaluated in this study depended on the chemical nature of the initial fuel. For hexane flames, it was observed that the addition of ethanol increased the aromatic character of SOF, although the average number of fused aromatic rings remains constant (Ra∼5) compared to that obtained from pure hexane. In contrast, the ethanol addition to benzene flame caused an increase in the aliphatic character of SOF samples, which corresponds to an increase of 2 aliphatic carbons by structural unit. The reason for this behavior is due to the dual effect of ethanol, which will depend on the chemical nature of fuel. One effect is the effect of the aliphatic fragment of ethanol in the formation of soot precursors in aromatic flames. The other effect is due to the OH radical which will play an important role in the oxidation process of the aliphatic flames. These results demonstrate that ethanol addition can change the chemical nature of soot and its precursors obtained in IDFs. A final question is ¿can ethanol produce the same effect on soot chemistry, when is it used as additive on traditional combustion devices?
Acknowledgements The authors are grateful to COLCIENCIAS and University of Antioquia for the financial
support
given
through
the
Project
1115-405-20283.
M.S.
thanks
COLCIENCIAS and the University of Antioquia for her PhD scholarship. M.V also thanks COLCIENCIAS for the economical support granted during his undergraduate studies.
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Oviedo ICCS&T 2011. Extended Abstract
References [1] Gaffney JS, Marley NA. The impacts of combustion emissions on air quality and climate – From coal to biofuels and beyond. Atmospheric Environment. 2009;43:23-39. [2] Cardona CA, Sánchez OJ. Fuel ethanol production: Process design trends and integration opportunities. Bioresource Technology 2007;98:2415-57. [3] Kohse-Höinghaus K, Oßwald P, Cool TA, et al. Biofuel Combustion Chemistry: From Ethanol to Biodiesel. Angewandte Chemie International Edition. 2010;49:3572-97. [4] Inal F, Senkan SM. Effects of oxygenate additives on polycyclic aromatic hydrocarbons(pahs) and soot formation. 2002;174(9):1-19. [5] Litzinger T, Colket M, Kahandawala M, et al. Fuel Additive Effects on Soot across a Suite of Laboratory Devices, Part 1: Ethanol Combustion Science and Technology. 2009;181(2):31028. [6] Lin Y, Tanaka S. Ethanol fermentation from biomass resources: current state and prospects. Applied microbiology and biotechnology. 2006;69(6): 627-42. [7] Galbe M, Zacchi G. A review of the production of ethanol from softwood. Applied microbiology and biotechnology. 2002;59(6): 628-38. [8] Wu J, Song KH, Litzinger T, et al. Reduction of PAH and soot in premixed ethylene–air flames by addition of ethanol. Combustion and Flame. 2006;144: 675-87. [9] Lapuerta M, Armas O, Herreros JM. Emissions from a diesel–bioethanol blend in an automotive diesel engine. Fuel. 2008;87: 31-7. [10] Therrien RJ, Ergut A, Levendis YA, Richter H, Howard JB, Carlson JB. Investigation of critical equivalence ratio and chemical speciation in flames of ethylbenzene–ethanol blends. Combustion and Flame. 2010; 157: 296-312. [11] McEnally CS, Pfefferle LD. The effects of dimethyl ether and ethanol on benzene and soot formation in ethylene nonpremixed flames. Proceedings of the Combustion Institute. 2007;31: 603-10. [12] Yao C, Yang X, Raine RR, Cheng C, Tian Z, Li Y. The Effects of MTBE/Ethanol Additives on Toxic Species Concentration in Gasoline Flame. Energy & Fuels. 2009;23: 354348. [13] Alfè M, Apicella B, Barbella R, Rouzaud J-N, Tregrossi A, Ciajolo A. Structure–property relationship in nanostructures of young and mature soot in premixed flames. Proceedings of the Combustion Institute. 2009;32: 697-704. [14] Wal RLV, Tomasek AJ. Soot oxidation: dependence upon initial nanostructure. Combustion and Flame 2003;134: 1–9. [15] Santamaría A, Eddings EG, Mondragón F. Effect of ethanol on the chemical structure of the soot extractable material of an ethylene inverse diffusion flame. Combustion and Flame. 2007;151: 235-44. [16] Santamaría A, Mondragón F, Molina A, Marsh ND, Eddings EG, Sarofim AF. FT-IR and 1H NMR characterization of the products of an ethylene inverse diffusion flame. Combustion and Flame. 2006;146:52-62. Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
An Analysis Of The Research Performed With The Argonne Premium Coals And Its Contribution To Coal Science Jonathan P. Mathews1, Yesica E. Alvarez1, Randall E. Winans2 1
John and Willie Leone Family Department of Energy & Mineral and EMS Energy Institute, 110 Hosler Building.# The Applied Research Laboratory, The Pennsylvania State University, University Park 16802, USA 2 X-ray Science Division, Argonne National Laboratory, Argonne, IL 60439, USA.
[email protected] Abstract The Argonne Premium Coal program collected eight USA coals of industrial relevance and provides well-mixed, -characterized, -preserved coal samples in borosilicate glass vials to the coal science community. Since being available in 1985, 33,235 ampules have been shipped. These ampules were sent around the globe and have been utilized in a wide variety of coal characterizations and studies. They are thus the most well evaluated coals in existence. The peer-reviewed web of science journal literature was examined for the terms ‘Argonne Premium Coal’, the individual coal names, and papers citing the seminal 1990 Vorres Argonne Premium Coal paper. The resulting 600+ Argonne Premium Coal journal articles were examined to determine which coals in the suite had been evaluated and to generally classify the research topic. Wordle and other analyses were used to determine the interest in various coals, topics, and author contributions.
The Illinois no. 6 bituminous coal was found to be the most well-studied coal appearing in >180 journal articles. Many of the other coals are well-studied with >130 papers. The “least well-studied” coals were the Lewiston-Stockton and the Blind Canyon bituminous coals with around 90 and 110 articles respectively. This is consistent with the vial shipment data evaluated by Vorres et al. in 1994.
A Wordle analysis of the article titles, an analysis of the general subject, and an analysis of the journal showed: 1) frequent use of the terms: analysis, structure, carbon, solvent, swelling, NMR, spectroscopy, and sorption. 2) The leading authors (determined by this approach) were Takanohashi, Kelemen, Suuberg, Kwiatek, Iino, Kandiyoti, Larsen, Wertz, Gorbaty, Saito, Schroeder among 140 authors. 3) When classed by subject,
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Oviedo ICCS&T 2011. Extended Abstract
structure, NMR, swelling, extraction, sorption, pyrolysis and liquefaction were frequent subject areas in a widely diverse field. The missing piece of this body of work is the databases, created here, to ease the discovery of the state-of-knowledge.
The creation of this sample suite has resulted in ad-hoc international-peer-reviewed literature collaboration. The success of the program has been immense and the field of coal science owes Argonne and the Office of Basic Energy Sciences (Chemical Sciences Division) a thank you for this exceptional gift of still expanding knowledge.
1. Introduction Coal is complex and highly variable in behavior and composition, within multiple aspects, making its systematic study extremely challenging. It is thus highly desirable to have coal standards that are available to the community, such that a body of knowledge can be produced and the field can more forward with scientific exploration building on the existing data suite. There have been several coal sample banks but in the U.S. the two most utilized sample banks are the Argonne Premium Coal suite [1] and the Pennsylvania State University Coal Sample Bank [2, 3]. They serve different purposes; the Argonne Premium Coal suite is 8 coals shipped in borosilicate glass vials (5g of 100 mesh or 10g of -20 mesh) while the Penn State Coal Sample Bank contains data for >1450 coals with many historic samples (>1100 available) and a suite of 38 wellpreserved Department of Energy Coal Samples (DECS) samples typically available in 300 g, 2.5 kg, or 12 kg foil multilaminate bags [4]. Thus the Argonne Premium Samples are well suited for advanced characterization and small-scale behavioral investigations.
For coal selection a cluster analysis of 200 samples in the Penn State Coal Sample Bank was used to determine desirable range of parameters for compositional selections of C, H, N, O, and S for coal mined in the United States. Five of the coals were selected based on C content across the rank range: lignite to low-volatile bituminous coal (anthracite although historically important and actively mined in small quantities was not included). Variation in S content and desire to include a coking coal (Pittsburgh) further expanded the selection. The inclusion of Blind Canyon was for its high liptinite content, similarly Lewiston Stockton for high sporinite and inertinite contents [1]. Table 1 shows the coal sample, rank and state. There are currently two extensive structural reviews available: Pocahontas No. 3 [5] and Illinois No. 6 [6]. Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
Table 1. Argonne Premium Coal Samples, State and Rank Number
Seam
State
Rank
1
Upper Freeport
PA
mvb
2
Wyodak-Anderson
WY
subbit.
3
Illinois No. 6
IL
high-vol bit.
4
Pittsburgh (No. 8)
PA
high-vol bit.
5
Pocahontas No. 3
VA
low-vol bit.
6
Blind Canyon
UT
high-vol bit.
7
Lewiston-Stockton
WV
high-vol bit.
8
Beulah-Zap
ND
lig.
Previously the status of the Argonne Premium Coal Sample Program was evaluated in 1987 [7], 1988 [8] and 1994 [9]. The goal here is to update that information and to create a database that will better serve the coal community by determining which coals have been studied throughout the broad literature. Furthermore, the paper evaluates the impact of this program on coal science.
2. Methodology An ISI Web of Knowledge (using the web of science database) evaluation of journal articles with “Argonne Premium Coal” in the title, as well as searching for individual coal names, was used to identify appropriate journal articles. A search was also performed to determine additional papers based on a citation evaluation of the 1990 Vorres paper “The Argonne Premium Coal Sample Program” [1]. The data was transferred into an Endnote library and supplemented by the lead authors Endnote library that had utilized Argonne Premium as keywords (to better enable coal data collection). The papers were examined to determine which of the suite had been utilized and non-Argonne suite coals were rejected and duplicate citations removed. Keywords were also added to the Endnote references to classify the general research area: combustion, pyrolysis, liquefaction, extraction, gasification, swelling, drying, characterization, adsorption, structure, among many others, as well as the individual coals studied. Search capabilities of Endnote allowed the determination of the number of journal articles that had evaluated which coal. The analysis was performed in December 2010 and so would miss more recent contributions. The ISI Web of Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
Knowledge and supplemented journal articles are certainly an underestimation of the work. The inclusion of the lead authors’ database may also bias the selection. However, it is an excellent starting point and considerations are underway for the sharing of the database and the inclusion of omissions.
Endnote was also used to generate an authors list and the journal titles listing for evaluation with Wordle [10]. Wordle creates a visual representation of text frequency. The terms “Argonne Premium Coal” and “Coal/Coals” were removed from the journal titles with Words’ find and replace function to allow a visual representation of the word frequency in the titles. Wordle removes authors’ initials thus the evaluation is at risk of over promoting authors with common surnames. Wordle distinguishes between the English spelling variations and also between singular and plural and so in titles solvents and solvent will appear as separate entitles. Subject classifications were also evaluated using Wordle.
3. Results and Discussion For the breakdown of the frequency of the coals evaluation in journal articles Illinois no. 6 bituminous coal was found to be the most well studied coal appearing in >180 journal articles. Many of the other coals are well studied with >130 papers. The “least well-studied” coals were the Lewiston-Stockton and the Blind Canyon bituminous coals with around 90 and 110 articles respectively. This is consistent with the vial shipment data evaluated by Vorres et al. in 1994 [9]. The Lewiston-Stockton coal is the liptiniterich coal sample while Blind Canyon is the high sporinite and inertinite content sample, which probably explains the reduced frequency. Much of the work has evaluated the rank range suite of which the 2 previous non-vitrinite maceral-rich coals are considered “spoilers”.
Figure 1 shows a Wordle view of the titles of journal articles that evaluated the Argonne Premium Coal suite. The color is an aid for viewing; it is the size of the word that provides an indication of the frequency of its occurrence. The frequent words were analysis, structure, carbon, solvent, swelling, NMR, spectroscopy, and sorption, etc. The figure is also word-rich indicating a wide variety of research topics.
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Oviedo ICCS&T 2011. Extended Abstract
Figure 1. A Wordle view of the frequency of terms in the titles of Argonne Premium Coal articles. Figure 2 shows a Wordle view of the authors. Among the 140 authors the prominent authors were: Takanohashi, Kelemen, Suuberg, Kwiatek, Iino, Kandiyoti, Larsen, Wertz, Gorbaty, Saito, Schroeder among others. This represents a range of international authors from academia, industry, and government entities.
Figure 2. A Wordle view of the frequency of authors in the titles of Argonne Premium Coal articles.
Figure 3 shows the Wordle view of the frequency of classification terms for Illinois no. 6, the most well-studied of the coals. As expected many of the papers deal with characterization as well as behavior with swelling, extraction, and with liquefaction.
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Figure 3. A Wordle view of the frequency of classification terms for Illinois no. 6 coal articles.
It is difficult to place a research value to the body of work that the Argonne Premium suite has enabled but it is likely in excess of $20 million (assuming conservatively that $100,000 could produce 3 journal articles). We have, at our fingertips a body of knowledge that spans chemical, physical and behavioral aspects. These contributions to our understanding of coal structure and reactivity are highly significant and could not be covered in a short discussion. However, select examples will be. Characterization of sulfur types has always been difficult until the X-ray spectroscopy technique (SXANES) was developed by Kelemen and Gorbaty [11]. They were able to determine the sulfur types in these coals and have expanded it to other coals and kerogens [12, 13]. However, they continually used the Argonne coals as a test when trying new approaches with this technique. Solids NMR has been an excellent tool to look directly at the coal structure but there was always uncertainties about the quantitation. A large number of groups around the world evaluated these coals and compared their results in a bookproducing ACS symposium [14]. More recently when carbon sequestration in coal seams became important, the Argonne coals where used to evaluate CO2 uptake in coals by a number of labs [15, 16]. The coal suite, and more importantly the data generated has enabled coal science to advance more rapidly, strategically, and with more meaningful results than would have otherwise been possible due to the ad-hoc collaboration of numerous universities, government agencies, and industry worldwide.
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Oviedo ICCS&T 2011. Extended Abstract
Conclusions The Argonne Premium Coal program was established to provide research samples of a small number of representative US coals. The samples are homogeneous and protected from oxidation allowing meaningful research from different labs and different times can be compared without worrying about the history of the sample. Since being available in 1985, 33,235 ampules have been shipped with their resultant data being produced in the peer-review literature. A ISI Web of Knowledge (using the web of science database) evaluation of journal articles with “Argonne Premium Coal” in the title, as well as searching for individual coal names, was used to identify appropriate journal articles in an Endnote database that were used to evaluate the coals studied, subject areas, and authors contributions. From the suite, Illinois no. 6 bituminous coal was found to be the most commonly shipped and most well studied coal appearing in >180 journal articles. Many of the other coals are well studied with >130 papers. Wordle was used to evaluate word frequencies from titles and authors contributions. The frequent words within the journal article titles were: analysis, structure, carbon, solvent, swelling, NMR, spectroscopy, and sorption, etc. Among the 140 authors the prominent authors were: Takanohashi, Kelemen, Suuberg, Kwiatek, Iino, Kandiyoti, Larsen, Wertz, Gorbaty, Saito, Schroeder among others. This represents a range of international authors from academia, industry, and government entities. The coal suite, and more importantly the data generated has enabled coal science to advance more rapidly, strategically, and with more meaningful results than would have otherwise been possible due to the ad-hoc collaboration of numerous universities, government agencies, and industry worldwide. Work is continuing and efforts are underway to further capture the body of knowledge and to share this database resource with coal scientists worldwide.
Acknowledgements. Many have contributed to the success of the Argonne Premium Coal Program. The original planners were: Randall Winans, Phil Horwitz, John Unik, Gary Dyrkacz and John Young (facility designer); Karl Vorres built the processing facility, acquired the samples and managed the facility; the staff of the U. S. Geological Survey who supervised the sample collection: C. Blaine Cecil, Ron Stanton and Peter Warwick. Over the last few decades the program manager has been: Karl Vorres, Ken Anderson, Jerry Hunt, Randall Winans with Deborah Vervack the facility administrator. Program Submit before May 15th to
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funding by the Office of Basic Energy Sciences (Chemical Sciences Division) was responsible for allowing this coal suite collection and operation. In addition many coal scientists have contributed to the success of the program and their individual efforts are also recognized.
References [1] Vorres KS, The Argonne Premium coal sample program. Energy Fuels 1990, 4, 4206. [2] Glick DC, Davis A, Operation and composition of the Penn State Coal Sample Bank and Data-Base. Organic Geochemistry 1991, 17, (4), 421-30. [3] The Pennsylvania State University Penn State Coal Sample Bank and Data Base. http://www.energy.psu.edu/copl/index.html [4] Glick DC, Mitchell GD, Davis A, Coal sample preservation in foil multilaminate bags. Int. J. Coal Geol. 2005, 63, 178-89. [5] Stock LM, Muntean JV, Chemical constitution of Pocahontas No. 3 coal. Energy Fuels 1993, 7, 704-9. [6] Castro-Marcano F, Mathews JP, Constitution of Illinois No. 6 Argonne Premium coal: a review. Energy Fuels 2011, 25, (3), 845-53. [7] Vorres KS, Preparation and distribution of Argonne Premium coal samples, Prepr. Pap.-Am. Chem. Soc, Div. Fuel Chem., 1987, New Orleans, LA, 32, 221-6 [8] Vorres KS, Sample preparation for, and current status of the Argonne Premium Coal sample program, 1988, Los Angeles, CA, 33, 1-6 [9] Vorres KS, Kruse CW, Glick DC, Davis A, Nater KA, A perspective on the status of coal research from shipments of samples, Prepr. Pap.-Am. Chem. Soc., Div. Fuel Chem., 1994, San Diago, CA, 39, 1-6 [10] Feinberg J Wordle. http://www.wordle.net/ [11] George GN, Gorbaty ML, Kelemen SR, Sansone M, Direct determination and quantification of sulfur forms in coals from the Argonne Premium Sample Program. Energy Fuels 1991, 5, (1), 93. [12] Gorbaty ML, Kelemen SR, Characterization and reactivity of organically bound sulfur and nitrogen fossil fuels. Fuel Process. Technol. 2001, 71, (1-3), 71-8. [13] Kelemen SR, Afeworki M, Gorbaty ML, Kwiatek PJ, Sansone M, Walters CC, et al., Thermal transformations of nitrogen and sulfur forms in peat related to coalification. Energy Fuels 2006, 20, (2), 635. [14] Botto RE, Sanada Y, Magnetic Resonance of Carbonaceous Solids. American Chemical Society: 1992; Vol. 229. [15] Goodman AL, Busch A, Bustin RM, Chikatamarla L, Day S, Duffy GJ, et al., Inter-laboratory comparison II: CO2 Isotherms measured on moisture-equilibrated Argonne Premium coals at 55 degrees C and up to 15 MPa. Int. J. Coal Geol. 2007, 72, (3-4), 153-64. [16] Goodman AL, Busch A, Duffy GJ, Fitzgerald JE, Gasem KAM, Gensterblum Y, et al., An inter-laboratory comparison of CO2 isotherms measured on Argonne Premium coal samples. Energy Fuels 2004, 18, (4), 1175-82.
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Porosity and Gas Absorption of Coals Studied by X-ray Scattering and Modeling Randall E. Winans1, Soenke Seifert1, Darren Locke1, Peter Chupas1, Karena Chapman1, Marielle R. Nariewicz2, Jonathan P. Mathews2, Joseph M. Calo3 1
X-ray Science Division, Advanced Photon Source, Argonne National Laboratory, Argonne, IL 60439, USA 2 Department of Energy & Mineral Engineering and EMS Energy Institute, 126 Hosler Building. The Pennsylvania State University, University Park, PA 16802, USA 3 Brown University, Division of Engineering, Box D, 182 Hope St., Providence, RI 02912, USA
[email protected] Abstract Small angle x-ray scattering (SAXS) and high energy wide angle scattering with pair distribution function (PDF) analysis has been used to study coal structure and to elucidate changes with exposure to high pressure CO2 at ambient temperatures. SAXS provides pore size, size distribution, shape and surface morphology over broad length scales, while PDF provides atom-atom correlations out to several nm. The studies were done on a series of coal pieces in a high pressure cell which was transparent to X-rays. This method allowed the real-time determination of porosity changes due to CO2 uptake or loss resulting in coal swelling and deswelling. The data suggested it may be possible to determine adsorption and pore filling quantitatively. Both SAXS and PDF calculated scattering data from large-scale molecular models of these coals and models with CO2 qualitatively described experimental observations. The series of Argonne Premium Coals where studied and a dramatic rank effect was evident.
1. Introduction Small angle X-ray scattering (SAXS) and high energy wide angle scattering with pair distribution function (PDF) analysis are used to follow changes in coal and oil shale structure as the solids are being exposed to reactive gases or solvents and high pressures.
For example, this approach is being used to better understand the
fundamental changes in coal structure when exposed to pressurized CO2, to model carbon sequestration[1]. SAXS provides pore size, size distribution, shape and surface morphology over broad length scales. PDF provides atom-atom correlations out to at least 2 nm. Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
In early work, SAXS was used to investigate the porosity of coals[2] and determine fractal properties of small pores[3]. More recently SAXS and SANS has been used to better understand the pore structure of coals.[4] It has been demonstrated that CO2 pore-fills and dissolves in coals and causing swelling. In addition the dissolved CO2 appears to plasticize the coal [5, 6]. This swelling however causes errors in quantifying CO2 capacity with common techniques coal swelling[7].
2. Experimental
For coal SAXS and PDF experiments one mm thick samples of the Argonne Premium Coal Samples and a New Zealand subbituminous coal, Ohai, shown in Table 1 were used. Table 1. Coal properties.
Coal
Rank
%C
%H
%N
%S
%O
%
(daf)
(daf)
(daf)
(daf)
(daf)
ash
%H2O
Ohai
Subbituminous
75.3
4.36
1.31
0.42
18.6
4.16
23.1
Illinois 6
hvC bituminous
77.7
5.00
1.12
2.38
13.5
15.5
7.97
Upper
mv Bituminous
85.5
4.70
1.55
0.74
7.51
13.8
1.13
Freeport
The coals were sealed in cells constructed of two conflat flanges with a 1 mm spacing between the kapton windows which were glued over holes drilled in the flanges to allow penetration by the X-ray beam. The cells where pressurized with an Isco syringe pump from 50 - 75 bar gas pressue. The SAXS data were obtained at the 12-ID beamline at the Advanced Photon Source [8]. The scattered X-rays (12 keV) are detected by a CCD detector (Mar 156 or mosaic APS). Data were obtained every two seconds as the samples were pressurized. The scattered intensities have been corrected for absorption, the empty cell scattering, and instrument background. The differential scattering cross section is expressed as a function of the scattering sector Q, which is defined as: Q =
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Oviedo ICCS&T 2011. Extended Abstract
(4π/λ)SinΘ, where λ is the wavelength of the X-rays and Θ is the half angle of the scattering. For the PDF experiments the same cell was used. The cell was mounted on the instrument at beamline 11-ID-B at the Advanced Photon Source, Argonne National Laboratory with the incident (and scattered) beam directed through the hole in the cell. High-energy X-rays (90.48 keV, λ = 0.1370 Å) were used in combination with a MAR345 image plate detector to record diffraction images[9] to high values of momentum transfer (Q ~ 20 Å−1). Corrections for multiple scattering, X-ray polarization, sample absorption, and Compton scattering were then applied to obtain the structure function S(Q). Direct Fourier transform of the reduced structure function F(Q) = Q[S(Q) − 1] up to Qmax ~ 20 Å−1 gave G(r), the pair distribution function. The PDF, G(r), gives the probability of finding an atom at a given distance r from another atom and for shorter distances can be considered as a bond length distribution. It is obtained from the powder diffraction (X-ray or neutron) via a Fourier transform of the normalized scattering intensity, S(Q):
where ρ(r) is the microscopic pair density, ρ0 is the average number density, and Q is the magnitude of the scattering vector (Q = (4πsin θ)/λ). Experimentally it is not possible to measure data up to infinite Q, and the cutoff at finite values of Qmax decreases the real space resolution of the PDF. This causes some aberrations in the form of “termination ripples” which propagate through G(r) as high frequency noise. For both X-ray scattering experiments, high energies (> 60keV) are required to access high values of Qmax to obtain the most accurate Fourier transform of the reduced structure function F(Q). For the large-scale coal model (50,000 atoms) the construction was based on: (1) HRTEM image analysis, (2) construction of the aromatic clusters, (3) inclusion of heteroatoms and functional groups via scripting, (4) addition of aliphatic side chains and generation of a cross-linked network structure, and (5) arrangement of cross-linked clusters into a simulation cell. This procedure enabled simplifying the model construction process, eliminating researcher structural bias and generating large-scale continuum molecular representations with improved accuracy. Molecular dynamics
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Oviedo ICCS&T 2011. Extended Abstract
(MD) simulations and energy minimizations were performed with Materials Studio 5.0 software from Accelrys Inc. [36]. Energetic interactions were modeled by the consistent-valence force field (CVFF) [37-39], which is intended for application to a wide range of organic systems. The non-bonded interactions were modeled with a Lennard-Jones (LJ) 12-6 potential for van der Waals interactions and a Coulomb potential for long-range electrostatic interactions. The LJ parameters for the cross interactions were determined by the standard Lorentz-Berthelot mixing rules. The velocity Verlet algorithm was used to integrate Newton's equations of motion, and the classical conjugate gradient and steepest descents algorithms were utilized for energy minimizations. The time step used in the MD simulations was taken as 1.0 fs.
3. Results and Discussion Preliminary experiments on a suite of North American and New Zealand coals established that SAXS could be used to observe directly the changes in coal structure caused by CO2 uptake. Subsequent experiments have shown that, at least for high rank coals, the process can be interpreted in terms of a simplified two-phase model beginning with shorter term void/pore filling and gas adsorption onto the solid matrix followed by longer term coal swelling.
1.2e-4 Uppper Freeport (APCS 1) Ohai Illinois No. 6 (APCS 3)
Invariant
1.1e-4
1.0e-4
9.0e-5
8.0e-5 0
100
200
300
400
Time (min)
Figure 1. Changes with time of the invariant at 900 psi of CO2.
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Oviedo ICCS&T 2011. Extended Abstract
Decrease in the invariant is the result of a decrease in difference in electron density, which occurs as the CO2 is filling the pores. ∞
Q = ∫0 q2dq·I(q)
These processes are consistent with an overall decrease in scattering intensity with increasing degree of coal swelling due to the disappearance of the smallest scatterers (voids/pores) accompanied by a shift of the normalized void/pore distribution to larger scatterers (Figure 1). The decrease is consistent with swelling of the solid matrix primarily via a Class II type diffusion process typically associated with a glassy polymer or gel structure.
The data suggested it may be possible to determine adsorption and pore filling in the coal samples quantitatively. Both SAXS and PDF calculated scattering data from large molecular models of these coals[10] and models with CO2[11] qualitatively described what is observed in the experiments. A model of a high rank Pocahontas coal filled with CO2 is shown in the top of Figure 3.
1e+9 3
A
1e+8
1e+7
rco2_150707_0111_s5_00432__bsub rco2_150707_0111_s5_00438__bsub rco2_150707_0111_s5_00444__bsub rco2_150707_0111_s5_00450__bsub rco2_150707_0111_s5_00456__bsub rco2_150707_0111_s5_00468__bsub rco2_150707_0111_s5_00480__bsub rco2_150707_0111_s5_00498__bsub rco2_150707_0111_s5_00534__bsub rco2_150707_0111_s5_00570__bsub rco2_150707_0111_s5_00624__bsub rco2_150707_0111_s5_00690__bsub rco2_150707_0111_s5_00780__bsub
0.01 9 8
6 5
-1
I(Q) (cm )
Intensity
7
1e+6
1e+5
1e+4
B
UpperFreeport APCS 1 Blank air subtracted
2
4
3
Coal Coal + CO2
2
0.001 9 8
1e+3
7 6 5
1e+2 0.01
4
0.1
1
Q(A-1)-1
Q(A )
3
4
5
6
7
8
9
2
3
4
5
6
7
0.1 -1
Q (Å )
-1
Q(A )
Figure 2. Modelling of a high rank bituminous coal. A - Calculated SAXS data for model with and without CO2; B – SAXS scattering for Upper Freeport coal (APCS 1) under 900 psi of CO2.
From this model and the model without CO2 the SAXS were calculated using CRYSOL[12]. Since his model has a finite and small size compared to coal samples the scatting does not change a lower Q which is normally observed at lower Q ( < 0.07). However, a decrease in scattering is clearly observed in the model at an intermediate Q
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Oviedo ICCS&T 2011. Extended Abstract
range with the addition of CO2, Figure 2A.
While in the coal the same effect is
observed as the coal swells with the diffusion of CO2 into the coal structure, Figure 2B. This gives one confidence with the validity of the model. The PDF was calculated from the model using the DISCUS method[13]. In Figure 3 the results of the calculated data from the model in compared to a high-rank coal measured PDF. The data is dominated by distances between carbons with the largest peak being the C-C aromatic bond. The second largest peak represents the first and third carbon in C-C-C. The fit is actually very good considered the complexity of the structure being studied. The calculated PDF fit for the CO2 expanded model has been difficult to obtain and is being further investigated.
Figure 3. Comparison of PDF from the high rank model (without CO2) with the PDF of the Pocahontas lv bituminous coal (APCS 5).
. With de-swelling, the scattering intensities increase with time as the CO2 progressively empties from the voids/pores, increasing the electron density contrast and then more slowly evaporates from the swollen solid matrix, re-opening voids/pores that were shrunken. However, the scattering curves of the de-swollen coal did not return to their initial early “unswollen” values. This suggests that, at least on the timescale of hours, there is a significant “memory effect” of the swelling process on the solid coal matrix.
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4. Conclusions The SAXS technique is able to directly observe changes in coal structure that accompany injection of CO2 at room temperature and 5,500 kPa (800 psi) and to provide new understanding of the process. For hig rank coals, the uptake of CO2 may be regarded as a combination of processes beginning with shorter term void/pore filling and gas adsorption onto the solid matrix, followed by longer-term coal swelling. Consideration of the Porod invariant data derived from the initial scattering plots suggests that the solid matrix swells primarily via a Class II type diffusion process typically associated with a glassy polymer or gel structure. Also, the swelling process shows some degree of hysteresis. Comparison of large-scale coal models with experimental scattering data shows reasonable agreement. The effect of CO2 promoted swelling can be observed with SAX Sand used to verify coal representations utility.
Acknowledgement. The authors would like to acknowledge the contributions of the late Tony Clemens, CRL who was a key contributor to the initiation and execution of this project. Use of the Advanced Photon Source, an Office of Science User Facility operated for the U.S. Department of Energy (DOE) Office of Science by Argonne National Laboratory, was supported by the U.S. DOE under Contract No. DE-AC02-06CH11357.
References [1] R.E. Winans, T. Clemens, S. Seifert, In situ SAXS studies on the effects of reactive solvents and gases on coal structure, Preprints of Symposia - American Chemical Society, Division of Fuel Chemistry, 51 (2006) 165. [2] M. Kalliat, C.Y. Kwak, P.W. Schmidt, Small-angle x-ray investigation of the porosity in coals, ACS Symp. Ser., 169 (1981) 3-22. [3] H.D. Bale, P.W. Schmidt, Small-angle x-ray-scattering investigation of submicroscopic porosity with fractal properties, Phys. Rev. Lett., 53 (1984) 596-599. [4] A.P. Radlinski, M. Mastalerz, A.L. Hinde, M. Hainbuchner, H. Rauch, M. Baron, J.S. Lin, L. Fan, P. Thiyagarajan, Application of SAXS and SANS in evaluation of porosity, pore size distribution and surface area of coal, International Journal of Coal Geology, 59 (2004) 245-271.
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[5] J.W. Larsen, The effects of dissolved CO2 on coal structure and properties, International Journal of Coal Geology, 57 (2004) 63-70. [6] A.L. Goodman, R.N. Favors, J.W. Larsen, Argonne Coal Structure Rearrangement Caused by Sorption of CO2, Energy & Fuels, 20 (2006) 2537-2543. [7] V.N. Romanov, A.L. Goodman, J.W. Larsen, Errors in CO2 Adsorption Measurements Caused by Coal Swelling, Energy & Fuels, 20 (2006) 415-416. [8] S. Seifert, R.E. Winans, D.M. Tiede, P. Thiyagarajan, Design and performance of a ASAXS (Anomalous Small Angle X-ray Scattering) instrument at the Advanced Photon Source, Journal of Applied Crystallography, 33 (2000) 782-784. [9] P.J. Chupas, X. Qiu, J.C. Hanson, P.L. Lee, C.P. Grey, S.J.L. Billinge, Rapidacquisition pair distribution function (RA-PDF) analysis, Journal of Applied Crystallography, 36 (2003) 1342-1347. [10] M.R. Narkiewicz, J.P. Mathews, Improved Low-Volatile Bituminous Coal Representation: Incorporating the Molecular-Weight Distribution, Energy & Fuels, 22 (2008) 3104-3111. [11] M.R. Narkiewicz, J.P. Mathews, Visual Representation of Carbon Dioxide Adsorption in a Low-Volatile Bituminous Coal Molecular Model, Energy & Fuels, 23 (2009) 5236-5246. [12] D. Svergun, C. Barberato, M.H.J. Koch, CRYSOL - a program to evaluate x-ray solution scattering of biological macromolecules from atomic coordinates, Journal of Applied Crystallography, 28 (1995) 768-773. [13] T. Proffen, R.B. Neder, DISCUS: a program for diffuse scattering and defectstructure simulation, Journal of Applied Crystallography, 30 (1997) 171-175.
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8
The role of sulfur in coals plastic layer formation L. Butuzova1, R. Makovskyi1, T. Budinova2 , Stefan Marinov2 1
Donetsk National Technical University, 58 Artema str., Donetsk 83000, Ukraine,
tel fax: +38(0622) 55-85-24,
[email protected] [email protected] 2
Bulgarian Academy of Sciences, Institute of Organic Chemistry, 9 Acad. G.Bonchev str., Sofia 1113, Bulgaria,
[email protected]
Abstract Coals blends with all possible combinations of high- and low-sulfur coals of the same rank and petrographic composition were pyrolysed in centrifugal field, allowing to separate solid, fluid and gaseous products. It has been shown that replacement of the low-sulfur by the high-sulfur coking coal of the same rank (≈83–88 % Cdaf) in blends leads to substantial increasing of fluid products yield which can be of great practical value. Using the EPR and DRIFT-spectroscopy methods the structural peculiarities of the obtained products have been studied. An apparent correspondence between the content of sulfur in coal, content of paramagnetic centers in plastic layer and coal coking ability has been discovered. Keywords: coal, blend, sulfur, plastic layer 1. Introduction The differences in coal thermoplastic properties have been attributed to higher hydrogen transfer reactions and to differences in the amount of tar produced during the plastic range [1]. The existing schemes of formation of coal plastic matter ignore a vital role of heteroatoms in coal organic matter (COM) and thus allow no definite prediction of the yield, composition and properties of pyrolysis products. Thermal transformations of COM heterorganic compounds, as well as the processes of donor hydrogen transfer and redistribution, are determined by the molecular structure of coal. Accordingly, studies of the chemical structure of coals with different sulphur and oxygen contents and their behavior during heating should play a crucial role in understanding of cokemaking as a process. A particular characteristic of seams from Donets coal basin is the occurrence of high- and lowsulfur coals, of the same rank differing by some physicochemical properties [1]. These differences are due to specified genetic types formed in alluvial or marine depositional environments during diagenesis processes [2]. The aim of this research is to study the effect of sulphur content in individual components of coal blends on the yield and characteristics of the plastic layer responsible for coking. 2. Experimental Two pairs of the isometamorphic Donets coals homogeneous by their petrographic composition, but formed under reductive (RC) or less reductive conditions (LRC) and different by their sulphur
content (Sdt=1,09-1,22 for LRC; Sdt=2,49-2,81 for RC) were used as objects of research. It was coals of J-Grade (Cdaf = 85,4 - 86,1; Vdaf = 30,5 - 32,7; Sdt = 1,1 - 4,1) and G-Grade (Cdaf =83,8 85,1; Vdaf=36,0 - 38,7; Sdt =1,22 - 2,49 according to Ukrainian classification. On the basis of these coals, blends (J:G=70:30) with all possible combinations of LRC and RC type coals were prepared. These coals and their blends were thermally treated up to 600°C at the rate of 1500 rev/min using the method of centrifugal thermal filtration (Ukrainian National Standard 17621-89). This method enables one to separate primary products that form the plastic mass, immediately, thus preventing their secondary transformations. The yields of the following products were found: the solid oversieve residue (OR), fluid non-volatile products (FNP) and vapour-gas compounds (VG). The amount and composition of FNP largely determine the processes of caking and coking. The theoretical plastic mass yield was calculated for the above blends using the rule of additivity. The EPR-spectra of the coals were recorded on a Bruker ER 200D SRC radiospectrometer at ambient temperature. Active coal with the content of paramagnetic centres (PMC) N=6.25 х 1016 was used as a standard. The IR-spectra were recorded on a Bruker FTS-7 spectrometer using the DRIFT technique. Semiquantitative processing of the IR-spectra was performed with the help of the software package Origin 6.1 using the basic-line technique. 3. Results and discussion Figure1 demonstrates a great difference in FNP yields and, therefore, in the caking capacity of coals of the same brand, but different coal-facies. This indicator is essentially higher (2.5 times) for LRC coals of G-Grade and RC coals of J-Grade as compared with their pairs of the same rank. theoretical
experimental
The
most
(maximum
35
advantageous FNP
yields)
composition is
a
blend
30
containing the reduced J-Grade (JRC) and the
25
low-reduced of G-Grade (GRC) coals. The
20
greatest deviation of the experimental values
15
from the calculated by the rule of additivity
10
is observed for the JRC + GRC blend, which
5
permits an assumption about the strongest
0 Jrc+Grc
Jrc+Glrc
Jlrc+Glrc
Jlrc+Grc
J : G = 70 : 30
Figure1. A comparative characteristics of the calculated and experimental values for plastic mass yields from samples under investigation
interaction between the components [3]. To understand the nature of this interaction it is useful to compare the parameters of the EPR-signals for the original coals, blends and their thermal filtration products.
As is seen from Table1, PMC concentration (N) in the samples under discussion essentially depends on the components genetic type by reductivity. When the GRC coal is transformed into a plastic
state, the basic amount of PMCs remains in the solid product, i.e. the over-sieve residue, whereas in FNP the concentration of PMCs is ≈ 35 times lower. Table 1. The results of EPR and IR-spectroscopy of coals and the products of thermal destruction Coals, Paramagnetic The results of IR-spectroscopy fluid mobile characteristics products, over-sieve N, ΔН, g-factor Relative intensity residues of coals and spin E Ix/I2920 Ix/I1440 Ix/I1600 blends -1 1190 1260 3040 1260 1600 1260 g x -17 10 G LRC 2,24 6,43 2,0039 0,37 0,34 0,30 0,45 1,89 0,24 origin G RC 64,25 6,79 2,0040 0,33 0,32 0,26 0,41 1,80 0,23 coals J LRC 64,28 7,02 2,0039 0,37 0,31 0,23 0,36 1,64 0,22 J RC 43,16 5,21 2,0040 0,31 0,25 0,19 0,31 1,32 0,24 G LRC 38,50 6,06 2,0040 0,28 0,24 0,36 0,45 1,46 0,31 G RC 1,78 6,01 2,0040 0,20 0,18 0,34 0,38 0,93 0,41 J LRC 114,9 6,25 2,0040 0,45 0,39 0,27 0,52 2,07 0,25 fluid J RC 159,9 6,26 2,0040 0,40 0,32 0,23 0,36 1,67 0,21 mobile products GRC+ JRC 110,5 7,07 2,0040 0,42 0,34 0,25 0,41 1,59 0,26 GLRC+JRC 111,4 6,92 2,0040 0,04 0,03 0,03 0,04 0,18 0,19 GLRC+JLRC 43,14 5,79 2,0040 0,55 0,48 0,48 0,70 2,42 0,29 GRC+JLRC 45,38 5,68 2,0040 0,65 0,55 0,42 0,65 2,25 0,29 G LRC 2,0040 2,78 13,5 0,21 G RC 61,14 6,64 2,0040 2,84 1,05 3,22 0,48 5,00 0,10 J LRC 6,5 6,92 2,0040 J RC 4,0 6,60 2,0040 oversieve GRC + JRC 2,0039 1,88 7,64 0,25 residues G +J 0,11 4,00 2,0040 6,03 20,3 0,30 LRC RC GLRC+JLRC 0,03 4,04 2,0040 4,18 2,17 1,93 4,42 GRC+JLRC 0,19 5,23 2,0040 17,2 0,26 The JRC coal is forming FNP with the PMC concentration ≈ 4 times higher than that in the original coal. At the same time, an increase in the yield of liquid thermodestruction products is observed, which is 2.5 times higher for the reduced coals as compared to the low-reduced ones (Fig.1). These results indicate that reactions resulting in the formation of the plastic layer occur with the participation of free radicals. The JRC coals characterized by narrower EPR signals (ΔН ~ 5.21 E), contain the most stable PMCs. The maximum rigidity of the polyconjugated areas is observed for solid residues of thermal filtration of the blends (ΔН ~ 4.00 - 4.04 E). The width and form of broad resonance lines in the EPR-spectra of coals are basically determined by hyperfine interaction with magnetic nuclei [4,5]. Thus, the observed differences in the paramagnetic properties of RC and LRC coals indicate greater
molecular rigidity of polyconjugated areas in the structure of reduced coals, primarily the JRC coal. When coal of JRC is added to coals of brand G (GLRC and GRC), the concentration of PMCs in FNP drastically goes up (by 2.5 times). In the FNP based on JLRC coal the concentration of PMCs is 2.6 times lower. These data obtained permit to attribute the optimal properties of the GLRC + JRC blend to the highest PMC concentration in FNP. The RC samples yield the plastic layer characterized by high content of aliphatic groups (2920сm-1 and 1440сm-1). When JLRC coal replacement by JRC in the blends the relative concentration of CHal groups in the FNP increases. The GLRC + JRC blend is characterized by a minimal relational proportion of -О-(-S-)/CHal groups (I1190,1260/ I2920) and proportion of СНаr/CHal (I3040/I2920 and I1600/I1440) groups. So, FNP of the blend with JRC coal is rapidly saturated with hydrogen and the solid residue - with aromatic and bridge segments, which conduces to the formation of the plastic layer and subsequently to caking coke. A comparison of the IR-spectra of GLRC and GRC coals and their over-sieve residues demonstrates that low-reduced coal and its OR are also distinguished by a much higher Har/Hal and –S– (–O–)/Hal ratio as compared to GRC. 4. Conclusions Generally, a replacement of one of the blend components by coal of the same rank but different sulfur content (different genetic type by reductivity) changes the PMC concentration, structuregroup composition and yield of thermal filtration products. The heteroatoms content in coals determines the quality and quantity of the plastic layer and the character of interaction between the blends components. The data obtained unambiguously indicate that it is necessary to consider the coal-facies when making coking blends. References [1] Nomura M., Kidena K., Hiro M., Murata S. Mechanistic study on the plastic phenomena of coal. Energ. Fuel 2000;14:904-9. [2] Bechtel, A., Butuzova, L., Turchanina, O., Gratzer, R. Thermochemical and geochemical characteristics of sulphur coals. Fuel Processing Technology 2002;77–78:45–52. [3] Zubkova, V.V. Investigation of influence of interaction between coals in binary blends on displacement of non-volatile mass of coal charge during carbonization. Fuel Processing Technology. 2002;76:105–119. [4] Butuzova, L., Krzton, A. and Kozlova, I. The paramagnetic characteristics of pyrolysis products for coals treated by alkali and acid, Proceedings 9th International Conference on Coal Science. Essen (Germany). 1997;1:91–4. [5] Butuzova, L., Rozhkov, S., Makovskyi, R., Rozhkova, N., Butuzov, G. The contribution of radical reactions during thermal processing of low-quality coals. GeoLines. 2009;22, № 5: 9–14.
Oviedo ICCS&T 2011. Extended Abstract
Understanding the effects of biomass addition to coking coals during carbonisation
M. Castro-Díaz1, A. Dufour2, N. Brosse3, R. Olcese2, C. Snape1 1
2
Department of Chemical & Environmental Engineering, Faculty of Engineering, University of Nottingham, Nottingham, NG7 2RD, UK E-mail:
[email protected]
CNRS, Nancy Université, Reactions and Processes Engineering Laboratory (LRGP), ENSIC, 1 rue Grandville, B.P. 20451, 54000, Nancy Cedex, France. 3
Nancy Université, Faculté des Sciences et Techniques, LERMAB, B.P. 239, 54506 Vandoeuvre les Nancy Cedex, France
Abstract The addition of biomass to coking blends has the potential benefits of reducing the amount of expensive coking coals and reducing carbon emissions. Therefore, an easily available biomass (miscanthus) has been chosen to study the effect of this additive on the fluidity properties of the coal. High-temperature 1H NMR and high-temperature rheometry were used to determine the fluidity of the miscanthus and some of its constituents (i.e. lignin, cellulose and xylan). The miscanthus was also pyrolysed at 250°C for 1 hour to produce torrefied miscanthus. The biomass samples were blended with high volatile (> 30 wt% daf) caking coals in the range of 5-15 wt%. It was found that the miscanthus constituent that develops more fluidity is lignin. Indeed, lignin developed 100% fluid material, followed by xylan (60%) and cellulose (< 40%). However, the miscanthus and the lignin produce chars at > 420°C that interact and destroy the fluid components in the coal. On the other hand, the torrefied miscanthus reduced the fluidity in the blend to a lesser extent than the pristine miscanthus. The addition of 5 wt% torrefied miscanthus did not cause a significant detrimental effect on the fluidity of the coal and hence this additive could be potentially be used in coking blends.
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1
Oviedo ICCS&T 2011. Extended Abstract
1. Introduction The reduced availability of prime coking coal has led to the blending of poor coking coals with carbonaceous materials to produce good coking coals. In this manner, the use of additives to partially substitute coal in coke-making and to improve the properties of the coal blends has been widely investigated by various authors [1-5]. However, there is little research related to the application of biomass as an additive in coking blends. High-temperature rheometry and 1H NMR are two powerful techniques that can monitor fluidity development in biomass and coal during pyrolysis. These techniques have been used in the past to elucidate the effect of different additives in coking blends [5]. The aims of this work are to identify the constituents in biomass that control fluidity development and elucidate the effect of biomass on the fluidity of coking blends.
2. Experimental section The miscanthus was used as a powder of 80-200 μm and the particle sizes for the two coals used was 53-212 μm. Coal A has a volatile matter content of 31.9 wt% daf, ash content of 9.8 wt% db and Gieseler maximum fluidity of 534 ddpm. Coal B has a volatile matter content of 35.7 wt% daf, ash content of 7.2 wt% db and Gieseler maximum fluidity of 8018 ddpm. The lignin of the miscanthus was separated using the Organosolv process. A description of this process can be found elsewhere [6]. A Doty 200 MHz 1H NMR probe was used in conjunction with a Bruker MSL300 instrument to determine fluidity development in the biomass samples. A flow of 25 dm3/min dry nitrogen was used to transfer heat to the samples and to remove the volatiles that escaped from the container. Below the sample region, a flow of 60 dm3/min of dry air prevented the temperature rising above 50°C to protect the electrical components. In addition, air was blown at 20 dm3/min into the region between the top bell Dewar enclosing the sample region and the outer side of the probe to prevent the temperature from exceeding 110°C. The sample temperature was monitored using a thermocouple in direct contact with the sample container.
The solid echo pulse
sequence (90o-τ-90o) was used to acquire the data. A pulse length of 3.50 μs was maintained throughout the test. Approximately 50 mg of sample was packed lightly into a boron nitride container, and 100 scans were accumulated using a recycle delay of 0.3 seconds. The samples were heated from room temperature to the final temperature at approximately 3°C/min. Duplicate analyses were carried out to ensure reproducibility in
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2
Oviedo ICCS&T 2011. Extended Abstract
the results. The spectra obtained were deconvoluted into Gaussian and Lorentzian distribution functions, which enabled the calculation of the fraction of total hydrogen that is mobile and its mobility (T2L). The higher the concentration and mobility of the fluid phase the higher the fluidity, and thus, fluidity depends on both concentration and mobility of the fluid or mobile phase. Rheological measurements were performed in a Rheometrics RDA-III high-torque controlled-strain rheometer. Coal, biomass and coal/biomass mixtures (1.5 g) were compacted under 5 tonnes of pressure in a 25 mm die to form disks with thickness of approximately 2.6 mm. The tests involved placing the sample disk between two 25 mm parallel plates which had serrated surfaces to reduce slippage. The coal and coal blends were heated quickly to 330°C and heated to 520°C at a rate of 3°C/min. The furnace surrounding the sample was purged with a constant flow of nitrogen to transfer heat to the sample and remove volatiles.
The sample temperature was monitored using a
thermocouple inside the furnace.
A continuous sinusoidally varying strain with
amplitude of 0.1 % and frequency of 1 Hz (6.28 rad/s) was applied to the sample from the bottom plate throughout the heating period. The stress response on the top plate was measured to obtain the complex viscosity (η*) as a function of temperature.
3. Results and Discussion Figure 1 shows the high-temperature 1H NMR results for the miscanthus and its constituents. Lignin develops 100% fluid material at 200°C and fluidity remains high up to 350°C. Xylan develops up to 60% fluid material at 275°C. The fluidity decreases for miscanthus (< 45% fluid material) and cellulose develops fluidity below 40% at high temperatures (325°C). The mobility of the fluid material in the samples, as indicated by T2L, show different trends. Xylan possesses the highest mobility although its maximum occurs at low temperatures (175°C) and this could be an artefact of the low NMR signal. The mobility of the fluid phase in lignin remains fairly constant between 225°C and 350°C, similarly to the trend for the concentration of fluid phase. The trends for miscanthus and cellulose show similar mobility values with temperature. These results suggest that lignin could be responsible for most of the fluid material developing in miscanthus. However, lignin is much less thermally stable than the other constituents and its mobility development originates from bond cleavages after glass transition.
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3
Oviedo ICCS&T 2011. Extended Abstract
Figure 1. Percentage of fluid phase (a) and mobility of the fluid phase (b) as a function of temperature for the miscanthus and its constituents.
High-temperature rheometry results for the miscanthus and the lignin are presented in Figure 2. These results are in agreement with the 1H NMR results, showing that the low the lignin is highly fluid above 150°C whereas the miscanthus has a minimum viscosity (or maximum fluidity) at 350°C.
The minimum in viscosity at 100°C could be
attributable to the presence of moisture in the sample. Figure 3 shows that lignin destroys the fluidity of coal when it is a added in concentrations of 5 wt%. This result suggests that the fluid material evolving from the lignin does not contribute to the fluid properties of the blend within experimental error, and that the char originating from the lignin above 420°C interacts with the coal and completely destroys its fluid phase. Similar results were found with other additives that were chars at the thermoplastic temperature of the coal [5]. The effect of miscanthus on coal fluidity (not shown) was also detrimental but to a lesser extent than that of lignin.
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4
Oviedo ICCS&T 2011. Extended Abstract
6
10
5
10
η* ( ) [Pa-s]
Miscanthus Lignin 4
10
103
2
10
0.0
100.0
200.0
300.0
400.0
500.0
Temp [°C]
Figure 2. Complex viscosity as a function of temperature for the miscanthus and its lignin constituent.
7
10
Coal Coal + lignin (5 %) 6
η* ( ) [Pa-s]
10
5
10
104
3
10 300.0
350.0
400.0
450.0
500.0
550.0
Temp [°C]
Figure 3. Complex viscosity as a function of temperature for coal A and the mixture of coal A and lignin (5 wt%).
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5
Oviedo ICCS&T 2011. Extended Abstract
The miscanthus was also torrefied at 250°C under nitrogen for 1 hour. The torrefied miscanthus was analysed in the rheometer and the results show that there is a small reduction in fluidity in the pre-treated sample as indicated by the smaller peak (Figure 4).
6
η* ( ) [Pa-s]
10
5
10
Miscanthus Torrefied miscanthus
4
10
0.0
100.0
200.0
300.0
400.0
500.0
Temp [°C]
Figure 4. Complex viscosity as a function of temperature for miscanthus and torrefied miscanthus.
The torrefied miscanthus was blended with coal B in the range of 5-15 wt% to elucidate the effect in the fluidity of the coal. The results are presented in Figure 5 and show a gradual reduction in fluidity as the concentration of biomass in the blend increases. This reduction seems to be directly related to the low fluidity of the torrefied miscanthus, especially after char formation, rather than any interaction with the coal. Moreover, the addition of 5 wt% torrefied miscanthus does not seem to cause a significant deleterious effect on the fluid properties of the coal and this additive may potentially be used in coking blends without affecting coke quality.
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6
Oviedo ICCS&T 2011. Extended Abstract
6
10
5
η* ( ) [Pa-s]
10
4
10
103
2
10 375.0
Coal Coal + torrefied miscanthus (5 %) Coal + torrefied miscanthus (10 %) Coal + torrefied miscanthus (15 %) 400.0
425.0
450.0
475.0
500.0
Temp [°C]
Figure 5. Complex viscosity as a function of temperature for coal B and mixtures of coal B and torrefied miscanthus (5-15 wt%).
4. Conclusions This work has shown that lignin develops high concentration of fluid material during pyrolysis. However, the fluid material in lignin and miscanthus evolves at temperatures lower than the thermoplastic temperature of coal, and the char forming in the biomass above 420°C interacts with the coal constituents causing a deleterious effect on fluidity development. The torrefaction of miscanthus removes the moisture and also reduces some fluid material, but the torrefied biomass does not interact to a great extent with the coal in blends containing 5 wt% biomass. Therefore, there is scope to use this pretreated material as additive in coking blends without altering the properties of the coking blend.
Acknowledgement The authors would like to thank Dr Paul Pernot from Centre de Pyrolyse de Marienau (CPM) in France and Drazen Gajic from DMT GmbH in Germany for supplying the coals.
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7
Oviedo ICCS&T 2011. Extended Abstract
References [1] Sakurovs R. Some factors controlling the thermoplastic behaviour of coals. Fuel 2000;79:379–389. [2] Nomura S, Kato K, Nakagawa T, Komaki I. The effect of plastic addition on coal caking properties during carbonization. Fuel 2003;82:1775–1782. [3] Sakurovs R. Interactions between coking coals and plastics during co-pyrolysis. Fuel 2003;82:1911–1916. [4] Uzumkesici ES, Casal-Barciella MD, McRae C, Snape CE, Taylor D. Co-processing of single plastic wastestreams in low temperature carbonisation. Fuel 1999;78:1697–1702. [5] Castro Diaz M, Steel KM, Drage TC, Patrick JW, Snape CE. Determination of the effect of different additives in coking blends using a combination of in situ high-temperature 1H NMR and rheometry. Energy & Fuels 2005;19:2423–2431. [6] Brosse N, Sannigrahi P, Ragauskas A. Pretreatment of Miscanthus x giganteus Using the Ethanol Organosolv Process for Ethanol Production. Ind Eng Chem Res 2009;48:8328–8334.
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8
Oviedo ICCS&T 2011. Extended Abstract
Influence of alkali additives on the swelling behavior of a high swelling bituminous coal 1
C.A.Strydom1, J.R. Bunt1,2 , Y. van Staden1 and J Collins1 Chemical Resource Beneficiation, North - West University, Potchefstroom Campus, South Africa 2 Sasol Technology (PTY) Ltd, Box 1, Sasolburg, 1947, South Africa Corresponding author:
[email protected]
Abstract The effect of the addition of some alkali compounds to a South African coal with a high swelling propensity was investigated. A vitrinite-rich bituminous coal from the Tshikondeni coal mine in the Limpopo province of South Africa was used in this study. CaCO3, NaCO3, K2CO3 and KHCO3 were added to the coal in mass percentages of 5%, 10%, 15%, 20%, 30% and 40%. The free swelling index (FSI), FTIR, Gieseler fluidity and dilatometry of the samples were measured. The FSI results showed that all additives significantly reduced the swelling properties and the order of effectiveness of the additives was found to be: KHCO3 >K2CO3>Na2CO3>CaCO3. A 40% addition of KHCO3 reduced the swelling index by a factor of three. The Gieseler fluidity results showed a narrowing in the plastic range with the addition of KHCO3. The dilatometry results
showed that the dilation reduced with addition of KHCO3 and the plastic
properties of the coal samples were changed from euplastic to subplastic.. It is concluded that the addition of alkali carbonates and bicarbonates to coal reduce the free swelling of coal, with alkali bicarbonates having the most impact on reduction of swelling.
1. Introduction Oxidation of coals exhibiting high swelling properties has been reported to influence the physical and chemical properties to various extends [1,2]. The deterioration of caking and coking abilities of coal after oxidation is well documented and ascribed to a loss of plastic properties of the coal [1-5]. Volatile bituminous coals have been known to undergo swelling and have been used as coking coals throughout the world [6]. The formation of cross-links within the macromolecular chemical structure of the coal at temperatures below that required for pyrolysis may reduce or eliminate swelling [7].
Oviedo ICCS&T 2011. Extended Abstract
Bexley and co-workers [8] claimed that alkali carbonates and bicarbonates decrease or completely destroy the dilatation of the coal, possibly through reaction with the phenolate and carboxylate groups of the studied Illinois No. 6 coal. The alkali compounds seem to act as catalysts, catalyzing the formation of ether cross-links at temperatures of approximately 573 K [8].
The Grootegeluk mine in the Waterberg area in South Africa produces 17 million tonnes per year of a high swelling coal. The mine has a resource of 75-billion tonnes of coal [9] and is the largest coal reserve in South Africa. Tshikondeni colliery is situated in the Limpopo province and has coal reserves of six-million tonnes and a resource of 32million tonnes of coal. 400 000 tonnes per year of premium hard coking coal is produced at this mine [10]. The paper described the study of the influence of CaCO3, Na2CO3, K2CO3 and KHCO3 additions on the swelling properties of the Tshikondeni coal during thermal treatment in air.
2. Experimental section A coal sample from the Tshikondeni colliery in South Africa was crushed and ground to sizes of <75 µm. Ultimate, proximate and petrographic analyses were performed according to standard methods on the coal and results are summarized in tables 1 – 3.
Table 1. Ultimate analysis
Carbon (%) Hydrogen (%) 90.4
4.9
Dry-ash-free (daf) basis Total Sulphur Nitrogen (%) (%) 1.8 1.1
Oxygen (%) 1.8
Table 2: Proximate analysis Air Dry basis Moisture (%)
Ash (%)
Volatile Matter (%)
Fixed Carbon (%)
0.9
12.3
23.9
62.9
5%, 10%, 15%, 20%, 30% and 40% (mass percentages) CaCO3, Na2CO3, K2CO3 and KHCO3 (analytically pure) were added to the coal. The resultant mixtures were subsampled to ensure constant fixed carbon content (no dilution effect). The samples were
Oviedo ICCS&T 2011. Extended Abstract
then further ground and sieved to a particle size of less than 250 µm.
Table 2. Petrographic analysis Maceral composition (percentage by volume) (mineral matter-free basis). Mean Reactive Inert random Liptinite Vitrinite (%) inertinite inertinite vitrinite Rank (%) (%) (%) reflectance (%) Medium 92 2 2 4 1.29 Rank B
The Free Swelling Index values (FSI) were determined according to the SABS standard method (ISO 501:2003). The procedure entails the heating of 1 g of sample in a silica crucible to a temperature of 820 ± 5°C within 21/2 minutes. The resultant coke button is then compared to a standard series of number profiles to determine the swelling index. The FSI is determined by comparing the size of the resulting button with a series of satndards and assigning a value between 1 and 9 to the button form. The value of 0 corresponds to the button falling apart and 10 to the button showing no deformation.
The additive resulting in the largest decrease in FSI values was determined and 10% and 40% (mass percentages) additions of this additive to the coal samples were further analysed using FTIR spectroscopy, Gieseler Fluidity and dilatometry. FTIR spectra were obtained on a Vertex 70 FTIR spectrometer in the MIR region. The background was run for 300 scans, and the sample for 400 scans with a resolution of 4 cm-1. KBr (200 mg) pellets containing the different samples (20 mg) were prepared.
The fluidity was measured using a Gieseler plastometer. A strirrer operation with a constant torque is inserted into the coal simple and the speed (revolutions) of the stirrer measured while heating the coal at 3 °C min-1. The relative values of dial divisions per minute at the various temperaturas are plotted to obtain a relative fluidity curve, from which the softening temperatura (Ts), temperature of máximum fluidity (Tm) and resolidification temperatura (Tr) are determined. The plastic range of the sample is given by Ts to Tr [11].
Oviedo ICCS&T 2011. Extended Abstract
Dilatometry identify four classes of plastic behaviour, i.e. subplastic, euplastic, oerplastic and fluidoplastic [12]. The change in volume of coal sample during heating is recorded using a dilatometer. The three parameters that characterize the volume change are: contraction, dilation and swelling [12].
3. Results and Discussion
Ultimate, proximate and petrographic analyses indicated that the coal is a vitrinite-rich (92%), medium ash (12.3%), medium rank B bituminous coal. The effect of the various additives on the free swelling index (FSI) of the coal is given in Figure 1. The results show that KHCO3 had the largest reducing effect on the swelling propensity of the coal. The raw coal (no additive) had a swelling number of eight, and with 40% KHCO3 addition to the coal, the swelling number was reduced by a factor of three. The order of decreasing effect on the swelling index of the additives was found to be: KHCO3>K2CO3>Na2CO3>CaCO3. Bexley et al. [8] found CaCO3 to have only a small effect on the swelling properties of the Illinois no. 6 bituminous coal they studied, with the effect of Na2CO3, K2CO3 and KHCO3 additions much larger. Their results thus followed the same overall trend as observed in this study. Further analyses were performed on the coal samples with KHCO3 additions of 10% and 40%. 9
CaCO3 8
Na2CO3 7
K2CO3
6
KHCO3
FSI
5
4
3
2
1
0 0
5
10
15
20
25
30
35
40
45
Additive Concentration [mass (%)]
Figure 1: FSI values of coal with additions of CaCO3, Na2CO3, K2CO3 and KHCO3.
Oviedo ICCS&T 2011. Extended Abstract
The Gieseler fluidity results for the coal samples with 10% and 40% additions of KHCO3 are given in figure 2.
4000
Raw Coal 3500
10% KHCO3 3000
Fluidity (dd/min)
40% KHCO3 2500
2000
1500
1000
500
0 399
409
419
429
439
449
459
469
479
489
499
509
Temperature ( C)
Figure 2: Gieseler fluidity results of the raw coal, and coal with 10% and 40% KHCO3. Gieseler fluidity results showed that the softening temperature of the coal sample and coal with 10% KHCO3 was 402°C and for the 40% KHCO3 coal sample 420°C. The resolidification temperatures of the raw coal, 10% and 40% KHCO3 additive was 503°C, 500°C and 491°C respectively. The increase in softening temperature and the decrease in re-solidification temperatures, results in a narrowing of the plastic range upon addition of KHCO3. Figure 3 represents the dilatation versus temperature results of the coal and coal with 10% and 40% KHCO3 additions. The coal behaves as a euplastic compound, but addition of KHCO3 changes the samples to behave as subplastic compounds, thus showing a decrease in volume.
Oviedo ICCS&T 2011. Extended Abstract
200
Raw Coal 10% KHCO3
Dilatation (%)
150
40% KHCO3
100
50
0
-50 330
350
370
390
410
430
450
470
490
Temperature ( C)
Figure 3: Dilatometry graphs of coal sample and coal with 10% and 40% additions of KHCO3. FTIR spectra for the coal and coal with 10% and 40% additions of KHCO3 are given in figure 4. The carboxylic C=O peak at 1605 cm-1 decreased substantially with the addition of KHCO3. Bexley et al. [8] suggested that a chemical reaction occurs between the additive and the carboxylate and phenolate groups of the coal. The intensity of the phenolic O-H stretching vibration at 3434 cm-1 decreased upon addition of KHCO3. This suggests that the additive may have reacted with the phenolic groups of the coal. The strong band at 1033 cm-1 is possibly a C-O phenol stretching vibration which decreased with addition of KHCO3. The carboxylic C=O peak at 1605 cm-1 also decreased with the addition of KHCO3. The peaks at 2924 and 1445 cm-1 (possibly aliphatic C-H groups) decreased with 10% KHCO3 addition and almost disappeared with 40% KHCO3 addition. This may be due to a reaction taking place between the additives and the coal, thus changing the structure of the coal. The same can be said for aromatic C-H bands at 3045, 880, 804 and 750 cm-1, which also showed a large decrease in peak intensity.
Oviedo ICCS&T 2011. Extended Abstract
1.6
Raw coal 1.4
10% KHCO3 40% KHCO3
Absorbance
1.2
1
0.8
0.6
0.4
0.2
0 0
500
1000
1500
2000
2500
3000
3500
4000
4500
-0.2
Wavenumber
Figure 4: FTIR spectra of raw coal and coal with 10% and 40% additions of KHCO3.
4. Conclusions
The order of largest decreasing effect on the swelling index of the additives was found to be: KHCO3>K2CO3>Na2CO3>CaCO3. The addition of KHCO3 changes the plastic properties of the sample from a euplastic to a subplastic coal, thus changing the coal sample into a coal for which no swelling occurs upon heating in air. The plastic range of the coal with KHCO3 additions was also decreased. According to literature alkali carbonates and bicarbonates may react with phenolate and carboxylate surface functional groups of coal, which then affects the swelling of coal through cross linking leading to the stabilization of the structure. A decrease in the number of these functional groups were observed on the FTIR spectra.
Acknowledgement. The authors would like to thank North-West University and SASOL Technology, Research and Development for financial support for the investigation.
Oviedo ICCS&T 2011. Extended Abstract
References [1] Iglesias MJ, De la Puente G, Fuente E and Pis JJ. Compositional and structural changes during aerial oxidation of coal and their relations with technological properties. Vib. Spectr. 1998:17:41-52. [2] Song Ch, Saini AK and Schobert HH. Effects of Drying and Oxidation of Wyodak Subbituminous Coal on Its Thermal and Catalytic Liquefaction.Spectroscopic Characterization and Products Distribution. Energy Fuels. 1994:8:301-312. [3] Larsen JW, Lee D, Schmidt T and Grint A. Multiple mechanisms for the loss of coking properties caused by mild air oxidation. Fuel. 1986:65:595-596. [4] Pis JJ, Cagigas A, Simon P, Lorenzana JJ. Effect of aerial oxidation of coking coals on the technological properties of the resulting cokes. Proc. Technol. 1988:20:307-316. [5] Khan MR and Jenkins RG, Influence of K and Ca additives in combination on swelling, plastic and devolatilization properties of coal at elevated pressure. Fuel. 1989:68:1336-1339. [6] Yu D, Xu M, Yu Y and Lui X. Swelling behaviour of a Chinese bituminous coal at different pyrolysis temperatures. Energy Fuels. 2005:19:2488-2494. [7] McCormick RL and Jha MC. Effect of catalyst impregnation conditions and coal cleaning on caking and gasification of Illinois No. 6 coal. Energy Fuels. 1995:9:10431050. [8] Bexley K, Green PD and Thomas KM. Interaction of mineral and inorganic compounds with coal. Fuel. 1986:65:47-53. [9] Njobeni, S. 2010. Looking for power that won't run out. Times Live, 21 Feb. http://www.timeslive.co.za/business/article317763.ece. [10] Creamer, M. 2009. Water, rail plans in place in Waterberg. Mining Weekly, 11 Dec. http://www.miningweekly.com/print-version/water-and-rail-plans-in-place-forgrowing-waterberg. [11] Speight JG. Handbook of Coal Analysis. Hoboken, New Jersey: John Wiley & Sons. 2005: p. 142-145. [12] Schobert HH. The Chemistry of Hydrocarbon Fuels. Londen: Butterworth Publishers. 1990.
Oviedo ICCS&T 2011. Extended Abstract
Estimation of the coking pressure in coke ovens by Koppers-INCAR test
R. Alvarez, C.Barriocanal, M.A. Díez Instituto Nacional del Carbón, CSIC, Apartado 73, 33080 Oviedo. Spain email:
[email protected] Abstract A laboratory test : Koppers-INCAR, designed and patented by INCAR has been used to control the coking pressure developed by coking coals during carbonization. Laboratory results were backed by tests carried out to semi-industrial and industrial scale. These results have been compared with semi-pilot and pilot scale tests. 1. Introduction There are several causes of “stickers” and “heavy pushes” at the end of the coking process [1]. The consequence of these problems can be costly, including reduced coke production, damage to machinery and oven walls and extra man power. However, coal or blend characteristics are the most important parameters determining coal expansion and coke contraction behaviour during carbonisation and this is the major source of coke pushing difficulties. It is well known that some coals are liable to damage the coke oven walls because of excessive pressure developed during carbonisation (coking pressure) or insufficient coke contraction at the end of the process. It has been suggested [2] that the existence of little or no contraction is a sign of high pressure during carbonisation and it seems probably that coals giving cokes with not enough contraction will develop dangerous pressures of sufficient magnitude to damage coke oven walls during carbonisation. The problem of coking pressure generation has been widely studied [3-5] and although it is not fully understood is generally accepted its relation with the development of internal gas pressure [3,6-8]. Many test methods have been used for the evaluation of coals with regard to their expansion and contraction behaviour in the coking process [9]. The use of the movable wall oven has been the most widely accepted and has also been fundamental to try to explain the mechanism of coking pressure generation [10]. The Spanish National Coal Institute (INCAR) has been largely involved in the study of the mechanism of coking pressure generation and the development and utilization of
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Oviedo ICCS&T 2011. Extended Abstract
tests to measure expansion and contraction during carbonisation. INCAR has developed and patented [11] the Koppers INCAR test, based on the early Koppers laboratory test and taking into account the Mott and Spooner modifications [12]. This Koppers test has been substantially modified by INCAR [13], contraction being the dominant parameter in determining the degree of danger associated with a coal. Laboratory results have been backed by tests carried out to a semi-industrial scale [14]. 2. Experimental Two movable wall ovens are available at INCAR at this moment : MWO 250( Figure 1) and MWO 15; the first one has a capacity of 250-300 kg of coal and the second one 15 kg. The MWO250 is a electrically heated gravity-charged oven with a length of 915 mm length, 840 mm height and 455 mm width. The oven works with a coking time of 18 h. the initial temperature is 880 °C and the final 1132 °C with a heating rate of 14 °C/h. During the test, wall pressure, temperature in the centre of the charge and in the wall, together with vertical contraction can be measured. The data are continuously recorded and displayed in a computer. The MWO15 is a gravity charging, electrically heated oven with a length of 250mm, a height of 800 mm and a width of 150 mm. The oven has a coking time of 2h45min, the temperature of the wall is maintained at 1010 °C and the final coke temperature is 950 °C. During the test, wall pressure, temperature in the centre of the charge and in the wall together with vertical contraction can be measured. The data are continuously recorded and displayed in a computer. Figure 2 shows a diagram of the equipment used to carry out the Sole heated oven (ASTM D 2014-85). A coal mass between 4,25 and 5,30 kg, crushed to no less than 70% and no more than 85% below 3,35 mm, with 1% moisture was enclosed in the carbonisation chamber at a bulk density of 881 kg/m3. A constant force of 15,2 kPa was applied on the top of the charge. Carbonisation temperature rose from 550ºC on the sole to 950ºC. The test was considered finished when the temperature of the top of the charge reached 500ºC. Figure 3 shows the Koppers-INCAR apparatus. Although the geometry of the most important parts of the early Koppers test was maintained, the heating system and some of the test constants (pressure on the charge, rate and duration of heating) were modified taking into account Mott and Spooner modifications and INCAR research. 3. Results and Discussion 3.1. Koppers-INCAR results Figure 4 shows typical curves obtained in this test for three different coals. The Submit before 31 May 2011 to
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Oviedo ICCS&T 2011. Extended Abstract
interpretation of the graph obtained may be summarised as: “coals giving a contraction greater than 10 mm are not dangerous during carbonisation”. Coals giving curves of type I are classified as dangerous and coals giving curves of type III as safes. Coals giving curves of type II represent the limit between dangerous and not dangerous coals. To establish this classification, laboratory results were backed by tests carried out to a semi-industrial scale [15] and to industrial scale [16]. 3.2. Semi-industrial scale tests This laboratory test was used to resolve the problem of coking a dangerous Spanish coal: Sabero (20.5%VM). The indications of this test were used to determine the minimum amounts of three different high volatile coals: Eskar, Carrocera and Hullasa (35.5, 37.3 and 41.1 VM respectively) that it was necessary to add, to reduce the risk of coking this dangerous coal. Laboratory results were backed by tests carried out on a semi-industrial scale, without losing sight of certain safety limits [14]. Using the laboratory data from the Koppers-INCAR tests (Figure 5) it was established that Sabero coal could be coked without risk by adding the following minimum amounts of the corrective coals : 10% Hullasa or 20% Carrocera or 30% Eskar. 3.3. Industrial scale tests During 1990 the Spanish Steel Industry, Ensidesa, was using blends and the KoppersINCAR curves from June to December are shown in Figure 6. At the beginning of 1991 due to problems with the coal received, a dangerous coal blend was coked. Stickers and heavy pushes were produced in the Veriña battery but not in the Avilés battery. This can be explained because the bulk density is 780 to 790 kg/m3 in Veriña (6.5 m high ovens) and 710 m to 720 m in Avilés battery (4.5 high ovens). The Koppers-INCAR curves of January 2 and January 3 of Figure 6 show the reliability of the Koppers-INCAR test. The Koppers- INCAR test has been used in several INCAR research works [13-22] and ECSC Projects [23-29] and also it has started in 2010 year a new one. The relationship among the Koppers-INCAR test and the other tests to measure the coking pressure available at INCAR are shown in Figures 7 (MWO 250-300 kg) [24], 8 (ASTM SHO) [24] and 9 (MWO 15-17 kg) [28]. 4. Conclusions The Koppers-INCAR test is a reliable method to predict the expansion and contraction behaviour of coals and blends during carbonization (coking pressure) and is also a rapid a cheap method compared with the used more frequently for coking pressure control at the coke oven plants. Since 2009 is being used for the control in advance of the blends Submit before 31 May 2011 to
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Oviedo ICCS&T 2011. Extended Abstract
prepared by the Spanish Steel Industry. References [1]
Kevin F. De Vanney. Coke plant pushing problems-Causes and troubleshooting guidelines. 2002 Ironmaking Conference Proceedings, 339-346
[2]
B.C.R.A. Technical Paper. December 1948
[3]
Loison R, Foch P, Boyer A. Coke Quality and Production. Butterworth, London, 1989 (p. 353-412).
[4]
Tucker J, Everitt G. 2nd International Cokemaking Congress, London, 1992 p. 40-49
[5]
Coke oven wall pressures. Measurement, Cause and Effect. Publication of the Iron and Steel Society,1990,ISBN 0-932897-525
[6]
Jordan P, Patrick JW, Walker A. A laboratory study of internal gas pressure generated during the coking of coals. Cokemaking Int.1992;4:12-15.
[7]
Te Lindert M, Van der Velden B. Research into internal gas pressure and shrinkage. Ironmaking Conference Proceedings,1994; 53:12-15.
[8]
Barriocanal C, Patrick JW, Walker A. The laboratory identification of dangerous coking coals. Fuel 1998; 77:881-884.
[9]
Meltzhein C, Buisine M. Etude de la poussée sur les parois des fours a coke. Rev.Gen.Therm. 1968;47: 147-165.
[10] Loison R, Foch P, Boyer A. Coke Quality and Production. Butterworth, London, 1989 p..336 [11] Procedimiento y sistema para evaluar el empuje de los carbones o mezclas coquizables. Patente de invención nº 524,258,1983 [12] Mott RC, Spooner C. The assessment of coals liable to damage oven walls Fuel 1939; 18:322-344. [13] Alvarez R, Pis JJ, Barriocanal C, Lázaro M. et al. Characterization of dangerous coals during carbonization. Effects of air oxidation and ash content of coals. Cokemaking Int.1991; 3:37-42. [14] Alvarez R, Miyar EA, Escudero JB. Application of a laboratory test to resolve the problem of coking a dangerous coal. Fuel 1990; 69:1151-1156. [15] J.B.Escudero, R.Alvarez. Influence of air oxidation on the pressure exerted by coking coals during carbonization. Fuel 1981; 60:251-253. [16] Alvarez R, Pis JJ, Barriocanal C, Sirgado M. Practical application of a laboratory test to measure expansion and contraction during carbonization. Cokemaking Int. 1992; 4:16-18. [17] Alvarez R,. Pis JJ, Lorenzana JJ. Characterization of dangerous coals during carbonization. Fuel Processing Technology 1990; 24:91-97. [18] Alvarez R, Pis JJ, Díez MA, Marzec A, Czajkowska S Studies on generation of Submit before 31 May 2011 to
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Oviedo ICCS&T 2011. Extended Abstract excessive coking pressure.1.Semicoke contraction versus thermoplastic properties. Energy and Fuels 1997; 11:978-991. [19] Marzec A, Czajkowska S, Alvarez R, Pis JJ, Díez MA, Schulten H. Studies on generation of excessive coking pressure.2.Field Ionization Mass Spectrometry of coals showing different contraction during Carbonization. Energy and Fuels 1997; 11:982-986. [20] Casal MD, Canga CS, Díez MA, Alvarez R, Barriocanal C. Low temperature pyrolisis of coals with different coking pressure characteristics. J Anal.Applied Pyrolisis 2005; 74:96-103. [21] Casal MD, Diaz-Faes E, Díez MA, Alvarez R, Barriocanal C.. Influence of porosity and fissuring on coking pressure generation Fuel 2008;87:2437-2443. [22] Barriocanal C, Díez MA, Alvarez R, Casal MD. Relationship between coking pressure generated by coal blends and the composition of their primary tars.J.of Anal.Applied Pyrolisis 2009; 85:14-520. [23] R.Alvarez. Coal Weathering. ECSC 7220-EB/755, 1994. [24] R.Alvarez.Caracterización de carbones con empuje peligroso. ECSC 72220-EB/756 1995. [25] R.Alvarez. Low cost of coke by increasing low volatile coals in blends ECSC 7220EB/344, 1997. [26] R.Alvarez. Coking pressure studies. 7220-EB/345, 2000. [27] R.Alvarez, C.Barriocanal. Study of parameters involved in coking pressure generation. ECSC 7220-PR/069, 2002. [28] R.Alvarez, C.Barriocanal. Laboratory and pilot scale tests to assess coke quality and coking pressure. ECSC 7220-PR/119, 2004. [29] R.Alvarez, C.Barriocanal. Coking pressure generation and moderation ECSC 7220PR/140, 2005. [30] C.Barriocanal. Generation of Swelling Pressure in a Coke Oven, Transmission on oven walls and consequences on wall degradation (SPRITCO) 2010.
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Oviedo ICCS&T 2011. Extended Abstract
Figure 1. MWO 250.
Figure 2. Sole Heated Oven (SHO).
Figure 3. Koppers-INCAR test.
Figure 4. Typical curves obtained in the Koppers-INCAR test. Submit before 31 May 2011 to
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Oviedo ICCS&T 2011. Extended Abstract
Figure 5. Koppers-INCAR curves of coal blends: a, Sabero-Eskar; b, Sabero-Carrocera; c, Sabero-Hullasa.
Figure 6. Koppers-INCAR curves of ENSIDESA blends. Submit before 31 May 2011 to
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Oviedo ICCS&T 2011. Extended Abstract
Figure 7. Relationship between KI and coking Figure 8. Relationship between KI and pressure in the MWO250.
S.H.O. contraction.
2
Wall pressure MWO15(kN/m )
90 80 70 60 50 40 30 20 10 0 -30
-20 -10 0 Contraction/expansion Koppers-INCAR (mm)
10
Figure 9. Relationship between KI and coking pressure in the MWO15.
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Oviedo ICCS&T 2011. Extended Abstract
High Performance Electric Double-Layer Capacitor using Activated Carbon from HyperCoal
K. Sato, K. Magarisawa, T. Takarada Department of Chemical & Environmental Engineering, Graduate School of Engineering, Gunma University, 1-5-1 Tenjin-cho, Kiryu, Gunma, 376-8515 Japan Contact address:
[email protected] A high performance electric double-layer capacitor (EDLC) was fabricated using activated carbon derived from ash-less HyperCoal. The activation was carried out at 500 - 800 ºC for 0.5- 4 h under flowing argon atmosphere using HyperCoal derived char. NaOH and KOH were used as activation agent. Specific surface area (SSA) of activated carbon and specific capacity of resultant EDLC are strongly depend on the activation conditions including variety and amount of activation agent, activation temperature and holding time. The largest SSA of 2560 and 3110 m2·g-1 was achieved for activated carbon when it was activated at 700 ºC for 3 h using KOH and NaOH, respectively. The resultant specific capacitance of EDLC fabricated from these activated carbons was 43.9 and 44.1 F·g-1 which are comparable or higher than EDLC fabricated from conventional phenol resin derived activated carbon.
1. Introduction Electric double layer capacitor (EDLC) is a promising energy storage device in the future because of their safety, environmentally benign and fast charge-discharge properties compared to those of other devices such as secondary ion batteries. [1, 2] Higher energy density and lower cost must be achieved for wide spreading of EDLCs in a number of applications. The capacity of EDLC can be described as; C=∫(ε/L)dS
(1)
where ε is the specific capacitance of the electrolyte within the electric double layer, L the thickness of the electric double layer, and dS the surface area of the electrodes available for the charge- discharge process, respectively. Activated carbons (ACs) are widely used for EDLC electrodes since it has very high specific surface area (SSA) and relatively high electrical conductivity.
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Oviedo ICCS&T 2011. Extended Abstract
The fabrication of ACs from the abundant natural resources such as coals and biomass has the advantage in material cost. However ash component involved in these resources may reduce the performance of EDLC, and inhibit the reuse of activation agents resulting in higher cost. A solvent extraction and condensation process can significantly reduce the ash content in coals. [3] The materials fabricated through the process is called as HyperCoal. The material cost of HyperCoal is expected to be still one order of magnitude lower than the synthetic phenol resin which is widely used to fabricate the ACs for EDLCs. In addition, much of the activation agents could be reused, providing lower total costs.[4] Here we try to fabricates ACs from HyperCoal, and investigate the its applicability for EDLC electrodes.
2．Experimental section A Gregory coal derived HyperCoal supplied from Kobe Steel Ltd. was used as starting material. Proximate and ultimate analysis results is shown in Table 1. The ash content was reduced to be 0.06 wt % in dry-bases through the HyperCoal fabrication. The HyperCoal was carbonized at 600 ºC for 7 min under flowing Ar of 3 L·min-1. The char yield was 69.2 %. The fabricated char was sieved to be < 105 μm and 1 g of the sieved char was spreaded onto alumina boat and then the activation reagent (KOH or NaOH) was put onto the char. The activation was carried out between 500 and 800 ºC for 0.5- 4 h under flowing Ar of 400 ml·min-1. The activated sample was washed by 2 M of hydrochloric acid aqueous solution and then filtered to remove alkaline species. The sample was rinsed by de-ionized water several times. Finally ACs were obtained by drying the washed and rinsed samples in the vacuumed oven at 200 ºC for 2 h. The SSA and pore size distribution are measured using N2 gas adsorption instrument. Table 1 Proximate and ultimate analysis results of the Gregory coal derived HyperCoal
Proxim ate analysis [w t%] (d.b.) M oist. V.M . Ash 0.38 39.45 0.06 U ltim ate analysis [w t%] (d.a.f.) C H N 82.14 6.39 1.59
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F.C .diff. 60.49 S 0.61
O diff. 9.27
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Oviedo ICCS&T 2011. Extended Abstract
A couple of EDLC electrodes was fabricated using the ACs. The ACs of 30 mg was mixed with 3.45 mg of acetylene black and 1.03 mg of Polytetrafluoroethylene using pestle and mortar, and then uni-axially pressed at 40 MPa for 20 min in a metal die. The diameter of the electrode pellets was 13 mm. An aluminum mesh was used as current correctors and attached to the electrodes by dry-pressing. A tetra-ethyl ammonium fluoroborate and propylene carbonate mixture was used as electrolyte. Infiltration of the electrolyte into the porosity of the electrodes was performed under vacuumed condition followed by set-up in a airtight grove box. The two electrode set-up was used for the performance evaluation of the cell. The cell was charged under the constant current density of 40 mA/g-ACs to be 2.5 V of cell voltage, and then discharged under the constant current density of 10 mA/g-ACs to be 0 V. The specific capacitance was determined using the following equation, Cg=I(dT/dV)
(2)
where Cg is specific capacitance, I the current density, dT/dV was gradient of timevoltage diagram between 2 and 1 V in discharge process.
3. Results and discussion Figure 1 shows the effect of holding time for activation at 800 ºC on the SSA of ACs and the Cg of EDLCs. The yield of ACs decreased with increasing holding time and was in the range of 61.0- 50.0 % relative to the char. The SSA increased during the first 1 h and then become constant up to 4 h. The Cg increased up to 3 h and then decreased. The highest Cg was 42.7 F·g-1 when holding time was 3 h. Figure 2 shows the effect of activation agent and temperature on the SSA of ACs and the Cg of EDLCs. The holding time was fixed to be 2 h. The yield of ACs decreased with increasing the temperature, and was in the range of 55.9-84.0 % and 39.1-76.7 % relative to the char for the case of KOH and NaOH as activation agent, respectively. The largest SSA of 2560 and 3110 m2·g-1 was achieved for activated carbon when it was activated at 700 ºC for 3 h using KOH and NaOH, respectively. The reduction of the SSA at 800 º C would be attributed to the pore coalescence by excessively proceeded activation. The Cg well followed the SSA change up to 700 ºC for both NaOH and KOH systems. The highest Cg was observed at 700 ºC and was 43.9 and
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Oviedo ICCS&T 2011. Extended Abstract
43.1 F·g-1 for NaOH and KOH, respectively. These performances were comparable or higher than that fabricated from conventional phenol resin (40.0 F·g-1). This means that HyperCoal is a promising source of ACs for high performance EDLC electrodes. Although the SSA was significantly decreased at 800 ºC, the Cg was not so much decreased. Figure 3 shows pore size distributions of ACs activated using NaOH and KOH at various temperatures. Average pore diameter increased with increasing activation temperature and reached to be about 1 nm at 800 ºC. Thus the only slight change of Cg between 700 and 800 ºC even though the significant SSA reduction can be attributed to the improved pore structure. These results indicated that both the SSA and the pore structure must be controlled for further enhanced performance.
2000 35
1000 0
0.5
1.0 2.0 3.0 Holding time / h
4.0 30
50
NaOH
40
KOH
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30 2000 KOH NaOH
40
Specific surface area / m2·g -1
3000
4000 KOH/C =4 in mass
1000 0
20 10
500 600 700 800 Activation temperature / ºC
Specific capacity / F·g-1
45
KOH/C =4 in mass
Specific capacity / F·g-1
Specific surface area / m2·g -1
4000
0
Figure 1 Effect of holding time on the
Figure 2 Effect of activation agent and
SSA of ACs and the Cg of EDLCs.
temperature on the SSA of ACs and the
The activation was carried out at 800 º
Cg of EDLCs. Holding time for the
C using KOH as the activation agent.
activation was 2 h.
Figure 3 Pore size distributions of ACs fabricated using (a) NaOH and (b) KOH at various temperatures. The holding time for activation was 2 h.
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Oviedo ICCS&T 2011. Extended Abstract
4. Conclusion It is demonstrated that the ash-less HyperCoal is a candidate starting materials of ACs for cost effective high performance EDLC electrodes. Both the increase of SSA and the pore structure must be controlled for higher performance EDLCs.
References [1] Sharma P, Bhatti TS, Energ Convers Manage 2010; 51: 2901-12. [2] Zhao XY, Cao JP, Morishita K, Ozaki J, Takarada T, Energy Fuel, 2010; 24: 188993. [3] Okuyama N, Komatsu N, Shigehisa T, Kaneko T, Tsuruya S. Fuel Process Technol 2004;85:947–67. [4] Sharma A, Takanohashi T, Morishita K, Takarada T, Saito I, Fuel, 2008; 87: 491-97
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Oviedo ICCS&T 2011. Extended Abstract
Degradation Characteristics of SOFC by Trace Elements in Coal Gasified Gas Y. Ueki1, T. Kobayashi2, R. Yoshiie2, and I. Naruse2 1
Energy Science Division, EcoTopia Science Institute, Nagoya University Furo-cho, Chikusa-ku, Nagoya 464-8603, JAPAN 2 Department of Mechanical Science and Engineering, Nagoya University Furo-cho, Chikusa-ku, Nagoya 464-8603, JAPAN
[email protected] Abstract Integrated coal Gasification Combined Cycle (IGCC) and Integrated coal Gasification Fuel Cell combined cycle (IGFC) are recognized as one of highly effective power generation systems for clean coal technologies. However, adverse effects of corrosion in a gas turbine and chemical and/or physical degradation of the fuel cell affect those performances since some trace elements exist in the coal gasified gas. In the present work, influences of trace elements of As and Se in the coal gasified gas on Solid Oxide Fuel Cell (SOFC) installing in the IGFC system were theoretically and experimentally elucidated. First, thermodynamic equilibrium calculations were carried out to estimate chemical compositions of the trace elements on anode of the SOFC. Second, actual reaction behaviors of the trace elements on the anode were tested, using a simulated coal gasified gas, which was supplied into a button-type SOFC. From the thermodynamic equilibrium calculations, The As converted to a solid phase NiAs, and 100 % of Se existed as H2Se gas at low temperature. As the experimental results, Se doping into the simulated gasified gas showed bad effect on power generation slightly, but the power generation performance came back to the initial performance after cutting the Se doping. For As doping, on the other hand, As reacted with Ni in the anode, so that the SOFC performance gradually decreased, and did not recovered after cutting the dope. Key words: Solid oxide fuel cell, Trace element, Coal gasified gas
1. Introduction Integrated coal Gasification Combined Cycle (IGCC) and Integrated coal Gasification Fuel Cell combined cycle (IGFC) have been recognized as one of the highly effective power generation systems for clean coal technologies. The IGFC is
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Oviedo ICCS&T 2011. Extended Abstract
mainly composed of a gas turbine, a fuel cell and a steam turbine. However, adverse effects of corrosion in a gas turbine and chemical and/or physical degradation of the fuel cell affect those performances since some trace elements exist in the coal gasified gas. Therefore, many researchers have studied the effect of trace elements on Solid Oxide Fuel Cell (SOFC). Currently, the effect of trace elements on the SOFC power generation has been studied, using hydrogen or methane as a fuel. However, the effect of trace elements in the coal gasified gas on the SOFC power generation has not been elucidated yet. In the present work, influences of the trace elements of As and Se in the coal gasified gas on the SOFC power generation were theoretically and experimentally studied. First, actual reaction behaviors of the trace elements on the anode were tested, using a simulated gasified gas, which was supplied into a button-type SOFC. Moreover, the anode surface before and after the experiments was observed by using a SEM/EDX. Second, thermodynamic equilibrium calculations were carried out to estimate chemical compositions of trace elements on the anode of SOFC.
2. Experimental 2.1 Sample of the button-type SOFC A photograph and a schematic diagram of the button type cell employed in this power generation experiment are shown in Figs. 1 and 2, respectively. A Ni-YSZ cermet (Ni-Yttrium Stabilized Zirconia composite materials) is used for an anode (Fuel side), and a YSZ (Yttrium Stabilized Zirconia) is used for an electrolyte. The anode diameter is 10mm, and the diameter of the electrolyte is 20mm, and the thickness of the anode and the electrolyte is about 1 mm. To collect the electricity generated, a Pt mesh is installed in both the anode and the cathode. Moreover, the electricity collected there connects with an electrochemistry measurement device through a lead wire and is measured continuously. On the other hand, the lead wire, that is called a reference electrode, is applied to the circumference of the button type cell to put an external load (constant voltage), and this is connected with the electrochemistry measurement device.
2.2 Experimental apparatus Figure 3 shows a schematic diagram of an experimental apparatus used for power generation experiments by the button type cell. The supply lines of gases in this experimental apparatus divides into a fuel gas (simulated coal gasified gas) and a Submit before 31 May 2011 to
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Pt mesh as current collector (Anode)
Fuel gas Lead wire
Anode (Ni-YSZ cermet)
Solid electrolyte (YSZ) Reference electrode (Lead wire)
Pt mesh as current collector (Cathode)
Lead wire
O2 in air
Fig. 1 Photograph of the button type
Fig. 2 Schematic diagram of the button type
cell
cell
gaseous trace element. The anode gas (fuel gas) and the cathode gas (Air) are supplied from the top and bottom, respectively. The trace element is added into the anode gas (fuel gas) as hydride by a hydride generation device shown in Fig. 3 (a). In this device, the solution of As and Se reacts with hydrogen, and the gas hydride is generated. For As, for example, the following chemical reactions occurs in this device. NaBH4 + HCl + 3H2O → H3BO3 + NaCl + 8H
(1)
As + 8H → AsH3 + 5/2H2
(2)
A standard solution of As and Se, a hydrochloric acid solution (HCl) and a sodium boron-hydride solution (NaBH4) are supplied to the device with a pump. Then, the hydride compound of the trace element is generated by the reactions above-mentioned. A humidifier for the anode gas is installed in the fuel gas line. The anode gas is kept at 333 - 353 K with a ribbon heater after a humidification of the anode gas. Air compressed in a gas cylinder is supplied as the cathode gas. The evaluation and examination device for the fuel cell, shown in Fig. 3 (b), is used during the power generation experiment. The supply lines of the anode gas, the cathode gas and an exhaust gas are made of alumina. The temperature of the button type cell is controlled by using an electric furnace. Electricity generated in the button-type cell is collected by a Pt mesh, and is measured by an electrochemistry weighing device, shown in Fig. 3 (c). In this experiment, method of the electric current measurement at the constant voltage is applied for the evaluation of the fuel cell, changing the power load to generate a constant voltage in the fuel cell.
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Oviedo ICCS&T 2011. Extended Abstract
b
a
c
Fig. 3 Schematic diagram of an experimental apparatus
2.3 Experimental procedure and conditions Figure 4 shows the experimental procedure. The button type cell is set up in the fuel cell evaluation and examination device. The temperature rises up to 1173 K in about 1.5 h. After that, N2 and air are introduced in the anode and the cathode, respectively. The anode gas is switched to H2 at 1173 K, and the anode is reduced for about 2 h under the condition of the infinite electric load (insulating state). The reason for this process is to remove the oxide film on the anode surface of the button-type cell. Afterwards, the measurement condition of the electrochemistry weighing device is changed to a constant voltage condition (0.75V). Finally, power generation by H2 begins. After attaining the steady state, the power generation experiments are conducted under each experimental condition. Experimental conditions and gas compositions are shown in Tables 1 and 2, respectively. Compositions in the simulated coal gasified gas, which refers to the gas compositions reported in the NEDO national project in Japan [1, 2]. The addition of trace element into the simulated gas is conducted after the current value attains the Submit before 31 May 2011 to
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Installation of the button type cell Anode: N 2, Cathode: Air Rising temperature to 1173K in 1.5 hour Reduction of the cell surface by H 2
Anode: H 2, Cathode: Air Open-circuit at 1173K in 2 hours Power generation of the cell by H 2 Anode: H 2, Cathode: Air Confirmation of stable current output under constant voltage (0.75V) condition Test for each experimental condition Fig. 4 Experimental procedure of power generation experiments of the button type cell
steady state. The trace element is added continuously for about 10 h. Afterwards, in order to observe the power generation recovery, the power generation experiment under the condition 2 is carried out again. To accelerate the degradation behavior of the cell by the trace element (As, Se), the high concentration of As and Se is added into the simulated coal gasified gas. After the power generation experiment, the anode surface of the cell is analyzed by a SEM/EDX for a deterioration evaluation by those trace elements.
Table 1 Common experimental conditions
Button type cell
Ni-YSZ/YSZ (10mm anode diameter)
Temperature
1173K
Gas flowrate in the anode
100mL/min
Gas flowrate in the cathode
100mL/min
Moisture in the anode
20% (Saturation vapor at 338K)
Pre-heating of anode gas
423K
Operation mode
Constant voltage 0.75V
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Table 2 Gas compositions for each experimental condition Anode gas (dry base) Condition 1 H2 power generation
H2:100%
Condition 2 Coal syn gas power generation
H2:20%, CO:50%, CO2:4%, N2:26%
Condition 3 Coal syn gas power generation with Se doping
H2:20%, CO:50%, CO2:4%, N2:26% H2Se:10ppm
Condition 4 Coal syn gas power generation with As doping
H2:20%, CO:50%, CO2:4%, N2:26% AsH3:10ppm
Cathode gas
Air:100%
3. Experimental result and discussion 3.1 Results of power generation experiments of the button type cell Results of the power generation experiments on condition 1 (H2) and condition 2 (simulated coal gasified gas) is shown Figs. 5 and 6, respectively. In these graphs, the horizontal axis and the vertical axis indicate an exposure period and a current output value at the constant voltage of 0.75V, respectively. The power generation period is about 20 h. From both the figures, the electricity generated is stable under the both conditions. Figures 7 (a) and (b) show the result of the power generation experiment under condition 3 and 4 (Se and As addition), respectively. For the Se addition, the current output value gradually decreases. The current output value of 20 % decreases in about 10 h. After stopping the Se addition, however, the current value almost recovers. For the As addition, on the other hand, the current output also decreases by adding As. The current value of 30 % decreases in about 10 h. However, the current value does not recover to the initial value. These results suggest that the degradation mechanisms for the Se addition will differ from that for the As addition.
3.2 Observation of anode surface of the button-type cell by SEM/EDX SEM images of the anode surface of the button-type cell are shown in Fig. 8. The anode surfaces after (b) H2 reduction, (c) H2 power generation and (d) the simulated coal gasified gas power generation are almost similar structure to that of (a) initial cell. However, the anode surfaces after (e) the Se addition and (f) the As addition change Submit before 31 May 2011 to
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0.30
0.30
0.25
0.25
0.20
0.20
Current [A]
Current [A]
Oviedo ICCS&T 2011. Extended Abstract
0.15 0.10
0.15 0.10
0.05 0.00
0.05 0
5
10
15
0.00
20
0
5
Time [h]
10
15
20
Time [h]
Fig. 5 Current output with time by H2
Fig. 6 Current output with time by the
(Condition 1)
the simulated gas (Condition 2). 0.30
0.30
Condition 2
0.25
Condition 3 (Se doping)
Condition 2
0.25 0.20
Current [A]
Current [A]
0.20 0.15 0.10 0.05 0.00
0.15
Condition 2
0.10
Condition 4 Condition 2 (As doping)
0.05
0
5
10
15
20
0.00
0
Time [h]
(a)
5
10
15
20
Time[h]
(b)
Fig. 7 Current output with time during power generation by the simulated coal gasified gas doped (a) Se (Condition 3) and (b) As (Condition 4)
greatly. For the Se addition, some pores in the anode are buried, compared with that after the H2 power generation and the simulated coal gasified gas power generation. For the As addition, on the other hand, the surface structure of the anode becomes disorder. This may be caused by the reaction of As with the anode material of Ni-YSZ cermet. Figure 9 shows the mapping analysis of Se and As on the anode surface by EDX. As the signal intensity of Se is very low, Se may not deposit on the surface. For As, on the other hand, As is detected on the anode surface. This suggests that an adhesion of As on the anode surface will affect the degradation of the cell performance.
4. Thermodynamic equilibrium calculation 4.1 Calculation condition To elucidate the chemical interaction between the trace elements of As and Se and Submit before 31 May 2011 to
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Oviedo ICCS&T 2011. Extended Abstract
15µm
15µm
(a) Before test
(b) After H2 reduction
15µm
15µm
(c) After H2 power generation (Condition 1)
(d) After coal syn gas power generation (Condition 2)
15µm
(e) After coal syn gas power generation with Se doping (Condition 3)
15µm
(f) After coal syn gas power generation with As doping (Condition 4)
Fig. 8 SEM images of the anode surface of the cell after each experimental condition Se
As
15µm
(a) After coal syn gas power generation with Se doping (Condition 3)
15µm
(b) After coal syn gas power generation with As doping (Condition 4)
Fig. 9 SEM-EDX mapping on the anode surface of the cell after each experimental condition
the anode materials, the thermodynamic equilibrium calculation was carried out, using
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FactSage version 6.1 [3, 4]. The calculation conditions are shown in Table 3, which is almost the same as the power generation condition in this work.
Table 3 Equilibrium calculation conditions for SOFC power generation condition by the simulated coal gasified gas -
Component
Input [mol]
Concentration
Anode material
Ni
2.556x10-4
-
CO
1.34
50%
CO2
0.107
4%
H2
0.536
20%
N2
0.696
26%
H2O
0.536
20%
Cathode gas
O2
0.01876
-
Trace elements
H2Se or AsH3
2.68x10-5
10ppm
Anode gas
4.2 Calculation results and discussion Figure 10 (a) and (b) show the calculation results of the H2Se and AsH3 addition, respectively. In these figures, the vertical axis indicates a mole fraction of the compounds including Se or As. From Fig. 10 (a), Se will mainly exist as a gas phase of H2Se since as SOFC generally operates at 1073 – 1173 K. Therefore, Se will not react with Ni on the anode surface. It means that the degradation by Se addition will be caused by physical adsorption of Se compounds on the anode surface. Therefore, the power generation performance recovered, as shown in Fig. 7 (a). For As, on the other hand, Ni will be able to react with As to form the solid phase of NiAs. This result suggests that the anode degradation shown in Fig. 7 (b) will be caused by the chemical reactions between As compounds and the anode materials. This is one of the reasons for the irreversible degradation obtained by the experiment.
5. Conclusions The power generation experiment by the button-type fuel cell (SOFC) with the simulated coal gasified gas, that contained the trace element (Se, As) as well as the thermodynamic equilibrium calculations were carried out to elucidate the degradation characteristics of SOFC by the trace elements. The following results were obtained. In the power generation experiment by the simulated coal gasified gas with Se and Submit before 31 May 2011 to
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0.8 0.6 0.4
Se (g) Se2 (g) H2Se (g)
0.2 0.0 1000
1100
1200
Mole fraction of As [-]
Mole fraction of Se [-]
1.0
1300
0.8 0.6 0.4
As (g) As2 (g) As4 (g) AsN (g) NiAs (s)
0.2 0.0 1000
Temperature [K]
1100
1200
1300
Temperature [K]
(a)
(b)
Fig. 10 Mole fractions of (a) Se compounds and (b) As compounds under the SOFC power generation condition by the simulated coal gasified gas
As, the degradation behavior of the anode by Se and As was reversible and irreversible, respectively. For the Se addition, the anode would be degraded by the physical process. For the As addition, on the other hand, As was chemically captured on the anode surface, based on the SEM/EDX observation and the thermodynamic equilibrium calculations.
References [1]
Suzuki E. Proceedings of 2003 Gasification Technologies Conference (SanFrancisco, October 12–15), Gasification Technologies Council, Arlington, USA, 2003, 11 pp.
[2]
Ohtsuka Y, Tsubouchi N, Kikuchi T, Hashimoto H. Recent progress in Japan on hot gas cleanup of hydrogen chloride, hydrogen sulphide and ammonia in coalderived fuel gas. Powder Technology 190 (2009) 340–347.
[3]
FACT, www.crct.polymtl.ca
[4] C.W. Bale, P. Chartrand, S.A. Decterov, G. Eriksson, K. Hack, R. Ben Mahfoud, J. Melançon, A.D. Pelton and S. Petersen. FactSage Thermochemical Software and Databases. Calphad Journal 62 (2002) 189-228.
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Oviedo ICCS&T 2011. Extended Abstract
Coupling gasification and solid oxide fuel cells: effect of tar on anode materials M. Millan1, E. Lorente2, J. Mermelstein1, C. Berrueco1, N.P. Brandon2 1
Department of Chemical Engineering, Imperial College London South Kensington Campus, London SW7 2AZ (UK) 2 Department of Earth Science and Engineering, Imperial College London South Kensington Campus, London SW7 2AZ (UK) email of corresponding author:
[email protected] Abstract The combination of gasification with solid oxide fuel cells (SOFC) has the potential to become an attractive technology for the production of electricity and heat. However the impact of tars, formed during gasification, on the performance and durability of SOFC anodes has not been well established experimentally. This study reports on an experimental study of the effects of carbon formation on anode materials of SOFC from synthetic model tars and real tars arising from the gasification of coal.
1. Introduction The combination of gasification with solid oxide fuel cells (SOFC) is a highly efficiency route of producing electricity and heat. Combined heat and power processes based on SOFCs and coal/biomass gasification have the potential to achieve efficiencies higher than 85% [1]. Solid oxide fuel cells operating at high temperature are able to internally reform a wide range of fuels, including syngas derived from coal gasification. The high catalytic activity of nickel-based anodes allows SOFCs to internally reform hydrocarbons such as methane and make use of CO as a fuel, making SOFCs the most fuel flexible of the different fuel cell types. However, when operating on hydrocarbon fuels (most significantly above C4), nickel-based SOFC anodes are susceptible to significant deactivation from carbon formation, which deteriorates the catalytic activity of the anode and as a consequence causes cell degradation [2]. In particular, aromatic hydrocarbons and tars, which is a complex mixture of organic compounds mostly of aromatic nature that derive from pyrolisis products and secondary reactions [3], may favour the development of carbon deposits. The allowable tar content in the fuel gas is
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one of the key research questions for upcoming fuel cell concepts with integrated gasification, since it determines the gas cleaning requirements. Although recent studies have shown the feasibility of using syngas in SOFCs, the effect of tars on the performance and durability of SOFCs has not been well established [4,5]. Experimental studies designed to obtain a clearer picture on how tars can lead to carbon formation and deposition on SOFC anodes have been carried out. An initial approach was based on the use of model compounds (benzene, toluene, naphthalene) as tars [6,7], which enabled the impact of operating conditions such as current density, steam and tar content and overall gas composition on anode degradation to be established. This work has been further expanded to investigate the impact of an actual gasification tar on two commercially available fuel cell anodes, Ni/YSZ (yttria stabilised zirconia) and Ni/CGO (gadolinium doped ceria). The effect of exposing the catalysts to real tars, as compared to model tars, will be assessed experimentally in terms of the amount of carbon deposition.
2. Experimental section 2.1. Materials Two types of anode catalysts, NiO/YSZ (yttria stabilised zirconia) and NiO/CGO (gadolinium doped ceria), were used in this study. The main characteristics of the powders, supplied from Fuel Cell Materials, are shown in Table 1. Prior to the carbon deposition tests, the anode materials were calcined at 1300 ºC in air for 1h. The calcined powders were then sieved to a particle size of 125-250 μm.
Table 1. Main properties of the SOFC anode materials Anode Material
NiO content (wt%)
BET surface area (m2/g)
NiO/YSZ
66
6.21
NiO/CGO
60
5.60
Toluene and benzene (HPLC grade, DBH, UK) were used as model tars. The real tar sample consisted in a tar obtained from an industrial coal gasifier. The profile of volatilisation temperature of the sample, which is shown in Figure 1, was obtained by a TGA-based weight loss determination [8]. A temperature of 350 °C was
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Oviedo ICCS&T 2011. Extended Abstract
chosen for the injection of tar, since a high percentage of the sample (around 96%) is volatilised at that temperature. 100
Fraction of tar volatilised (%)
90 80 70 60 50 40 30 20 10 0 50
100
150
200
250
300
350
400
450
500
550
600
Temperature (°C)
Fig. 1. Profile of volatilisation temperature of coal gasification derived tar
2.2. Carbon deposition experiments In order to test the carbon deposition characteristics of the nickel oxide based anode materials due to tars present in syngas, a fixed bed flow reactor under atmospheric pressure was used. Typically, 40 mg of powder material was placed on a piece of quartz wool in the 6 mm OD quartz tube reactor. The material was heated to a typical SOFC operating temperature of 765 °C at a rate of 10 °C/min in dry nitrogen. Reduction was carried out for 90 min, by exposing the catalysts at temperature in a 2.5% H2O/H2/N2 mixture, with increasing concentrations of H2, from 5% to 25%, over the reaction period. The gaseous mixture was then changed to the experimental operating conditions of 15% H2 and 2.5% H2O (N2 balance). Tar was injected via a syringe pump (KD Scientific) at a rate of 100 μL/h, producing a tar concentration of 15 g/m3. The feed line was heated to an adequate temperature (around 150 °C for toluene experiments, and around 350 °C for the experiments with real tar) to allow for vaporization of the tar species into the gas phase, and subsequent mixing with the incoming gas. The total flow rate during heating, reduction and reaction was fixed at 100 ml (STP)/min. Anode catalyst samples were exposed to tar for 1h. The amount of deposited carbon (Wcarbon) on the anode materials was measured by weight difference of the samples before and after reaction (Winitial and Wfinal,
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respectively), taking into account the weight loss due to the reduction of NiO to Ni (ΔWreduction), according to Eq. 1. Wcarbon = Wfinal - (Winitial - ΔWreduction)
(1)
3. Results and Discussion The degree of carbon formation on the anode materials, Ni/YSZ (yttria stabilised zirconia) and NiO/CGO (gadolinium doped ceria), exposed to real tar was examined at the experimental conditions described in section 2. The samples were treated at 765 °C in a gas stream of 15 % H2, 2.5% H2O and 15 g/m3 tar for 1 h. To isolate the carbon formation from tars, other gases present in gasification syngas (CO, CO2, methane) were not used in this study. For comparison purposes, the impact of toluene and benzene as model tars were also studied at the same conditions. Figure 4 shows the results obtained in the carbon deposition experiments, in terms of amount of carbon formed over the catalysts materials. It can be observed that Ni/CGO presents a better performance (less carbon formation) than Ni/YSZ in the presence of both model and real tars. This result is in agreement with the expected behaviour of ceria-based anodes, which have been recognised to be effective in suppressing carbon deposition due to the redox nature of ceria [9]. Regarding the comparison between the model compounds and real tar, it can be seen in Figure 4 that the use of the real gasification tar resulted in an intermediate value of carbon deposition as compared with the experiments performed with toluene and benzene. Furthermore, in the case of Ni/YSZ, a sharp decrease is observed in the degree of carbon formation due to the real tar, in comparison with the amount of carbon formed over the catalyst exposed to toluene. The fact that benzene has considerably less impact on the nickel-based anodes degradation than toluene can be related with the higher stability of this compound [10, 11]. As it has been stated before a real tar is a complex mixture of organic compounds mostly of aromatic nature. Therefore, the effect of tar on carbon deposition over the anodes is expected to have a more complex influence than that of model compounds such as toluene and benzene. The results of the present study show the relevance of testing the anode materials with real tars, as the tolerance level of anode materials is highly
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Oviedo ICCS&T 2011. Extended Abstract
influenced by the model tar selected for the study.
1.00
Toluene Real tar Benzene
Carbon deposited (mg C/mg reduced sample)
0.90 0.80
0.08 0.06 0.04 0.02 0.00
Ni/YSZ
Ni/CGO
Fig. 2. Amount of carbon deposited over Ni/YSZ and Ni/CGO, exposed to 2.5% humidified steam, 15 % H2, and 15 g/m3 tar for 1 h at 765 °C 4. Conclusions The impact of toluene and benzene as model tars and a real tar present in coal gasification syngas on the performance of two commercially available SOFC anode materials has been investigated. Carbon deposition measurements indicate that less degradation of the anode catalysts by carbon formation occurs when the anodes are fed with humidified hydrogen gas containing the real gasification tar or benzene, as compared with toluene. The results presented here are considered as a first step in the detailed studies required for full understanding of the influence of gasification tars on short- and long-term SOFC performance.
References [1] Zhang X, Chan SH, Li G, Ho HK, Li J, Feng Z. A review of integration strategies for solid oxide fuel cells. J Power Sources 2010;195:685-702. [2] Offer GJ, Mermelstein J, Brightman E, Brandon NP. Thermodynamics and Kinetics of the Interaction of Carbon and Sulfur with Solid Oxide fuel Cell Anodes. J Am Ceram Soc 2009;92:763-780.
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[3] Milne TA, Evans RJ, Abatzoglou N. NREL/TP-570-25357, 1998 [4] Mermelstein J, Millan M, Brandon NP. The impact of carbon formation on Ni-YSZ anodes from biomass gasification model tars operating in dry conditions. Chem Eng Sci 2009;64:492-500. [5] Singh D, Hernandez-Pacheco E, Hutton P, Patel N, Mann MD. Carbon deposition in an SOFC fuelled by tar-laden biomass gas: a thermodynamic analysis. J Power Sources 2005;142:194-9. [6] Mermelstein J, Brandon NP, Millan M. Impact of steam on the interaction between biomass gasification tars and nickel-based solid oxide fuel cell anode materials. Energy Fuel 2009;23:5042-8. [7] Aravind PV, Ouweltjes JP, Woudstra N, Rietveld G. Impact of biomass-derived contaminants on SOFCs with Ni/Gadolinia-doped ceria anodes. Electrochem Solid St 2008;11:B24-8. [8] Adegoroye A, Paterson N, Li X, Morgan T, Herod AA, Dugwell DR, Kandiyoti R. The characterisation of tars produced during the gasification of sewage sludge in a spouted bed reactor. Fuel 2004;83:1949-1960. [9] Zhu WZ, Deevi SC, A review on the status of anode materials for solid oxide fuel cells, Mater Sci Eng 2003;362:228-239. [10] Ellig DL, Lai CK, Mead DW, Longwell JP, Peters WA. Pyrolysis of volatile aromatic hydrocarbons and n-Heptane over calcuium oxide and quartz. Ind Eng Chem Process Des Dev 1985;24:1080-7. [11] Coll R, Salvado J, Farriol X, Montane D. Steam reforming model compounds of biomass gasification tars: conversion at different operating conditions and tendency towards coke formation. Fuel Process Technol 2001;74:19-31.
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6
Explosions in Coal Mines due to Emission of Molecular Hydrogen via Atmospheric Weathering Processes 1,2
Haim Cohen 1-
,
Department of Biological Chemistry, Ariel University Center at Samaria, Ariel, 40700 Israel, phone: 00972-524306878, fax: 00972-8-9200749, email:
[email protected]
2-
Chemistry Department, Ben-Gurion University of the Negev, Box 653 Beer Sheva , Israel, email:
[email protected]
Abstract Explosions in coal mines are attributed to methane gas concentrations above the LEL or accumulation of fine coal dust which can undergo fast radical reactions. In order to avoid occurrence of such conditions in deep coal mines, methane detectors are installed and good ventilation is required. Still the annual death toll due to explosions in coal mines is ~ 5,000, mainly in China. Bituminous coals in contact with atmospheric oxygen undergo autocatalytic heating via exothermic oxidation of atmospheric oxygen to produce carbon dioxide, carbon monoxide, low molecular weight organic molecules (C1-5) and water. It has been established that molecular hydrogen is also produced. The hydrogen is produced via oxidative decomposition of formaldehyde groups with surface hydroperoxides and is catalyzed by the coal surface. Thus accumulation of hydrogen in a crack in the coal seam in a underground mine above 4.1% (LEL in air) might initiate an explosion. Indeed in the last decade several unexplained explosions have been reported in coal mines. We have checked if accumulation of hydrogen in a deep coal mine might reach the LEL and found that indeed such a situation might occur. These observations indicate that molecular hydrogen accumulation in confined spaces might be another source for unexplained explosions in coal mines and that installation of hydrogen detectors in coal mines is essential to reduce the risk of occurrence of explosions in coal mines.
Keywords: heating
coal mines, explosions, coal weathering, hydrogen explosions, self
1
1. Introduction Coal serves today as the largest source of fossil fuel and is expected to continue as a major energy supply in the next century. Thus, coal mining is an important and vital industry in the global economy. One of the major problems in underground coal mining is the occurrence of explosions in coal mines. The death toll due to these explosions accounts for thousands of human lives annually1. The main source observed for explosions in coal mines (and also confined storage facilities containing coal) stems from 2 sources: (i) Accumulation of methane gas (coal bed methane) released from the inner pores of the coal upon mining and exposure to atmospheric oxygen. The LEL (Lower Explosion Limit) of methane in air is 5.3% and (ii) Accumulation of fine coal dust in the air of the mine because of the mining. This dust has a very large surface area and thus it is available for radical reactions. Both processes emerge because of fast radical reactions that can be initiated by several sources (such as static electrical charge, random spark etc.). As these reactions are very fast (diffusion controlled rate, rate constants approaching 1x1012 M-1s-1 order)) and very exothermic the result can be a devastating explosions. In order to minimize the occurrences of such explosions in underground coal mines several measures are taken. Methane monitors are installed in the coal mines to indicate accumulation of methane gas which might approach dangerous levels. As well as efficient ventilation of the atmosphere of the mine in order to dilute dangerous gas concentrations and filter the fine coal dust which is emitted during mining. Still the death toll approaches 5,000 per year, the majority of which are in the Peoples Republic of China which is both the biggest consumer and producer of coal in the world. The main reason could be human negligence in the maintenance and operation of the mines. However, in several cases the operation of the mine was according to all safety measures and still unexplained explosions did occur. In the last 30 years, Israel has started to use bituminous coal as the major fossil fuel for power production in utility plants and at present 13 Mtonns of coal are consumed annually at 4 power stations producing ~55% of the electricity consumed in Israel. As Israel has no coal of it's own, it has to import the coals mainly from South Africa but also from Russia, Indonesia, Australia and Columbia which is transported via large ships. Also large coal piles (more than 1 Mtonnes are stored in two storage facilities in Ashkelon and Hadera. Coal, like any other organic material in contact under atmospheric conditions, undergoes surface oxidation/weathering processes which might result in autocatalytic heating of the coal pile if the heat formation in the pile is faster than the heat dissipation from the pile2. The process is dependant on coal rank and properties As a result, lignite coals undergo deterioration due to weathering processes in a matter of days while it can take bituminous coals several weeks to reach the same state. This multistage mechanism is quite complicated and even today it is not fully understood. The first stages involve physical adsorption and chemisorption of atmospheric oxygen. The second stage is the formation of surface oxides and hydroperoxides which can partially decompose to yield low molecular weight inorganic gasses like carbon oxides (CO, CO2), water, hydrogen (H2) and some organic gases (C1-5)3. The observation that hydrogen gas is evolved during the weathering process is quite novel and was extensively studied in our group in the last 2
30 years. It is observed that the release of H2 (that is considered as an reduction product) is linearly dependant on the atmospheric O2 consumed by the coal (which is an oxidative process!!!)3,4. The mechanism of this complex reaction is under study and involves reaction of formaldehyde CH2O release from the coal (via the weathering process) with hydroperoxide –COOH groups at the coal surface to yield dioxirane, H2CO2 as intermediate which further decomposes to yield molecular hydrogen, H2 as a product5. We have decided to check the possibility that accumulation of molecular hydrogen in mines could be the source of unexplained explosions in underground mines and to suggest means to monitor that possibility in confined spaces containing coal.
2. Experimental All chemicals and gases used throughout the study were of AR grade and supplied by Aldrich, Fluka, Merck or Maxima. The water used throughout this study was purified water (via ion exchange columns). Coal. Experiments in this work were carried out with three classes of coals; bituminous, sub-bituminous and lignite. The bituminous coal was from South Africa (denoted as SA) and the sub-bituminous coals from Indonesia (denoted as INA) and the USA (denoted as BAI). The South African bituminous coal used in this work serves as the major fossil fuel in coal fired power plants in Israel (more than 60% of the coal consumption). However the Sub-Bituminous Indonesian coal and the BAI USA coals are also fired in the Israeli utilities. The properties of all the coals are presented in Table 1. The experiments were performed in sealed glass vials (40 ml) used as batch reactors. The reactors were charged with coal (particle size 74µ ≤ X ≤ 250µ) in an air &&ve oven model FT 300. atmosphere and heated at 55-95C for various periods in a nu The effect of the oxygen concentration has on the processes was also studied under an atmosphere of pure argon or pure oxygen. All the coals in the present study were prepared by grinding them down and separating them by grain size via sieves. The coal samples were then dried in a Heraeus vacuum oven model VT6060 for 24 hours at 60C. Gas Chromatography. The amount of the gases (CO2, CO,, N2, O2, hydrocarbons) in the reactors was determined using a gas chromatograph (Varian model 3800) equipped with a thermal conductivity detector & a flame ionization detector connected in series. The gases were separated on a Carbosieve B 1/8”, 9’ ss column using a temperature programmed mode. The experimental error in the G.C. determination is ±5%. The gaseous atmosphere was sampled (1ml samples) after the reaction, with gas tight syringes (Precision Syringes, model A2) and the composition was determined in the gas chromatograph. The gases that could be determined are hydrogen, nitrogen, oxygen, carbon dioxide, carbon monoxide, methane & ethylene. The argon gas present is not separated from oxygen in the GC column, thus the value determined for oxygen includes ~0.93% argon gas. As the reactions studied are gas/surface reactions, the reproducibility of the results is not good. Therefore, each experiment has been carried out with duplicates in order to reduce the total experimental error. However, the error is ±15%, mainly due to the nature of the heterogeneous reactions studied in the experiments. 3
Table 1. Properties of Coalsa type
a b
proximate analysis (wt %) ashdb
VMdaf
SA
moisture 1.20
13.60
IN
2.61
BAI
5.87
ultimate analysis (wt %, daf) CV( J ⋅ g
H 4.05
O
S
27.93
C 73.3
5.52
0.46
28,416
10.37
35.43
71.3
4.51
8.81
0.66
28,564
7.78
37.20
78.1
5.18
NA
1.50
28,898
VM = volatile matter; CV = calorific value; daf = dry ash free. db – dry basis SA = South Africa; IN = Indonesia; BAI- USA.
3. Results and Discussion Release of Molecular Hydrogen- Effect of Coal Type We have decided to measure the effect of coal type on the release of molecular hydrogen via the weathering process. Thus, simulation experiments with the three coals have been carried out in small (100ml) glass reactors containing 5 grams coal at the temperature range 55-95C heated for 24 hours under isothermal conditions were performed, the results are given in Table 2. Table 2. Emission of molecular hydrogen – simulation weathering Rate of H2 emission at 55C Rate of H2 emission at 95C b coal ppmv/gram coal x hour 0.615 SA 12.9
a b
IN
0.826
15.0
BAI
1.51
24.5
carried out in 49ml sealed glass reactors heated at different temperatures for 24 hours SA = South Africa; IN = Indonesia; BAI- USA.
As can be clearly seen the emission of molecular hydrogen is dependent on the properties of the coal. However all the bituminous coals studied do release H2 upon weathering and the process is temperature dependent. In order to asses the possibility of accumulation of molecular hydrogen in an underground coal mine we have envisaged the following scenario: A 2.0 cubic meters volume of a confined space (crack) occurring in contact with a coal seam where the temperature of the coal is 55C and the coal is under atmospheric conditions. At these conditions the emission is at a rate of 0.615; 0.826 or 1.51 ppmv H2/gram coal x hour for SA, IN or BAI coal respectively. If the size of the surface coal approaches 1,000 Kg around the crack (probably this is only a lower limit) and the weathering process endures for 30 days at least than the amount of molecular hydrogen release assuming total accumulation (without ventilation in the crack) to the crack from the surface of the coal walls into the crack 4
−1
)
is 44 ; 56 and 102 L for the three coals. This corresponds to 2.2 ; 2.8 and 5.1 %vol which is higher (for the BAI coal) and very close (for the South African and the INIndonesian coals) to the 4.0% LEL level (lower explosion limit) in air of molecular hydrogen!!!. Thus this very simple estimate indicates that accumulation of molecular hydrogen, H2, produced via weathering low temperature oxidation of bituminous coals might be the main cause for unexplained explosions in underground coal mines.
4. Conclusions The results of this approximation and the conclusions are very simple. (i)
(ii)
(iii)
In order to prevent the risk of initiation of explosions via the accumulation of molecular hydrogen H2 in underground coal mines due to the low temperature oxidation (weathering) of bituminous coals it is recommended to install in addition to methane gas monitors also H2 detectors which are simple and cheap in order to follow up such possibility and to have alarms if H2 levels are rising in the mine. Furthermore it is also reasonable to install such monitors in confined spaces such as coal ship holds or bunkers in which coals are being stored for period of weeks and longer. This operation will increase the safety measures in the coal mining industry and will probably reduce the death toll appreciably in this industry.
5. References 1. . a. A. W.Davies and A. K. Isaac, Coal dust explosions: a continuing menace, Inst. Mining and Mettallurgy Trans. Sect. A: Mining Industry, 108, 85, 1999. b. G. L. Smith and J. J. du Plessis, J. South African Inst. Mining and Metallurgy, 99, 117, 1999. c. R. McGregor, Miners pay high price for China’s coal, Financial Times, July 17th, 2002. d. Coal mine explosion kills 48 in China, nd www.annanova.com/news/story/sm268418, 22 April, 2001. 2. C.R. Nelson, Chapter 1, “Chemistry of Coal Weathering”, C. R. Nelson editor, Elseiver (1989). 3. a. S. L. Grossman Ph.D. Thesis, “Low Temperature Atmospheric Oxidation of Coal”, Chemistry Department, Ben-Gurion University of the Negev, Beer-Sheva, Israel, (1994). b. S. L. Grossman, S. Davidi and H.Cohen, Fuel, 70, 897 (1991). c. S. L. Grossman, I. Wegener, W. Wanzl, S. Davidi and H.Cohen, Fuel, 73, 762 (1994). 4.
S. Davidi, S. L. Grossman and H.Cohen, Fuel, 74, 1357 (1995).
5.
U. Green and H. Cohen, Energy & Fuels 9 23 (6), pp 3078–3082 (2009).
5
Oviedo ICCS&T 2011. Extended Abstract
Pyrolysis and Combustion kinetics using the Distributed Activation Energy Model
F. Saloojee1, S. Kauchali2, N. Wagner
University of Witwatersrand, Johannesburg 1
[email protected]
2
[email protected]
Oviedo ICCS&T 2011. Extended Abstract
Abstract In this work, the kinetics of pyrolysis and combustion of coal have been studied. Thermo-gravimetric experiments were simulated for these reactions at constant and variable heating rates. The results were used with an appropriate model in order to determine the reaction kinetics. The Distributed Activation Energy Model (DAEM) is commonly used to describe the coal pyrolysis process [1]. The model states that coal devolatilizes according to a number of first order reactions, each with unique activation energy. An algorithm has been developed to calculate the kinetic parameters of each reaction using the DAEM [1]. The algorithm was tested on simulated TGA data for pyrolysis reactions at constant and variable heating rates. Results show that this is a robust method for calculating the kinetic parameters of first-order reactions. Further scrutiny of the inversion algorithm has shown that the calculation of the activation energy is a model-free method. The algorithm was applied to simulated TGA data for coal combustion following the shrinking core model. Results show that the DAEM can be used as a model-free method to calculate the activation energy of coal combustion. However, the calculation of the pre-exponential factor requires the use of an appropriate structural sub-model.
Oviedo ICCS&T 2011. Extended Abstract
1. Introduction The conversion of coal to provide energy is one of the major causes of South Africa’s contribution to greenhouse gas emissions. These emissions can be reduced by correct design and efficient operation of coal conversion equipment. In order for this to be achieved, kinetics of coal conversion reactions need to be easily available. This was the motivation for the work which has been presented here. The Distributed Activation Energy Model (DAEM) has been used to model pyrolysis reactions of coal [1,2,3]. This model describes pyrolysis as a series of first-order Arrhenius reactions, each uniquely characterised by its activation energy. The rate equation can be expressed as follows:
!" #$%,
, exp exp
&1(1
Where M(t) is the mass of the fuel at time t, M0 is the initial mass of the fuel and w is the fraction of ash in the fuel. For each component in the fuel which reacts, fi,0 is the initial mass fraction of that component in the fuel, Ei is the activation energy of the reaction and Ai is the pre-exponential factor of the reaction. Based on this model, Scott et al. (2006) [1] developed a method of calculating the Ei, Ai and fi,0 for each reaction. This calculation requires data from TGA experiments at two different heating rates. For each devolatilization reaction occurring, the rate equation can be written as: exp
2
Integration of the above expression gives:
, exp exp B
3
where B is the heating rate and the exponential term can be expressed as Ψ. At a given conversion, the value of fi is always the same. Therefore, the right hand sides of Equation 3 can be equated for any two heating rates to calculate Ei. It is then assumed
Oviedo ICCS&T 2011. Extended Abstract
that each reaction reaches a conversion of 63.21 % [1] and the value of Ei is used to calculate the Ai for each reaction. If Equation 1 is written in Matrix form as: Ψ- Ψ. 1 Ψ 1
-, $ 1 ,- / 0Ψ- - Ψ. - 1 Ψ$ - 12 3 , ., / 1
4, . Ψ- . Ψ. . Ψ$ . 1
4
then inversion of the matrix equation provides the values of fi,0. The aim of this work was to use the method developed by Scott et al. (2006) as a modelfree method to calculate the activation energy of coal char combustion. 2. Pyrolysis Modelling The first step in this work was the testing of the algorithm on simulated TGA curves. Curves were created with specific E’s and A’s for different types of reactions. The DAEM algorithm was applied to these curves to test whether the parameters used for the simulations could be regenerated. The calculated kinetic parameters were then used to model the reactions for comparison with the original curves. 2.1 Simulated single, first-order reaction with constant heating rate For pyrolysis, a curve was created for a single first-order reaction with the kinetic parameters provided in Table 1. The DAEM was able to correctly calculate the kinetic parameters used and model the curves at the two heating rates used for the calculation. It was also able to extrapolate these predictions to higher and lower heating rates. The kinetic parameters calculated by the DAEM are presented in Table 1 as well. The simulated curves and model fits for the case of the single reaction are presented in Figure 1. Table 1: Comparison of original kinetic parameters and those recovered by the DAEM for a single, first-order reaction with constant heating rate Parameter
Recovered Value
Original Value
% Error
f0
0.9999
1
0.01
E (KJ/mol)
135.1974
135.2
0.0019
A(min-1)
1.13×1010
1.1001×1010
2.8
Oviedo ICCS&T 2011. Extended Abstract
1 0.001 K/min 10 K/min 1000 K/min 10000 K/min 0.001 K/min model 10 K/min model 1000 K/min model 10000 K/min model
0.9
Mass Fraction Remaining
0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0 300
400
500
600
700 800 900 Temperature (K)
1000 1100
1200
Figure 1: Simulated TGA curves for a single, first-order reaction at different heating rates, with model predictions for each of these heating rates 2.2 Seven simulated first-order reactions with constant heating rates For the second pyrolysis test, TGA curves were created for seven first-order reactions, using the kinetic parameters in Table 2. The DAEM again proved capable of calculating the correct kinetic parameters, as shown in Table 2, and of modelling the curves at the different heating rates. The simulated curves and model fits are presented in Figure 2. Table 2: : Comparison of original kinetic parameters and those recovered by the DAEM for seven first order reactions at constant heating rates
Reaction
Sum of f0
Original f0
Average E
Original E
Average A
Original A
1
0.1436
0.142857
150.1595
150
1.035E+15
1E+15
2
0.1432
0.142857
176.6007
175
1.552E+15
1E+15
3
0.1398
0.142857
192.6182
190
1.619E+15
1E+15
Oviedo ICCS&T 2011. Extended Abstract
4
0.145
0.142857
200.2781
200
1.056E+15
1E+15
5
0.1431
0.142857
225.9348
225
1.21E+15
1E+15
6
0.1426
0.142857
251.253
250
1.277E+15
1E+15
7
0.143
0.142857
274.7396
275
9.675E+14
1E+15
1 0.001 K/min 10 K/min 1000 K/min 10000 K/min 0.001 K/min model 10 K/min model 1000 K/min model 10000 K/min
0.9
Mass Fraction Remaining
0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0 300
400
500
600
700 800 900 Temperature (K)
1000 1100
1200
Figure 2: Simulated TGA curves for seven first-order reactions at different heating rates, with model predictions for each of these heating rates 2.3 Simulated first-order reaction with non-constant heating rate It was noted that in real TGA experiments, the heating rates of the sample are not always constant. Curves were then created for the heating rate profiles shown in Figure 3. The algorithm was modified to use instantaneous values for the heating rate at each temperature, as opposed to a constant value. In this case, the DAEM algorithm was again capable of calculating the correct kinetic parameters, as shown in Table 3.
Oviedo ICCS&T 2011. Extended Abstract
70
45 Constant Heating Rate Variable Heating Rate
60
Constant Heating Rate Variable Heating Rate
40 35 Heating Rate (K/min)
Heating Rate (K/min)
50 40 30 20
30 25 20 15 10
10 0 300
5 400
500
600
700 800 900 Temperature (K)
1000 1100 1200
0 300
400
500
600
700 800 900 Temperature (K)
1000 1100 1200
Figure 3: Heating rates vs. temperature profiles used for simulation of TGA data Table 3: Original and Recovered Kinetic Parameters from DAEM for a first-order reaction with variable heating rates Parameter
Recovered Value
Original Value
f0
0.9965
1
E (KJ/mol)
135.212
135.2
A(min-1)
1.1×1010
1.1001×1010
3. Model-free kinetics using the DAEM Examination of this method indicates that the calculation of E is a “model-free” method, while the calculation A is based in the assumption of a first order reaction. The aim of this work was to use the method developed by Scott et al. (2006) as a model-free method to calculate the activation energy of coal char combustion. The rate equation for a particular component in coal can be expressed as: exp 6
5
where g(fi) is a function describing the reaction model of component i. If a heating rate B is used, Equation 5 can be written as: = exp < =; 89.; 6 89
6
Integrating both sides of Equation 6: ? ?@ , A ln ΨB, T
7
Oviedo ICCS&T 2011. Extended Abstract
Integrating both sides of Equation 6: ? ?@ , A ln ΨB, T
7
Where G(fi) – G(fi,0) represents the integral on the right hand side of Equation 7 between limits fi and fi,0, and Ψ is as described previously [1]. At a specific conversion, the values of fi for any two heating rates are equal [1], so: ln ΨB- , T- ln ΨB. , T.
8
From Equation 8, the value of Ei can be calculated, proving that the method is indeed a model-free method. 3.1 Model-free combustion kinetics A combustion reaction was simulated according to the shrinking core model [4] and the data was fed into the algorithm. The DAEM was able to recalculate the activation energy used for the simulation. However, the value of A calculated by the algorithm was incorrect. The correct value was calculated by integrating the shrinking core rate expression and using the calculated value of E. Using these parameters, the shrinking core model was able to predict the reaction at the two heating rates used in the algorithm and at higher and lower heating rates, as seen in Figure 4.
Oviedo ICCS&T 2011. Extended Abstract
1 1 K/min 10 K/min 1000 K/min 10000 K/min 1 K/min model 10 K/min model 1000 K/min model 10000 K/min model
0.9
Mass Fraction Remaining
0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0 300
350
400
450 500 550 Temperature (K)
600
650
700
Figure 5: Simulated TGA curves for shrinking core reaction at different heating rates with model predictions for each of these heating rates
4. Conclusion The simulation of TGA data provides a test of how well the algorithm is able to perform for different types of reactions. The effects of multiple reactions, non-constant heating rates and a non first-order reaction mechanism were investigated. It has been found that the DAEM inversion algorithm is a model-free method of calculating the activation energy of coal pyrolysis and combustion. If the reaction mechanism is not first-order, the calculated value of E must be used in conjunction with an appropriate model to calculate the pre-exponential factor of the reaction.
Oviedo ICCS&T 2011. Extended Abstract
Acknowledgement University of Witwatersrand, Johannesburg Eskom, South Africa References [1] Scott, S.A., Dennis, J.S., Davidson, J.F. & Hayhurst A.N., 2006. An algorithm for determining
the
kinetics
of
devolatilisation
of
complex
solid
fuels
from
thermogravimetric experiments, Chemical Engineering Science, 61, p. 2339-2348 [2] Pitt, G.J., 1962. The kinetics of the evolution of volatile products from coal, Fuel, 41, p. 267-274 [3] Donskoi, E. & McElwain, D.L.S., 2000. Optimisation of coal pyrolysis modelling, Combustion and Flame, 122, p. 359-367 [4] Everson, R.C., Neomagus, H.W.J.P., Kasaini, H. & Njapha D, 2006a. Reaction kinetics of pulverized coal chars derived from inertinite-rich coal discards: Characterisation and combustion, Fuel, 85, p. 1067-1075
Oviedo ICCS&T 2011. Extended Abstract
A Graphite Furnace Atomic Absorption Spectrometer as an Experimental Platform for Studying Matrix Effects in Trace Element Vaporization during Coal Combustion E. I. Kozliak, O. V. Klykov, A. A. Raeva, D. T. Pierce and W. S. Seames The Sustainable Energy Research Initiative Program (SUNRISE), University of North Dakota, Departments of Chemistry and Chemical Engineering, 151 Cornell St., Mail Stop 9024, Grand Forks, ND 58202 USA,
[email protected]
Abstract To date, no direct method is available to assess the partitioning of toxic trace elements (TTE) between the gas and condensed phases during coal combustion. Such a method was developed by us based on the use of a graphite furnace atomic absorption spectrometer (GFAAS) as the experimental platform, with the capability of generating data for temperatures up to 2800 °C. While our prior work reported on the previous ICCS&T meeting introduced to GFAAS small amounts of pure TTE standards to emulate the coal particle pyrolysis (reducing environment, high in carbon), the current study addressed the partitioning of TTEs entrapped in inorganic inclusions, introduced as TTE-doped aqueous solutions. The TTE (arsenic, antimony and selenium) atomization activation energies were determined with and without matrices to assess the matrix effects. Two cationic matrices, iron(III) and, particularly, calcium(II), significantly increased the atomization/vaporization activation energies indicating the increased TTE retention in the solid phase. By contrast, anionic matrices, e.g., acetate and aluminate, did not alter the activation energies, as expected. Unexpectedly, one more anionic matrix, silicate, decreased the atomization/energy for selenium while reducing its atomic absorption signal. Apparently, this matrix non-specifically blocked the TTE access to carbon resulting in its evaporation in the molecular rather than elemental form. Thus, several inorganic matrices were shown to alter the concentration and occurrence of TTEs in the gas phase at high temperatures characteristic for coalfired furnaces.
1
Oviedo ICCS&T 2011. Extended Abstract
1. Introduction Accurate modeling of the partitioning of trace elements (TE; e.g., As, Se and Sb) during pulverized coal combustion is an important step in understanding the impacts these TEs have on the environment for traditional PC combustors1 and on downstream process systems for gasification2-4 and oxy-coal combustion systems. Previous methods are all based on using average or bulk phase conditions and thus do not account for the effects of the microenvironment/matrix surrounding TEs. In our previous work, we used a graphite furnace atomic absorption spectrometer (GFAAS) to simulate in situ the microenvironment similar to that observed during coal combustion (i.e. reducing, high temperature microenvironment). Aqueous solutions of As, Se and Sb oxides were introduced into the furnace to model organically associated TEs.5 Then, the activation energies of TE vaporization/atomization in the furnace were calculated. It has been shown that the activation energies determined this way are independent from operational parameters of the instrument (TE concentration, Ar gas flow rate, ashing temperature, and atomization temperature).5 In the current study, TEs present as mineral inclusions were modeled by introducing the TE solutions together with various organic and inorganic matrices into GFAAS. The use of water soluble matrices allowed for a homogeneous mixing of the given TE and matrix, yet enabling the inherent accessibility of carbon. Matrices such as Ca(OAc)2, Fe(NO3)3, K2SiO3 and NaAlO2 were used. Upon heating, Ca(OAc)2,and
2
Oviedo ICCS&T 2011. Extended Abstract
Fe(NO3)3 decompose into CaO and Fe2O3 respectively, which are the major components of inorganic fraction of coal.6 K2SiO3 and NaAlO2 were also used as matrices to determine if there is any retention of TEs by mineral deposits formed on the walls of the boilers; their major components include potassium and sodium aluminosilicates.7
2. Experimental section Temperature measurements and instrument modification. A Modline 5 infrared (IR) thermometer (Model 52-3024 with 2B lens, IRCON, Niles, IL, USA) was positioned above the furnace to measure temperatures of the furnace wall. The thermometer provided a RS-485 digital output of factory-calibrated temperatures to a host computer. Factory calibration was traceable to NIST blackbody standards. Operated in this manner, the IR thermometer was capable of measuring accurate temperatures over a range of 750 – 3000 °C within a focal distance of 152 – 305 mm, and with a response of 6.6 ms to 60 s. The Hitachi Z-2000 AA spectrometer was modified (as described elsewhere5) to provide simultaneous and precisely synchronized absorbance and temperature measurements. The precise synchronization of absorbance and temperature measurements was achieved by recording and aligning (in time) optically-encoded signals embedded within the AA output and recorded along with the thermometer analog current output. Calculation of activation energies. Activation energies were calculated using the method described by Smets.8 This method is based on the calculation of rate constants
3
Oviedo ICCS&T 2011. Extended Abstract
of atom formation:
k1 =
n(t ) ∞
=
A(t ) ∞
(1)
∫ n(t )dt ∫ A(t )dt t
t
where k1 is the first-order rate constant, n(t) is the number of TE atoms at time t, and A(t) is the absorbance value at time t. Absorbance vs. time profiles from the Hitachi ∞
AA software were used to find A(t). The integrated absorbances ( ∫ A(t )dt ) were t
calculated using OriginPro 8.1. The portion of the absorbance vs. time curve with a positive slope (i.e., the atom formation part of the peak as opposed to its dissipation) was used to construct the Arrhenius plot. The points taken were from the part of the curve between 10×noise to 0.9×Amax. When an absorbance profile contained double peaks (which was the case for some of the experiments with matrix), the multiple peak fit function in OriginPro 8.1 was used for peak deconvolution. The fit used was Gaussian. Temperature vs. time data were obtained using a Modline 5 IR thermometer as described above. Activation energies were calculated from the slope and pre-exponentials were calculated from the intercept of Arrhenius plots (log k1 vs. 1/T).
3. Results and Discussion The effect of two representative inorganic matrices, NaAlO2 and Ca(OAc)2, on the activation energies for As, Sb, and Se atomization was determined (Table 1).
4
Oviedo ICCS&T 2011. Extended Abstract
Table 1. The effect of two representative inorganic matrices on the atomization activation energies for As, Sb, and Se.
Activation energy, kcal·mol-1 0.01 – 0.1 M
TE
No
0.01 M
matrix
NaAlO2
Ca(OAc)2* Peak 1
Peak 2
As
60 ± 5
61 ± 2
51 ± 2
70 ± 8
Sb
37 ± 2
34 ± 3
41 ± 3
65 ± 5
Se
69 ± 5
57 ± 7
136 ± 9
* Ca(OAc)2 concentration was 0.01 M for As and Se and 0.1 M for Sb
0.01 M NaAlO2 did not exhibit any statistically significant effect on the activation energies for all three elements. Figs. 1 - 3 show that neither the peak shapes nor temperatures at which the peaks start rising depend on whether this matrix is present. This observation indicates that aluminates are not expected to influence the partitioning of these three TEs. 0.35
0.25
No matrix
A
NaAlO2
-0.05
0
t, s
2
No matrix
A
NaAlO2
-0.05 0
1.5
t, s
Figure 1. Absorbance profile
Figure 2. Absorbance profile
for 100 ppb As with and
for 100 ppb Sb with and without
without 0.01 M NaAlO2.
0.01 M NaAlO2. 5
Oviedo ICCS&T 2011. Extended Abstract
0.2
0.35
NaAlO2
0 -0.05
No matrix
No matrix
A
2
t, s
Figure 3. Absorbance profile
A
Ca(OAc)2
-0.05 0
1.5
t, s
Figure 4. Absorbance profile
for 100 ppb Se with and without for 150 ppb Sb with and without 0.01 M NaAlO2.
0.1 M Ca(OAc)2.
In contrast to aluminate, calcium showed a significant effect on the activation energies for As, Sb, and Se. The largest effect was observed for Se; its activation energy increased by 86 kcal·mol-1 in the presence of this matrix. A higher activation energy indicates significant chemical interactions between the TE and this matrix, which is supported by the literature. Note that Ca(OAc)2 readily decomposes upon heating to CaO. Thus, CaO is expected to retain Se in mineral exclusions unless high temperatures (around 2300 °C) are reached. Then, if this TE is released to the vapor phase, it will be in its elemental form rather than its oxide. For As and Sb, two absorbance peaks instead of one were observed. Activation energies were calculated separately for each peak upon their deconvolution, as described in the Experimental section. Their occurrence suggests the existence of two competing atomization mechanisms for each TE. The activation energy calculated using the data points of the first peak matched that obtained without a matrix, thus indicating that this peak is likely due to the interaction of TE vapors with carbon. By contrast, the activation energy calculated using the second peak’s data points was significantly higher, apparently reflecting the TE’s interaction with the given inorganic matrix, i.e., calcium. This increase in Ea indicates a significant TE retention by CaO. However, the retention for As and Sb was not as strong as in the case of Se. Perhaps, 6
Oviedo ICCS&T 2011. Extended Abstract
this effect is due to the difference in preferred forms of occurrence of TEs in the vapor phase. Since Se evaporates in the form of its acidic oxide, its retention by basic CaO is expected. By contrast, if other TEs, e.g., As, vaporize mostly as atoms, a less pronounced influence of CaO is expected. Indeed, even though two peaks were observed for As in the presence of calcium, the activation energies calculated with and without this matrix were similar (Table 1). The slight influence of CaO on the evaporation of Sb, along with the occurrence of two peaks, indicates that this TE may still evaporate in both atomic and molecular forms, although the atomic form is dominant. An explanation for this effect is that vaporization combined with atomization occurs via multiple evaporation/ condensation cycles, some of them being coupled with atomization. This would explain the occurrence of two peaks for Sb, as in Figure 4. The main peak is similar to that obtained using the unlined graphite tube, i.e., it is due to both the atoms’ sorption on any surface and the oxides’ sorption on carbon. The second, smaller, peak observed only in the presence of the matrix, is due to the sorption of oxides on basic CaO, resulting in hindered TE atomization. The effect of iron on Ea was similar to that of calcium, thus proving that the increase of Ea was due to the TE oxide interactions with a catión. However, replacement of aluminate with another anionic matrix, silicate, unexpectedly resulted in a significant decrease of Ea for Se and Sb to 26-30 kcal/mol, whereas that for As remained unchanged. Based on this observation, the mechanism of As atomization and, therefore, its partitioning during coal combustion do not appear to be affected by the presence of 0.01 M K2SiO3. By contrast, for Se and Sb, silicates appear to either facilitate the rate-limiting step of their atomization or alter the atomization path resulting in a different rate limiting step of a lower activation energy. Additional mechanistic information was obtained when analyzing the atomic absorbance signal intensity. For all three TEs, the analytical signal was significantly suppressed in the presence of this matrix. The signal suppression of the elemental TE was most pronounced for Se; this feature provided an opportunity to obtain information on the TE speciation in the vapor phase. Apparently, in the presence of silicate, compared to the matrix-free case, the vaporized Se was produced in larger amounts in its molecular form, not ‘seen’ by the GFAAS detector. So, its sizable fraction left the 7
Oviedo ICCS&T 2011. Extended Abstract
furnace prior to atomization (only the target atoms are detected by GFAAS). The importance of this observation to coal combustion is that the presence of silicates appears to promote selenium evaporation without its atomization, thus stabilizing molecular species, e.g., selenium oxides. Blocking the selenium access to elemental carbon as a reducing agent could be the chemical reason for this effect.
4. Conclusions According to the developed method allowing for in situ measurements, exclusive evaporation in the elemental form occurs for selenium only in the presence of calcium whereas arsenic is expected to evaporate predominantly in its elemental form, either with or without matrix. By contrast, selenium (in the absence of a cationic matrix) and, particularly, antimony (with or without matrices) may evaporate as oxides followed by their subsequent atomization upon reactions with carbon. Cationic matrices bind selenium and antimony oxides thus significantly reducing the partitioning of these TEs into the gaseous phase.
Acknowledgement Funding for this study was provided by the U.S. Department of Energy (Grant DEFG02-6ER46292).
References [1] Seames WS, Wendt JOL. Partitioning of arsenic, selenium, and cadmium during the combustion of Pittsburgh and Illinois #6 coals in a self-sustained combustor. Fuel Process Technol 2000;63:179-96. [2] Frandsen F, Dam-Johansen K, Rasmussen P. Trace elements from combustion and gasification of coal-An equilibrium approach. Prog Energy Combust Sci 1994;20:115-38. [3] Kalfadelis CD, Magee EM. Ind Eng Chem Fund 1977;16(4): 489. 8
Oviedo ICCS&T 2011. Extended Abstract
[4] Erickson TA, Galbreath KC, Zygarlicke CJ, Hetland MD, Benson SA. Trace Element Emissions, Final Technical Progress Report, Prepared for U.S. Department of Energy, DE-AC21-92MC28016, October 1998. [5] Raeva AA, Pierce DT, Seames WS, Kozliak EI. A method for measuring the kinetics of organically associated inorganic element vaporization during coal combustion. Fuel Process Technol 2011;92:1333-9. [6] Bool LE, Helble JJ. A laboratory study of the partitioning of trace elements during pulverized coal combustion. Energy Fuels 1995;9:880-7. [7] Thompson D, Argent BB. Mobilisation of sodium and potassium during coal combustion and gasification. Fuel 1999;78:1679-89. [8] Smets B. Atom formation and dissipation in electrothermal atomization. Spectrochim. Acta 1980:35: 33–41.
9
Oviedo ICCS&T 2011. Extended Abstract
Organic sulphur alterations in consecutively chemically and biotreated lignites 1
1
2
2
L.Gonsalvesh1, S.P.Marinov , M.Stefanova , R.Carleer , J.Yperman 1
Institute of Organic Chemistry, Bulgarian Academy of Sciences, bl. 9, Sofia 1113, Bulgaria,
[email protected] 2
Research group of Analytical and Applied Chemistry, CMK, Hasselt University, Agoralaan – gebouw D, B-3590, Diepenbeek, Belgium,
[email protected]
Abstract There are a variety of chemical, physical and biological desulphurization methods for inorganic sulphur removal. Recently, biological approaches received special attention as a potential technique to reduce organic sulphur under mild experimental conditions at lower operating and capital costs. The aim of the present study is to apply combination of desulphurization methods (chemical and microbial) toward inorganic and organic sulphur functional forms removal in coal. In the current study, one Bulgarian coal sample from “Maritza East” lignite deposit, which is with a significant role in the total energy supply for the country, is used. In order to improve the organic desulphurization effect and to concentrate efforts on a deeper research of organic sulphur changes, investigated sample is preliminary chemically treated. The following products are subjected to biodesulphurization with Pseudomonas Putida bacteria: initial coal, preliminary oxidized coal under atmospheric air and high temperature, demineralized coal, demineralized and oxidized coal, demineralized and depyritized coal, as well as demineralized, depyritized and oxidized coal. Maximum total (71.0 %), pyritic (90.6 %) and organic (49.4 %) sulphur desulphurization effects are achieved for the last sample. Temperature programmed reduction at atmospheric pressure coupled with mass spectrometry (AP-TPR/MS) is used to specify the organic sulphur forms in coal and to assess the changes in organic sulphur that occur as a result of applied treatments. The following organic sulphur types are specified in comparison with model compounds: thiols, organic sulphides, thiophenes, organic sulphonic acids, sulphoxides and sulphones. 1
Oviedo ICCS&T 2011. Extended Abstract
1. Introduction Clean coal utilization is a challenge for higher energy consumption and global ecological impact. Coal is under particular inspection, as during combustion a huge amount of pollutants is produced. Burning of sulphur-containing compounds in fossil fuels evolves sulphur oxides, provoking harmful effects on health, environment and economy. Since there are a variety of chemical, physical and biological desulphurization methods for inorganic sulphur removal, biotreatment received special attention as a potential approach to reduce organic sulphur in mild experimental conditions at lower operating and capital costs. Maritza East lignites are the main Bulgarian deposit with low rank coal. They are characterized by high ash and sulphur contents but a great part of the energy production of the country is based on them. The aim of the present study is to apply a combination of desulphurization methods (chemical and microbial) with an expressed desulphurization potential toward inorganic sulphur and organic sulphur functional forms removal in lignites. In order to increase the total desulphurization effect initial sample is preliminary chemically treated as follow: i) demineralization; ii) demineralization and depyritization; iii) oxidation under heating in air flow. Temperature programmed reduction at atmospheric pressure (AP-TPR) coupled with mass spectrometry (AP-TPR/MS) is used to specify the organic sulphur functional forms in lignites and to evaluate the changes in organic sulphur occurring as a result of applied treatments.
2. Experimental 2.1. Coal samples preparation Bulgarian lignites from Maritza East deposit (Trojanovo-North mine) is selected due to its high inorganic and organic sulphur contents and energy significance. Technological sample of freshly mined lignites is air dried, milled and sieved, < 0.25 mm.
2.2. Chemical treatments
2
Oviedo ICCS&T 2011. Extended Abstract
Demineralization of initial lignite sample is carried out by the Radmacher method – consecutively treatment with 5% HCl, 40 % HF and 35 % HCl at ambient temperature. Subsequently the sample is treated by diluted nitric acid (17 %) for depyritization [1]. Oxidation with atmospheric air is carried out in an electrical oven at 150 oC, for 45 h [2]. Samples placed in thin layer in quartz crucible are oxidized with air flow of 150 ml min-1.
2.3. Microbial desulphurization It is known that Pseudomonas putida (PP) bacteria are effective microbial culture capable to decrease coal organic sulphur [3]. In the present study PP bacterial strain is isolated from soils polluted with crude oil. These microorganisms are grown on Raymond nutrient medium at pH 6.8, temperature 28 oC and 30 days duration. The following products are subjected to biodesulphurization with PP bacteria: initial coal (In), preliminary oxidized coal under atmospheric air and high temperature (In-oxy), demineralized coal (AF), demineralized and oxidized coal (AF-oxy), demineralized and depyritized coal (APF), as well as demineralized, depyritized and oxidized coal (APF-oxy). After biotreatment, "PP" is added to the abbreviation of the corresponding sample. Laboratory scale shake-flask experiments are carried out at pulp density 10% (w/v), at 28 oC for 15 days. The biotreated samples are separated from the media by filtration and consequently washed with 5% HCl solution, and hot distilled water. The dried samples at 105 o
C are stored in inert atmosphere for analyses.
The data characterizing coal samples are given in Table 1.
2.4. AP-TPR/MS technique To specify the organic sulphur forms in coal and to assess the changes in organic sulphur that occur as a result of applied biotreatments, AP-TPR coupled “on-line” with mass-spectrometry (AP-TPR/MS) in different gas media (H2 and He) is used. Apparatus and experimental procedure of the AP–TPR are described elsewhere [4]. In each experiment, approximately 40 mg of sample mixed with 25 mg of fumed silica are placed in the reactor under a 100 ml min-1 flow of pure hydrogen or helium. A linear temperature program of 5 °C min-1 from ambient temperature up to 1025 °C is followed. TPR reactor is coupled “on-line” with a mass 3
Oviedo ICCS&T 2011. Extended Abstract
spectrometer (FISONS-VG Thermolab MS) through a capillary heated at 135 °C. The mass spectrometer equipped with a quadruple analyzer is set at an ionizing voltage of 70 eV. The MS signals m/z 10 ÷ 160 are “on-line” monitored.
3. Results and discussion 3.1. Bulk characteristics Ultimate analysis of the initial lignite sample in Table 1 shows its high organic and inorganic sulphur contents. In order to increase total desulphurization effect and to concentrate efforts on organic sulphur study, initial sample is subjected to chemical (demineralization, depyritization) and thermochemical (oxidation under heating in air flow) treatments prior to biotreatment with PP bacteria. It is known that mild oxidation with atmospheric air under heating is suitable for desulphurization of low rank coal with high pyrite content [2]. This dry oxidative method is based on pyritic sulphur transformation into sulphate sulphur by air oxidation of coal heated before combustion. Advantages of the technique are the lack of liquid phase, absence of expensive oxidative reagents and simplicity. It is necessary to emphasize that the generated sulphatic sulphur is retained by the slag and does not transformed into SOx during combustion. In the present study maximum total (71.0 %), pyritic (90.6 %) and organic (49.4 %) sulphur desulphurization effects are achieved for APF-oxy-PP sample. Total sulphur removal for InPP and In-oxy-PP samples is respectively 28.4 and 39.5 %. There is no appreciable difference in pyritic sulphur desulphurization effect registered for the last two samples, but in the case of In-oxy-PP organic sulphur is more attacked. Concerning demineralized samples, i.e. AF-PP and AF-oxy-PP, total sulphur removal is 43.9 %. Differences in pyritic sulphur and organic sulphur desulphurization for AF-PP and AF-oxy-PP samples are negligible. Table 1. Characteristics of the coal samples. Proximate analysis (%)
S content (%), daf
Desulphurization (%)
Sample W
VM
Cfix
Ash
St
Ss
Sp
So
∆St
∆Ss
∆Sp
∆So
In
8.40 31.30
28.20
32.10
9.48 1.65
3.19
4.64
-
-
-
-
In-PP
7.90 31.90
29.20
31.00
6.79 0.21
2.45
4.13
28.4
87.3
23.2
11.0
In-oxy-PP
5.00 28.50
31.10
35.40
5.74 0.16
2.37
3.21
39.5
90.3
25.7
30.8
4
Oviedo ICCS&T 2011. Extended Abstract AF-PP
5.90 47.10
43.10
3.90
5.32 0.08
2.35
2.89
43.9
95.2
26.3
37.7
AF-oxy-PP
6.20 38.80
49.40
5.60
5.32 0.10
2.30
2.92
43.9
93.9
27.9
37.1
APF-PP
3.40 49.50
44.70
2.40
2.87 0.02
0.31
2.54
69.7
98.8
90.3
45.3
APF-oxy-PP
8.00 35.30
54.90
1.80
2.75 0.10
0.30
2.35
71.0
93.9
90.6
49.4
St – total sulphur; Ss – sulphatic sulphur; Sp – pyritic sulphur; So – organic sulphur
3.2. AP-TPR experiments coupled “on-line” with MS detection 3.2.1. In H2 atmosphere The H2S kinetograms of AP-TPR/MS in H2 atmosphere of samples under consideration are visualized in Fig.1. Only m/z 34 ion profile is shown as m/z 34 (H2S+) and m/z 33 (HS+) demonstrate the same evolution. There are always two dominant peaks in m/z 34 profiles of In, In-PP and In-oxy-PP samples: the first one with Tmax at about 400 °C is attributed to dialkyl and alkyl-aryl sulphides; second peak at about 600 °C refers to the presence of di-aryl sulfides and more complex thiophenic structures. This assumption is based on the model compound approach and also on AP-TPR/MS profiles of typical aliphatic and aromatic fragments for investigated samples [4]. Structures discussed in AP-TPR/MS kinetograms are illustrated in Fig. 2. Despite of high pyrite content of In sample, individual peak corresponding to pyrite presence is not observed. Nevertheless it is noteworthy that the peak with maximum at about 650 °C is slightly asymmetric and starts at about 500 °C. This could be related to pyrite reduction/hydrogenation. In m/z 34 profile of In sample a weak shoulder at 300 °C can be recognized. It is more pronounced in the In-PP sample. This is explained by aliphatic mercaptans hydrogenation. The shoulder under consideration disappears in m/z 34 profiles of In-oxy-PP sample inasmuch as mercaptans are oxidized. Since the shoulder for thiols hydrogenation is observed in m/z 34 profile of In-PP sample, the oxidation of thiols is rather due to preliminary oxidation with atmospheric air than to biotreatment. It is known that thiols can be oxidized to disulphides under mild air oxidation, while under more severe condition, they can be oxidized to sulphonates [5]. Two peaks are observed in AP-TPR/MS m/z 34 kinetograms of AF-PP and AF-oxy-PP samples. The peaks are attributed in the same way as the peak assignment in m/z 34 profile for In sample. Since a well shaped shoulder (in the first peak) is observed at 300 ºC in m/z 34 profile of AF-PP sample, the aliphatic mercaptans presence is more pronounced. The 5
Oviedo ICCS&T 2011. Extended Abstract
emergence of this emphatic shoulder is explained by the reduced content of mineral components. It is proven that more reliable information with AP-TPR/MS can be achieved for demineralized samples. It is because mineral components (such as siderite) capture H2S and obtained H2S profile is less detailed [6]. The shoulder at 300 ºC again disappears in m/z 34 profile of AF-oxy-PP sample. The presence of pyrite influences the second peak in different manner: for AF-PP an asymmetric profile is generated towards the lower temperature region; for AF-oxy-PP a well expressed shoulder is recognized around 550 °C.
In In-PP In-oxy-PP
Intensity (mbar)
3.00E-012
2.00E-012
1.00E-012
0.00E+000
100
200
300
400
500
600
700
800
900
1000
Temperature (°C)
AF-PP AF-oxy-PP
Intensity (mbar)
3.00E-012
2.00E-012
1.00E-012
0.00E+000
100
200
300
400
500
600
700
800
900
1000
Temperature (°C)
6
Oviedo ICCS&T 2011. Extended Abstract
APF-PP APF-oxy-PP
Intensity (mbar)
1.20E-012
8.00E-013
4.00E-013
0.00E+000
100
200
300
400
500
600
700
800
900
1000
Temperature (°C)
Figure 1. AP-TPR/MS (H2), m/z 34 kinetograms for investigated coal samples. O R1,2
Thiols
SH
R1,2
S
O
Sulphonic acids
OH
R1,2
S
R1,2 Sulfides R1,2 S R1
S
O
R1,2
Thiophenes R1,2
O S
R1 - alkyl R2 - aryl
Sulphoxides
R1,2
Sulphones
O
Figure 2. Structural formulae of organic sulphur functional forms discussed in the text.
A broad signal (in the range of 200-800 ºC) with indications for apexes at 300 ºC, 400 ºC, 550 ºC and 650-700 ºC can be seen in AP-TPR/MS m/z 34 profile of APF-PP. No well formed peaks can be observed in this profile. This is probably due to low organic sulphur content and the modified coal matrix. Nevertheless interpretation of the apexes at 300 ºC, 400 ºC and 650700 ºC can be attributed in the same way as the interpretation of H2S profile of In sample. The apex at 550 ºC is probably related to hydrogenation of “new” organic sulphur species formed during depyritization and biodesulphurization of ash pyrite free coal. Other explanation is that some organic sulphur compounds could be more visible due to changes in the coal matrix as a result of the treatments. In m/z 34 profile of APF-oxy-PP again a broad 7
Oviedo ICCS&T 2011. Extended Abstract
signal is observed. It starts at higher temperature (~300 ºC) compared to APF-PP because of thiols oxidation. The maximum of the signal is at 550 ºC. Similar to APF-PP it is attributed to hydrogenation of new formed organic sulphur species of oxidative origin which are more dominating. 3.2.2. In He atmosphere AP-TPR pyrolysis in inert atmosphere is performed to receive more information for the nature of oxidized sulphur species. Fig. 3 gives SO2 evolution in He for samples under consideration. In it, only m/z 64 profiles are shown since they exhibit the same evolution pattern as m/z 48 profiles. Two peaks appear in m/z 64 kinetograms of In, In-PP and In-oxy-PP samples. The first broad peak is registered at lower temperature up to 400 ºC. Based on model compounds approach and also on AP-TPR/MS (He) profiles of typical aliphatic and aromatic fragments it can be attributed to organic sulphonic acids. The second peak maximizing at 490 ºC can be assigned to iron sulphates, sulphoxides and sulphones. Presently to distinguish last three types of sulphur containing compounds additional research is needed. Inasmuch as In-PP and Inoxy-PP samples have low content of sulphatic sulphur (see Table 1), the second peak could be mainly referred to sulphoxides and sulphones. It should be noted that in the case of In-oxyPP sample, the intensity of the second peak increases compared to the first one, confirming the increased presence of oxidized organic sulphur compounds. In AP-TPR/MS (He), m/z 64 profiles of AF-PP and AF-oxy-PP samples again two peaks are registered. These peaks are interpreted in the same way as the assignment of m/z 64 profiles for In-PP and In-oxy-PP samples. Again, due to the chemical treatment, the coal matrix is altered and thus the shape of the registered profiles is influenced. For the APF-PP profile a further shift towards lower temperature compared to the AF-PP profile is noticed, due to the additional chemical treatment. Further oxidation did occurred for APF-oxy-PP sample, resulting in the removal of already oxidized organic sulphur compounds and thus a clear change in the obtained profile compared with the one for APF-PP is registered.
4. Conclusions A combination of chemical and microbial desulphurization methods are applied for inorganic and organic sulphur removal in low rank coal, Bulgarian lignites. Maximum total (71.0 %), pyritic (90.6 %) and organic (49.4 %) sulphur desulphurization effects are achieved for 8
Oviedo ICCS&T 2011. Extended Abstract
preliminary chemically treated (demineralized, depyritized and oxidized) biodesulphurized sample APF-oxy-PP. Qualitative specification of a broad range sulphur and oxygen-sulphur functional forms is done by using AP-TPR/MS technique. Some changes are depicted as well: thiols disappear in all oxidized and biotreated samples; new formed sulphur containing compounds are observed after some treatments. Semi-quantitative appraisal of current desulphurized coals will be a target of future study.
1,20E-011
Intensity (mbar)
In In-PP In-oxy-PP
8,00E-012
4,00E-012
0,00E+000 100
200
300
400
500
600
700
800
900
1000
Temperature (°C)
3.00E-012
Intensity (mbar)
AF-PP AF-oxy-PP
2.00E-012
1.00E-012
0.00E+000 100
200
300
400
500
600
700
800
900
1000
Temperature (°C)
9
Oviedo ICCS&T 2011. Extended Abstract
6.00E-012
Intensity (mbar)
APF-PP APF-oxy-PP
4.00E-012
2.00E-012
0.00E+000 100
200
300
400
500
600
700
800
900
1000
Temperature (°C)
Figure 3. AP-TPR/MS (He), m/z 64 kinetograms for investigated samples.
Acknowledgements The study was performed within the framework of Cooperation agreement for joint supervision and award of a doctorate between Hasselt University, Belgium and Bulgarian Academy of Sciences, BAS-Bulgarian bilateral project with Hasselt University, FWOFlanders.
References [1] Marinov SP, Stefanova M, Stamenova V, Carleer R, Yperman. Sulphur functionality study of stream pyrolyzed “Maquinenza” lignite using reductive pyrolysis technique coupled with MS and GC/MS detection systems. Fuel Process Technol 2005;86:523-34. [2] Rustchev D, Bekyarova E, Patent №36265 MPK CL 10 9/06 (1983). [3] Rai,C. & Reyniers, J.P., Biotechnol Prog 1988;4: 225-30. [4] Mullens S, Yperman J, Reggers G, Carleer R, Buchanan III AC, Britt PF, et al. A study of the reductive pyrolysis behaviour of sulphur model compounds. J Anal Appl Pyrolysis 2003;70:469-91. [5] Tsai CS. Fundamentals of Coal Beneficiation and Utilizations. Coal Science and Technology 2, Elsevier, Amsterdam, 1982. 10
Oviedo ICCS&T 2011. Extended Abstract
[6] Maes II, Gryglewicz G, Yperman J, Franco DV, Haes JD, D`Olieslaeger M, et all. Effect of siderite in coal on reductive pyrolytic analyses. Fuel 2000;79:1873-81.
11
Oviedo ICCS&T 2011. Extended Abstract
Rank-dependent formation enthalpy of coal
M. Sciazko Institute for Chemical Processing of Coal, Zabrze, Poland,
[email protected]
Abstract The final state of the geochemical transformation of plant debris is coal, which is a complex organic matter contaminated with dispersed mineral substances. These chemical changes led to the formation of a specific coal structure, dependent mainly on original substance composition and temperature-pressure history of its transformation. Coal organic matter is not described as a defined chemical compound but rather as a model form is used depending on the rank of coal investigated. Thus, coal may be characterized by a defined thermodynamic state at standard conditions of pressure and temperature. That quantity is called enthalpy of formation and is easily determined for pure species by using the proper thermodynamic tables or measuring suitable heats of reactions. This is not the case for coal due to the complexity of its composition. Considering the particular chemical structure of coal as a result of a specific reaction, one can assume that it may be also characterized by a defined thermodynamic state such as enthalpy of formation. In this work, a method to calculate enthalpy of formation for coal was developed over the entire rank range. Here, enthalpy of formation for a complex chemical compound was defined as the difference between the experimentally determined heat of combustion and thermodynamically calculated heat of combustion reaction of elementary substrates. A boundary condition for that approach is defined by the value of enthalpy of formation of graphite. The above-described method should produce a value of zero for graphite. Using the correlation for enthalpy of formation developed, this manuscript proposes a model of coal classification via a thermodynamic quantity reflecting its structure and technological suitability. Based on the analysis of enthalpy of formation with respect to coal composition, enthalpy of formation may have both negative and positive values, depending on the type of fuel. Furthermore, a change in formation enthalpy is continuous but corresponds to different chemical structures of coals.
1
Oviedo ICCS&T 2011. Extended Abstract
Comparing previous coal classification data with their corresponding values for enthalpy of formation suggests that the enthalpy of formation can be used to classify solid fuels. Various types of fuels are characterized, and the following values for enthalpy of formation are reported: anthracite (+250 kJ/kg), peat (< -3200 kJ/kg) and medium volatile bituminous coal ca. zero.
1. Introduction Established classification systems guide the interpretation of scientific research results and allow for the comparisons of operating conditions and coal selection for chemical processing. However, the analysis of such classification and the resulting parameters indicate that the system is inconsistent. In addition, when classifying coal types for processing in the chemical and power industries, the technological needs are not directly reflected, and the focus is rather on operational parameters such as caking ability, ash content or caloric content [1], [2]. Analysis of existing classification systems suggests that they are not related to the thermodynamics that characterize the work potential. This paper identifies a key thermodynamic parameter (enthalpy of formation) to characterize coals and relates it to coal properties. The enthalpy of formation represents the energy resulting from the reactions that created coal. For pure substances, the enthalpy of formation is a single defined value. Due to the various chemical structures of coal, however, the enthalpy of formation will depend on the rank of coal and on the structure specific for a given type of coal. The enthalpy of formation may thus characterize the coal with an accompanying classification system. To maintain consistency with chemical thermodynamics, a nascent model should consider that coals of increasingly high metamorphism approach the structure of graphite, in which the heat of calorimetric combustion and the thermodynamic enthalpy of combustion reaction are equal.
2. Experimental section The objective of this study is to determine the influence of coal (described by elemental composition) on the change of the enthalpy of its formation. This assumes that the enthalpy of formation for coal, which is a complex fuel, will be defined as the difference between the measured heat of combustion and thermodynamically calculated heat of combustion of coal elements. The difference between the measured heat of combustion 2
Oviedo ICCS&T 2011. Extended Abstract
and the calculated thermal effect of combustion reaction is based on 224 sets of data describing properties of Russian [3], American [4] and Polish [5] coals and chars/cokes. For computational and comparative purposes, all the elemental composition data have been converted into a dry and ash-free condition; the heat of combustion was treated similarly. The oxygen content is obtained by subtracting the determined contents from the other elements, where sulphur exists as combustible sulphur. The database comprised a very wide spectrum of fuel properties, and their ranges are presented in Table 1. Table 1. The range of elemental composition and of heat of combustion values. Parameter
Min. value
Max. value
C
69.87
91.26
H
2.61
6.66
O
2.08
23.08
N
0.07
2.53
S
0.39
9.37
Qsdaf, MJ/kg
27.27
36.98
Content, % (daf)
3. Results and Discussion Assuming that the enthalpy of coal formation is equal to the difference between enthalpy of combustion reaction of coal component elements in the standard state and the determined heat of combustion [6], the consistency of both methods for the determination of fuel calories should lead to the same result. In the case of elemental carbon represented by graphite, the standard enthalpy of formation is zero. Unfortunately, all known relationships for the calculation of the heat of combustion do not have such a feature [7]. Calculating the enthalpy of combustion reaction of coal as a dry and ash-free acc. to the general relationship (1) we have: ∆c H
0 , daf
C daf H daf S daf 0 0 0 = ∆ f H CO2 + ∆ f H H 2O ( l ) + ∆ f H SO 2 MC 2M H MS
(1).
Here, the symbols of elements stand for their weight fraction in coal organic matter divided by their atomic mass, which are multiplied by the appropriate molar enthalpies of formation for the individual products of element combustion as adopted from thermodynamic tables.
3
Oviedo ICCS&T 2011. Extended Abstract
After introducing the value of standard enthalpy of suitable components formation and assuming that the element fraction is a percentage, the equation becomes the following: ∆ c H 0 = −327,633C daf − 1417,892H daf − 92,768S daf kJ/kg(2) c Taking into consideration the variation between the values of enthalpy of combustion reaction and the calorimetric heat of combustion, it is critical to determine the key differences between them. Based on elemental analysis, the effect of hydrogen and oxygen on the heat of combustion of coal suggests that the latter has a stronger and more consistent effect. Therefore, applying the definition of the enthalpy of coal formation expressed by equation (3), it is necessary to correlate the thermodynamic heat of combustion with the enthalpy of combustion reaction. Analyzing equation (3), the convention is that the heat of combustion is always positive and the enthalpy of combustion reaction is negative due to an exothermic effect. ∆ f H 0 = Qs + ∆ c H 0
(3)
In general, the relationship between the heat of combustion and calorimetric heat of combustion may have the following form: Qs = (−1)∆ c H 0 f (θ)
(4).
Here, f(θ) is a function that modifies the enthalpy of combustion reaction, defining the ratio of absolute values of the heat of combustion and the enthalpy of combustion reaction. This is a normalized heat of combustion. Based on the elemental analysis, the influence on a change of the heat of combustion, an assumption was made that the correcting function f(θ) is defined by the oxygen content in coal (Oddaf ), expressed in wt.%. Analogous to an ideal gas virial equation of state, in the first approach, the following form of the correcting function was assumed:
f (θ) = 1 + a1θ + a2θ 2 + a3θ 3 + a4θ 4 ...
(5).
It is necessary to consider that the enthalpy of combustion reaction has a sign opposite to the thermodynamic heat of combustion, and thus, the factor of -1 is added in (4). The experimental data lead to the results presented in Table 2.
Table 2. Coefficients in correcting equation (5). Coefficient Value
a1 4.852·10
a2 -3
-1.437·10
A3 -3
8.467·10
A4 -5
-1.66·10-6
4
Oviedo ICCS&T 2011. Extended Abstract
The heat of combustion calculated this way has a mean error of +/-502 kJ (1.5%). Using a Student’s t test, the mean was x = 1.000, and the standard deviation s(x)=0.015. The correctness of assuming the zero hypothesis H0:x = 1 has been confirmed at the significance level of 0.05 by the result t = 0.128. The equation should be used with fourth-degree virial terms because the accuracy of the heat of combustion is computed in relation to the enthalpy of combustion reaction at an oxygen content below 5%, in the area of the enthalpy of formation sign change (Fig. 1). The correcting function was derived for the (daf) coal condition, and for this, reason it is valid only for this condition. 1,04
experimental data calculated data
Normalized heat of combustion
1,02 1
0,98 0,96 0,94 0,92 0,9
0,88 0
5
10
15 Oxygen content, %
20
25
30
Fig. 1. Normalized heat of combustion vs. oxygen content. As a result of the model, the enthalpy of any coal formation ∆fH0w is described by equation (6) resulting from the combination of equations (3) and (4). ∆ f H w0 = ∆ c H w0 [1 − f (θ)]
(6)
Because the enthalpy of formation values in the positive area are lower than 200 kJ/kg (with a mean correlation error for the heat of combustion of 500 kJ/kg), the Student’s t test offers additional credence to the results. All results of the (Qs/∆cH0) ratio that are higher than one were analyzed. Overall, this comprised 40 data points corresponding to anthracites, cokes and chars produced by coal pyrolysis. For this population, the mean was 1.004. Assuming the level of significance is 0.05 and checking the hypotheses H0:x =µ, H0:x<µ and H0:x>µ consecutively for values of µ <1, 1 and >1, respectively, the results support the following hypothesis: the positive values do not result from a measurement or calculation error but are a physical feature of coal or its solid products. These results were also confirmed by bond enthalpy analysis in model coals [8].
5
Oviedo ICCS&T 2011. Extended Abstract
4. Coal classification This model for enthalpy of coal formation calculation has a generalized nature because it unifies calorimetrically the determined heat of combustion and thermodynamic heat of coal combustion. Because of the continuity of coal composition changes and the presence of all individual element contents in the range studied, the relationships derived allow explicit determination of the enthalpy of formation for a given coal. In addition, the standard enthalpy of coal formation is determined for the reference state of temperature and pressure, which allows for comparisons of the results to the enthalpies of formation of other chemical compounds tabulated in thermodynamic handbooks. In turn, their tabulation enables comparison and ranking of individual coals. Calculations of the enthalpy of formation conducted for reference fuels adopted in this paper according to Jasienko [9] with the use of (6) lead to the results presented in Fig. 2. These are in line with the American classification (Fig. 2). The following information results: characteristic ranges of the enthalpy of formation may be distinguished depending on the solid fuel type; and range differentiation allows for the assumption that the enthalpy of formation may be a classification parameter for solid fuels. Considering the above, individual fuel types feature the following values for the enthalpy of formation according to the Polish classification: •
anthracite: 250 kJ/kg,
•
semi-coking coal, type 37: 125 kJ/kg,
•
ortho-coking coal, type 35: 0 kJ/kg,
•
gas coal, type 33: -600 kJ/kg,
•
flame coal, type 31: -1300 kJ/kg,
•
brown coal: -2500 kJ/kg,
•
peat: below -3200 kJ/kg.
The above specification is a type of thermodynamic classification of coals and chars, i.e., pyrolysis products. Chars and cokes usually contain less than 2% oxygen, which is less than that contained in anthracites. Thus, an assumption may be made that they constitute a group of fuels for which the enthalpy of formation is lower than 250 kJ/kg and higher than zero.
6
Oviedo ICCS&T 2011. Extended Abstract 500
Enthalpy of formation, kJ/kg
Lignite
Subb. B
HVB C
HVB B
HVB A
0 MVB
LVB
Anthracite
-500
-1000
-1500
-2000
-2500 Type of coal
Fig.2. Thermodynamic classification of solid fuels referred to the American Standards classification. The heat effect of coking coal formation, in which the enthalpy of formation is close to zero, can explain their ability to form good coke. The process of coke formation in general is not accompanied by any internal thermal effect. This is an important indication in the preparation of a blend for coal coking. A positive influence of the lack of thermal effect may be due to the fact that the coke formation basically starts in the coal plastic phase. Both the endothermic and exothermic effects are harmful to the formation of a homogeneous coke structure because of the shortening of the plastic phase. At the exothermic effect, the pyrolysis and volatiles evolution are more sudden. The residence time of coal shortens the plastic phase. In the case of the endothermic effect, the plastic layer is internally cooled, and its duration is reduced. In the coking conditions, this time translates to diminished thickness of the plastic layer. Concurrently, it is necessary to bear in mind that the thermal effect of the conversion is opposite to the calculated enthalpy of coal formation. When the enthalpy of formation has a negative value, this means that in the process of coal organic matter formation, an equivalent amount of heat is released as a result of formation of bonds between oxygen and hydrogen. In the case of anthracite coals, a positive result means that a supply of heat was necessary for their formation. An equivalent amount of heat will be released in their decomposition.
5. Conclusions Analysis of existing classification systems suggests that they fail to consider crucial issues in the reaction thermodynamics of coal evaluation. This paper deals with the selection of a key thermodynamic parameter, enthalpy of formation, to characterize coal and to relate it to coal properties. Enthalpy of formation was evaluated for the entire range of solid fuels, and a correlation proposed. The value of enthalpy of formation 7
Oviedo ICCS&T 2011. Extended Abstract
indicates the state of coal metamorphism and is thus directly related to coal structure and bond characteristics. Enthalpy of formation can be used for coal classification and is also necessary for the energy balance of the coal conversion systems.
References [1] Speight JG: Handbook of Coal Analysis. Jon Wiley&Sons, Inc., Hoboken. New
Jersey; 2005. [2] International codification system of hard coals by type; 1956
http://www.unece.org/energy/se/pdfs/coal6/coedhard.pdf. [3] Gagarin SG, Gladun TG. Evaluation of enthalpy of formation of coal organic matter
(in Russian). Chemia Tvierdovo Topliwa 2003;4:3-23. [4] Coal Conversion Systems Technical Data Book, section IA.50.1; 1982, vol.2. [5] Sciazko M, et al. Balancing of coking process. IChPW Report 2010;30. [6] Annamalai K, Puri IK. Combustion science and engineering. CRC Press, Taylor & Francis Group, Boca Raton, FL; 2005, p.169. [7] Channiwala SA, Parikh PP. A unified correlation for estimating HHV of solid, liquid and gaseous fuels. Fuel 2002;81:1051-1063. [8] Sciazko M. Models of coal classification in terms of thermodynamics and kinetics (in Polish). Krakow: AGH; 2010. [9] Jasienko S. Chemistry and physics of coal (in Polish). Wroclaw: University Press.; 1995.
8
Oviedo ICCS&T 2011. Extended Abstract
Ion beam tomography for coal characterization *
A. Bhargava1, P.J. Masset1, N. Gordillo2, C. Habchi2, P. Moretto2
1
TU Bergakademie Freiberg, Centre for Innovation and Competence – Virtuhcon, Freiberg -Germany, D-09599 2
Université Bordeaux 1, CNRS/IN2P3, Centre d’Etude Nucleaires de Bordeaux Gradignan, CENBG, Chemin du Solarium , BP 120, 33175 Gradignan, France Abstract 3D Scanning Transmission Ion Microscopy-Tomography (STIM-T) of German brown coal was carried out using a proton beam as a nuclear probe at the energy of 2.8 MeV at the nanobeam line of CENBG. Each scan produced series of images which were later stacked together numerically using computer algorithms to produce virtual cut sections across the scanned volume, revealing the internal structure. The STIM-T images were obtained with a spatial resolution in the sub-micrometer range. The analysis showed local density variations due to the presence of pores, minerals, cleats and fractures. The colour image analysis enabled the visualization of the internal microstructure of coal showing fair contrast among fractures and pores. These experiments were complemented by 2D Particle Induced X-ray Emission-Tomography (PIXE-T) for rapid non destructive chemical analysis showing spatial distribution of mineral matter in the selected coal slices. All experiments were done in vacuum (~ 5x10-6 mbar) to maintain a high degree of sensitivity for light elements.
Keywords: Coal characterization, Ion beam analysis, PIXE and STIM tomography
1. Introduction The world has witnessed rapid consumption of fossil fuels over last several decades to mitigate global energy demands. Among major fossil fuels, coal continues to hold a dominant share in the world’s energy spectrum. There are an estimated 275 billion tonnes of coal reserves in USA alone as compared to 4 trillion tonnes of coal resources [1]. Because of high abundance of coal reserves in many countries of the world, and enormous plant replacement costs as compared to other alternatives, coal fired power plants are here to stay for long time. Amidst rising energy demands and increasing oil prices, clean coal *Corresponding author:
[email protected]
Oviedo ICCS&T 2011. Extended Abstract
technologies have gained significant attention. Of particular importance is, Integrated Gasification Combined Cycle (IGCC) power plants that have attracted considerable interest to provide clean energy from coal. The present challenge before the power generation industry is to meet emission norms and to reduce green house gases by CO2 sequestration methods. The heart of such a process unit lies in the gasifier reactor where multiphase reactions between coal (solid phase), oxygen (gas phase) and steam take place at elevated pressure (up-to 45 bars) and temperature (up-to 1500 ºC) under partial oxidation conditions. In last several decades a number of experimental, modelling and simulation studies have been reporting these dynamic heat and mass transfer processes at the particle level [2] [3]. However, few of them have been able to show the physiochemical aspects of gas-solid reactions at the reacting inter-phase through experiments. Previous attempts using X-ray computed tomography (X-ray-CT) were made by Mathews [5] and Yao et.al [6]. Also, Van Geet et. al. [7] used similar techniques with SEM –EDX to validate the presence of minerals like liptinite and vitrinite. The scope of this work aims to track and to explore the physio-chemical properties of only solid phase (coal) involved in the reaction, using ion beam technologies such as particle induced X-ray emission (PIXE) and scanning transmission ion microscopy-tomography (STIM-T). As far as we know, few attempts were made in early 1990s with ion beam techniques for coal and coal ash analysis [4]. Like, X-ray-CT, the STIM-T is also a non destructive method which allows the 3D sample visualization with a sub-micrometric lateral resolution [8]. Hence, very accurate spatial positioning and material characterization analysis is possible to ascertain coal matrix slice composition. 2. Experimental methods
2.1
CENBG nanobeam line and the analysis chamber
All experiments were performed at the nano-beamline of the AIFRA (Applilcation Interdisciplinaire des Faisceaux d’Ions en Aquitaine) facility at the CENBG using ultra – stable single–ended HVEE 3.5 MV SingletronTM accelerator. A proton beam at 2.8 MeV was used as a probe for all the analysis. The beam line attains a resolution of 300 nm in the STIM mode with a current of few thousands of counts per second and about 1μm when working in PIXE configuration with a current of few pA upto hundreds of pA. Fig. 1 shows the overview of the experimental setup.
Oviedo ICCS&T 2011. Extended Abstract
a)
b)
c)
FC STIM
RBS
microscopes goniometer
Fig. 1. a) The analysis chamber. b) The goniometer, microscopes and detectors. The red line indicates the beam direction. c) The top-view of the mounted sample on the goniometer.
The target chamber is equipped with three microscopes, one viewing the incoming beam (front side) and the others viewing the sample (backside). The sample holder is driven by a X, Y, Z motorized stage (NewportTM) allowing careful sample positioning in three directions. Additionally, a goniometer is provided to enable rotational motion of the sample. Detailed description of the experimental facility is beyond the scope of this paper and indicated references may be consulted [8], [11].
2.2
Sample Preparation
For this work two types of samples were chosen. Sample A grains were German lignite coal particles obtained from feed used in the local coal fired power plant. While, sample B were coal char particles, prepared by heating sample A to 450 ºC in the presence of nitrogen to remove the surface moisture. Sample B was then diluted in formvar (liquid polymer) solution and later, wires were stretched out of this solution and dried in air during one day. The size of these particles was about 25 μm. These formvar wires containing coal particles were inserted into a glass capillary of diameter ~150 μm and further inserted into a needle of ~400 μm. The whole assembly was mounted on the goniometer sample holder. Similarly, sample A of about 120 µm was glued using AralditeTM on the tip of the capillary glass of about 200 µm and mounted on the goniometer in the similar way.
Oviedo ICCS&T 2011. Extended Abstract
2.3
Scan settings
For 3D-STIM-T experiments a rectangular scan of 160 x 80 µm2 with a resolution of 1.25 µm/pixel was chosen for sample A, meanwhile the area of 100 x 50 µm2 with a resolution of 0.79µm/pixel was selected for sample B. Scan rate was about ~70 scans/projection for both samples and, a total of 100 projections were recorded in 180 degrees. For PIXE-T analysis two slices of 160 µm in width were selected from sample A, wherein the two large grains were including and, one slice in the case of sample B of 100 µm. This, allowed us to study the chemical distribution in the coals. The scan rate was 3000 scans per projection. And, the scan speed during the analysis was 100 μs/pixel for PIXE-T experiments and 200 μs/pixel for STIM-T experiments. Fig. 2 shows the sample images obtained by the camera and the microscope after mounting on the goniometer. a)
b)
Analyzed region
Analyzed region
50 µm
Optical Image
100 µm Sample A
Optical Image
Sample B
Fig. 2. a) A coal grain of about 200 μm was glued using AralditeTM on top of the tip of a glass tube of diameter 200 μm, which was further inserted in the needle of diameter 400 μm. b) Sample B: The analytical coal grains of about 20-30 μm were included in a formvar capillary (~150 μm diameter) which was then inserted into a glass tube of 200 μm and further into the needle of 400 μm.
After all the adjustments were made and the precise beam spot size on the sample was obtained, H+ beam as a microprobe was used to scan the chosen areas. The transmitted ions were collected in passivated implanted silicon lithium detector (Canberra PIPS detector. 25mm2, 12 keV energy resolution). The X-rays emitted from the sample were detected using a Si-Li X-Ray detector (Sirius/80/Be/PIXE, 140 eV energy resolution) placed at 45º backwards from the ion direction and at 22 mm distance from the sample.
Oviedo ICCS&T 2011. Extended Abstract
3. Results
3D – STIM Reconstruction and PIXE analysis
3.1
Series of 2D projections were stacked to create the complete 3D volume matrix. Tomorebuild code, developed at CENBG, was used to convert this data (energy sinogram) into 3D images using a filtered back projection algorithm [9]. These energy sinograms were reconstructed into 3D images as shown in Fig. 3 a) and Fig. 4 a) for sample A and B, respectively.
a)
3D STIM-T reconstruction: Sample A STIM Slice n° 4
Selected slice for PIXE-T
3 2 1
10 µm
b)
STIM slice n°:
10 µm
1
10 µm
2
3
4
max
g/cm3
10 µm
c)
0
STIM (black and white) + PIXE (color) overlapped
1.20
STIM slice
1
Ca
0.14
0.021
Mn
Fe
0.43
Ti
g/cm3
Ratio Z/Ca
50 µm 0
50 µm
50 µm 0
50 µm 0
50 µm 0
0
Fig. 3. a) 3D-STIM -T reconstruction of sample A where different cross sections are indicated as STIM-T slices and also the selected PIXE-T slice for the chemical is shown. b) Corresponding cross section from the selected STIM-T slices. c) PIXE-T slice (colour) overlapped with the STIM-T slice (black and white) in order to show the element distribution. Colour scale bar are normalized to the Ca element.
The colour contrast shows the local density variations in the 3D image post reconstruction. Several lateral planes show the virtual slicing of the selected grain. The qualitative nature of the various pores and fractures is visible from each cut away section using colour
Oviedo ICCS&T 2011. Extended Abstract
contrast, also shown in Fig. 3 (b). The darker regions show areas having lower densities possibly related with the fractures and pores, while the lighter areas show higher density indicating presence of mineral matter, such as calcium based compounds and other elements as shown in Fig. 3. (c). In this figure, PIXE-T slices are overlapped with the STIM-T slices. All gas – solid reactions take place on the surface and in the interior of these particles and these pores and fractures often allow the channeling effect of the reactant gas during the mass transport (diffusion) within the coal particle. The average density calculated for the coal grain (sample A) was 1.35 g/cm3, which is in fair agreement with the values published by Bell et. al [1].
a)
3D STIM-T reconstruction: Sample B Selected slice for PIXE-T Selected slice for n° 1 PIXE-T
n° 2 20 µm
20 µm
STIM (black and white) + PIXE (color) overlapped: slice n° 1
b)
1.17
0.15
Fe
0.57
S
1.00
Ca
g/cm 3
Ratio Z/Ca
20 µm
20 µm
0
20 µm 0
20 µm 0
0
STIM (black and white) + PIXE (color) overlapped: slice n° 2
c) 1.18
0.09
Fe
0.60
S
1.00
Ca
g/cm 3
Ratio Z/Ca
20 µm 0
20 µm
20 µm 0
20 µm 0
0
Fig. 4. a) 3D-STIM-T reconstruction of sample B where the named coal grains are represented in green and the two PIXE-T slices selected for the analysis are shown. b) STIM-T (black and white) and PIXE-T (colour) slice overlapped for the slice nº1. b) STIM-T (black and white) and PIXE-T (colour) slice overlapped for the slice nº2. The red circles are to stand out the high density of the two big grains chosen. Colour scale bar are normalized to the Ca element.
The 3D reconstruction corresponding to coal char (sample B) is shown in Fig. 4 a). The small green particles may be related with the various impurities (included in the formvar solution) as well as crushed grains. In this case the average density of the grains plus
Oviedo ICCS&T 2011. Extended Abstract
formvar solution was 1.08 g/cm3. The reason for deviation from these values is due to low density of both grains and formvar which, makes it difficult to have a better image contrast. Also, shown in Fig.
4 b) and Fig.
4 c) are series of images showing a
combination of STIM-T and PIXE-T results displaying elemental distribution with the density contrast, in the selected slices. The higher density material is indicated by the light colour areas (in STIM-T slices) showing mineral matter or hard fractures. The chemical element distribution obtained with the PIXE-T analysis verifies the presence of the grains when they are overlapped with the STIM-T slices. 4. Discussion In view of the heterogeneous composition of the coal matrix, several samples and experiments are required to ascertain the effect of randomly distributed mineral matter in mass transport of the reacting species in the solid phase. Since, the mineral matter is nonporous; it acts as a hindrance in the transportation of diffusing chemical species to the active carbon sites. As the temperature increases the solid phase gets consumed vis a vis bulk gaseous reactants. The inherent mineral matter acts as a barrier in the diffusion of chemically reactive species. It often, alters the trajectory of the diffusing gaseous specie (oxygen/carbon dioxide), delaying their availability at the reacting inter-phase resulting in incomplete reaction, or sluggish kinetics, also termed as diffusion limitation. Thus considering rather a practical case, in which the reaction between coal and a reactive gas is controlled by not only by a single transport phenomenon but rather a mix of several phenomenon (such as adsorption, chemical reaction and desorption). The spatial location of the mineral matter may significantly affect the kinetics of high temperature reaction (gasification/pyrolysis). Several existing models do not incorporate the effect of these mineral matter constituents in heat and mass transport due to their random presence in the coal matrix. In fact many of them make an assumption of non – existence of any hindrance to such pathways inside coal particle to maintain the simplicity of the mathematical nature of the model, overlooking the impact it could have on gasification kinetics. The authors do believe that the complexity of such a model may increase upon incorporating such assumptions; however it will certainly give illuminating reasons for lower conversions caused due to diffusion limitation at elevated temperatures. Thus, such effects should not be neglected and due attention should be paid to incorporate inhomogeneous nature of coal.
Oviedo ICCS&T 2011. Extended Abstract
5. Conclusions
The 3D-STIM –T and PIXE-T experiments provide accurate elemental maps of the analysed coal samples showing various components. Also, examined and corroborated with XRD experiments, in previous study [12]. The ion beam techniques can be used to characterize the local cleat, fractures and density variations along with the elemental mapping in the solid reactant phase. The results show the spatial distribution of mineral matter in the analysed coal slice. The images obtained using this technique will provide an additional experimental validation tool for several gas solid heterogeneous multiphase reactions in which the reaction rate is controlled by the processes occurring in the solid phase. Also, it is to be noted is that due to experimental limitation extending the scalability of such a technique from particle to bulk sample is questionable. This will also be discussed in the forth-coming publications, wherein quantification of various physical parameters controlling the mass transfer phenomenon in coal matrix is likely to be presented. In view of the present study, it is proposed that if the results of such experiments are accompanied with high temperature in-situ analytical techniques such as those with scanning confocal laser microscopy and hot stage microscopy it would be prove to be an important step in the validation of tracking inter-phase gas solid reactions.
Acknowledgements
The authors are grateful for the support provided by the European Communities as an integrating activity for the “Support of Public and Industrial Research Using Ion Beam Technologies (SPIRIT)” under transnational access activities. Also, the authors thank the BMBF (Fedral ministry of higher education, Government of Germany) to provide financial support under the framework of project Virtuhcon.
References [1] David A. Bell. (First edition 2011). “Coal Gasification and its Applications” Elsiever [2] Sadhukhan, A. K., P. Gupta, et al. (2010). "Modelling of combustion characteristics of high ash coal char particles at high pressure: Shrinking reactive core model." Fuel 89(1): 162-169.
Oviedo ICCS&T 2011. Extended Abstract
[3] Zhu, W. C., C. H. Wei, et al. (2011). "A model of coal-gas interaction under variable temperatures." International Journal of Coal Geology 86(2-3): 213-221. [4] Wang, X., X. Shen, et al. (1993). "PIXE study on effects of coal burning in a coal-fired power station on atmospheric environmental pollution." Nuclear Instruments and Methods in Physics Research Section B: Beam Interactions with Materials and Atoms 75(1-4): 273-276. [5] Mathews, J. P., J. D. N. Pone, et al. (2011) “High-resolution X-ray computed tomography observations of the thermal drying of lump-sized subbituminous coal.” Fuel Processing Technology 92(1): 58-64. [6] Yao, Y., D. Liu, et al. (2009). “Non-destructive characterization of coal samples from China using microfocus X-ray computed tomography.” International Journal of Coal Geology 80(2): 113-123. [7] Van Geet, M., R. Swennen, et al. (2001). “Quantitative coal characterization by means of microfocus X-ray computer tomography, colour image analysis and back-scattered scanning electron microscopy.” International Journal of Coal Geology 46(1): 11-25. [8] Barberet, P., S. Incerti, et al. (2009). "Technical description of the CENBG nanobeam line." Nuclear Instruments and Methods in Physics Research Section B: Beam Interactions with Materials and Atoms 267(12-13): 2003-2007. [9] Michelet-Habchi, C., S. Incerti, et al. (2005). "TomoRebuild: a new data reduction software package for scanning transmission ion microscopy tomography." Nuclear Instruments and Methods in Physics Research Section B: Beam Interactions with Materials and Atoms 231(1-4): 142-148. [10] Barberet, P., L. Daudin, et al. "First results obtained using the CENBG nanobeam line: Performances and applications." Nuclear Instruments and Methods in Physics Research Section B: Beam Interactions with Materials and Atoms doi:10.1016/j.nimb.2011.02.036. [11] N. Gordillo et al., Technical developments for computed tomography on the CENBG nanobeam line, Nuclear Instruments and Methods. B (2011), doi:10.1016/j.nimb.2011.02.032 [12] A.Bhargava, P.J. Masset, “Ash fusibility studies in high temperature oxidising/reducing atmospheres”, 4th International Freiberg Conference on IGCC & XtL Technologies - 3-5th May, 2010.
The potential to upgrade petroleum cokes using high temperature processing Mohamed Ismail, John W Patrick, Ed Lester School of Chemical and Environmental Engineering, The University of Nottingham, University Park, NG7 2RD Nottingham, United Kingdom
[email protected]
Abstract Metallurgical coke is used primarily as a reducing agent for the reduction of iron in the blast furnace. Due to the high cost, high demand and reduced availability of the high quality coking coals used in the production of this coke, alternative resources are being sought fulfill the role of this coke. One possibility is to make use of petroleum coke. The encouraging aspects for this use of petroleum coke include higher calorific value than traditional coke, relatively low cost, low ash content and ready availability. On the other hand, petroleum coke may not meet the requirements as regards mechanical strength and reactivity due to the influence of in both the macro and micro structures. In this study three samples have been examined for potential use of petroleum coke as a replacement for blast furnace coke. Two of these samples were green petroleum cokes and third one was a good quality metallurgical coke used for the purpose of comparison. These samples were subjected to microstructure analysis before and after heat treatment by X-ray computed tomography which is a non destructive test to explore the internal structure of these samples. In this preliminary work heat treatment up to 900 C led to increased densities of the petroleum cokes from about 1.4 to 1.8g/cm3 and these changes were reflected in the decreased voidage as demonstrated by the X ray computed tomography. The heat treated petroleum coke had a similar voidage to the good quality blast furnace coke indicating that heat treated petroleum coke may be a possible substitute for blast furnace coke. 1. Introduction Petroleum coke is a residual by-product of the cracking process in oil refineries. It may be described as a carbonaceous solid that serves various applications after suitable processing [1]. In recent years, crude oils have become heavier resulting in heavier coker lower case feeds [2]. There are many factors which affect the structure of petroleum coke such as the properties of crude oils and the processing conditions of the cokes such as temperature and pressure. Petroleum coke is mainly used as fuel in power plants and cement factories as a construction material for anodes for aluminum production, lime kiln firing, and specialty
applications, such as dry cells and electronics. It has been used as an additive to coal blends which are used in the production of metallurgical coke. However, it has not been used directly for the reduction of iron ore in a blast furnace. Petroleum coke offers some advantages over traditional metallurgical coke such as higher calorific value, relatively low cost, lower ash, ready availability, and for some types of petroleum coke, low sulfur content [1]. The objective of this study is to gain a thorough understanding of the structure of petroleum coke before and after heat treatment with the aim of assessing the possibility of changing the structure through heat treatment to enable the coke to be considered as a blast furnace coke substitute. 2. Experimental work
Coke materials used
One sample of a good metallurgical coke (BFC) was used as a standard for the purpose of comparison with two samples of petroleum coke, cokes K and RG. These three samples were prepared in the form of disc lumps with the dimensions of 5 to 7 mm diameter and 3 to 4 height.
Techniques
In order to assess the feasibility of petroleum cokes for conventional heat treatment, their physical properties were characterized using helium pycnometry and X ray computed tomography. Chemical properties were assessed using thermal gravimetric analysis (TGA), elemental analysis and bomb calorimetiry.
Procedure
In order to produce a means of distinguishing the microstructure of petroleum cokes from that of industrial metallurgical cokes, the three samples were scanned by X-ray computed tomography before and after treatment. XRCT is a non destructive technique which can be used to obtain digital information about the microstructure of the specimen in three dimensions [3], the images then being analyzed by using ImageJ software [4]. The specimens of petroleum coke were subjected to the heat treatment in an inert atmosphere (N2) using TGA with a variety of ramp rates and maximum temperatures. After initial heating rate of 50 C/min to 105C, the samples were held for 5 minutes and then heated again to a final temperature of 700C or 900C using a ramp of 10C/min, 30 C/min or 50C/min. the samples where then held at the chosen temperature for 20 minutes.
3. Results and Discussions Table 1 and 2 show that the carbon content in the petroleum cokes (K and RG) is greater than that of the traditional metallurgical coke (BFC). As a result, of that the heating value of K and RG is greater than that of BFC; this is shown in Table 3. Sample Name
Moisture VM, C, wt% wt% wt% 1.5 8.6 79.8 1.4 9.7 90.8 1.8 10.6 86.0 Table 1 Proximate analysis of cokes
BFC K RG
Ash, wt% 10 1.3 1.7
Sample Name N ,wt% C ,wt% H, wt% BFC 0.00 79.45 2.36 K 1.47 90.84 3.76 RG 0.00 83.96 3.66 Table 2 Elemental analysis of coked
S ,wt% 0.9 0.0 0.0
Sample Name BFC K Calorific Value, MJ/kg 28.82 36.20 Table 3 Gross calorific values of cokes
GR 36.48
Table 4 shows that for petroleum coke K there is a clear effect due to the increase in temperature between 700C and 900C but the rate of heating between 10C/min and 50C/min had little effect. The percentage weight loss at 700C was of the order of 4.4% and 7.25% at the final temperature of 900C. Ramp, Maximum Initial Wt, g Wt loss % C/min Temp, C K1 10 700 134.47 4.35 K2 30 700 98.40 4.53 K3 50 700 127.84 4.20 K4 10 900 132.03 7.11 K5 30 900 107.60 7.32 K6 50 900 127.88 7.31 Table 4 the effect of temperature and heating rate on the specimens of petroleum coke K Sample
As a result of these values petroleum coke sample RG was heat treated to 900C and a ramp rate of 50C/min. the data for this coke and the corresponding data for the petroleum coke K are given in Tables 5 and 6. The thermogravimetric analyser measured the weight change with temperature and the density of the samples was obtained by helium pycnometry. The data in these tables represented samples of the two cokes and provide an indication of repeatability. Tables 5,6 show the change in density of petroleum cokes K and RG. Before Treatment
Sample K7 K8 K9 K10 K11
After Treatment
Wt, g
, g/cm3
Wt, g
, g/cm3
wt%
0.377
1.447
0.344
1.873
8.775
0.326
1.455
0.297
1.897
8.812
0.402
1.452
0.370
1.858
8.085
0.398
1.447
0.366
1.862
8.073
0.294
1.461
0.270
1.873
8.325
Table 5 Green Petroleum Coke (K1) before and after heat treatment
In comparison of petroleum coke after heat treatment with good quality of blast furnace coke from Tables 5 & 6 and Figures 1, 2 & 3 show that the density is increased consequently as voidage is decreasing. Before Treatment
Sample RG1 RG2 RG3 RG4 RG5
After Treatment
Wt, g
, g/cm3
Wt, g
, g/cm3
wt%
0.316
1.492
0.293
1.873
7.378
0.348
1.463
0.322
1.860
7.533
0.399
1.448
0.370
1.871
7.336
0.334
1.482
0.310
1.868
7.269
0.362
1.452
0.335
1.847
7.386
Table 6 Green petroleum coke (RG) before and after heat treatment
Table 7 shows Image J data analysis for K1 specimen before and after heat treatment. Table 7 K1 petroleum coke Before heat treatment After heat treatment Blast Furnace Coke BFC
Total Area, mm2 15.5 8.1 12.1
Voids Area Fraction % 21.6 11.5 13.0
Figure 1a shows the scanned specimen of petroleum coke by X ray CT. X ray scanning parameters for a good quality image were determined by adjusting the energy of X rays, intensity of the beam and primary and secondary filters. Figure 1b shows the threshold of the image using software to analyze the data, the threshold being applied to characterize the components according to density and then converts the representative grey scale of a component into value. 1c is the analysis of image which is the extraction of void areas in pixels.
a
b
c
Figure 1. XRCT Image of Petroleum coke k1 before heat treatment
a
b
Figure 2 XRCT Image of petroleum coke K1 after heat treatment
c
a
b
c
Figure 3 XRCT Image of a good quality of blast furnace coke BFC 4. Conclusions Heat treatment of the petroleum coke samples results in a relatively small weight loss of approximately 4-8% depending on the maximum temperature. The heating rate appears to have little effect on the final weight loss. The void area fraction of pet coke K, from XRCT, appears to decrease to a similar value to that seen in the standard blast furnace coke BFC. The weight loss for Green petroleum coke RG appears to be higher, in all cases, than for pet coke K. the increase in density (from approximately 1.5 g/cm3 to 1.9 g/cm3), as a result of that heat treatment, also appears to be similar regardless of heating rate or maximum temperature. Understanding the microstructure of petroleum coke is vital in order to assess the feasibility of heat treating petroleum coke lumps. Acknowledgements This project is funded by the IDB bank for developing countries, under the Development Solutions Program and the University of Nottingham and also special thanks to Mr Michael for X-ray CT. References [1] Ellis, P.J., and Paul, C.A., ‘Petroleum Coke Calcining and Uses,’ Proc. 3rd Int.Conf. on Refining Processes, Delayed Coking, Atlanta, Ga , (2000) [2] Adam H. A. In: Marsh H., Heintz E., Rodiriguez-Reinoso F., editors.’ Introduction to carbon Technology’. Alicante 1997. Ch. 10. [3] Ketchum, R.A, Carlson, W.D,’Acquisition, Optimisation and Interpretation of X-Ray Computed Tomographic Imeagery: Applications to the Geoscience’ Computers and Geosciences 27,2001,pg 381-400 [4] ImageJ, http://rsb.info.nih.gov/ij/ software download and manual, accessed 30/10/2008
Oviedo ICCS&T 2011. Extended Abstract
A study of the feasibility of an anthracene oil-based pitch for isotropic carbon fibres preparation N. Díez, P. Álvarez, R. Santamaría, C. Blanco, R. Menéndez and M. Granda. Instituto Nacional del Carbón, CSIC. P.O. Box 73, 33080-Oviedo, Spain
[email protected] Abstract A new environmentaly friendly pitch, obtained from anthracene oil, is used for the preparation of isotropic carbon fibres. The pitch exhibits an adequate thermal behaviour and is free of solid particles. Green carbon fibres were obtained by means of a meltspinning process with no filtering step, and subsequent stabilization and carbonization. For the optimization of the melt-spinning process, the influence of the spinning temperature, extrusion pressure, spinneret hole size, winding speed and the interrelationship of these factors upon the microstructure and diameters of the fibres was studied.
1. Introduction Isotropic pitch-based carbon fibres are usually produced by melt-spinning [1]. One of the most important factors to be taken into account in their preparation is the specific characteristics of the precursors, in particular their softening point, carbon yield or the presence of solid particles -including mesophase spheres- [2,3]. The use of commercial pitches (coal-tar pitches and petroleum pitches) as carbon fibres precursors usually requires certain pre-treatments in order to adjust some of their characteristics [4]. This is because these pitches are mainly produced for use as binder and impregnating agents in aluminium and graphite industries. Industrial anthracene oil, obtained as a distillation fraction from coal tar, can be employed as raw material for the preparation of pitches on an industrial scale, since it is readily available (it represents ~ 30 wt. % of coal tar) and chemically consistent (polycyclic aromatic hydrocarbons of 3-5 rings [5]). This raw material is considered nowadays a good alternative for producing pitches, with a controllable softening point, a high carbon yield and no solid particles [6]. Among the unique characteristics of these pitches is their low environmental impact, derived from the low content in polycyclic aromatic hydrocarbons (PAH) catalogued as carcinogenic [7]. Furthermore, the preparation process (oxidative thermal condensation followed by thermal treatment) is
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Oviedo ICCS&T 2011. Extended Abstract
very versatile as it enables parameters such as the softening point of the pitch, to be easily controlled during the final steps of the preparation procedure [8]. In this paper we report on the feasibility of using novel anthracene oil-based pitches for the preparation of isotropic carbon fibres at laboratory scale. The main goals of the study are: (i) To characterize the anthracene oil-based pitch in order to determine the parameters that are relevant to its subsequent transformation into a carbon fibre (e.g., composition and pyrolysis behaviour) and (ii) to determine and optimize the main variables (spinning temperature, extrusion pressure, spinneret hole size and winding speed) that most affect the melt-spinning process. After stabilization and carbonization, the mechanical properties of the fibres (i.e., tensile strength) are determined.
2. Experimental section The carbon fibres precursor used in this work was an isotropic anthracene oil-based pitch (AOP) supplied by Industrial Química del Nalón, S.A. This pitch was obtained from anthracene oil [6] by a recently reported procedure consisting of oxidative thermal condensation and subsequent thermal treatment / distillation. This second processing step was adjusted until the desired softening point was reached (~ 250 °C). AOP pitch was characterized in terms of elemental composition, solubility tests, softening point and Fourier-transformed infrared spectroscopy. The thermal stability of the pitch was studied by means of a standard industrial test. The pitch was heated up to 45 ºC above its softening point for 20 h in an airtight device. The softening point of the product was then measured and compared with that of the parent pitch. AOP was melt-spun into fibers in a stainless steel reactor, equipped with stainless steal spinnerets of diameter 300 μm and 500 μm. The spinning temperatures chosen ranged btween 260 °C and 285 °C. Once the spinning temperature was optimized, different nitrogen extrusion pressures (from 1 bar to 5 bar) and winding speeds (from 50 cm s-1 to 250 cm s-1) were used to obtain fibres of different diameters. The green fibres were stabilized in an oven under an air flow of 20 L h-1, using the following multi-step program: heating at 1 °C min-1 from room temperature to 150 °C, maintaining this temperature for 4 h, and then heating at 1 °C min-1 to 160 °C, 180 °C, 200 °C, 220 °C, 250 °C and 270 °C, with 1 h of residence time at each of these temperatures. The stabilized fibres were then carbonized in a horizontal furnace, under a
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Oviedo ICCS&T 2011. Extended Abstract
nitrogen atmosphere, at 2 °C min−1 to 900 °C and 30 min of soaking time at this temperature. The texture and the diameter of the green, stabilized and carbonized fibres were studied by scanning electron microscopy (SEM). The tensile strength of the fibres was measured according to the ASTM D3379-75 Standard for single-filament materials.
3. Results and Discussion The anthracene oil-based pitch used comprises a series of characteristics that makes it highly suitable as a precursor for the preparation of carbon fibres (Table 1). The pitch is totally ash free. Therefore, its toluene and N-methyl-2-pyrrolidinone insolubles contents (TI and NMPI, respectively) are due exclusively to the presence of molecules with a high condensation degree, which makes them insoluble in these solvents. Table 1. Characteristics of the anthracene oil-based pitch.
AOP
Elemental Analysis (wt.%) C H N O S
IAr1
Ash 2
TI3
NMPI4
SP5
93.3
0.68
0.0
58
23
247
4.1
1.4
0.8
0.4
1
Aromaticity index determined by FTIR. Ash content ( wt. %). 3 Toluene insolubles content ( wt. %). 4 N-methyl-2-pyrrolidinone insolubles content ( wt. %). 5 Mettler softening point (ºC) 2
AOP is mainly composed of carbon (> 93 %) and, to a lesser extent, hydrogen, nitrogen and oxygen. The hydrogen is mainly aromatic, as indicated by its high aromaticity index. The low oxygen content (0.8 wt. %) proves that the oxygen groups introduced during the oxidative thermal condensation step were successfully removed during the subsequent thermal treatment. Another important feature is the low sulphur content of this pitch (less than 0.4 wt. %). Thermogravimetric analysis shows that below ~ 350 °C, weight loss is negligible (< 3 wt.%) as could be expected from a pitch with low volatiles content. Weight loss which mainly occurs between ~ 350 ºC and ~ 600 ºC yields a carbonaceous residue at 1000 ºC of 62 wt. %. These results indicate that the pitch could be heated above its softening point (required step for melt-spinning the pitch) up to ~ 350 °C without undergoing any significant loss of volatiles. In addition, the high carbon residue obtained is indicative of the high carbon yield of this pitch at high temperature.
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Oviedo ICCS&T 2011. Extended Abstract
The thermal stability of the precursor was studied according to the industrial procedure described in the experimental section. After 20 h at 292 °C, the softening point of the pitch increased by only 0.5 ºC. Furthermore, the thermogravimetric curve of the pitch after the test almost exhibited the same pattern than that of AOP, indicating that the pyrolysis behaviour of the pitch had not undergone any significant changes during the experiment. From these results it can be inferred that AOP would be especially suitable for use as a carbon fibre precursor. Having established that anthracene oil-based pitches do not require any pre-treatment prior to melt-spinning, we next evaluate the feasibility of this pitch for melt-spinning. For this purpose, a laboratory-scale apparatus that uses nitrogen pressure to extrude the pitch through a mono-hole spinneret was employed. Among the parameters involved in the pitch melt-spinning process, spinning temperature, extrusion pressure, winding speed and spinneret hole size are the most important, requiring optimization and analysis in detail [9,10]. Spinning was carried out at different temperatures using a 300 µm monohole spinneret, a nitrogen extrusion pressure of 5 bar and a winding speed of 250 cm s-1. The extrusion of the pitch only occurs at temperatures above 260 ºC, but spinning temperatures higher than 265 ºC are necessary to achieve a continuous flow of pitch through the spinneret. Examination of the green fibres by SEM (Figure 1) revealed that only spinning temperatures higher than 280 ºC (~30 ºC above their softening point) lead to green fibres with smooth and defect-free surfaces (Figure 1a). Below this temperature (Figure 1b, position A) the presence of defects on the surface of the fibres was detected. a
b
20 µm
Figure 1. SEM images of green fibres spun at (a) 275 and (b) 280 ºC. In order to study the effect of the extrusion pressure and winding speed on the diameter of the green fibres, several experiments were carried out using different nitrogen pressures (3 to 5 bar) and winding speeds (50 to 250 cm s-1), while the spinning temperature (280 ºC) and spinneret hole size (300 µm) were kept constant (Figure 2).
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Oviedo ICCS&T 2011. Extended Abstract
100
Diameter (µm)
80 60
5 bar 4 bar 3 bar
40 20 0 0
50
100
150
200
250
-1
Winding speed (cm s )
Figure 2. Variation of green fibres diameter with winding speed and nitrogen pressure. The results show that only nitrogen pressures between 3 and 5 bar allow a continuous flow of pitch resulting in correctly spun fibres. Thus, whereas nitrogen pressures lower than 3 bar fail to produce extrusion of the pitch, pressures higher than 5 bar do not allow the pitch to be stretched and wound correctly. Variation in the diameter of the green fibres with the winding speed confirms that the fibre diameter decreases with the increase in winding speed, possibly because the filament is more easily stretched. It was also observed that an increase in the extrusion pressure does not exert any significant influence on the average diameter of the extruded fibres. The thinnest fibres (average diameter of ~25 µm) were obtained at a nitrogen pressure of 3 bar and a winding speed of 250 cm s-1. The average diameter of the fibres obtained with the 500 µm spinneret (under a nitrogen pressure of 3 bar and a winding speed of 250 cm s-1) is slightly higher than that achieved with the 300 µm spinneret. Thus, at 3 bar and 250 cm s-1 the pitch that is spun through the 500 µm spinneret shows an average diameter of ~30 μm, ~10 μm larger than that spun with the 300 µm spinneret diameter. However, when a spinneret of a larger size was used, green fibres with diameters as low as ~ 20 μm were obtained because it was possible to apply lower extrusion pressures (up to 1 bar). Once spun, the green carbon fibres must be stabilized prior to carbonization in order to render the fibre infusible. The green fibres obtained at a spinning temperature of 280 °C,
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Oviedo ICCS&T 2011. Extended Abstract
a nitrogen pressure of 1 bar and a spinneret hole size of 500 μm (conditions that led to the green fibres with the smallest diameters) were stabilized, using the temperature/time program given in the experimental section, and then carbonized at 900 °C for 30 min. SEM observations showed that neither stabilization nor carbonization produced any defect on the surface of the fibres. Moreover, the microstructure of the carbon fibre remains completely isotropic whereas the morphology of the fibres, especially their diameter, undergoes substantial changes. Thus, stabilization causes a slight increase in the diameter of the fibres, which may be related to the uptake of oxygen during the process. Subsequent carbonization produces shrinkage that results in a decrease in the diameter of the fibre. This shrinkage leads to carbon fibres that have an even smaller diameter than that of the green fibre (~ 15 μm). The mechanical properties of the carbon fibres were evaluated in terms of their tensile strength (Figure 3).
Tensile strength (MPa)
1200 1000 800 600 400 200 0 10
20
30
40
50
60
Diameter (µm)
Figure 3. Variation of the tensile strength with carbon fibre diameter. As expected, tensile strength increases exponentially as the carbon fibre diameter decreases, from ~ 200 to > 1100 MPa (for carbon fibres with an average diameter of ~ 40 μm and 15 μm, respectively). These values are comparable to those reported in the literature for isotropic carbon fibres, even for other prepared from petroleum derivatives [11,12] which confirms the suitability of the pitch for high quality isotropic fibres preparation.
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Oviedo ICCS&T 2011. Extended Abstract
4. Conclusions A novel anthracene oil-based pitch of high softening point was successfully transformed into isotropic carbon fibres by melt-spinning. Analysis of the mechanical strength of the carbon fibres demonstrated that the isotropic precursor (anthracene oil-based pitch) and the procedure used in this study leads to carbon fibres with a tensile strength comparable to that of typical pitch-based isotropic carbon fibres produced from standard pitches. This demonstrates that the production of carbon fibres from this novel precursor represents an alternative to the ones available nowadays in the market.
Acknowledgement. The research leading to these results has received funding for the European Union’s Research Fund for Coal and Steel research programme under Grant Agreement number RFCR-CT-2009-00004. Dr. Alvarez also gratefully acknowledged Spanish Ministry of Science and Education by her Ramon y Cajal contract.
References [1] Derbyshire F, Andrews R, Jacques D, Jagtoyen M, Kimber G, Rantell T. Synthesis of isotropic carbon fibers and activated carbon fibers from pitch precursors. Fuel 2001;80:345-56. [2] Zeng SM, Maedaa T, Tokumitsua K, Mondoria J, Mochida I. Preparation of isotropic pitch precursors for general purpose carbon fibers (GPCF) by air blowing—II. Air blowing of coal tar, hydrogenated coal tar, and petroleum pitches. Carbon 1993;31:41319. [3] Mora E, Blanco C, Prada V, Santamaría R, Granda M, Menéndez R. A study of pitch-based precursors for general purpose carbon fibres. Carbon 2002;40:2719-25. [4] Alcañiz J, Cazorla D, Linares Solano A, Oya A, Sakamoto A, Hoshi K. Preparation of General Purpose Carbon Fibres from coal tar pitches with low softening point. Carbon 1997;35: 1079-87. [5] Fernandez JJ, Alonso M. Anthracene oil-based pitches. Light Metals 2004;449-50. [6] Álvarez P, Granda M, Sutil J, Santamaría R, Blanco C, Menéndez R, Fernández J J, Viña JA. Preparation of Low Toxicity Pitches by Thermal Oxidative Condensation of Anthracene Oil. Environ. Sci. Technol. 2009;43:8126–32.
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Oviedo ICCS&T 2011. Extended Abstract
[7] USEPA, Provisional Guidance for Quantitative Risk Assessment of PAH, EPA/600/R-93/089, United Status Environmental Protection Agency (1993). [8] Álvarez P, Granda M, Sutil J, Menéndez R, Fernández JJ, Viña JA, Morgan TJ, Millán M, Herod AA, Kandiyoti R. Characterization and Pyrolysis Behavior of Novel Anthracene Oil Derivatives. Energy & Fuels 2008;22: 4077–86. [9] Eddie DD, Dunham MG. Melt spinning pitch-based carbon fibres. Carbon 1989;27: 647-55. [10] Kase S, in: A.Ziabicki and Kawai H (Eds.), High-Speed Fibre Spinning, Wiley Interscience, New York, 1985, pp. 67-113. [11] Yang KS, Lee DJ, Ryu SK, Korai Y, Kim YJ, Mochida I. Isotropic carbon and graphite fibres from chemistry modified coal-tar pitch. Korean J. Chem. Eng. 1999;16: 518-24. [12] Wazir AH, Kakakhel L. Prepartion and characterization of pitch-based carbon fibres. New Carbon Materials 2009;24: 83-88.
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Oviedo ICCS&T 2011. Extended Abstract
Advanced characterisation of liquid hydrocarbons from South African high volatile bituminous coal M.H. Makgato1, H.W.J.P Neomagus1, R.C Everson1, J.H.L. Jordaan1, H.H. Schobert2 1
Chemical Resource Beneficiation (CRB), School of Chemical and Minerals Engineering, North-West University, Private Bag X05, Potchefstroom 2531, South Africa 2 The EMS Energy Institute, Penn State University, University Park, PA 16802 USA Abstract Detailed characterisation of liquid products from the first step of a two stage non catalytic direct liquefaction of a South African vitrinite-rich bituminous coal is reported. Elucidation of the composition, structural properties and molecular mass distribution of coal-derived liquids was carried out by GC/MS, NMR spectroscopy and Maldi-MS. Methylated-decane and cosane derivatives are the most predominant aliphatic hydrocarbons whereas alkylated derivatives of naphthalene, phenanthrene, pyrine, anthracene and pyrilene are the backbone aromatic structures. Of the five hydrocarbon classes identified, major changes were observed on the distribution of aliphatics and aromatics whereas the more polar compounds remain unchanged on solvent extraction.
Corresponding Author: +27 (0) 18 299 1991; E-mail address:
[email protected]
1. Introduction When coal is subjected to catalytic hydrogenation in direct liquefaction, it yields valuable organic products that could be used as feedstock for the preparation of petrochemicals, stable jet fuels and automotive fuels [1; 2]. To make coal a valuable feedstock for fuel engine and useful chemical addition of hydrogen or carbon removal is required [3, 4] to increased the H/C ratio from 0.8 in bituminous coals to 1.3-2.0 in petroleum crude oil, gasoline and diesel [5] preferably by direct coal liquefaction [6].
Previous studies [7, 8] on coal extraction of South African coals were mainly aimed at elucidating the structural properties of coal relative to its extracts or studying the swelling behavior of the coals without detailed characterization of the extracts. The data presented in this paper is results from the technical experiments on elucidation of the
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Oviedo ICCS&T 2011. Extended Abstract
chemical and structural compositions of the coal liquid derived extraction of coal with residue oil.
2. Experimental section 2.1 Material studied Proximate compositions of the vitrinite-rich Waterberg coal of interest were: 2.4% moisture, 37.1% volatile matter, 52.0% fixed carbon and 10.5% ash on dry basis whereas the ultimate analysis were 74.4% C, 5.3% hydrogen and 7.3% oxygen whereas nitrogen and total sulfur compositions were 1.5% and 0.99% on dry ash free basis. The tar distillate oil from gasification plants in Sasol Secunda, South Africa used as solvent showed the chemical class type analysis as 11% paraffins, 89% aromatic and less 0.1% cycloparaffins with the boiling range between 110-451°C at 0.5-95% weight loss.
2.2. Methods and techniques 2.2.1. Procedure for coal extraction Batch extraction experiments were carried out in a 2500 ml stirred autoclave at a temperature of 370°C, 7 bar final pressure, 2 hours resident time and solvent and coal ratio of 5:1 and 10:1 under initial N2 pressure and a constant 400 rpm stirring. After separation of residue coal from the coal extract, the degree of coal conversion was determined by weight loss between the feed coal and the residue coal on dry ash free basis using equation (1) and (2).
% Total conversion (daf) = 100 × %x Extraction (db) = 100 ×
1−Residue weight (g)/Coal weight (g) 1−Ash (wt %, db )/100
Coal weight (g) ∙ % x C −Residue weight (g)∙ % 𝑥𝑥 𝑅𝑅𝑅𝑅 Coal weight (g) ∙ % x C
(1)
(2)
2.2.2. Advanced characterization of products All solid materials, including the raw coal were analysed for proximate and ultimate compositions according to the South African Bureau of Standards (SABS) and ISO standards for coals and cokes analysis. In addition, spectroscopic techniques such as XRD, laser desorption mass spectroscopy and solid state NMR were used to examine the structural composition. The liquid products after fractionation into solubility class
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Oviedo ICCS&T 2011. Extended Abstract
distribution were characterized by GC-MS and liquid NMR spectroscopy.
3. Results and Discussion 3.1. Coal conversion The total coal conversion at constant 7 bar final pressure, 2 hours reaction time and 10/1 solvent to coal ratio increased from 45% at 350°C to 63 % at 360°C. Compared to other studies [9, 10, 11] that gave higher yield at either high solvent/coal ratios or catalytic hydrogen pressure, the yield reported here is from solvent extraction with low solvent composition and no catalyst of hydrogen pressure.
3.2. Advanced characterisation 3.2.2.1. Solubility fractionation Fractionation of residue oil and coal liquids extract with pentane, toluene and THF gave 80% pentane soluble fraction and no toluene insoluble fraction for residue oil while the coal liquid contains 50% of pentane soluble and 39% of toluene soluble while the toluene insoluble fraction was composed of 10% tetrahydrofuran soluble and only 4% insoluble residue.
2.2.2.2. Gas chromatography/Mass spectroscopy (GC-MS) Of the twenty most abundant aromatic compounds identified by GC-MS, pure compounds such as naphthalene, dibenzofuran and phenanthrene-derivatives are the most predominant, while cosine and decane-derivatives are the most predominant saturated hydrocarbons in both the residue oil and coal extract. Low molecular mass compound predominated phenols, benzene derivatives and high molecular mass region dominated C20 to C27 saturated hydrocarbons. A decrease in molecular mass from 522 m/z in the residue oil to 474 m/z in the extract were observed.
2.2.2.3. Nuclear magnetic resonance (NMR) Figure 3 shows summarises the structural parameters from 13C NMR spectra of residue oil and coal extract.
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Oviedo ICCS&T 2011. Extended Abstract
0.9 0.8
Relative distribution
0.7 0.6 0.5 0.4
RO
0.3
CEXT
0.2 0.1 0 fa*
fal
faP
faS
faH
faB
falH
Structural parameter
Figure 3. Distribution of structural chemical compositions of coal on solvent extraction
The coal liquid shows increased aromatic content (fa*), phenolic content (faP), alkyl substitution (faS). As with GC-MS, an increased fraction of alkyl-substituted aromatics in CEXT together with lower aliphatic fraction (fal) ring joining CH2 and CH due to depolymerisation was observed. Occurrence of depolymerisation and aromatisation was accounted in this case by the dissolution of more aromatic compounds from coal by residue oil while the substitution reactions occur between radicals from coal and residue oil as a result on breakage of CH2 and CH linkages. Khan, et al. [12] showed a decrease in the contribution of hydroaromatics and polyaromatics as well as an increase in substitution on extraction of Pakistani coal with benzene at 400°C using NMR as a characterisation tool.
2.2.2.4. Matrix-laser desorption ionisation mass spectroscopy (Maldi-MS) The Maldi-MS reflectron spectra analysis of residue oil and its coal extract showed that both samples contain a low mass envelope below 340 m/z and three maxima at 260, 370-400 and 550 m/z with the residue oil giving large abundance of low molecular weight compounds at 260 m/z than the extract.
4. Conclusions Advanced characterisation techniques such as NMR, GC-MS and Maldi-MS techniques were used to elucidate composition, structural properties and molecular mass distribution on coal extraction with residue oil. GC-MS analysis showed naphthalene, phenanthrene, pyrine, anthracene and pyrilene as backbone structures for alkyl-
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Oviedo ICCS&T 2011. Extended Abstract
aromatics and polycyclic hydrocarbons (PAH), with methyl-naphthalene been the most dominant aromatic hydrocarbon with up to 3 methyl substitutions. The study has established that although depolymerisation occurs on solvent extraction the composition of aromatic compounds also increases, which accompanied by a decrease in the molecular mass from 522 m/z in residue oil to 474 in coal extract. It was also found that long chain aliphatic hydrocarbon rather than aromatic compounds form major composition of the high molecular mass species, whereas phenols are the most abundant low molecular mass compounds.
Acknowledgement The authors wish to acknowledge financial support for this research from the South African National Research Institute (SANERI). Exxaro R&D and Sasol R&D for supplying the coal samples and residue oil provided respectively.
References [1] Schmid, B.K. Jackson, D.M. 1981. The SRC-II process. Phil. Trans. R. Soc. Lond. A, 300: 129-139. [2] Burgess, C.E., Schobert, H.H. 2000. Direct liquefaction for production of high yields of feedstock for specialty chemicals or thermally stable jet fuels. Fuel Process. Technol. 64: 57-72. [3] Whitehurst, D.D., Mitchell, T.O., Farcasiu, M. 1980. Coal Liquefaction: The chemistry and technology of thermal processes, Academic Press, London. P. 206-270. [4] Gorin, E.G. 1981. “Fudamentals of coal liquefaction” Chapter 27 of Chemistry of coal utilization, 2nd supplementary volume. Ed. Elliott, M., John Wiley & Sons. [5] Williams, R.H. & Larson, E.D. 2003. A comparison of direct and indirect liquefaction technologies for making fluid fuels from coal. Energy for Sustainable Development, 7: 103-129. [6] Huang, H., Wang, K., Wang, S., Klein, M. T., Calkins, W. H. 1996. Kinetics of Coal Liquefaction at Very Short Reaction Times. Energy Fuels, 10: 641-648. [7] Larsen, J.W., Green,T.K. Kovac, J. 1985. The nature of the macromolecular network structure of bituminous coals. J. Org. Chem. 50: 4729-4735. [8] Iino, M. 2000. Network structure of coals and association behaviour of coal-derived materials. Fuel 62: 89-101.
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[9] Okuyama, N., Komatsu, N., Shigehisa, T., Kaneko, T., Tsuhuya, T. 2004. Hypercoal process to produce the ash-free coal. Fuel Process. Technol., 85: 947-967. [10] Miura, K., Nakagawa, H., Ashida, R & Ihara, T. 2004. Production of clean fuels by solvent skimming of coal at around 350°C. Fuel, 83: 733-738. [11] Burgess Clifford, C., Schobert, H.H., Escallón, M. & Griffith, J. 2007. Effects of coal rank and reaction conditions on production of two-ring compounds for coal-based jet fuel. (Paper presented at International Conference on Coal Science and Technology held in August 2007.) Nottingham, UK. [12] Khan, M.A. Ahmad, I. Ishaq, M., Shakirullah, M. (2003) Spectral characterisation of liquefied products of Pakistani coal. Fuel Process. Technol. 85, 63-74.
6
A thermo-petrographic method to identify coals prone to self-oxidation Claudio Avila and Edward Lester Department of Chemical and Environmental Engineering, University of Nottingham. University Park, Nottingham (NG7 2JT). United Kingdom.
[email protected] Abstract The self oxidation of coal at low temperature is a significant problem for the coal industry. It is important not only because the economical losses associated to the reduction of the calorific value and the emission of noxious gases to the atmosphere; it is also relevant because it can lead to a bigger problem: the spontaneous combustion of a coal stockpile and a major fire. Although there have been created a number of methods to forecast this unwanted reaction, these are not reliable enough for many in the coal industry. Most of these tests are based on the measurement of heat release and oxygen consumption rate, with a significant variability in the results, depending of the testing procedure. This paper presents an alternative test based on the visual quantification of thermal alteration by means of combined petrographic and image analysis techniques. The experimental test heats up small coal samples at a low ramp rate (0.5oC min-1) under an oxidative atmosphere, starting from room temperature up to 250oC. Then, changes in the optical properties of fresh and oxidized material such as light reflectance, morphology and textural attributes are assessed by oil immersion microscopy and image analysis processing. Results showed a notable difference between the reflectance histograms of fresh and oxidized samples that relate directly to the reactivity of samples and to the coals susceptibility to spontaneous combustion. Also some physical alterations were detected and quantified such as oxidation rims, micro fractures, and changes in porosity that are key features with prone samples. Finally, a standardized procedure is proposed for propensity assessment.
1. Introduction The natural oxidation of coal is a phenomenon that begins immediately when coal come into contact with an oxidative atmosphere. This oxidation process is a set of several chemical reactions (in series and in parallel), with an overall net energy output. Under specific conditions, a thermal runaway can take place when the heat released during oxidation exceeds the dissipation rate of the material into the environment [1,2]. It is a slow process that takes days or months, and it is a particular problem when the material is stored at large
scale. The causes of spontaneous combustion are well known [3,4,5,6], and several methods have been developed, using these mechanisms, to predict the spontaneous combustion potential. However, the self-oxidation reaction is a complex series of reactions and accurate prediction is still difficult. Despite the advances in the understanding of the problem, the different experimental tests can deliver contradictory results [7,8]. Among the methods developed to predict this reaction, thermal methods are the most commons. In this case, these are based in the measurement of changes in the temperature of the sample, the heat released, and the change in the gas concentration produced by the emission of gaseous products and due to the oxygen adsorption [9,10,11]. However, these tests do not consider some relevant coal properties such as the weight evolution, the maceral composition and the optical properties of coals (random reflectance). This is relevant because these properties are frequently used by the coal industry to assess the coal reactivity, representing an opportunity to use a well known source of information available to predict this unwanted phenomenon. The use of petrographic information obtained from samples before, during, and after a thermal ‘event’ could result in a relevant contribution to improve the current tests, or to lead to a new generation of testing procedures. This paper presents the experimental results obtained from several coals, including well known spontaneous combustion coals, which have been exposed to different non isothermal programs in a thermal reactor. The paper evaluates changes in reactivity, based on the different morphotypes generated after the heating process, and the change in reflectance of the sample as a whole.
2. Materials and methods For this study, 25 coal samples from different parts of the world have been used, including at least 3 well-known spontaneous combustion coals. All experiments used fresh sample that were pulverized and then sieved into the size range 53-106µm. Each sample was characterized petrographically (random reflectance, morphologic characterization and maceral content), before and after thermal treatment. Finally, the optical characteristics are compared for the raw and treated material, which are also linked to the thermal response observed during the thermal treatment.
2.1 Coal thermal treatment A reactor has been designed to study the thermo-chemical behaviour of the samples under a controlled slow heating environment. In this reactor, the weight loss of the sample was measured simultaneously along with the temperature at different positions (diagram shown in Figure 1). The experimental procedure is described as follows: 100g of sample is placed into the sample holder and connected to the digital balance. Thermocouples are connected to the main PC unit which recorded all measured variables including weight and temperatures at each location. A gas flow of 80 ml/min through the centre of the reactor was set up to flow into the gas analyzers. After that, a heating ramp of 0.5oC min-1 starts to heat the furnace from room temperature to 250oC. When temperature reaches 250oC, all recorded signals are stopped, the sample holder is removed from the furnace, and the remaining material is prepare for petrographic analysis. 2.2 Sample analysis An automated light microscopy system consisting of a Zeiss Leitz Ortholux II POL-BK microscope, with oil-immersion objectives in the range of 10 to 32X magnification attached to colour digital camera Zeiss AxioCam is used for petrographic analysis. For these, polished blocks were prepared using carnauba paraffin wax to embed coal particles. After that, petrographic blocks were polished following the ASTM standard, at a final resolution of 0.04 µm. Finally, random reflectance measurements and maceral analysis were performed on all samples following the ISO for petrographic analysis of bituminous coals. Mosaic images were also obtained for all coal samples in order to identify and quantify the different morphotypes generated after the heating.
Figure 1: Diagram of experimental reactor designed to determinate experimentally the crossing point temperature values (CPT), to observe the transient thermal profiles of coals in a bulk sample, and to obtain the material needed for petrographic analysis.
3. Results 3.1 Thermal profiles Characteristic thermal profiles were recorded using the experimental reactor. In Figure 2, shows the data for two well known coals that are prone to spontaneous combustion. In this case, the high water content present in both samples (~15%) produced an inflection in the temperature profiles around 100oC. As can be seen by comparison with the Figure 3, in which are from a medium reactive coal (El Cerrejon, ~3.5% moisture content) and a low reactive coal (Pocahontas, ~0.6% moisture content), the effect of the water content in the thermal profiles is evident. The effect in the water content during the reaction also has been previously considered as a key self-heating mechanism [6]. From the experimental profiles, it is possible to observe that the temperature in the centre of the sample exceeds the temperature in the surroundings slightly (inside the sample), when the profile is close to 100oC. It is relevant because this suggests that heat production in the centre is not dissipated by the system fast enough to keep a flat profile, and this can be the beginning of a hotspot in a larger system.
Figure 2: North Dakota coal (left) and Fenosa coal (right). Both are extremely high reactive at low temperature.
Figure 3: El Cerrejon coal (left) and Pocahontas coal (right). The influence of the water content in the thermal profiles is clear.
Figure 4 shows a comparison of the Ignition temperature as predicted by various methods, including the classical crossing point temperature test [8] and the new crossing point temperature test, which is an improvement of the previous test introduced by Chen [10]. From this figure it is possible to see the discrepancies between the different approaches.
However, from previous historical evidence from these coals, the CPT value obtained using the Chen approach provide the closest indication of self heating potential.
Figure 4: Comparison of thermal parameters used to identify coals prone to spontaneous combustion. Among these, the most accepted is the CPT value obtained using the Chen correction [10], which delivers the minimum crossing temperature.
3.2 Reflectance comparison for fresh and oxidized coals As it can be seen from Table 1, the change in the reflectance value depends of the original sample reflectance. Low rank coals show larger changes than high rank coals. However, these variations are not directly related to the crossing point values recorded, and also they do not agree with previously reported tests of reactivity at low temperature [12]. Nevertheless, the change in the overall reflectance has a close relationship with the concentration of morphotypes presented in the sample (these are described in the next section) i.e. particles from highly reactive coals show a homogeneous change in reflectance whereas high rank or unreactive coals tend to show large amounts of “oxidation rims” and “high reflectance particles”. Clearly the reflectance change of the sample is linked to the kind of alteration formed, which is also related to the thermal profiles.
Table 1: Reflectance comparison for fresh and oxidized coals Coal
Reflectance Raw coal
Reflectance Treated coal
Change
North Dakota
0.41
0.59
0.18
% Change 45.0
Nadins
0.68
0.90
0.23
33.4
Fenosa
0.39
0.53
0.13
32.5
Hambach
0.32
0.42
0.10
31.9
Indo
0.62
0.78
0.16
25.6
Daw Mill
0.55
0.67
0.12
22.0
Kaltim Prima
0.54
0.65
0.11
20.4
Kleinkopje
0.76
0.89
0.13
17.8
La Loma
0.54
0.64
0.09
17.4
Asfordby
0.52
0.61
0.09
17.1
Zondag 1
0.61
0.71
0.10
15.5
Hunter Valley
0.77
0.89
0.11
14.3
Goedehoop
0.59
0.67
0.08
12.9
Cerrejon
0.57
0.63
0.06
11.2
Ironbridge
0.56
0.62
0.06
10.5
Lea Hall
0.60
0.66
0.06
10.3
Indiana EDF
0.68
0.75
0.07
10.1
Illinois 6
0.41
0.45
0.04
10.1
Littleton
0.65
0.71
0.06
10.0
Bentink
0.76
0.82
0.07
8.6
La Jagua
0.56
0.61
0.05
8.1
Yanowice
0.67
0.73
0.05
8.1
Pocahontas
1.33
1.36
0.03
2.4
Blue Creek
0.96
0.98
0.02
2.3
Deep Navigation
1.64
1.68
0.03
2.1
3.3 Morphologic characterization After heating, several characteristics in the coal particles were identified. The most common morphotypes are classified and described as follow: Homogeneous change of reflectance: This is the most common morphotype identified. These are produced when the oxidation reaction takes place over the whole particle, changing the reflectance uniformly. Whilst this kind of particle can be present in low rank and high porosity coal samples, there are significantly more with spontaneous combustion coals. Oxidation rims: This is the second most common morphotype identified. These are produced by a strong oxidation over the coal surface but not internally within the particle. In this case, these alterations are occur when the oxidation reaction is strongly controlled by oxygen diffusion. Cracks and micro fractures: These fractures were less significant or evident, and were produced by the shrinkage of particles induced by the temperature change and the volatile release. These particles were characterized by their perpendicular edges, which could also be perpendicular to a main fracture. High reflectance particles: These morphotypes are produced when the oxidation reaction take place over the whole particle changing the reflectance uniformly, but also inducing a plastic transformation, which increase the reflectance of the whole particle. These kinds of particles are mainly produced by high rank coals.
3.4 Spontaneous combustion assessment A standard assessment is proposed to predict the spontaneous combustion propensity of a coal sample, using automated image capture and manual post processing of images. Grey scale histograms from a specific fresh sample is compared with the oxidized material after the thermal treatment. The fresh histogram is subtracted from the oxidised histogram to highlight the material that has disappeared and the new material that has been generated during heating. Figure 5 shows an example of the results obtained for a highly reactive coal that is prone to self-heating. In this case, this sample is characterized for a high percentage of particles with a homogeneous change in reflectance. The histogram of the treated sample is similar to that of the fresh material profile, but with a higher reflectance peak position (Figure 5, graph i). Additionally, it also can be seen a slight change in the peak magnitude, which is produced by a percentage of particles with oxidations rims in the surface (Figure 5, graph i). Subtracting
the two profiles produces graph ii), to identify the position of the changes. The magnitude of this change is a clear indication of the scale of the change, which is a potential indicator of propensity.
i)
ii)
Figure 5: Histogram analysis of Fenosa coal, a well known coal prone to spontaneous combustion. In this case, the reflectance of the treated sample changes about 80% with regards to the original material.
4. Conclusions An experimental reactor has been designed and implemented in order to quantify thermochemical response of samples under a slow heating ramp rate, simulating a self-heating process. This instrument has enabled the measurement of the crossing point temperature values for a set of coals, as well as several temperatures inside the sample holder (at different positions), in order to estimate the reactivity at low temperature of the samples under study. At the same time, the reactor has been used to produce oxidized material in order to study the change of optical properties of coals as a result of the heating process. The change of reflectance has been studied using petrographic methods, and several different morphotypes have been identified, some of which appear to be characteristic of self-oxidized coals. Finally, the reflectance change measured between the fresh and the oxidized material has been linked with the thermal profiles during heating. This information has been also related to the
concentration of specific thermal alterations (morphotypes) present in each coal. Results indicate that an experimental procedure could potentially estimate the potential of coals to develop a thermal runaway using petrographic analysis and image analysis techniques.
5. References 1 Feng, K., Chakravo, R., and Cochrane, T. Spontaneous Combustion - Coal Mining Hazard. Canadian Institute of Mining Bulletin, 66 (1973), 75-84. 2 Parr, S. and Kressman, F. The spontaneous combustion of coal. The Journal of Industrial and Engineering Chemistry (1911), 152-158. 3 Bhattacharyya, K. The role of sorption of water vapour in the spontaneous heating of coal. Fuel , 50 (1971), 367-380. 4 Carpenter, D. and Sergeant, C. The initial stages of the oxidation of coal with molecular oxygen III-effect of particle size on rate of oxygen consumption. Fuel, 45 (1966), 311-327. 5 Gethner, J. The mechanism of the low temperature oxidation of coal by O2: observation and separation of simultaneous reactions using in situ FT-IR difference spectroscopy. Applied Spectroscopy, 41 (1987), 50-63. 6 Wang, H., Dlugogorski, B., and Kennedy, E. Coal oxidation at low temperatures: oxygen consumption, oxidation products, reaction mechanism and kinetic modelling. Progress in Energy and Combustion Science, 29 (2003), 487-513. 7 Gouws, M. and Wade, L. The self-heating liability of coal: Predictions based on composite indices. Mining Science and Technology, 9 (1989), 81-85. 8 Gouws, M. and Wade, L. The self-heating liability of coal: Predictions based on simple indices. Mining Science and Technology, 9 (1989), 75-80. 9 Jones, J. Recent developments and improvements in test methods for propensity towards spontaneous heating. Fire and Materials, 23 (1999), 239–243. 10 Chen, X. On basket heating methods for obtaining exothermic reactivity of solid materials: The extent and impact of the departure of the crossing-point temperature from the oven temperature. Trans IChemE, 77 (1999), 187-192. 11 Nugroho, Y., Mcintosh, A., and Gibbs, B. Using the crossing point method to assess the self-heating behaviour of Indonesian coals. Twenty-Seventh Symposium on Combustion, The Combustion Institute (1998), 2981-2989. 12 Avila, C. Study of spontaneous combustion of coals by means of Thernogravimetric analysis. (Berlin 2010), Proceedings of the Second International Conference on Coal Fire Research.
Monitoring Hot Spots in Bituminous Coal Stored at Atmospheric Conditions 1,2
,1
Haim Cohen , Uri Green , Franz Gildemeister 4b Shay Wasserman 12345-
1,3
5
4a
Lionel Metzger , Moshe Pesimberg
and
Department of Biological Chemistry, Ariel University Center at Samaria, Ariel, Israel Chemistry Department, Ben-Gurion University of the Negev, Beer Sheva, Israel TU Bergakademie Freiberg, Fakultät 4, Institut für Energieverfahrenstechnik und Chemieingenieurwesen 09599 Freiberg, Germany. a. Rutenberg Power Station, b. Orot Rabin Power Station, Israel Electric Environmental Protection & Licensing Unit, Israel Electric
Abstract Large coal piles (50-150,000 tons) undergo weathering processes during long term storage in open air. Exothermic chemisorption of atmospheric oxygen, formation of surface oxides and the decomposition of inherent coal structure result in release of organic and inorganic gases (e.g. Methane- CH4, Ethylene- C2H4, Ethane-C2H6 Carbon dioxide- CO2, Carbon monoxide- CO and Hydrogen - H2). Some of these reactions are exothermic and when the rate of heat dissipation in the pile is less than that of its formation, a significant increase in temperature can be registered at the locality. These localities are termed "hot spots" where in extreme cases can result in a fire. Monitoring of hot spots at the 2 main coal storage sites of bitumineous coal, Orot Rabin (Hadera) and Rutenberg (Ashkelon), of the Israel Electric Company have been conducted. A unique monitoring unit that can penetrate up to 8 meters into the coal pile and sample gases and measure the temperature was used. The study has shown that the hot spots are formed only in isolated spots and in extreme cases, open fires have been observed. The maximal temperature measured in the hot spots was 330C. Keywords: Coal Weathering, Self Heating, Coal Fires
Introduction In Israel, bituminous coal serves as the main fossil fuel for power generation (>60% electricity production in 2010) in utilities. At present the annual coal consumption is ~13 M tons of coal in these two power stations. As Israel has no coal available the coal is imported by large ships from South Africa (main source), Indonesia, Russia, Australia and Columbia. Furthermore, more than 1 M tons of coal serve as strategic supply in two storage facilities. Like any organic fossil fuel, large coal piles stored under open air undergo weathering processes during long term storage prior to combustion in utility plants. The weathering process is relatively fast and is dependent on the rank of the coal and storage conditions (wind direction, coal pile structure etc.). The geographical location of the mine also proves an important factor as the composition of the coal can differ from mine to mine and weathering is affected appreciably by the coal properties. Weathering processes involve physical adsorption and chemisorption of atmospheric oxygen subsequently forming surface oxides and hydroperoxides which can partially decompose to yield low molecular weight inorganic gases
like carbon oxides (CO, CO2), water, hydrogen (H2) and some organic gases (C1-5) . If the heat formation (due to exothermic processes) is greater than the heat dissipation, self heating of the pile might occur to form hot spots in the pile in which in extreme cases results in fire eruptions(3) as seen in Figure 1.
Figure 1: Photo of stockpiles in Ashkelon storage site affected by open fire and several hot spots
The major product that is released from coals undergoing low temperature aerial oxidation processes is carbon dioxide. This poster presents the monitoring results of temperatures and gases emitted from the storage piles in Ashkelon, in the Rutenberg Power Station of The Israel Electric Company
Experimental Monitoring and measurement of hot spots in Ashkelon Storage piles were conducted in order to obtain a deeper understanding about size and development of hot spots, as well as the determination of existing temperatures and generated gases. Temperature measurement. was conducted both with an Extech IR thermometer (model 42515) coupled with an 8 meter long thermocouple for depth measurements. In order to determine temperatures from within the coal piles, stainless steel pipes with a tailored membrane at the tip were pushed inside the pile and the thermocouple was inserted inside the pipe to measure the temperatures. In order to sample gases from the pile a vacuum pump connected by rubber tubing to a glass sampling cylinder was attached to the protruding stainless steel, the developed measuring unit for monitoring coal piles is presented in Fig 2.
Fig. 2: measuring system for monitoring stockpiles The work at the coal piles is usually divided into two steps. The first step is surface temperature measurements in order to get an overall picture of the apparent thermal status of the coal pile. Smoke releasing chimneys were clear indication of a forming hot spot. After plotting the apparent surface temperature of the pile penetration measurements were conducted in order to determine a more accurate temperature profile of the pile. At several localities the gas phase was sampled by the method described above and analyzed using the following methods. Gas Chromatography. The concentrations of the gases (CO2, CO,, N2, O2, hydrocarbons) in the reactors were determined using a gas chromatograph (Varian model 3800) equipped with a thermal conductivity detector & a flame ionization detector connected in series. The gases were separated on a carbosieve B 1/8”, 9’ ss column using a temperature programmed mode. The experimental error in the G.C. determination is ±5%. The gaseous atmosphere was sampled (1ml samples) after the reaction, with gas tight syringes and measured in the gas chromatograph.
Results A temperature profile of a coal pile in Ashkelon is shown in Fig 3. The temperature line of 55°C was added in order to dist inguish between cold areas and hot spots (T 55°C). A short summary of occurring temperatures i s presented in table 2-3 .
Fig. 3: Temperature profile of coal pile in Ashkelon (2.12.2010)
Distance from epicenter
Fig. 4: Temperature profile of coal pile in Ashkelon (23.1.2011)
Distance from epicenter
Table 1: Occurring temperatures from monitored hot spots
highest median median temp. temp. temp. date hot spot spread hot spot total [m] [°C] [°C] [°C] Ashkelon Dec/10 west 1 105 300 93 east 20 100 240 Jan/11 21,7 160 353 136 On both dates hot spots were detected and the gas content of these spots were sampled and analyzed. The results are presented in Tables 2,3 Table 2: Composition of the gas samples from two hot spots - Ashkelon (2.12.2011)
Sampling depth Temperature [m] [°C] hot spot 1 0.20 167 hot spot 2 0.20 135 1 n.d. = not detected name
CO [vol.-%] 8.97 1.27
CO2 [vol.-%] 12.47 9.75
H2 [vol.-%] 0.30 n.d.
Table 3: Composition of gas samples in Ashkelon taken with at increasing depth in the pile(23.1.2011)
depth temperature O2 CO CO2 H2 [m] [°C] [vol.-%] [vol.-%] [vol.-%] [vol.-%] 0 136 13.1 2.40 13.0 0.14 0.9 179 15.9 1.01 10.7 n.d. 1.5 140 9.6 1.11 12.7 0.03 3.5 136 4.3 1.39 15.8 n.d. 1 n.d. = not detected It is interesting to notice that the highest CO2 concentration of approximately 16 % is found at a depth of 3.5m at 136°C. It can b e assumed that diffusion of gases from the surface to deeper colder layers of the stockpile occurs because of temperature gradients or fluctuations in compaction. The higher concentration of remaining oxygen near the surface is reasonable because of the higher partial pressure of oxygen due to quick setting equilibrium between pile and surrounding air and wind. The highest CO amounts were measured at the pile surface. The Ashkelon stockpiles exhibited a large number of hot spots and several open fires. Such propensity to LTO has not been observed in the Hadera coal storage piles. As a climate comparison (table 4) shows no appreciable difference the reason for this propensity for the Ashkelon piles cannot be attributed to climatic factors. Table 4: Climatic characteristics in Ashkelon and Hadera
Precipitation Dec-Jan 2010/11 [cm/month] Average Temperature Dec-Jan 2010/11 [°C]
Ashkelon
Hadera
4.01
3.25
8 – 32
10 – 30
One of the most noted differences between the coal piles is their relative degree of compaction. The density of a less – compacted stockpile as found in Ashkelon is 850 whilst in Hadera the piles are compacted to within 950 to 1,250 .
[1]
This difference in compaction degrees is due to the different
storage operation equipments used: stacker-reclaimer in Ashkelon PS and bulldozers in Hadera PS. Thus, oxygen diffusion is appreciably slower in the Hadera coal pile is as compared to the Ashkelon storage site.
Conclusion It is clear that besides other existing parameters the different level compaction of the stockpiles has a significant high effect on coal stockpiles undergoing LTO. Furthermore, it is obvious from the above results that the behavior of the carbon oxide formation and equilibrium in the coal piles is extremely interesting and warrants further investigation.
Acknowledgment Financial support acknowledged
1
by
the
Israeli
Electric
Corporation
is
gratefully
. L. Grossman Ph.D. Thesis, “Low Temperature Atmospheric Oxidation of Coal”, Chemistry Department, Ben-Gurion University of the Negev, Beer-Sheva, Israel, (1994)
REDUCING THE ENVIRONMENTAL IMPACT OF SPONTANEOUS COAL COMBUSTION IN COAL WASTE GOBS X. Querol1, X. Zhuang2, Jing Li2, O. Font1, M. Izquierdo1, A. Alastuey1, B.L. van Drooge1, T. Moreno1, J. O. Grimalt1, F. Plana1 1
Institute of Environmental Assessment and Water Research, CSIC, C/ LLuis Solé Sabarís s/n, 08028 Barcelona, Spain, 2Institute of Sedimentary Basin and Mineral, Faculty of Earth Resources, China University of Geosciences, Hubei, 430074, People's Republic of China
[email protected]
ABSTRACT The environmental characteristics of burning coal gangue dumps Shanxi Province, one of the major coal producing area in China and the world, were investigated in this study with the aims of obtaining relevant conclusions on: a) the major air and water quality impacts, as well as b) on remediation strategies. The mineralogy and composition of the original coal gangue, burned wastes and condensates from gaseous emissions, as well as their leaching properties are characterized in 4 different gobs with very diverse degree of reclamation. The results show that the combustion temperature could reach 1200oC in the burning core. Around this core a degasification aureole is formed with temperatures gradually being reduced down to around 100oC near the vents. During combustion in the burning core and degasification of the aureole, elements such as S, F, C, Cl, F, S, N, As, Cd, Hg, Pb, Sn, Ge and Se, and a number of organic pollutants are emitted into the atmosphere. Condensation processes account for the partial trapping of gaseous emissions of PaH, As, S, N, Hg, Pb and Se, among others. Thus, condensates of tar and mineralizations of elemental sulfur and a large variety of salts, including ammonium bearing species, enriched in Hg, Se, As and other trace elements are frequent in the gas vents. Leachates arising from the condensates reach strong acidity or mild alkalinity, depending on the condensate material, posing a serious threat to the environment. The findings show a condensation sequence model for coal gangue fires that may be generalized for most spontaneous combustion of waste gobs, and probably coal dumps and natural coal fires. It was found that covering the coal waste dumps with a layer of compacted soils of around 25 to 50 cm appears to be an excellent cost-effective method to reduce frequency and magnitude of spontaneous combustion. Furthermore, this cover is scavenging or trapping pollutants from gaseous emissions, and minimizing risks associated with the leaching of readily soluble salts condensed on the surface. KEY WORDS: Shanxi coal, China. Geochemistry, spontaneous combustion, leaching
INTRODUCTION The main coal-bearing units in the Datong and Yangquan mining districts belong to Shanxi and Taiyuan (Early Permian) and Datong (Middle Jurassic) formations. During Early Permian time this area was located in the northern part of the North China interior epicontinental sea basin[1]. This was the cause of the relatively high sulfur contents in the Permian coal from the districts[2]. The mining of these coals initially focused mainly on the relatively shallow-lying Jurassic deposits, but these have already been mined out in most mines. Current activity therefore focuses on the deeper and high-sulfur Permian carbonaceous deposits, which calls for special measures to prevent coal fires. Spontaneous combustion of coal may take place when the rate of heat generated by the oxidation of organic matter exceeds the rate of heat dissipation[3]. Most studies point to oxidation of organic matter as the main cause of coal self-ignition (seams or waste), but also other factors such as the heat from the oxidation of inorganic coal-bearing phases e.g. pyrite, could be a key factor in attaining the necessary heat for self-ignition[4]. Prior studies[4-7] have shown that the factors playing a major role in promoting coal oxidation are (a) increase in ambient air temperature, (b) thermal conductivity of coal; (c) grain size of coal and (d) the coal rank. Spontaneous combustion in coal waste dumps from Shanxi was common in the past, however currently most coal waste dumps had been compacted and covered with a layer of compacted soil to the point that it was not easy to find active vents. The land reclamation measures have drastically reduced the occurrence of spontaneous combustion, being thereafter restricted to sporadic and deep fires in the sliding faults developed at the front of coal-waste dumps covered with soil. The spontaneous combustion may take place in coal and coal waste dumps, coal mines, in coal outcrops, in ships transporting coal, and even in peat bogs. These fires constitute an important source of emissions of a large variety of volatile organic and inorganic atmospheric emission of pollutants, mainly CO2, and CO, NO, NO2, SO2, H2S, HF, NH3, HCl, n-alkanes, n-alkenes, sugars, alcohols, PAH, Hg, As, Pb and Se. After volatilization from coal gangue and owing to the sharp fall in temperature, these gaseous organic and inorganic species may condense or chemically interact with the overlying material giving rise to condensates enriched in the above pollutants. Thus, spontaneous coal combustion in coal waste gobs concentrates a number of toxic components in the surface of the gobs by mobilizing them by combustion (in the burning core) and de-gasification (in the thermally altered aureole) of coal and by subsequent condensation on the surface when temperature drops drastically. These pollutants are ready to be leached by rain with the consequent impact on water quality. With this in mind, in this study, coal fires at for different and representative coal mining waste dumps covering a variety of reclamation conditions were investigated concerning the
degree of land reclamation to identify possible threats to the environment and to devise remediation strategies. METHODOLOGY Four coal gobs, with (2) and without (2) top soil reclamation were sampled in terms of the original coal gangue, burned wastes, condensates from gaseous emissions in vents and top soils, and the chemical and mineralogical composition of the samples was determined by ICP-MS, ICP-AES and XRD. Furthermore, leaching experiments using the European Standard leaching test EN-12457 were carried out to all the above samples to determine the leaching potential of major, and trace elements. The pH and ionic conductivity of the leachates were determined by conventional methods. The content of major and trace elements of the leachates were determined by ICP-AES and ICP-MS. The content of Hg was determined directly on leachates by the same procedure as for the solid coal samples using GA-AAS. The content of ammonium was determined in the leachates by using a specific electrode. MAIN RESULTS 1. Based on the mineralogy of the burned wastes (mullite, cordierite, mullite, augite, diopside, amphiboles) in the fire core, we conclude that the combustion temperature could reach up to 1200oC in all 4 cases. 2. Around the core a degasification aureole is formed with temperatures gradually being reduced down to around 100oC near the vents. Thus, thermally altered or degasified coal wastes are found between the fire core (with oxidized combustion wastes) and the surface. 3. During combustion in the burning core and degasification of the aureole, elements such as S, F, C, Cl, F, S, N, As, Cd, Hg, Pb, Sn, Ge and Se, and a number of organic pollutants are emitted into the atmosphere. 4. Mercury tends to be highly enriched (up to 100 ppm) in high ammonium sulfate, ammonium chloride and Al-sulfate condensates, together with Se. 5. Close to the surface of the gob, condensation processes account for the partial trapping of gaseous emissions of PaH, As, S, N, Hg, Pb and Se, among others. Thus, condensates of tar and mineralizations of elemental sulfur, a large variety of Ca, Al-K-Fe sulphates, ammonium sulfate and ammonium chloride, highly enriched in Hg, Se, As and other trace elements are frequent in the gas vents. 6.
Organic condensates have extremely high concentrations of PaH and n-alkanes.
7. After deposition and condensation these unstable organic and inorganic species, these suffer important oxidation giving rise to gypsum crusts with low contents of organic matter.
8. In the two gobs with soil cover, both organic and inorganic condensates are found in the top surface of the buried gob, but in the soil top only ca sulfate (gypsum) was detected as a condensate material. This means that the most of volatile species have condensate in depth and that only SO2 is reacting with the top soil. Consequently, a 25 to 50 cm top soil is probably not only diminishing spontaneous combustion by diminishing the oxygen flux, but also reducing the emissions of gaseous pollutants. 9. Al, K, and Fe-bearing sulfates condensing at gas vents are regarded as the main concern in terms of leaching due to strongly acidic leachates (down to 1.6 pH) and environmentally relevant releases of potentially harmful elements (Al3+, Hg, As, Se, among others). Locally, the occurrence of ammonium chloride yields alkaline leachates up to 8.6 pH. The overall leaching of trace pollutants and sulfate would account for severe inputs to surface and groundwater. Minimizing the exposure of condensates to direct rainfall should be a major target of reclamation activities, thus the top soil also diminish leaching of the condensates. 10. It was found that covering the coal waste dumps with a layer of compacted soils of around appears to be an excellent cost-effective method to reduce frequency and magnitude of spontaneous combustion. Furthermore, this cover is scavenging or trapping pollutants from gaseous emissions, and minimizing risks associated with the leaching of readily soluble salts condensed on the surface. ACKNOWLEDGEMENT This study had financial support from the National Natural Science Foundation of China (Nos. 40572089 and 40972104). REFERENCES (1) Han, D., Yang, Q. Coal Geology of China. Second volume, , 1980, Coal Industry. (2) Mao, J., Xu, H. Evaluation and prediction of Chinese coal resources. Beijing: Science Press, 19991–465 (In Chinese). (3) Misra, B.K., Singh, B.D. Int. J. Coal Geol., 1994, 25, 265-286. (4) Pone, J.D.N., Hein, K.A.A., Stracher, G.B., Annegarn, H.J., Finkelman, R.B., Blake, D.R., McCormack, J.K. Schroeder, P. Int. J. Coal Geol., 2007, 72, 124-140. (5) Querol, X., Izquierdo, M., Monfort, E., Alvarez, E., Font, O., Moreno, T., Alastuey, A., Zhuang, X., Lu, W. Wang, Y. Int. J. Coal Geol., 2008, 75, 2, 93-104.
Plasma Supported Coal Ignition and Combustion V.E. Messerle1, E.I. Karpenko2, A.B. Ustimenko3 1 2
Research Department Plasmotechnics, Almaty, Kazakhstan
Research Institute of Experimental and Theoretical Physics al-Farabi Kazakh National University, Almaty, Kazakhstan 2
Ulan-Ude Division of the Institute of Thermophysics of SB RAS Ulan-Ude, Russia E-mail:
[email protected]
Abstract This work presents new plasma technology for solid fuel ignition and combustion. It promotes more effective and environmental friendly low-rank coal ignition and combustion. To realise this technology at coal fired power plants plasma-fuel systems (PFS) were developed. PFS is pulverized coal burner equipped with arc plasmatron. Temperature of the flame from the plasmatron is varied from 5000 to 6000 K. The base of the PFS technology is plasma thermochemical preparation of coal for burning. It consists of heating of the pulverized coal and air mixture by arc plasma up to temperature of coal volatiles release and char carbon partial gasification. In the PFS coal-air mixture is deficient in oxygen and carbon is oxidised mainly to carbon monoxide. As a result, at the PFS exit a highly reactive mixture is formed of combustible gases and partially burned char particles, together with products of combustion, while the temperature of the gaseous mixture is around 1300 K. Further mixing with the air promotes intensive ignition and complete combustion of the prepared in the PFS fuel. PFS have been tested for boilers start up and pulverized coal flame stabilization in different countries at 30 power boilers of 75 to 950 t/h steam productivity. They were equipped with different types of pulverized coal burners: direct flow, muffle and swirl burners. At PFS testing power coals of all ranks (lignite, bituminous, anthracite and their mixtures) were incinerated. Volatile content of them was in range of 4 to 50%, ash varied from 15 to 48% and heat of combustion was from 1600 to 6000 kcal/kg. To show the advantages of the plasma technology before conventional technologies of coal combustion numerical investigation of plasma ignition, gasification and thermochemical preparation of a pulverized coal for incineration in a power boiler was fulfilled. Two computercodes were used for the research. The numerical experiments were conducted for low-rank
bituminous coal of 40% ash content incinerated at the boiler of 420 ton per hour steam productivity. Comprehensive image of plasma activated coal combustion processes in a furnace of pulverized coal fired boiler was obtained. Both analysis of the numerical experiment and experience of PFS industrial use showed ecological efficiency of the plasma technology. When the plasmatron operates in the regime of plasma stabilization of pulverized coal flame, NOX emission is reduced twice and amount of unburned carbon is reduced four times. 1. Introduction To improve efficiency of solid fuels use, to decrease fuel oil rate in fuel balance of thermal power plants (TPP) and to minimize harmful emissions a plasma technology of coal ignition, gasification and incineration was developed [1, 2]. This technology is plasma thermo-chemical preparation of coal for burning. In the framework of this concept some portion of pulverized solid fuel (pf) is separated from the main pf flow and undergone the activation by arc plasma in a special chamber with plasmatron – PFS (Figs.1 and 2). The air plasma flame is a source of heat and additional oxidation, it provides a high-temperature medium enriched with radicals, where the fuel mixture is heated, volatile components of coal are extracted, and carbon is partially gasified. This active blended fuel can ignite the main pf flow supplied into the furnace. This technology provides boiler start-up and stabilization of pf flame and eliminates the necessity for additional highly reactive fuel.
Figure 1. Sketch of the plasmatron.
Figure 2. Sketch of the Plasma-Fuel System (PFS).
2. Plasma-Fuel System The plasma thermo-chemical preparation of coal is schematically illustrated in Fig. 2. The arc plasmatron (Fig. 1) consists of copper water-cooled electrodes (cathode and anode)
through which the plasma forming air is blown. The plasmatron power is varied from 100 to 200 kW. Its height is 0.4 m, diameter – 0.25 m, and its weight is 25 kg. The measured energy conversion efficiency of the plasmatron is some 85%. Features of fuel-air mixture interaction with arc plasma in the PFS are given in Fig. 3. Across the plasma flame, coal particles with an initial size of 50-100 µm experience ‘heat shock’ and disintegrate into fragments of 5-10 µm. This increases the active interface of the particles, significantly accelerating the devolatilisation (CO, CO2, H2, N2, CH4, C6H6 and others) and 3-4 times accelerates the process of oxidation of fuel combustibles.
Figure 3. Features of arc plasma interaction with air-fuel mixture in the PFS. 3. PFS Industrial Tests PFS have been tested for boilers plasma start-up and pf flame stabilization in different countries at 30 power boilers steam productivity of 75 to 950 ton per hour equipped with different type of pulverized coal burners [2]. At PFS testing power coals of all ranks (lignite, brown, bituminous, anthracite and their mixtures) were used. Volatile content of them varied from 4 to 50%, ash - from 15 to 48% and calorific values - from 6700 to 25100 kJ/kg. For example, the PFS have been implemented in the furnace of a 640 t/h steam full-scale steam raising boiler (Gusinoozersk TPP, Eastern Siberia, Russia). A schematic view of the furnace equipped with the PFS, along with its main dimensions, is shown in Fig. 4. The furnace consists of two symmetrical combustion chambers (semi-furnaces), each provisioned with 8
tangentially directed pf burners in two layers. The combustion chambers are interconnected by a central section. Each burner comprises a primary air/pf delivery section and a secondary air section. Four PFS take the place of the four lower layer burners as shown on the right side of Fig. 4. The plasmatrons operate during the boiler start-up period and in the case of an unstable flame. When the boiler performance is stabilised, the plasmatrons are switched off and the PFS continue to function as conventional pf burners. In the case of flame instability, the plasmatrons are restarted. The fuel was Tugnuiski bituminous coal of 20 % ash content and 35 % of volatile matter. In total, four of the combustors of this TPP were equipped with sixteen PFS. It is estimated that, since 1995, more than 20000 tons of fuel oil has been saved in this facility. This corresponds to a reduction in the emissions of nitrogen and sulphur oxides, carbon monoxide and
7.744 m
33.596 m
vanadium pentoxide of some 13000 tons per year.
18.176 m
Figure 4. Scheme of the industrial furnace of BKZ 640-140 boiler and the boiler furnace equipped with four PFS (top view). Fig. 5 illustrates the scheme of arrangement of the PFS on the boiler combustor BKZ-420 in Ulan-Bator TPP-4 (Mongolia). According to the boiler construction twelve corner-fired burners are placed at three elevations. Two PFS were mounted cornerwise on the lower layer. All eight boilers of the power plant were equipped with PFS for fuel oil free boiler start-up. In 2-3 seconds after light-up with the PFS, the temperature of both pulverised coal flames increased up to 1100-1150 OC. In one hour, the temperature of the flames had achieved 1260-1290 OC and their length reached about 7 - 8 m. In accordance with the operating instructions, the total duration of the boiler start-up was 4 hours.
Fig. 6 demonstrates a scheme of arrangement of three plasma torches on a direct-jet flat-flame pulverised-coal burner of the low layer of BKZ-640 boiler at Gusinoozersk TPP (from the left it is the top view; from the right it is the cross section).
Figure 6. Scheme of arrangement of burners and PFS on BKZ-640 boiler.
1.1
4.0
1.0
3.5 Unburned carbon, %
NOx concentration, g/Nm
3
Figure 5. BKZ-420 boiler furnace equipped with two PFS (top view).
0.9 0.8 0.7 0.6 0.5 0.00
3.0 2.5 2.0 1.5 1.0
0.05
0.10
0.15
0.20
Specific power consumptions, kW h/kg of coal
0.00 0.05 0.10 0.15 0.20 Specific power consumptions, kW h/kg of coal
Figure 7. Specific power consumption influence Figure 8. Specific power consumption onto reduction of nitrogen oxides concentration influence onto reduction of unburned carbon at at plasma aided pulverised coal combustion. plasma aided pulverised coal combustion. Knowledge of the specific power consumption of a plasmatron is required to estimate PFS efficiency. This parameter is defined as the ratio of plasmatron electric power to pf consumption in the PFS. Figs. 7 and 8 present experimental results for NOX reduction and the decrease of unburned carbon during PFS operation versus specific power consumption for the plasmatron. It is seen that the NOX concentration is halved, and the amount of unburned carbon is reduced by a factor of 4. The NOX decrease is caused by the fact that the fuel nitrogen, released from the coal inside the PFS in conditions of oxygen deficiency, forms molecular nitrogen in the gas phase.
Since the fuel nitrogen is evolved inside the PFS and converted to molecular nitrogen there, mainly thermal nitrogen oxides are formed within the combustor volume. However, fuel nitrogen is the main source of nitrogen oxide emission from conventionally-fired pf combustors [3]. As to unburned carbon (Fig. 8), its decrease indicates a fuel reactivity increase which is explained by enlargement of the coal particles reactive surface due to ‘heat explosion’ and fragmentation resulting their interaction with arc plasma. 4. Numerical experiment Results of PFS application at a boiler BKZ-420 of 420 t/h steam productivity of Almaty TPP-2 (Kazakhstan) are presented in this section (Fig. 9). PFS for the boiler BKZ-420 are based on three main burners: two outer burners of the lower layer and a middle burner of the upper layer. PFS are placed in the burner instead of the channel of the primary air-fuel mixture (the inner channel of air-fuel mixture) (Fig. 10). PFS were designed and engineered using two computer-codes, one-dimensional Plasma-Coal [4], that takes into account plasma source and the detailed kinetics of the thermochemical transformations of fuel in two-phase flow, and three-dimensional Cinar ICE [5], that takes into account the geometry of the combustion chamber, the turbulence environment, radiation heat transfer and combustion of coal particles by the model of fast chemistry. Two modes of the boiler operation were chosen for the numerical studies. The first one was traditional regime, using six pulverized-coal burners, and the second one was regime with plasma activation of combustion, using the replacement of three pulverized-coal burners onto PFS. Parameters of highly reactive two-component fuel, derived from the air-fuel mixture in the PFS, were calculated using Plasma-Coal code. They were taken as initial parameters for the three-dimensional calculation of boiler’s furnace equipped with PFS, which were carried out by the program Cinar ICE. Dust of Ekibastuz bituminous coal of 40 % ash content, 24 % volatile, 5 % wet and calorific value of 4000 kcal/kg is burned in the boiler. Fineness of the coal grinding is R90=15 %. It means that averaged diameter of the particles is 60 µm. Initial data for calculation of the PFS by Plasma-Coal program are given in Table 1. As a result of the calculation distribution of temperature and velocity of gas and particles, concentrations of gas-phase components, degree of gasification, and carbon concentration in the coke residue were calculated.
Figure 9. Layout of the pf burners (1) and PFS (1) in the boiler BKZ-420 of Almaty TPP-2. Figure 10. Scheme and layout of two stage PFS for the boiler BKZ-420 of Almaty TPP-2: 1 – channel of the external flow of pf, 2 – secondary air duct, 3 - inlet of pf external flow, 4 – inlet of pf internal flow, 5 - plasmatron, 6 – chamber for pf flow turning, 7 – chamber for plasma chemical preparation of fuel for combustion, 8 - chamber for mixing and thermochemical preparation of fuel, 9 - furnace. Table 1. Initial parameters for PFS computation. Parameter Plasmatron power, kW Air-coal mixture temperature, К Consumption of coal through PFS or internal channel of burner, kg/h Primary air rate, kg/h PFS length, m Pulverized coal composition, mas. % Ash C H2 H2O CO CO2 40.0 46.18 2.63 1.84 3.95 1.4
Value 200 362 6000 8955 3.687 CH4 0.55
C6H6 3.45
It can be seen (Fig. 11) that in the initial part of PFS (X < 0.3 m) gas temperature exceeds the temperature of the particles due to the initial heat exchange of the plasma source with the gas phase. In this case, gas and particles velocities do not increase, almost without distinguishing between them. Later on the heated coal particles devolatilization with simultaneous gasification of carbon coke residue is observed (Fig. 12). Due to oxidation of carbon on the surface of the particles their temperature increases to 1350 K, exceeding the temperature of gas on 400 degrees
(X = 0.5 m). At the PFS exit between gas and particles thermal equilibrium is reached at a temperature of 1025 K and the gas flow velocity reaches 49 m/s, exceeding the particle velocity on 1 m/s (Table 2). Note that the flow velocity at the exit of PFS is much higher than the velocity of air-fuel mixture at the exit of traditional pf burners. Concentration of combustible components (CO, H2, Н, CH4, C6H6) increase with the PFS length, reaching its maximum (10%) at the PFS exit. The concentration of oxidizing agents (CO2, H2O, O2) at the PFS exit is 19.2%. The degree of gasification of coal along the length of PFS increases, reaching 48% at the exit, which is sufficient to produce highly reactive fuel. 100
N2
1400 1200
Ci, %
T, K
1000 1 800
H2
1
CO H2O
O2
C6H6 CH4
600 400 0.0
CO2
10
2
0.1 0.5
1.0
1.5
2.0
2.5
3.0
3.5
X, m
Figure 11. Gas (1) and particles (2) temperature (T) distribution along the PFS (X).
0.01 0.0
0.5
1.0
1.5
2.0 X, m
2.5
3.0
3.5
Figure 12. Gas components concentration (Ci) distribution along the PFS (X).
Table 2. Composition of highly reactive fuel at the PFS exit. Composition of gaseous phase, vol. % H2 CO CH4 C6H6 CO2 H2O N2 1.05 7.75 0.3 0.77 15.6 3.55 70.84 Gas temperature, К Solids temperature, К 1025 1025
O2 0.15
Ash, kg/h Carbon, kg/h 1518 261 Flow velocity, m/s 48.2
The integral characteristics of plasma-chemically prepared fuel for combustion at the exit of PFS (Table 2) were used as initial parameters for the three-dimensional numerical simulation of pf and highly reactive fuel co-combustion in the furnace of power boiler BKZ-420 (Fig. 9) with the aid of the program Cinar ICE. The results of the calculation of the furnace are shown in Fig. 13. There is a difference of temperature fields in two modes of coal combustion in the furnace. At conventional incineration
of coal six pf flames with a maximum temperature of 1852 OС are formed. In Fig. 13 (right) PFS is located on top layer in the center (longitudinal section of the furnace). Effect of PFS appears to change the shape of the flame of highly reactive fuel. High temperature cores of the flames with a maximum temperature of 1588 OС are shifted closer to the burner embrasures and the PFS, as well as to the upper part of the furnace.
Figure 13. Temperature field in the plane of the central burners and the PFS at conventional incineration of coal (on the left) and using three PFS (on the right). Concentration of unburned carbon, which characterizes the completeness of coal burnout, at the outlet of the furnace 16% less in the case of plasma activated regime of pf combustion using three PFS in compare with traditional coal burning. Fig. 14 shows average concentration of carbon dioxide distribution along the furnace height. While using three PFS concentration of CO2 is higher over the entire height furnace and at the furnace exit this excess amounts to 1%. It confirms high efficiency of coal combustion because of its more complete burnout. PFS also improves the environmental characteristics of pf combustion due to 33% reduction of nitrogen oxide emissions. 5. Conclusions Developed, investigated and industrially-tested plasma-fuel systems improve coal combustion efficiency, while decreasing harmful emission from pulverized-coal-fired Thermal Power Plants. NOx and unburned carbon concentrations decrease improves eco economic indexes of TPP.
PFS eliminate the need for expense gas or oil fuels on start-up, stabilisation of pulverized-coal flame and stabilization of liquid slag output in furnaces with liquid slag removal. Power consumption for PFS does not exceed 2% from heat capacity of the reequipped pulverized coal burner and payback period is not more than 18 months.
Figure 14. Averaged СО2 concentration distribution along the furnace height of the boiler BKZ-420: 1 – regime with PFS, 2 – conventional regime of coal incineration. References [1] Karpenko EI, Messerle VE, Ustimenko AB. Plasma-Aided Solid Fuel Combustion // Proceedings of the Combustion Institute, 2007, V.31, Part II, P.3353-3360 [2] Karpenko EI., Karpenko YuE, Messerle VE, Ustimenko AB. Using Plasma-Fuel Systems at Eurasian Coal-Fired Thermal Power Stations // Thermal Engineering, 2009, V.56, N 6. – P.456-461 [3] Tike DH, Slater SM, Sarofim AF, Williams JC. Nitrogen in Coal as a Source of Nitrogen Oxide Emission from Furnace. // Fuel, 1974, 53, P. 120-125 [4] Kalinenko RA, Levitski A.A., Messerle V.E., Polak L.S., Sakipov Z.B., Ustimenko A.B. Pulverized of Coal Plasma Gasification // Plasma Chemistry and Plasma Processing. V. 13. N 1, 1993, P. 141-167. New-York, London, Paris. [5] Messerle VE, Ustimenko AB, Askarova AS, Nagibin AO. Pulverized Coal Torch Combustion in a Furnace with Plasma-Coal System // Thermophysics and Aeromechanics, 2010, V. 17, N 3, P. 435-444
Oviedo ICCS&T 2011. Extended Abstract
Diffuse soil CO2 flux to assess the reliability of CO2 storage in the Mazarrón-Gañuelas Tertiary basin (Spain) J. Rodrigo-Naharro1 *, O. Vaselli2,3, B. Nisi4, M. Lelli4, R. Saldaña1, C. Clemente-Jul5, L. Pérez Del Villar1 1
Environmental Department. CIEMAT. Avda. Complutense, 22. 28040 Madrid (Spain) Department of Earth Sciences, Via G. La Pira, 4. 50121 Florence (Italy) 3 CNR-IGG Institute of Geosciences & Earth Resources, Via G. La Pira, 4. 50121 Florence (Italy) 4 CNR-IGG Institute of Geosciences and Earth Resources, Via G. Moruzzi, 1. 56124 Pisa (Italy) 5 Department of Chemical Engineering and Fuels. ETS Ingenieros de Minas. Universidad Politécnica de Madrid (UPM), Alenza 4. 28003 Madrid (Spain) 2
Abstract Geological storage of CO2 is nowadays internationally considered as the most effective method for greenhouse gas emission mitigation, in order to minimize its effects on the global climatology. One of the main options is to store the CO2 in deep saline aquifers at more than 800 m depth, because it achieves its supercritical state. Among the most important aspects concerning the performance assessment of a deep CO2 geological repository is the evaluation of the CO2 leakage rate from the chosen storage geological formation. Therefore, it is absolutely necessary to increase the knowledge on the interaction among CO2, storage and sealing formations, as well as on the flow paths for CO2 and the physico-mechanical resistance of the sealing formation. Furthermore, the quantification of the CO2 leakage rate is essential to evaluate its effects on the environment. One way to achieve this objective is to study of CO2 leakage on natural analogue systems, because they can provide useful information about the natural performance of the CO2, which can be applied to an artificial CO2 geological storage. This work is focused on the retention capacity of the cap-rock by measuring the diffuse soil CO2 flux in a site selected based on: i) the presence of a natural and deep CO2 accumulation; ii) its structural geological characteristics; and iii) the nature of the cap-
Oviedo ICCS&T 2011. Extended Abstract
rocks. This site is located in the so-called Mazarrón-Gañuelas Tertiary Basin, in the Guadalentin Valley, province of Murcia (Spain)
Therefore the main objective of this investigation has been to detect the possible leakages of CO2 from a deep saline aquifer to the surface in order to understand the capability of this area as a natural analogue for Carbon Capture and Sequestration (CCS).
The results obtained allow to conclude that the geological sealing formation of the basin seems to be appropriate to avoid CO2 leakages from the storage formation.
1. Introduction The scientific community has general accepted that long-term extrapolation in terms of safety of a deep geological storage of toxic industrial wastes, such as high activity radioactive wastes, industrial and mining wastes and even greenhouse gases, can not be satisfactorily done on the basis of short term researches in the laboratory [1]. Therefore, countries affected by these problems have developed methods of investigation which include both short-term tests in the laboratory, where the variables are controlled, as the study of natural analogues.
Although the studies about CO2 natural accumulations are not yet sufficiently developed, some authors [2,3] have included in their works the existing CO2 reservoirs in the world and the experimental reactions between CO2 and the storage formations [4]. Moreover, in the last decade there are many works focused on the evaluation of the safety of a CO2 geological storage by means of the study of CO2 leakage natural analogues [5-11]. Regarding Spain, there’s one current important project cofunded by the Ministry of Science and Innovation and FEDER European Funds, whose main objective is the global study of the several CO2 natural analogues in all over the country. Among them, the natural analogue of storage, and natural and artificial leakage of CO2 located in the – Gañuelas-Mazarrón Tertiary basin (Province of Murcia) is being studied by the CIEMAT reseach team (Fig. 1). The CO2 diffuse flux in the soil by means of a WEST-
Oviedo ICCS&T 2011. Extended Abstract
SYSTEMS fluxmeter has been performed in the above-mentioned site, in order to know whether the cap-rock is able to retain possible escapes of CO2 at the surface.
Fig. 1. Geographical location of the study area (red square)
2. Experimental section In the Gañuelas-Mazarrón Tertiary basin, according to the structural geological features [12,13], four areas were selected for a comprehensive CO2 flux study. They are located at the intersection of high density lineaments (Fig. 2) that should likely correspond to preferential leakage paths of deep-seated CO2. These areas are: Las Moreras, La Majada and Leiva (Fig. 3), which are at the contact between the Tertiary basin and the Triassic surroundings mountains, and the El Saladillo place, situated inside the GañuelasMazarrón basin.
The equipment used for CO2 flux measurements is that licensed by West-System and consists in an accumulation chamber from where the soil gas is forced to be pumped through an IR cell set at the wavelength of CO2. The increase of CO2 with time allows the measurement of the flux by means an algorithm that takes into account the pressure and temperature data collected in the field [14].
Oviedo ICCS&T 2011. Extended Abstract
The CO2 soil fluxes were carried out in September 2009 and March 2010 during dry and meteorologically stable periods in order to avoid the possible influence of variations induced by environmental parameters on soil degassing. Laboratory experiments were performed to assess both, the reliability of CO2 flux measurements and the calibration of the instrument [14].
QUATERNARY. Conglomerates, sands, slimes and clays
Figs. 2 and 3. Location of the areas where it has been measured the CO2 flux (left) [13] and their respective schematic geological situation (right)
NEOGENE. Volcanic Rocks: dacites and andesites NEOGENE. Carbonates and marls NEOGENE. Limestones
3. Results and Discussion
In September 2009 the CO2 flux soil was computed for a surface of ~52,700 m2 in Las Moreras; ~86,800 m2 in La Majada; ~179,600 m2 in Leiva; and ~136,000 m2 in El Saladillo. In these areas, 127, 277, 257 and 187 evenly distributed measurements were done, respectively. In March 2010, the investigation in La Majada and Leiva areas was enlarged with 93 and 94 measurements, covering additional surfaces of ~39,000 m2 and 30,000 m2, respectively. The measured φCO2 at Las Moreras oscillates from 0.007 to 0.929 moles m-2 day-1, with an average value of 0.262 moles m-2 day-1, while at El Saladillo they were spanning between 0.020 and 1.103 moles m-2 day-1, with an average value of 0.353 moles m-2 day-1. At La Majada a large interval of variation was observed in September 2009, ranging from
Oviedo ICCS&T 2011. Extended Abstract
0.007 to 7.503 moles m-2 day-1, with an average value of 0.877 moles m-2 day-1; whereas, in March 2010, a lower interval, between 0.025 and 1.425 moles m-2 day-1, was observed, being its average value of 0.456 moles m-2 day-1. Finally, at Leiva the φCO2 values varied between 0.024 and 1.490 moles m-2 day-1, with an average value of 0.391 moles m-2 day-1 (September 2009) and between 0.041 and 1.074 moles m-2 day-1 with an average value of 0.310 moles m-2 day-1 (March 2010). In order to better constrain the total φCO2 and the CO2 spatial distribution overall the investigated areas, the values are divided in populations according to the method proposed by Sinclair [15]. The diffuse φCO2 values in the four investigated areas were lower than 1.0 moles m-2 day-1, whereas values up to 7.5 and 1.49 moles m-2 day-1 were measured in September 2009 at the La Majada and Leiva areas, respectively. It is worthy to mention that φCO2 values higher than 1 moles m-2 day-1 were only sporadically recorded. 4. Conclusions On the basis of the diffuse soil CO2 degassing surveys carried out in September 2009 and March 2010, the general picture emerging from the present study is that in the area under study, although characterized by a complex geological setting, the efficiency of the cap-rock, as sealing formation, in the Gañuelas-Mazarrón Tertiary basin does not allow any relevant CO2 leakages at the surface. That is, in terms of CO2 soil flux, the Tertiary sedimentary deposits filling the basin act then as an impermeable layer through which the escape of CO2 is not jeopardized. This is strongly supported by the measurements of the φCO2 carried out by means of the accumulation chamber method. The investigated areas have generally low φCO2. They are basically comparable to those observed in cultivated areas worldwide, with very few exceptions that can possibly be related to structural weakness or fault zones. Nevertheless, this statement is not sufficiently supported by the available data. It is however matter of fact that the geological sealing formation results to be effective and efficient in case of any leakage of CO2.
Oviedo ICCS&T 2011. Extended Abstract
Acknowledgements We wish to thank to the Ministry of Science and Innovation of Spain and the European Union FEDER Funds for supporting this study. This work was carried out within the Project PSE-CO2, whose objective has been the development of the Carbon Capture and Storage (CCS) Technologies. References [1] Petit JC. Reasoning by analogy : rational foundation of natural analogue studies. Appl. Geochem 1992; Supplementary Issue 1:9-12. [2] Czernichowski-Lauriol I, Sanjuan B, Rochelle C, Bateman K, Pearce J, Blackwell P. The underground disposal of carbon dioxide. In: Holloway S, editor. Inorganic Geochemistry. Final Report of Joule II Project Nº CT92-0031. [3] Pearce JM, Holloway S, Wacker H, Nelis MK, Rochelle C, Bateman K. Natural occurrences as analogues for the geological disposal of carbon dioxide. Energy Converse and Management 1996;37:6-8. [4] Pearce JM, Rochelle C. CO2 storage: mineral reactions and their influences on reservoir permeability. A comparison of laboratory and field studies. Elsevier; 1999. [5] Czernichowski-Lauriol I, Pauwels H, Vigouroux P, Le Nindre YM. The french carbogaseous province: an illustration of natural processes of CO2 generation, migration, accumulation and leakage. Greenhouse Gas Control Technologies. Vols I and II, Proceedings 2003;411-416. [6] Hawkins, DG. No exit: thinking about leakage from geologic carbon storage sites. Energy 2004;29:1571-1578. [7] Beaubien SE, Lombardi S, Ciotoli G, Annuziatellis A, Hatziyannis G, Metaxas A et al. Potential hazards of CO2 leakage in storage systems-Learning from natural systems. Greenhouse Gas Control Technologies 2005;7:551-560. [8] Nordbotten JM, Celia MA, Bachu S, Dahle HK. Semianalytical solution for CO2 leakage through an abandoned well. Environmental Science & Technology 2005;39: 602-611. [9] Oldenburg CM, Lewicki JL. On leakage and seepage of CO2 from geologic storage sites into surface water. Environmental Geology 2006;50:691-705. [10] Riding JB. The IEA Weyburn CO2 monitoring and storage project - Integrated results from Europe. Advances in the Geological Storage of Carbon Dioxide: International Approaches to Reduce Anthropogenic Greenhouse Gas Emissions 2006;65:223-230.
Oviedo ICCS&T 2011. Extended Abstract
[11] Lewicki JL, Birkholzer J, Tsang CF. Natural and industrial analogues for leakage of CO2 from storage reservoirs: identification of features, events, and processes and lessons learned. Environmental Geology 2007;52:457-467. [12] Pérez del Villar L, Pelayo M, Recreo F. Análogos Naturales del Almacenamiento Geológico de CO2 (Fundamentos, Ejemplos y Aplicaciones para la Predicción de Riesgos y la Evaluación del Comportamiento a Largo Plazo). CIEMAT; 2007. [13] Pérez del Villar L. Memoria Científico-Técnica del periodo 2008-2009 del PSE120000-2008-6 (PSS-120000-2008-31). Línea de Análogos Naturales: "Resultados preliminares del estudio de los análogos naturales estudiados en: la región de La Selva (Girona), Valle del Alto Guadalentín (Murcia-Almería), Alicún de las Torres (Granada), Alhama de Aragón-Járaba (Zaragoza) y Castilla León” CIEMAT; 2009. [14] Nisi B, Vaselli O, Lelli M, Tassi F, Rodrigo-Naharro J, Pérez del Villar L. Diffuse CO2 flux and dissolved gases in the Mazarrón-Gañuelas area (Guadalentin Valley). Report PSE; 2010. [15] Sinclair AJ. Selection of threshold values in geochemical data using probability graphs. Geochem. Expl. 1974;3:129-149.
Carbon and Storage by pH swing aqueous mineralisation using a mixture of ammonium salts Aimaro Sannaa,b *, Marco Dri a,b, Xiaolong Wang a, Matthew R Hall b, Mercedes Maroto-Valer a a
National Centre for Carbon Capture and Storage, The Sir Colin Campbell Building, Innovation Park, University of Nottingham, Nottingham, NG7 2TU, UK
b Nottingham Centre for Geomechanics, Division of Materials, Mechanics and Structures, Faculty of Engineering, University of Nottingham, Nottingham, NG7 2RD, UK
Corresponding author: Aimaro Sanna, Energy and Sustainability Research Division, Faculty of Engineering, University of Nottingham, Nottingham, NG7 2RD, UK, Tel: +44 (0)115 9514198,
[email protected] ; Other authors: Marco Dri, Energy and Sustainability Research Division, Faculty of Engineering, University of Nottingham, Nottingham, NG7 2RD, UK, Tel+44 (0)115 9514198,
[email protected] ; Xiaolong Wang, Energy and Sustainability Research Division, Faculty of Engineering, University of Nottingham, Nottingham, NG7 2RD, UK, Tel+44 (0)115 9514198,
[email protected] ;Matthew R Hall, Division of Materials, Mechanics and Structures, Faculty of Engineering, University of Nottingham, Nottingham, NG7 2RD, UK, Tel: +44 (0)115 846 7873, Fax: +44(0)115 951 3159,
[email protected] ; Mercedes Maroto-Valer, Head of Energy and Sustainability Research Division, Faculty of Engineering Director of Centre for Innovation in Carbon Capture and Storage (CICCS), University of Nottingham, University Park, Nottingham NG7 2RD, UK, Tel: +44 115 846 6893, Fax: +44 115 951 4115,
[email protected].
Abstract
Carbon capture and storage by mineralisation (CCSM) focuses on mixing the carbon dioxide (CO2) in the flue gases with rocks rich in magnesium oxide. The oxides react with CO2 producing solid mineral carbonates, which are stable solids and can provide safe storage capacity on a geological scale. The aim of this work was to optimize a pH swing mineralisation processes developed at the University of Nottingham that uses recyclable ammonium salts to widespread implementation of the technology. Carbonation using a mixture of NH4HCO3 and (NH4)2CO3 under different temperatures was investigated considering that these solids are the expected intermediate solid products of the chilled ammonia capture process. The highest carbonation efficiency was 61.5% and this value is lower than that obtained when using only NH4HSO4 (70-80%).
1. Introduction
Carbon capture and storage (CCS) by geological storage has the potential to sequester about 50% of the CO2 emission per year in Europe by 2030 [1]. The state of the art indicates that this process can be applied mainly to large emitters, while it is less appropriate for smaller emission sources [2]. CCS by mineralisation (CCSM) can sequester CO2 by mixing the CO2 from flue gases with industrial solid waste rich in oxides or rocks rich in magnesium or calcium oxides. The oxides react with CO2 producing solid mineral carbonates, which are
stable and can provide safe storage capacity on a geological scale. It is estimated that the global magnesium silicate rock deposits are enough to sequester the CO2 generated by all the fossil fuels resources. CCSM can therefore contribute to decrease of CO2 emissions in areas where geological storage cannot be deployed and also can be applied to small industrial emitters [2]. The aim of this work is to investigate the technical barriers for the deployment of CCSM, such as the high energy demand for pre-treatment of the minerals and the slow kinetics of direct gas-solid reactions. Indirect CCSM by pH swing using ammonium salts has been recently investigated to enhance the efficiency of both dissolution and carbonation, resulting in 70-80% CO2 sequestered [3]. The process shown in Figure 1consists in 3steps: first, the dissolution of minerals rich in magnesium bringing it in the solution as MgSO4. Second, the removal of the impurities (Fe, Al, Mn etc.) by increasing the pH from acidic to basic and finally, the carbonation of the MgSO4 with ammonium carbonates coming from the CO2 capture stage producing hydromagnesite. Furthermore, this method can recycle most of the chemicals used during the process. In this study, further research is conducted in this process with the objective of optimise the operating variables. CO2
NH3
Capture
NH4HSO4
NH4HCO3 (NH4)2CO3 Serpentine H2O
Dissolution Silica
pH swing
MgSO4
Impurities (Fe, Al, Mn, etc.)
Regeneration
(NH4)2SO4
Carbonation Hydromagnesite
Figure 1 Scheme of the pH swing process used in this work.
The capture of CO2 by ammonia-based wet scrubbing is similar to the capture process using amines. Ammonia and its derivatives react with CO2 and water to form different solids at temperature below 10°C [4]. This process can precipitate several ammonium carbonate compounds in the absorber, mainly ammonium bicarbonate (NH4HCO3) and ammonium carbonate ((NH4)2CO3). Under CO2 loadings higher than 0.5 both ammonium salts are
present, while when the CO2 loadings is lower than 0.5 the only solid is ammonium carbonate [5]. Therefore, the chilled ammonia process is likely to produce a mixture of ammonium carbonate and bicarbonate. Therefore, this study wants investigate the carbonation of CO2 in presence of both salts at different temperatures and to compare with previous studies carried out using ammonium bisulphate. A series of dissolution and carbonation experiments were performed in a batch reactor under different temperatures (50, 70, 100°C) to evaluate its effect on the dissolution of serpentine mineral (1st reaction limiting step) and on the silica product layer diffusion (2nd reaction limiting step). The products of the reaction were then analysed to establish the mass balance and the overall CO2 sequestration efficiency compared with previous work [3,6]. 2. Experimental section
200g of serpentine with particle size ranging from 75 to 150μm were added into 4000mL 1.4M NH4HSO4 solution for a 1:20 solid liquid ratio after the solution temperature was stabilised at 100°C. A solution sample was extracted after 5, 10, 15, 30, 60, 120 and 180 minutes to establish the content of Mg and other ions in the solution by ICP-MS. After dissolution, the flask content was filtered with a 0.7μm Pall syringe filter and the solid was dried for 24 hours at 105°C and then analysed at the TGA. After the dissolution experiments, the impurities were removed by adding ammonia-water to rise the initial acid pH of the solution to neutral state and then to precipitate all the impurities such as iron, manganese and aluminium [3]. The carbonation experiments were carried out at 50, 70 and 100°C using the set-up shown in Figure 2. 200mL of the solution produced in the dissolution experiments was poured into a 250mL 3-necks flask and heated up at the required temperature by a silicon-oil bath under continuous stirring at 800rpm. As soon the temperature was stabilised and the pH was higher than 8, the stoichiometric mixture of 50% (NH4)2CO3 and 50% NH4HCO3 was added to the solution to start the carbonation reaction. Sampling and solutions analysis were done as for the dissolution experiments [3]. The carbonate content in the starting serpentine and final products was investigated by a TGA analyses were performed using a thermal gravimetric analyzer (TGA Q500, TA Instrument). The weight loss in the temperature region of 350°C to 500°C was determined to be carbonates [6]. The elemental composition of the serpentine used in this work was obtained from the literature [7, 8].
5
6
4 3
7
2 1 Figure 2 Carbonation set-up. 1 Temperature controller, 2 silicon-oil bath, 3 3-necks flask, 4 pH electrode, 5 thermo-couple, 6 cooling system, 7 sampling point.
3. Results and Discussion
Figure 3 shows the extraction of magnesium and iron carried out in triplicates. The higher magnesium extraction was obtained after 3hrs with 85% of the Mg removed from the serpentine particles. The three runs present the same dissolution trend indicating that a fast extraction occurs in the first 30 minutes with sequent slower dissolution afterwards. In fact, 70% of the magnesium is extracted in 30 minutes and only 15% more is extracted after 2 hours. The 3rd hour does not enhance the serpentine dissolution in a sensible way as can be seen in Figure 2. Therefore, 2 hours is the optimal dissolution time to reach the higher Mg extraction. However, 30 minutes might be preferred for a full scale mineral carbonation plant considering the trade off between the lower capital costs associated with faster kinetics [6] and the lower efficiency (70%) after 30 minutes extraction. It should be noted that the leaching of 200g of serpentine resulted in a lower extraction efficiency compared to the previous dissolution experiments carried out using 20g of serpentine [3]. This may be due to the different particle size distribution compared to previous work. The second stage of the overall process was the pH swing from acidic (pH 0.2) to neutral (pH 7) (reaction ii) and then basic (pH ≥8) by addition of ammonia water to remove all the impurities (Fe, Al, Zn, Ni, Cu, Mn etc.) from the solution.
Mg‐1
Mg‐2
Mg‐3
Fe‐1
Fe‐2
Fe‐3
90 80
Mg, Fe, %
70 60 50 40 30 20 10 0 0
40
80
120
160
200
Time, min
Figure 3 Magnesium and iron extraction efficiency during the 3hrs experiment.
Figure 4 shows the concentration of Fe, Al and Mn as a function of the pH. About 20% of Fe and Mn and 10% of Al precipitated when the pH increased from 0.2 to 2.5. The concentration of Fe, Mn and Al remained stable until pH approached 8. Then, at pH higher than 8 all the other Fe, Al and Mn precipitate as hydroxides leaving the solution of MgSO4 ready for the carbonation reaction. Fe
Al
Mn
80 70
Fe, Al, Mn, %
60 50 40 30 20 10 0 0
2
4
6
8
10
pH
Figure 4 Precipitation of iron, manganese and aluminium in function of the pH.
The last step of the process was the carbonation of the MgSO4 solution using a 50/50 mixture of NH4HCO3 and (NH4)2CO3.
Figure 5 shows how the concentration of magnesium in the solution decreases during the carbonation experiments due to the precipitation of the final product (MgCO3 Mg(OH)2 H2O). Only 10% and 30% of the total Mg precipitated in the first 5 minutes of the reaction in the experiments at 50 and 70°C, respectively. On the contrary, just after 5 minutes 80% of the total Mg in solution was precipitated at 100°C. This indicates that the temperature significantly influence the carbonation reaction. Figure 5 was also used to calculate the carbonation efficiency at the different temperatures. An efficiency of 51.6%, 58.5% and 76.5% was calculated considering the dissolved and precipitated magnesium. However, not all the magnesium was converted into carbonate. 50°C
70°C
100°C
120
Mg, wt%
100 80 60 40 20 0 0
10
20
30 Time, min
40
50
60
Figure 5 Variation of the magnesium concentration during carbonation at different temperatures.
The final carbonation efficiency (Mg in MgCO3 (after carb.) – Mg in MgCO3 (before carb.) / Mg in Serpentine *100) at 100°C was 61.5% that was much higher than at 50°C (39%) and at 70°C (41%). Therefore, the mineral carbonation at 100°C would require a low amount of starting serpentine (4.8t/tCO2) compared to that required at 50 and 70ºC. However, the carbonation efficiency using only NH4HCO3 under the same conditions was found higher (70-80%) [6], indicating that the carbonation using (NH4)2CO3 might require higher process conditions to reach similar efficiencies.
4. Conclusions
The results indicate that the pH swing process can be used to carbonate the CO2 trapped in the mixture of NH4HCO3 and (NH4)2CO3 produced during the chilled ammonia capture
process. However, the efficiency of the overall capture and storage process using the salts mixture resulted lower compared to that obtained by using only NH4HCO3.
Acknowledgements The work presented here was funded by the Centre for Innovation in Carbon Capture and Sorage (EPSRC grant EP/F012098/1). References
[1] Stangeland A, A model for the CO2 capture potential, International Journal of Greenhouse Gas Control, 1, 2007, 418-429. [2] IPCC, 2005: IPCC Special Report on Carbon Dioxide Capture and Storage. Prepared by Working Group III of the Intergovernmental Panel on Climate Change [Metz, B., O. Davidson, H. C. de Coninck, M. Loos, and L. A. Meyer (eds.)]. Cambridge University Press, Cambridge, United Kingdom and New York, NY, USA, 442 pp. [3] Wang X and Maroto-Valer M.M, Dissolution of serpentine using recyclable ammonium salts for CO2 mineral carbonation, Fuel 90 (2011) 1229–1237. [4] Figueroa J.D, Fout T, Plasynski S, McIlvried H, Srivastava R.D, 2008, Advances in CO2 capture technology—The U.S. Department of Energy’s Carbon Sequestration ProgramS, International journa l of greenhous e gas control , 2, 9-20. [5] Darde V, Thomsen K, van Well W and Stenby E, Chilled ammonia process for CO2 capture, Preprint – ICPWS XV Berlin, September 8–11, 2008 [6] Wang X and Maroto-Valer M, Integration of CO2 capture and storage based on pH– swing mineral carbonation using recyclable ammonium salts, Energy Procedia 4 (2011) 4930–4936. [7] Gerdemann S J, Dahlin D C, O’Connor W K and Penner L R, Carbon dioxide sequestration by aqueous mineral carbonation of magnesium silicate minerals, DOE-ARC, 2003, 018, http://www.osti.gov/bridge/ [8] Teir S, Revitzer H, Eloneva S, Fogelholm K-J, Zevenhoven R, Dissolution of natural serpentinite in mineral and organic acids, Int. J. Miner. Process. 83 (2007) 36–46.
New equipment for characterization of rocks for geological CO2 storage in coal seams Cienfuegos, P. & Loredo, J. Mining and Exploration Department – University of Oviedo - SPAIN
[email protected] Abstract The geological storage of CO2 in coal seams is an emerging option in the portfolio of mitigation actions for reduction of atmospheric greenhouse gas concentrations. A background study focused to the selection of favorable sites for CO2 geological storage are necessary steps, and in the selection of reservoirs for CO2 sequestration a complete petrophysical characterization of the sample is necessary. To complement the classical petrophysical parameters measured on the rocks of the geological formation with potential to be used to store the injected CO2, a new equipment has been designed and constructed to simulate at a laboratory scale the interaction between the rock and the injected CO2, at different pressure conditions simulating depths of the geological formations up to 1 000 meters. Essays focused to study the alterability of the rock in contact with CO2 either in subcritical or supercritical state, as well as essays for CO2 injectivity on the rock can be accomplished. 1. Introduction Tests and studies on characterization of rocks have been used to understand the interaction between physical properties, the chemical composition of the rock and the industrial use. The aim is to determine the measurable properties of the rock with special interest in this new industrial application. Two more interesting aspects of the geometry of space are porosity and the study of the interaction of the rock with supercritical CO2 (storage conditions). Therefore, we must consider the status of the rock before and after its contact with CO2, and its behavior, as it is important to determine the new conditions. The first phase consisted on the characterization of technology of fresh rock samples, and in the second phase, will be identical characterization changes after the test. The research began with a petrological study indicates that rock your petrographic analysis, microfractographic, morphological and chemical. Subsequently, we tests of water and mechanical properties parallel. These tests allow us to obtain the value of open porosity of the sample and its degree of saturation, deducted as well as important information geometry its empty spaces (pores and fissures) of great interest in the geological storage of any fluid. The period for the completion
of these tests can vary from one to three months, depending on the type of rock. Mechanical tests are faster and they get information that allows determining the elastic constants of tested rocks. New test equipment was necessary to develop the interaction of the rock with CO2. 2. 2. Experimental section: Design of a new equipment for complementary petrophysical characterization of rocks New equipment, called RockTestCO2, has been designed to carry out the study of the interaction of CO2 in rocks at high pressure conditions, simulating the process that occurs during the injection of CO2 in underground geological formations (figure 1).
Figure 1. Photogram and diagram of operation of test equipment named ROCKTESTCO2. The central part of the equipment consists of a AISI 304 stainless steel chamber with internal dimensions of 320x320x320 mm, 40 mm thick and an approximate weight of 200 kg. A sample rock is introduced with approximate dimensions of 270x270x270 mm. It also provides two 280x280 mm racks of stainless steel AISI 304 for the fixing of prismatic test specimens of 50 mm square and cylindrical samples with a diameter of 54 mm (figure 2).
Figure 2. Details of stainless steel chamber.
The CO2 is introduced into the chamber by a rod directly embedded into a drill that will have been previously done on the rock (figure 3, left), and it is inserted through the central part. The water enters the bottom of the camera. There are two metering pumps Dosapro mark MILTON ROY, with a flow rate up to 7.5 L / h, a pressure up to 300 bar and a AISI 316 L stainless steel dispenser (figure 3, right). These two pumps are lubricated in oil in a waterproof cap and speed reducers are built into the mechanics.
Figure 3. Details for two metering pumps “Dosapro”. The closing of this camera is achieved through the action of a hydraulic cylinder, and the end of this cylinder has a tray on which the rock is placed to study, so that the action of hydraulic cylinder allows both entering the rock inside the camera and closing it. As a safety feature, there is a switch that lets you keep closing pressure of the chamber throughout the procedure, preventing accidental opening of the system. It also has a ruptured disk and additional elements for protection and security. A camera 10 is coupled 180 W resistors and two 200 W resistors to heat the interior of it until the desired temperature, using a temperature sensor and a digital controller. The hydraulic closure is capable of exerting a force of over 95 tons, and it consists of a hydraulic cylinder of 200x140 mm and 500 mm in length, a tank with its accessories, filter screens, levels, etc. There is a motor-group with two pumps, and the low pressure gives the flow required for rapid movement and a pressing cylinder for 320 bar, as well as control valves and the associated control. The entire assembly approximately weighs 1 300 kg. This equipment includes a cabinet where both switches activation of different elements by the user, such as electrical control elements and protections necessary to ensure the safety of the people and the equipment according to current regulations (figure 4).
Figure 4. ROCKTESTCO2 test equipment. Basically, the new test consists of contacting the rock sample with CO2 in the natural reservoir conditions of pressure and temperature. The operational parameters of this equipment allow simulating in the laboratory the geological storage of CO2 in the natural conditions of pressure, temperature and salinity. 3. Results and Discussion Five samples of rocks (carbonates and sandstones) are used in initial tests. It is necessary to know the physical properties of these rocks in their natural state. For this study performed petrological and petrophysical studies and chemical and morphological analysis. Subsequently, subjected to interaction with CO2 for a variable duration from weeks to months, and finally, it is necessary to re-make all petrophysical tests chemical and morphological analysis to obtain differences by comparison of the physical properties and chemical composition of rocks (table 1). The physical properties are possible to observe the variation in short term tests of several months. Possible variations in chemical composition need more time. Table 1. Data indicate changes in parameters of physical properties as density and porosity. PARAMETER Density mean , ρd Desv. (Kg/cm3) Open mean porosity, no (%) Desv.
A 2638.24
ACO2 2629.73
B 2491.10
SAMPLES BCO2 C 2487.37 2614.37
CCO2 2577.11
D 2771.86
DCO2 2756.24
7.19
8.79
8.26
8.70
39.68
33.96
4.46
4.40
4.62
4.99
7.71
8.89
4.54
6.42
0.56
1.80
0.43
0.48
0.21
0.73
0.82
1.30
0.06
0.08
Preliminary results of petrophysical properties confirm those obtained by other laboratories [1], [2]. It shows a slight increase in effective porosity and a decrease in the mechanical strength.
4. Conclusions The design and construction of this equipment ROCKTESTCO2 allows us to investigate known physical and chemical processes that occur between the rocks store/seal and the CO2 injected into geological storage of CO2. Simulation of storing CO2 deep is essential to know the behavior of rocks (store or seal) in its interaction with the CO2. Finally, besides the mathematical modelling, it is necessary to develop a new petrophysical characterization equipment to simulate the pressures and temperatures at which rocks are the target of a geological storage of CO2. References 1.
Benson, S.M., L. Tomutsa, D. Silin, T. Kneafsey and L. Miljkovic: Corescale and Porescale Studies of Carbon Dioxide Migration in Saline Formations, Proceedings of 8th International Conference on Greenhouse Gas Control Technologies (GHGT8), IEA Greenhouse Gas Programme, Trondheim, Norway, June 19-22, (2006).
2. Hovorka, S.D., C. Doughty, S.M. Benson, K. Pruess, and P.R. Knox: “The Impact of Geological
heterogeneity on COB2B Storage in Brine Formations: A Case Study from the Texas Gulf Coast,” Geological Storage of Carbon Dioxide, S.J. Baines and R.H. Worden (eds.) Geological Society, London Special Publications, 233, p. 147-163. (2004).
Oviedo ICCS&T 2011. Extended Abstract
The Current State of Affairs of Coal Research in U.S. Universities Jonathan P. Mathews1, Bruce G. Miller1, Chunshan S. Song1, Harold H. Schobert1, Francois Botha2, and Robert B. Finkleman3 1
John and Willie Leone Family Department of Energy & Mineral and EMS Energy Institute, 126 Hosler Building, The Pennsylvania State University, University Park 16802, USA.
[email protected] 2 Illinois Clean Coal Institute, 5776 Coal Drive, Suite 200, Carterville, IL 62918, USA 3 University of Texas at Dallas, Richardson, TX 75080, USA Abstract An ISI Web of Knowledge (using the web of science database) evaluation of journal articles with “coal” in the title (in English-language journals) was performed, for the periods 1970 to 2010 and 2000 to 2010, as one approach to evaluate the historically and currently active research centers in coal science. The approach underestimates the contributions but provides a basis for comparisons. Contributions were broken down by research institutions and by countries of contributing authors using analysis tools within ISI Web of Knowledge. The United States has 30% of the publications (between 1970 to 2010), with Japan contributing 7.8% and the People’s Republic of China 7.5%. England, Australia, India, Canada, Poland, Spain, Germany, and France contributed between 5 to 2% each. However, China has been leading the publication production since 2006. Japan’s publications are relatively steady while U.S. publications have declined after a high point during 1984-2002 when a small resurgence occurred. This paper reports on the coal science contributions of U.S. universities and determines journal article publication activity level and focus. The leading academic institutions’ publication records were evaluated using Wordle to visually determine research focus areas through word frequency analysis of the journal article titles. Similarly, active authors (for the period of analyses) were identified using this approach. There has been a significant decline in the U.S. academic institutions that are active in coal science. Among the academic entities, the leading institutions (quantity of journal articles) were The Pennsylvania State University (Penn State), Kentucky, West Virginia, Southern Illinois, MIT, Utah, Brigham Young, Pittsburgh, Illinois, Ohio State, Virginia Tech, Wyoming, Auburn, Carnegie Mellon, North Dakota, Iowa, California at Berkeley, Tennessee, Texas, Purdue, Texas A&M, Missouri, Georgia, and Western Kentucky. Of these, about half are still “active” in “coal science” as defined as 15 coal science publications (by this search approach) between 2000 to
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Oviedo ICCS&T 2011. Extended Abstract
2010 where the focus is science and engineering research on coal but does not include coal-related areas such a post-combustion pollution control or the various catalytic aspects of coal gasification products. 1. Introduction The U.S. is second only to China in the amount of coal produced and consumed, with about 1.1 billion short tons consumed annually. Projections estimate a 6% increase in coal use between 2008 and 2035 [1]. Even if constraints on CO2 emissions were to be mandated in the U.S., substantial amounts of coal will still be used for the foreseeable future. Coal remains the only strategically secure major source of energy in the U.S. If energy security concerns are high or oil prices rise, higher projections are likely [1]. Although coal science is a relatively mature field new perspectives and more detailed understanding of coal characteristics are necessary to address new or alternative mining approaches, beneficiation, clean(er) combustion technologies, growing environmental concerns, the use of lower quality coals, and various emerging societal issues. In short, coal science can help to ensure the efficient, cost effective, environmentally compatible use of the abundant coal resources domestically and internationally. Coal research is faced with very significant challenges: substantial retirements of senior scientists, the cyclic nature of funding; traditionally conservative attitudes of the mining industry; and shifts in government funding towards renewable energy research. Many coal researchers, who started their careers in the energy crisis of the 1970s are now retired or soon will be. The mass influx of these coal researchers in the late 1970s and early 1980s has resulted in a ‘missing generation’ of coal scientists as most available coal-related research and academic positions have been filled by these researchers for the past 30 years. This imminent loss of expertise is significant for the continued progression and retention of the institutional memory without which the ‘wheel will be reinvented’. Funding cycles also strain the talent pool as researchers move into the “hot” funding areas without returning to the coal arena. This paper examines the current state of affairs for U.S. coal science academic centers. Here coal science is defined as the science and engineering research on coal – uses (primarily combustion and coking), occurrence, distribution, properties, technology options, environmental and health impacts, beneficiation, and ash chemistry but does not include coal-related areas such a post-combustion pollution control or the various catalytic aspects of coal gasification products.
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Oviedo ICCS&T 2011. Extended Abstract
2. Methodology An ISI Web of Knowledge (using the web of science database) evaluation of journal articles with “coal” in the title (in English-language journals) was performed, for the periods 1970 to 2010 and 2000 to 2010, as one approach to evaluate the historically and currently active research centers in coal science. The ISI Web of Knowledge tools were used to evaluate the contribution of countries as well as the academic entities in the United States. The U.S. data were transferred into an Endnote library that was used to generate an authors list and a journal titles listing for evaluation with Wordle [2]. Wordle creates a visual representation of text frequency. Note: the publications of U.S. federal and state agencies are not included in this survey. 3. Results and Discussion Table 1 provides one indication of global institutional historic productivity, which consists of the number of journal records for each institution with “coal” in the title of journal articles using the ISI Web of Knowledge [3] selecting only the Web of Science (Science Citation Index Expanded) database for English language publications between 1970 and 2010. While this will also include coal mining, mine reclamation, CO2 climate change, energy policy, etc., those are expected to be a relatively small portion of the coal research for most institutions. This approach however, is one measure to identify those centers that have produced copious coal-related English-language research publications. The search resulted in >22,700 journal publications. The top 100 entities were analyzed (shown in Table 1) and a breakdown of the total publications by country of origin obtained is shown in Fig. 1. Those entities listed in Table 1 that are shown in bold had >15 publication in the last decade (2000 to 2010). These values are certainly an underestimation of research productivity (this research approach found only 10 of the lead authors 18 mostly-recent applicable papers); however it is a useful means of comparing institutions. Several notable institutions have clearly moved away from coal research during this period. Fig 1. also shows the analysis of publication frequency per year for the United States (30%), Japan (7.8%), and China (7.5%) the three leading countries publishing English language coal research papers. For the U.S., there was a rapid increase in coal articles after the first energy crisis, peaking in 1984 with a general decline until 2002 where publication quantities were on par with Japan and China. Since 1983, Japans publication rate has been relatively steady. China’s publications, however, have rapidly increased since the late 1990s and have now significantly surpassed those
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Oviedo ICCS&T 2011. Extended Abstract
of the U.S. Thus, what were/are the active U.S. academic centers and their areas of focus? Table 1. Universities and Institutions Sorted by Frequency of “Coal” in Journal Article Titles (1970-2010)*
*English language, excluding DOE and Geologic Surveys. Brackets represent article count (estimates). Those in bold had >15 papers in the last decade (2000-2010) (date of analysis December 2010)
Figure 1. Journal article by country of origin (note authors of collaborative works are counted in the percentage of each country) between 1970 and 2010 and the publication frequency per year for the United States, China, and Japan The Pennsylvania State University is one of the few academic centers where there are formal education opportunities in coal science. Historically this has been through the Fuel Science program. The current iteration of this program is the Fuel Science option Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
within the graduate program of the Department of Energy and Mineral Engineering [4]. Over 200 research degrees (M.S. and Ph.D.) have been awarded with coal in the title between 1970 and 2010 in various departments. Much of the coal research is within the Earth and Mineral Sciences (EMS) Energy Institute [5]. Figure 2 shows the Wordle analysis of those journal titles with the relative frequency of occurrence being illustrated by font sizes (color use is decorative). The analysis shows the expected trend of activity in nearly all aspects of coal science, notably coal liquefaction (including coal to jet
Figure 2. Wordle generated image of the frequency of journal article titles (coal has been omitted) for Penn State (1970-2010). Frequency of “liquefaction” = 30. fuel), combustion (bench and pilot-scale) and gasification, mineral matter or ash characterization and behavior, coal characterization, structure and behavior, coal structural modeling, coking, petrology, coal cleaning and preparation, CO2 sequestration in coal, and coal-water slurry research. All coal ranks are represented and the EMS Energy Institute houses the Penn State Coal Sample Bank and Database [6, 7]. Over the review period the most active coal scientists were Drs. H.H. Schobert (author and H.H. Storch Award recipient), P. Painter, A. Davis (retired), P.L. Jr. Walker [8-10] (deceased), P. Given [11] (deceased), R.G.L. Austin (retired), P.G. Jenkins (retired), A.W. Scaroni (no longer active—administration), J.P. Mathews (author), M.M. Coleman (retired), W. Spackman (retired), G.D. Mitchell, P.G. Hatcher (currently at Old Dominion University), M. Sobkowiak, S. Falcone Miller, C.S. Song (author, EMS Energy Institute Director, and H.H. Storch Award recipient), and C. Burgess-Clifford, among many others. The University of Kentucky houses the Center for Applied Energy Resources (CAER) [12] and several noted coal scientists: Drs. J. Hower (former editor of International Journal of Coal Geology), G.P. Huffman (retired), F.E. Huggins, C.F. Eble, B.H. Davis, Submit before May 15th to
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N. Shaw, and F. Derbyshire (deceased, former CAER director) [13, 14] , among others. A concentrated location of coal science that has focused on coal petrology and geology (often in collaboration with the state Geologic Survey), coal liquefaction, combustion, coal processing/upgrading, mineral matter and ash issues, carbon materials, and waste combustion product uses. Still very active, CAER has produced 106 journal articles addressing coal in the last decade. These articles cover a wide range of activities. West Virginia University houses the National Research Center for Coal and Energy [15]. The interests are in coal liquefaction, combustion, gasification (both traditional and underground) and coal analysis, coal waste use, and coal-derived carbons. The university also has a large mining program with a publication concentration in mining health, as well as coal beneficiation. Recently coal-bed methane and CO2 sequestration in coal are also active research areas. The active coal scientists (not including authors involved in coal workers health-related research) during the period of analysis were: D. B. Dadyburjor, J.W. Zondlo, M.S. Seehra, A.H. Stiller, C.Y. Wen, and R.G. Ames, among others. Southern Illinois University houses the Coal Research Center formed to stimulate and coordinate the efforts to improve coal mining and coal use [16]. The Department of Geology also houses the Coal Characterization Laboratory (mostly petrology orientated). Illinois has the largest of the U.S. bituminous coal fields but faces the very significant challenge of utilizing high-sulfur coal. There is a wide range of activity with the obvious focus of coal cleaning, desulfurization, combustion, and petrology. The active individuals during the period of evaluation were noted coal petrologist J.C. Crelling (semi-retired), S. Lavani, M.K. Mohanty, C.B. Muchmore, R. Honakar, and S. Harpalania, among others. A recent addition is the petrologist S. Rimmer. Massachusetts Institute of Technology was very active in coal research between 1980 and 1995. The Wordle analysis shows a concentration in combustion and (rapid) pyrolysis [17]. This is no surprise given the notable coal scientists: A. Sarofim (no longer at MIT), J.B. Howard (deceased) [18], W.A Peters, J.M. Beer, among others. MIT published the influential future of coal report [19], and is very active in clean coal aspects, specifically carbon capture and sequestration [20]; however, it has not been as active as a coal research center, in coal science as defined here, beyond a few individuals. University of Utah is notable for both pilot/lab-scale coal combustion (often in collaboration with Brigham Young University) and quantitative NMR applications to Submit before May 15th to
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coals and coal-related solids/liquids, pyrolysis mass-spectroscopy, among other areas. They run an active Clean Coal Program [21]. There is an active educational component in the former Department of Chemical and Fuels Engineering [22] but now the renamed Department of Chemical Engineering. There remains an active interest in energy and fuels. The coal scientists active during the period of the analysis were R. Pugmire (active in both research and administration), H.L.C. Meuzelaar (not recently active), W.H. Wiser (retired), D.M. Grant (retired?), L.L. Anderson, J.M. Veranth, A.F. Sarofim, and M.S. Solum, among many others. This institution is particularly noteworthy due to its broad collaborative record. Brigham Young University is relatively small in comparison to a number of the other universities but has a coal combustion focus housing the Advanced Combustion Engineering Research Center [23], often in collaboration with the University of Utah. Interests are pulverized coal combustion, pyrolysis, gasification, and modeling, among a wide range of topics. The authors active in the period of evaluation were: the noted combustion scientists: L.D. Smoot and T.H. Fletcher, P.O. Hedman, and M.L. Lee, among others. The University of Pittsburgh: was the home to very notable coal research in the “golden era” under Dr. I. Wender [24]. Currently the university is noted for running the successful Pittsburgh Coal Conference [25]. There was a focus on liquefaction from the 1980s to 1990s, along with coal-dust characterization and coal processing, among others. Active authors were Y.T. Shah, G.E. Klinzing (no longer active – administration), J.W. Tierney (retired), and I. Wender (retired and H.H. Storch Award recipient), among many others. There has been only limited activity recently. University of Illinois at Urbana-Champaign’s history in coal research started in 1901 when S. Parr founded the now Department of Chemical and Biomolecular Engineering. UI also houses the state’s Geological Survey, where the ISGS Coal Section is primarily concerned with the study of coal bearing Pennsylvanian rocks in the Illinois coal fields [26]. UI was active in combustion, coal geology, gasification, ash, petrology, and characterization, broadly defined. Recent activity has included CO2 sequestration in coal. Ohio State University: The state of Ohio is one of the six leading states for coal production. Much of the historic work in the period evaluated was related to combustion, which is defined broadly due to the presence of noted scientist Dr. R. Essenhigh (formally of Penn State). Work includes generation and use of coal bySubmit before May 15th to
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products [27]. Clean coal research has continued under Dr. L.S. Fan on coal gasification and chemical looping combustion, while much of the other coal-related publications are focused on environmental, watershed, and mine reclamation issues. Virginia Polytechnic Institute and State University: Virginia Tech houses the Virginia Center for Coal and Energy Research [28] and the Department of Minerals and Mining Engineering [29] and has a long history of coal research. The Wordle analysis shows a concentration in ash, combustion, coal cleaning, solvent-refined coal, and a variety of environmental interests. For example, much of the ash interests are heavy metals and watershed issues. This is another institution where coal-science work has dwindled to only a few applicable papers (as defined here) in the last decade. Those who were active: D. S. Cherry, L.T. Taylor, and R.K. Guthrie, among others. University of Wyoming: Wyoming produces more coal than the next five leading states combined [30]. As such, it would be expected to be a center of coal research. The University of Wyoming [31] and the Western Research Institute [32] are separate entities but have a strong history of collaboration with a clear and expected focus on the strategic use of the large subbituminous coal reserves with a focus on liquefaction, gasification, coal liquids characterization, coalbed methane, among others. Those active in the period evaluated here are: R.J. Hurtubise, H.F. Silver, H.W. Haynes, and others. The university has a strategic partnership with the university of Queensland Australia and houses biannual conferences alternating between those locations [33]. The state has made strategic investments in subbituminous coal broadly with the creation of the Clean Coal Research Account [34] and thus coal research activities are expected to increase. Dr. J. G. Speight also generated a number of coal books here during this period [35, 36]. Auburn University: houses the Coal and Energy Laboratory [37] and was active in liquefaction, co-processing, combustion. Recent activity (last decade) has mostly been focused on environmental exposure issues. Those active in the time frame evaluated were: C.W. Curtis, J.A. Guin, and A.R. Tarrer among others. Carnegie Mellon University (Pittsburgh) is active in coal research although much of the current activity is system analysis and economic modeling with a focus on energy policy. The university houses the relatively unique Engineering and Public Policy department in the College of Engineering [38]. There are also interests in coal tar and water interactions, as well as mathematical modeling of devolatilization and heat and mass transfer. Leading authors during the period of evaluation were: E. Rubin (policy related) and R.G. Luthy, among others. Submit before May 15th to
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University of North Dakota is located near the lignite deposits of the Fort Union Region. UND houses the Energy and Environmental Research Center (EERC) [39]. They are active in gasification, combustion, ash chemistry, and emissions control, with an emphasis on mercury capture. UND’s research started by studying the state’s vast lignite resources and investigating potential uses for them; however, over the years they became a leading institution on low-rank coal science and technology, and currently are active in studying all ranks of coal. Active authors during this period were: J.W. Nowok (retired), S.A. Benson, among others. Iowa State University [40] was very active in coal research in the 1980s and somewhat active in the 1990s but has had little recent coal-science activity. The Wordle evaluation showed a concentration in combustion, mostly considering emissions and waste material, with a wide range of interests from microwave applications, fine coal agglomeration, to exploring new characterization techniques. Those active during the period of analysis were: R. Markuszewski, T.D. Wheelock, and C.D. Chriswell among others. Ames Laboratory [41] is also located here and specializes in materials for energy application. University of California at Berkeley has been active in coal research in a variety of fields. The coal interests have ranged widely from coal liquids, particulates/surfaces, and trace elements. Those active in the period of analysis were: D.W. Fuerstenau, L. W. Tian (now in Hong Kong), and A.T. Bell, among others. There has been recent activity in the areas of particulate matter and permeability. University of Tennessee [42] is close to the Appalachian bituminous coal basin but the production of coal from this state is small, producing the least amount of coal in the basin, only 2% of the production of the leading state (West Virginia) in this region [42]. However, the state is a significant user of coal and is located close to Oak Ridge National Lab [43], which was also active in coal research. The Worlde analysis showed a range of interests including PAH, toxicity, catalytic gasification, among others. Those active in the period of analysis were: G. Mamantov, E.L. Wehry, and J.W. Larsen (now retired from Lehigh University), among others. University of Texas is a multi-campus university with the coal research being mostly in Austin [44] with some work in Dallas [45]. Texas has both lignite and bituminous coal deposits, is the leading coal producing state for the Interior basin, is the state that uses the most coal in the U.S. [46], yet the quantity of coal science research is not in proportion to its extensive coal use. The Wordle analysis showed an interest in Submit before May 15th to
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underground gasification (modeling), ash chemistry (water contamination and trace elements), among others. Those active in the period of analysis were: T.F. Edgar, R.K. Guthrie (retired), and D.D. Cherry among others. A relatively new addition to the Dallas campus is R. B. Finkelman (author, formally with USGS). Purdue University houses the Center for Coal Technology Research [47]. It was formed as a state agency in 2004 to “promote the use of Indiana's coal reserves in an economically and environmentally sound manner”. The bulk of the coal work was in the 1980s with only a few publications in the 1990s. Activity has increased however in the last decade. The Wordle analysis showed that there was a concentration in coal structure characterization (often solvent swelling related). Those active in the period of analysis were: N.A. Peppas, L.M. Lucht, and C.K. Chao, among others. Texas A&M: The Wordle analysis shows an interest in co-firing (biomass and other material), coal-water slurry, and coal liquids, among others. Those active in the period of analysis were: K. Annamalai and J.M. Sweeten, among others. The Missouri Education System has two universities. The Missouri University of Science and Technology (Missouri S&T), formally The University of Missouri-Rolla [48]) and the University of Missouri at Columbia (Mizzou) both active in coal research, There was activity in coal-log transportation (Mizzou), ash and inorganics, with an emphasis on trace elements (Missouri S&T). The University of Georgia the university [49] has coal-research centered around trace elements and exposure. As most of the work is removed from coal-science, as defined here, it is not a coal center but is included for completeness. University of Western Kentucky: This university is of note for the minor in coal chemistry and a Masters of Science option within the chemistry department and the Institute for Combustion Science and Environmental Technology [50]. The bulk of their activity has been in the last decade. Their focus is on combustion and mercury speciation/emissions/capture and co-firing biomass (often chicken litter) with coal, among other research interests. Those active in the period of analysis were: W.P. Pan, Y. Cao, C.W. Chen, and J.T. Riley, among others. Others academic centers: Colorado State University was active in fly-ash chemistry and NMR spectroscopy with G.E. Maciel, D.F.S. Natusch, and A. Jurkiewicz being active among others. Lehigh University has been a center for coal science due to the combination of the work of noted coal scientist J. Larsen (retired—former editor of Energy & Fuels) and their Energy Center [51] with a focus on process optimization, Submit before May 15th to
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combustion, and processing. Brown University has a long history of coal science with several coal scientists: E. Suuberg (editor of the journal Fuel), R. Hurt, and J. Calo. The University of Ohio houses the Ohio University Coal Research Center in the Russ College of Engineering and Technology [52]. The focus there being air quality issues rather than coal science (as defined here). The University of Chicago was active in coal science under L. Stock and his colleagues (Dr. Stock’s journal articles alone numbered 50 publications between 1970 to 1997). 4. Conclusions There has been a notable rise and decline of coal-science publications from the U.S. Currently, the U.S. generates fewer journal articles than China and no more than half of the historically active coal science centers are still active. Wordle analyses were used to determine focus areas and authors for U.S. academic entities. Among the U.S. academic entities the leading institutions (quantity of journal articles) were Penn State, Kentucky, West Virginia, Southern Illinois, MIT, Utah, Brigham Young, Pittsburgh, Illinois, Ohio State, Virginia Tech, Wyoming, Auburn, Carnegie Mellon, North Dakota, Iowa, California at Berkeley, Tennessee, Texas, Purdue, Texas A&M, Missouri, Georgia, and Western Kentucky. References [1] EIA/DOE Annual Energy Outlook. http://www.eia.doe.gov/oiaf/aeo/index.html [2] Feinberg J Wordle. http://www.wordle.net/ [3] Thiomson Reuters ISI Web of Knowledge. http://apps.isiknowledge.com/ [4] The Pennsylvania State University Department of Energy and Mineral Engineering. http://www.eme.psu.edu/ [5] The Pennsylvania State University The EMS Energy Institute. http://www.energy.psu.edu/ [6] The Pennsylvania State University Penn State Coal Sample Bank and Data Base. http://www.energy.psu.edu/copl/index.html [7] Glick DC, Davis A, Operation and composition of the Penn State Coal Sample Bank and Data-Base. Organic Geochemistry 1991, 17, (4), 421-30. [8] 1969 Storch award: Dr Philip L. Walker, Jr. Fuel 1970, 49, (1), 102-3. [9] Philip L. Walker Jr. Publications. Carbon 1991, 29, (6), 693-701. [10] Marsh H, A tribute to Philip L. Walker. Carbon 1991, 29, (6), 703-4. [11] Derbyshire FJ, Obituary: Dr Peter Given. Fuel 1988, 67, (8), 1167-. [12] University of Kentucky Center for Applied Energy Research. http://www.caer.uky.edu/ [13] 1996 Storch Award—Frank J. Derbyshire. Fuel 1997, 76, (2), 99-. [14] 1997 Henry H. Storch Award—Frank Derbyshire. Fuel 1997, 76, (1), 1-. [15] West Virginia University National Research Center for Coal and Energy. http://www.nrcce.wvu.edu/about.cfm [16] Southern Illinois University at Carbondale Coal Research Center. http://www.crc.siu.edu/ [17] Mechanical Engineering at MIT Energy Science and Engineering. http://web.mit.edu/ese/ (Accessed February 22nd, 2006), [18] 1983 Storch Award—Jack Howard. Fuel 1983, 62, (9), 1101-. [19] MIT The future of coal, an interdisciplinary MIT study; 2007.
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Oviedo ICCS&T 2011. Extended Abstract [20] Massachussetts Institute of Technology (MIT) Carbon Capture & Sequestration Technologies. http://sequestration.mit.edu/research/index.html [21] The Utah State University Clean Coal Program. http://www.cleancoal.utah.edu/ [22] The University of Utah Department of Chemical Engineering (formaly Department of Chemical and Fuels Engineering). http://www.che.utah.edu/index.php [23] Bringham Young University Advanced Combustion Engineering Research Center. http://www-acerc.byu.edu/ [24] Wender I, Coal science in a changing world. Fuel 1985, 64, (8), 1035-8. [25] IPCC International Pittsburgh Coal Conference,. http://www.engr.pitt.edu/pcc/ [26] Illinois Geological Survey Illinois Geological Survey. http://www.isgs.illinois.edu/aboutisgs/about.shtml [27] Ohio University Ohio Coal Research Center. http://www.ohio.edu/ohiocoal/ [28] Virginia Center for Coal and Energy Research Virginia Tech. http://www.energy.vt.edu/ [29] Virginia Tech Department of Mining and Minerals Engineering. http://www.mining.vt.edu/ [30] EIA Domestic distribution of U.S. coal by origin state, consumer, destination and method of transportation. http://www.eia.doe.gov/cneaf/coal/page/coaldistrib/a_distributions.html [31] Clean Coal Technology Center University of Wyoming. http://www.uwyo.edu/ser/info.asp?p=3737 [32] Western Research Institute History. http://wri.uwyo.edu/about.aspx [33] Com. ACCD Advanced Coal Technologies Conference. http://www.advancedcoalconference.com/ [34] The University of Wyoming Clean Coal Technology Fund. http://www.uwyo.edu/ser/research/clean-coal/index.html [35] Speight JG, Knovel (Firm), Handbook of coal analysis. In Chemical analysis v. 166, Wiley-Interscience: Hoboken, N.J., 2005; pp x, 222 p. [36] Speight JG, The chemistry and technology of coal. 2nd ed.; M. Dekker: New York, 1994; p xi, 642 p. [37] Auburn University Coal and Energy Laboratories. http://www.eng.auburn.edu/research/dept-research/chen.html [38] Carnegie Mellon University Engineering and Public Policy. http://www.epp.cmu.edu/ [39] University of North Dekota Energy & Environmental Research Center. http://www.undeerc.org/ [40] Iowa State University http://www.vpresearch.iastate.edu/institute/ [41] Ames Laboratory (DOE) Ames Laboratory. http://www.ameslab.gov/research [42] Freme F Energy Information Administration U.S. coal supply and demand 2009. [43] Oak Ridge National Laboratory Oak Ridge National Laboratory. http://www.ornl.gov/ [44] University of Texas at Austin University of Texas at Austin. http://www.utexas.edu/ [45] University of Texas at Dallas University of Texas at Dallas. http://www.utdallas.edu/ [46] DOE/EIA Annual coal report 2008. http://www.eia.doe.gov/cneaf/coal/page/acr/acr_sum.html [47] Purdue University Center for Coal Technology Research. http://www.purdue.edu/dp/energy/CCTR/ [48] University of Missouri University of Missouri. http://www.missouri.edu/ [49] University of Georgia http://www.uga.edu/ [50] University of Western Kentucky Institute for Combustion Science and Environmental Technology. http://www.wku.edu/ICSET/comblab.htm [51] Lehigh University Energy Research Center. http://www.lehigh.edu/energy [52] Ohio Coal Research Center Ohio University. http://www.ohio.edu/ohiocoal/
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The Current State of Coal Research In The United Kingdom, Germany, Australia, and South Africa Jonathan P. Mathews1, Bruce G. Miller1, Chunshan S. Song1, Harold H. Schobert1, Francois Botha2, Robert B. Finkleman3, Alan Chaffee4, 1
John and Willie Leone Family Department of Energy & Mineral and EMS Energy Institute, 126 Hosler Building, The Pennsylvania State University, University Park 16802, USA.
[email protected] 2 Illinois Clean Coal Institute, 5776 Coal Drive, Suite 200, Carterville, IL 62918, USA 3 University of Texas at Dallas, Richardson, TX 75080, USA 4 Monash University, Victoria 3800, Australia 1. Introduction As indicated in an earlier paper [1] there is about to be a significant decline in coal science capability, for certain countries, as key experienced individuals retire and institutions refocus. It was therefore desirable to determine those locations with a coal focus, their research interests and level of current activity based on journal article publication record. Of the 22717 journal articles with coal in the title from an ISI Web of Knowledge English language evaluation using the web of science database: the leading publishing countries between 1970 and 2010 were: United States (30%), Japan (7.8%), China (7.5%) and England (5%). Australia (5%), India (4%), Canada (4%), Poland (4%), Spain (3%), Federal Republic of Germany (3%), France (2%), Turkey (2%), Russia (2%), South Africa (1%) and the Netherlands (1%). These countries have very different general trajectories in the frequency of coal-science focused publications: declining (USA, UK, Germany), consistent (Japan), or rapid rise (China). This paper focuses on the United Kingdom, Germany, Australia and South Africa.
2. Methodology An ISI Web of Knowledge (using the web of science database) evaluation of journal articles with “coal” in the title (in English-language journals) was performed, for the periods 1970 to 2010 and 2000 to 2010, as one approach to evaluate the historically active and currently active research centers in coal science. The ISI Web of Knowledge tools were used to evaluate the contribution of countries and research entities. The data were transferred into an Endnote library that was used to generate an authors list and the journal titles listing for evaluation with Wordle [2]. Wordle creates a visual
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representation of text frequency for the determination of key phrases and topics in titles and contributing authors.
3. Results and Discussion Fig. 1 shows the frequency of journal articles, with coal in the title of English-language journal articles (web of science database), versus year for the UK (here England, Scotland, and Wales), Australia, Germany, and South Africa. There is a general peak for all countries between 1980 and 1985, a decline from the peak until around 1991. For both the UK and Germany current output of articles is half of the historic high. Australia has a more prominent cyclic nature peaking in 2000 and 2009 presumably in response to funding opportunities. South Africa produced an average of only 2 papers per year in the 1970’s, 8 in the 80’s, 6 in the 90’s, with a significant increase after 2003.
Figure 1. Publication frequency of journal articles with coal in the title by country of author contribution(s) United Kingdom The United Kingdom was the birthplace of the coal-fueled industrial revolution. It is of no surprise that it was also a center of excellence for coal science over much of the history of coal science research. The UK has produced many noted coal scientists such as P. Given [3], F. Derbyshire [4], H. Marsh, and C. Snape [5]. Previously very active coal centers such as Newcastle and Sheffield Universities have refocused their interests as coal became of marginal interest regionally. Coal production had dropped to 12%
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(1999) of the 1913 peak output with a relatively steady decline since 1960 [6]. Those universities of note for coal science centers are now Nottingham, Leeds, and Imperial College London. There is a generally reduced level of interest in coal science concurrent with reduced domestic coal production.
Imperial College London has coal research within the Energy Engineering group of the Chemical Engineering and Chemical Technology department [7]. There activities are focused on the range of solid fuel processing technologies (combustion, gasification, pyrolysis, liquefaction) for the main utilization industries (power generation, fuel production, and chemicals production). Key individuals are: R. Kandiyoti, P. Fennell, M. Millan, N. Paterson, A. Herod (retired) [7]. One particular specialty of this group has been the characterization (advanced analytical analyses) of heavy coal liquids.
Nottingham University [8] has one of the largest concentrations of coal scientists in the UK, although not all are currently active in the field, and several are relatively recent additions (thus their research productivity is dissipated among their previous institutions and universities by the approach utilized here). Activities have traditionally been focused on combustion, liquefaction, structural analyses (coal, trace elements, and mineral matter), processing, and beneficiation. Of particular note is their work on coal petrology (including chars and image analysis) and microwave applications to coal (housing the National Center for Industrial Microwave Processing [9]). The coal scientists include: M. Cloke (administration), K. Steel (now at the University of Queensland, Australia), C. Snape [5] (administration - Director Energy Technologies Research Institute [10]) but still active in research, N. J. Miles, J. P. Wright, E. Lester, J. Patrick (special professor status, semi-retired but still active), S. Kingman, among others (such as M. Maroto-Valer, and J. Andrésen who are not currently active in coal). This latter group, along with C. Snape is also known for applying NMR analyses to coal although much of that work occurred in other institutions. Nottingham hosted the 2007 International Conference on Coal Science and Technology.
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Figure 1. Wordle view of the frequency of journal article titles 1970-2010 for Nottingham University Leeds University used to have the Department of Fuels and Energy. It changed and merged with other departments around 2004 into the School of Process, Environmental and Materials Engineering and offers Chemical Engineering and Energy Engineering degrees (among others) with research endeavors within the Energy and Resources Research Institute [11]. Their focus is combustion, pollution abatement, and biomass energy with strong ties still to coal science. Their stated vision is "to be recognized for internationally leading research in the sustainable development of natural resources, the sustainable use of fossil fuels and the development of renewable and future fuels."
Others: While not an academic institution, the IEA Clean Coal Center is also located in the UK and is worthy of mention, specifically for the high quality publications addressing many coal issues [12]. Strathclyde University, has P. J. Hall, (formerly C. Snape) and is also active in coal-based sequestration research.
Germany An industrial powerhouse, coal fueled the Germany economy and industrial complex. Hard coal production has declined from a peak in 1958 to producing 17% of that value
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in 2005 [13]. Underground coal mining is potentially slated to end in 2018 with the removal of coal subsidies. There are however large lignite reserves and Germany has been the leading producer of lignite coal doubling the U.S. production and almost doubling China’s [14]. Much of the coal science work is related to brown coal, coal-tar, gasification, and combustion.
Bergbau Forschung GMBH: An industrial company now named CarboTech AC GmbH specializing in activated carbon, activated cokes, and carbon molecular sieves from coal [15]. Activity stopped in the 1990’s under this name but with little activity since. A review of the petrology work at this laboratory is available [16].
The Max-Planck Institute for Coal Research: The Institute was formed in 1912 to study the chemistry and uses of coal. While it currently still retains the coal in the title, the focus is basic research in organic and organometallic chemistry, catalysis, and theoretical chemistry [17]. An evaluation of papers published between 2005-2008 showed that out of about 700 papers less than 10 contained coal in the title. However, historically the Institute was certainly an important center of coal research with pioneering gasification based coal-liquefaction research and the well-known FisherTropsch synthesis [18]. With the retirement of M. Haenel (2009) it is likely the end of coal research at this “coal” institution.
University of Stuttgart: The university houses the Institute of Combustion and Power Plant Technology [19], which was established in 1958 (as the Institute of Process Engineering and Power Plant Technology until a name change in 2009) with research focusing on combustion and gasification of solid fuels, air quality control, and various aspects of power plant technology. Recent activities include CO2 mitigation through oxyfuel, pre- and post-combustion capture processes, and renewable energy research including biomass co-firing and gasification. A key individual has been K. Hein (emeritus and past director of the Institute) now in Australia.
Others Institute of Geology and Geochemistry of Petroleum and Coal at RWTH Aachen University, specializes in coal petrology, coal-based sequestration and other coal aspects [20]. B. M. Kroos and several others are active. Technical University of Freiberg Submit before May 15th to
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(Saxonia) coal combustion and gasification research in the Institute for Energy Process Engineering and Chemical Engineering (Prof. Bernd Meyer) [21]. With core themes of geosciences, materials, energy and environment as well as a strategic partnership with the University of North Dakota coal-related science may increase. Engler-BunteInstitute of the Technical University of Karlsruhe [22]: coal and biomass conversion Prof. Rainer Reimert (retired).
South Africa South Africa ranks 6th in terms of global coal production (2009), is a significant coal exporter, and user of coal [14]. Coal is the major source for electricity and a coal-toliquid fuels and coal-to-chemicals industry (Sasol via gasification and Fisher-Tropsch synthesis) uses about 20% of the coal consumed [23]. After many years of relatively low-productivity of published coal research, relative to the importance of coal, South African coal science has seen a rebirth and increased activity. Much of this work is centered around Witwatersrand University and increasingly at University of the North West “Potch” campus. Cape Town held the 2009 International Conference of Coal Science and Technology, with a strong South African presence. There has also been expertise in gasification, combustion, and coal beneficiation due to the presence of high “ash” coals [24].
Witwatersrand University (Witts) located in Johannesburg has its origins as the South African School of Mines. While hard rock mining is still prominent at this institution, the bulk of the coal research has focused on coal fires, petrology, and coal quality issues of the inertinite-rich coals. There was also some coal mining: productivity, safety and health-related work. Those active in the period of evaluation were: D. Glasser, and R. M. S. Falcon (South African National Energy Research Institute Chair in Clean Coal Technology), among others. N. Wagner (petrologist) has recently (2007) joined the university’s Coal and Carbon Research Group [25] and thus the expansion of coal science is on going.
The University of Cape Town [26] is active in coal preparation, ash/waste, and coal cleaning issues. Academic offerings include a course on Fuels and Chemicals from Coal and Syngas. Active authors were: J. P. Franzidis, M. C. Harris, and P. Stonestreet among others. Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
Sasol is the major producer of coal-to-liquids and does so via in-direct liquefaction approaches. Thus as expected, there is considerable expertise in Fisher-Tropsch chemistry that is not covered in this analysis. Sasol also has been increasingly active in publishing since 2003 (gasification issues, mineral matter/ash transformation etc., notably J. C. van Dyk, among others).
Potchefstroom Campus of the University of the North West [27], while this university does not appear in Table 1 it has undergone considerable expansion/investment in coal research recently. Currently active in beneficiation, combustion/gasification (especially kinetics), coal handling, and coal chemistry. Those active are J. R. Bunt (formally with Sasol), R. C. Everson, and Q. P. Campbell among others. Graduate students currently active in coal science number around 8. Thus expectations are for a significant increase in coal science contributions from this institution.
Australia Australia has very significant national reserves of both brown and black coal [28] providing > 80% of its electricity. Australia is the world’s leading coal exporter and black coal is also its largest commodity export, ~23% of total exports [29]. As such, research associated focussed towards the efficient recovery and environmentally sustainable utilistation for this resource is of particular national significance. Organizations active in coal research have included the Commonwealth Scientific Industrial Research Organisation (CSIRO), several universities, a small number of industry research laboratories and, in respect of the brown coal in Victoria, the Herman Research Laboratory (HRL). There was a peak of activity during the 1980s, through the ‘oil crises’ of that period, as extensive funding was directed towards coal liquefaction. Whilst coals from around the country were investigated extensively, the work on Victorian brown coal was carried through to the successful demonstration of the 50 ton dry-coal/day BCL (Brown Coal Liquefaction) pilot plant at Morwell [30]. Victoria. Research activity has accelerated again in the last approximately 5 years with a focus on reducing the carbon emissions associated with coal-based electricity generation. Research has blossomed in aspects that can be related (a) to the development of improved process configurations for electricity generation and (b) to technologies for carbon capture and sequestration. Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
Federal and state government funding foster these activities. In the late 1970s the federal government introduced a research levy of 5 cents per tonne of saleable coal and established the National Energy Research, Development and Demonstration Council (NERDDC) to allocate the research funds.. In 1992, the State Electricity Commission of Victoria (SECV), of which HRL was a part, was privatized. The considerable downsizing that occurred as a result of industry restructuring led to considerable loss of Brown coal expertise and significantly reduced funding. Australia’s unique system of cooperative research centres (CRC) - joint ventures between government, industry and universities – was able to foster much of the coal research around the turn of the century. Black coal research (2001-2008) was carried out as part of the CRC for Coal in Sustainable Development [31]. Brown coal research was advanced in the CRC for New Technologies for Power Generation from Low-Rank Coal, which later became known as the CRC Clean Power from Lignite (1993-2006). Although they carried out some research ‘in-house’, these organizations supported a large proportion of the work occurring at the universities through this period. Currently (since 2003) the CRC for Greenhouse Gas Technologies (CO2CRC) funds a number of research and pilot projects associated with CO2 capture and sequestration. The Australian National Low-Emissions Coal Council Research and Development (ANLEC R&D) Ltd. was established (2009) by the federal government to foster research to assist with the deployment of Low Emission Coal Technology [32]. In 2010, the Victorian government established Brown Coal Innovation Australia (BCIA) [33] to invest in the development of technologies and people that broadens the use of brown coal for a sustainable future. Other state governments have more general funding schemes that encompass R&D funding for coal-based projects. Most now have primary emphasis on lowering the emissions associated with its use.
CSIRO [34] has been intensively involved in coal R&D since the establishment of its Coal Research Section soon after the conclusion of the second world war. During the 1980s, its Divisions of Fossil Fuels and Energy Chemistry (which later united to form the Division of Coal Technology) were very active in research on coal conversion to liquids by both direct and indirect means, but interest in this topic waned as the oil crises of that decade subsided. Specialists in coal petrology (Shiboka), mineralogy Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
(Swaine), geochemistry (Wilson), chemistry (Durie, Wailes, Lynch, Schafer) and combustion (Mulcahy, Smith, Duffy) made major contributions to understanding the variability in coal structure and properties. Coalscan was developed to provide real time information regarding the mineral, moisture and “ash” content of the coal and its products during the mining and recovery. Projects for producing ‘ultra-clean’ coal developed. More recently, the current Division of Energy Technology has focused on carbon capture and storage systems, coal gasification for higher efficiency energy production and the direct use of coal in engines. Fundamental work regarding the nature of methane interactions with coal seams has relevance to coal mine safety in addition to CO2 sequestration. Industry Research Laboratories: During the 1980s, the corporate research laboratories of BHP Ltd. (both in Melbourne and Newcastle), CSR Ltd. as well as the The Australian Coal Industry Research Laboratory (ACIRL) built up impressive facilities within their respective laboratories and substantially contributed to national sponsored studies of coal liquefaction. Herman Research Laboratories (HRL): As part of the SECV prior to 1992, HRL was particularly well equipped and played a preeminent role in research and development regarding all aspects Victorian brown coal characterization and use. Most of the prominent specialists at HRL at the time contributed chapters to a monograph entitled ‘The Science of Victorian Brown Coal’ [35] where their many significant contributions in relation to brown coal structure, mineralogy, dewatering, liquefaction, combustion and industrial applications are well presented. From 1992 HRL diversified to provide consultancy and contract research services to a broader industry base. Currently, it promotes a high efficiency process configuration for electricity production from brown coal, known as IDGCC (integrated, driving and gasification combined cycle).
Monash University: This university, named after Sir John Monash under who’s leadership the SECV began the development of Victoria’s enormous brown coal reserves, has played a major role in brown coal research. The steam fluidized bed drying process, originally developed by Potter in the 1970s, has seen further development and growing applications over recent years. In the 1980s, both the School of Chemistry (Jackson, Larkins) and the Department of Chemical Engineering (Agnew, Sridhar) were involved in coal liquefaction studies. Later, research into non-evaporative methods for Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
dewatering and remediation of the product water (Chaffee, Hoadley) as well as on catalytic aspects of pyrolysis/gasification/combustion (Li, then Bhattacharya) developed [36]. An updated monograph, ‘Advances in the Science of Victorian Brown Coal’ was produced by Li 2004 [37]. A new wave of activity is now underway with the recent appointment of two professorial BCIA Research Leadership Fellows [33] and the commencement of substantial new programs addressing oxy-fuel combustion, spontaneous combustion, etc.
University of Newcastle: For over 30 years Prof Terry Wall, in collaboration with Gupta, Lucus, Bryant, Moghtaderi and many others, has played a preeminent role in combustion research, principally on bituminous coals and ash chemistry [38]. He led the development of significant studies into coal blending which has been important for Australia’s export industry. Claus Diessel, (now emeritus) made substantial contributions to the petrology of Australian coals.
University of Melbourne [39]: Through the 70s and 80s, the Department of Chemical Engineering became well known for its fundamental work on the swelling characteristics of brown coal (Evans, Allardice), the rheological behaviour of coal-water slurries (Boger) and coal drying (the Evans-Siemon process). In the School of Chemistry, an evaporative drying process (densified coal) and fundamental work on the organic geochemistry of brown coal was pursued. Both departments also pursued liquefaction studies. Johns, Verheyen, Larkens, and others were also active.
University of Queensland: Active in Fluidized bed gasification (Rudolph). Stanmore was also active in combustion-related work. A program on dynamic fluid transport properties and permeability of in-situ coal (Massarotto and colleagues) is being applied to reservoir simulation studies for coalbed methane (CBM) extraction, coal mine demethanation (CMM) and carbon dioxide geosequestration in coal seams. The university is active in mineral processing [40].
Others: The geology/petrology lab of Colin Ward at the University of New South Wales [41] has worked extensively on the occurrence and transformation of mineral phases during coal utilization. Much of this has been in collaboration with CSIRO (French) making use of the powerful QEMSCAN analysis system. At the University of Adelaide, Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
Ashman and Mullinger have led a group investigating the gasification potential and combustion characteristics of South Australian lignites with a view to their eventual utilisation. At Curtin University, Li (who moved from Monash) and colleagues have recently established the Fuels and Energy Technology Institute (FETI) which is investigating the production of fuels and chemicals from lignite as well as gasification and oxyfuel combustion of Western Australian lignites [42, 43]. Acknowledgements The authors thank the following individuals who contributed to our ability to evaluate the current state of coal capabilities: Drs. Haenel (Max-Planck Institute for Coal Research), and J. M. Jones (University of Leeds) and others for helpful conversations. This project was funded by the Illinois Clean Coal Institute with funds made available through the Office of Coal Development of the Illinois Department of Commerce and Economic Opportunity. References [1] Mathews JP, Miller B, Song C, Schobert HH, Botha F, Finkelman RB, The current state of affairs of coal research in U.S. Universities, International Conference on Coal Science and Technology, 2011, Oviedo, Spain, [2] Feinberg J Wordle. http://www.wordle.net/ [3] Derbyshire FJ, Obituary: Dr Peter Given. Fuel 1988, 67, (8), 1167-. [4] 1996 Storch Award—Frank J. Derbyshire. Fuel 1997, 76, (2), 99-. [5] 2006 Henry H. Storch Award in Fuel Chemistry goes to Colin Snape. Fuel 2007, 86, (1-2), 1-2. [6] Hicks J, Allen G House of Commons Research Paper: A century of change: Trends in UK statistics since 1900. www.parliament.uk/commons/lib/research/rp99/rp99-111.pdf [7] Imperial College London Energy Engineering Research Group, Department of Chemical Engineering and Chemical Technology. http://www3.imperial.ac.uk/chemicalengineering/research/researchthemes/researchfocusareas/e nge [8] Nottinham University Chemical and Environmental Engineering. http://www.nottingham.ac.uk/Engineering/Departments/Chemenv/People/index.aspx [9] Nottinham University National Center for Industrial Microwave Processing. http://www.nottingham.ac.uk/ncimp/research.php [10] Nottingham University Energy Technology Research Institute. http://research.nottingham.ac.uk/NewsReviews/ExpertiseResults.aspx?id=3417 [11] University of Leeds Energy and Resources Research Institute. http://www.engineering.leeds.ac.uk/erri/ [12] IEA Clean Coal Center http://www.iea-coal.org.uk/site/ieacoal/about/history [13] Energy Watch Group Coal: Resources and future production. www.energywatchgroup.org/.../EWG_Report_Coal_10-07-2007ms.pdf [14] EIA Coal Statistics (2008 data). http://www.eia.gov/coal/data.cfm#reserves [15] CarboTech AC GmbH Company history. http://www.carbotech.de/en/history.php [16] Steller M, Arendt P, K¸hl H, Development of coal petrography applied in technical processes at the Bergbau-Forschung/DMT during the last 50 years. Int. J. Coal Geol. 2006, 67, (3), 158-70.
Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract [17] Max-Planck Institute for Coal Research http://www.mpimuelheim.mpg.de/kofo/english/mpikofo_home_e.html [18] Haenel MW, History of the Max-Planck-Institut für Kohlenforschung. 2008. [19] University of Stuttgart Institute of Combustion and Power Plant Technology. http://www.ifk.uni-stuttgart.de/index.en.html [20] University RA Institute of Geology and Geochemistry of Petroleum and Coal. http://www.lek.rwth-aachen.de/cms/index.php?id=12&L=2 [21] Technical University of Freiberg (Saxonia) Institute for Energy Process Engineering and Chemical Engineering. http://tu-freiberg.de/index.en.html [22] Karlsruhe Institute of Technology Engler-Bunte-Institute. http://ceb.ebi.kit.edu/english/index.php [23] Ebergard A The future of South African coal: markets, investment, and policy changes. http://pesd.stanford.edu/publications/ [24] de Korte GJ, Coal preparation research in South Africa. Journal of the South African Institute of Mining and Metallurgy 2010, 110, (7), 361-4. [25] Witwatersrand University Center for Coal Research. http://web.wits.ac.za/Academic/EBE/ChemMet/ResearchUnits.htm [26] The University of Cape Town http://www.uct.ac.za/ [27] North West University PC Energy Systems, The Coal Group. http://www.nwu.ac.za/pfe/currentres.html [28] Geosciences Australia (Australian Government) Coal resources. http://www.ga.gov.au/energy/coal-resources.html [29] Australia Coal Association The Australian Coal Industry- Coal Exports. http://www.australiancoal.com.au/the-australian-coal-industry_coal-exports.aspx [30] Okuma O, Sakanishi K, Liquefaction of Victorian Brown Coal. In Advances in the Science of Victorian Brown Coal, Li, C.-Z., Ed. Elsevier: New York, 2004; pp 85-133. [31] Cooperative Research Center for Coal in Sustainable Development http://www.ccsd.biz/index.cfm [32] ANLEC R&D http://www.anlecrd.com.au/ [33] Brown Coal Innovation Australia http://www.bcinnovation.com.au/ [34] CSIRO Energy from coal. http://www.csiro.au/science/Coal.html [35] Durie RA, The Science of Victorian brown coal: structure, properties, and consequences for utilization. Butterworth-Heinemann: Oxford, 1991. [36] Monash University Clean Energy Technology. http://www.eng.monash.edu.au/research/strengths/clean-energy-technologies/ [37] Li C-Z, Advances in the Science of Victorian Brown Coal. Elsevier: New York, 2004. [38] University of Newcastle Low emission coal. http://www.newcastle.edu.au/researchcentre/energy/research/low-emission-coal.html [39] University of Melbourne University of Melbourne. http://www.unimelb.edu.au/ [40] University of Queensland http://www.chemeng.uq.edu.au/Mineral-Processing-InterfacialProcesses [41] University of New South Wales http://www.unsw.edu.au/index.html [42] Curtin Univeristy Curtin Univeristy. http://www.curtin.edu.au/ [43] Curtin Univeristy Curtin Centre for Advanced Energy Science and Engineering. http://energy.curtin.edu.au/
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Oviedo ICCS&T 2011. Extended Abstract
Adsorption Behavior and Biogasification of Soma Lignite Mustafa Baysal1, Sedat İnan2, Fırat Duygun2 ,Yuda Yürüm1 1
Faculty of Engineering and Natural Sciences, Sabanci University, Orhanli, Tuzla, Istanbul 34956, Turkey 2 TÜBİTAK Marmara Research Centre, Earth and Marine Sciences Institute, GebzeKocaeli, Turkey Corresponding author: e-mail:
[email protected] Abstract Coal bed methane (CBM) can arise from both thermogenic and biogenic activity on the coal beds and adsorb on the porous matrix of the coal. Therefore, investigation of pore structure and gas capacity of the coal is essential for accurate estimations of coalbed gas potential. Coal samples of lignite to sub-bituminous rank were obtained from different depths of Soma basin and were characterized by low pressure CO2 adsorption isotherms at 273 K. Micropore surface areas of the samples were calculated by using D-R model, changed from 232,653-274,73 m2/g. Micropore volume and capacity were determined by D-R equation to vary between 0.075 cm3/g and 0.92 cm3/g and between 40.63 cm3/g to 47.92 cm3/g, respectively. Pore widths of all samples were below 1 nm; suggesting that micropore ratios of the samples are very high. On the other hand, high pressure (up to 1 MPa) nitrogen and methane adsorption isotherms were determined by using Hiden Isochema Intelligent Gravimetric Analyzer (IGA-001) at room temperature. Results showed that methane adsorption on the samples increased with increasing micropore ratio. Effects of outgas temperature, organic carbon content on gas adsorption capacity of the samples were determined. Carbon isotope analyses of the coal gas desorbed from coal core samples of the Soma lignite basin in Turkey suggest bacterial origin. For purpose of better understanding of secondary biogenic gas potential of the samples biogasification experiments have been started. Firstly, coal samples were solubilizied by using Lewis bases. At moderate pH levels at moderate pH (9≥pH≥5) level carbonate and phosphate systems could solubilize coal efficiently. In the biogasification process, coal samples which were ground properly were incubated in anaerobic media with carbonate solution with microorganisms. After determination of optimal gasification parameters, methane generation due to the microbial activity will be calculated on daily basis.
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Oviedo ICCS&T 2011. Extended Abstract
1. Introduction Energy demand of the world is increasing constantly. At present, coal still keeps its value as one of the primary sources of energy to supply this demand. However, utilization of coal as an energy source has lots of negative impacts on the environment. For this reason, scientists have investigated alternative processes to produce clean energy from coal. In order to achieve that, extraction and production of natural gas from coal has become a more significant subject for the energy providers. Coal bed methane (CBM) production is a large and clean energy source with many advantages. In the USA, %10 percent of natural gas demand has been supplied by CBM [1]. CBM may have thermogenic or biogenic origin and the coal gas is adsorbed in the porous coal surface. In the last two decades, production of the secondary biogenic methane by utilization of additional microorganisms has been studied by scientists aiming to obtain more of clean energy from coal. In this study, our primary objective is to understand CBM capacity of the Soma coal basin. For this reason, porosity of the coal samples must be determined [2]. Usually, surface area and the porosity of the materials can be calculated through the N2 physical sorption experiment, in this method entire relative pressure range (10-8 to 1) can be analyzed without using high pressure equipments [3]. However, for microporous materials like carbon materials and zeolites physical sorption occurs at very low relative pressure ranges (10-8 to 10-3) and experiments that are conducted with N2 are less reliable due to the low diffusion rate and adsorption equilibrium in the pores between 0,5 to 1 nm at 77 K. It is also known that specifically for carbon materials experiments that are conducted at low temperatures such as N2 sorption causes pore shrinkage that leads to the low sorption equilibrium ([4,5,6]). Most important factors that affect physical interaction between absorbent and the absorbate are dynamic radius of the absorbate, temperature and solubility parameters of the materials. In the literature, there are many examples where carbon dioxide gas was used for microporous materials instead of nitrogen [7]. Since, dynamic radius of the CO2 is relatively smaller than that of N2 (CO2: 3,3 angstrom, N2: 3,6 angstrom [8,9]), also solubility parameter of the CO2 is far greater than nitrogen ( for CO2 δ=6.1 cal0.5cm1.5
, for N2 δ=2.6 cal0.5cm-1.5 ). Owing to these superior properties, interaction between
coal and the CO2 is better than N2-coal interaction [10, 11]. The last and the most important parameter is the temperature, for physical adsorption of the CO2, measurement
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Oviedo ICCS&T 2011. Extended Abstract
temperature of the isotherm can be 273 K or 298 K which means that we can avoid slow adsorption equilibrium, diffusion limitations at 77 K, also pore shrinkage of the coal at low temperatures can be overcome by using CO2 for the micropore characterization of the coal. Therefore, CO2 can reach narrow and wavy micropore structure of the absorbates due to the high diffusion rate which is called activated diffusion [12, 13]. With all these advantages, coal micropore characterization has been determined by CO 2 since 1964 [14, 15]. In 1984, Smith and Williams reported a relation between high pressure methane adsorption capacity and low pressure CO2 adsorption of coal by comparing the results of these experiment and observed close results which means that low pressure CO2 adsorption also gives idea of the CBM potential [16]. In 1982, Cohen & Gabrielle published the first report on the biological conversion of the coal by microorganisms [17]. Since that time, biological conversion of the coal has been a major area of interest for scientists. Biological treatment of the coal can be divided into two categories; first one is the removal of the sulfur, nitrogen, metals and other unwanted components of the coal and the second one is the conversion of the coal like liquefaction, microbial gasification and microbial pretreatment [18]. Usually, biological treatment of the coal takes place under mild conditions at low temperature and pressure unlike the classic thermo-chemical processes. For instance, during the thermo-chemical processes, formation of the gas products and liquid hydrocarbons from the coal have been carried out by the thermo catalytic breakdown of deeply buried organic matter at relatively high temperatures (> 80oC). On the other hand, in the anoxic biogasification processes, microorganisms cause degradation of the organic content (aromatic hydrocarbons) of the coal to produce gas and other hydrocarbons. 2. Experimental section For this study, coal samples collected from different depths of Soma basin were used. In order to understand the basic characteristic of our samples ultimate and proximate analysis were performed in TÜBİTAK Marmara Research Center (MRC) Energy Institute; results are shown in Table 1. Rock-eval pyrolysis was conducted for determination of level of maturity and type of the organic matter contained. Petrographic analysis yielded maceral types and huminite/vitrinite reflectance was also measured. All these results are given in Table 2. To understand origin of the coal bed gas,
13
C isotope
analyses were conducted by using Continuous Flow Gas Chromatography—Isotope Ratio Mass Spectrometer (GC-IRMS) at TÜBİTAK Marmara Research Center (MRC)
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Oviedo ICCS&T 2011. Extended Abstract
Earth and Marine Sciences Institute (EMSI). Low pressure CO2 micropore surface area and micro porosity experiments were conducted at 273 K by using Quantachrome Autosorb Automated Gas Sorption System. Samples were outgassed at 373 K for 6 h prior to measurements, this temperature was chosen due to the fact that high temperature could cause the collapse of the organic matrix of the coal and also low temperature cannot remove water molecules from pores. High pressure gas adsorption experiments were performed by using Hiden Isochema Intelligent Gravimetric analyzer (IGA-003) with nitrogen and methane. The IGA has a fully automatic microbalance system that allows measuring the weight change as a function of time, gas pressure and the sample temperature. The precision of the measurement can be controlled by a PC. Long term stability of microbalance is 0.1µg with a weighting resolution of 0.2 µg and temperature stability is 0.1oC. For nitrogen adsorption experiments samples were outgassed at 105oC 3 hours under 10-6mbar vacuum, for methane experiments samples were only outgassed under vacuum without any heat treatment. In order to understand the effect of outgas temperature to methane adsorption capacity of the coal, only one sample was outgassed at 105oC 3 hours. For nitrogen experiments, linear driving force mass transfer model was used to get asymptotic uptake for every pressure point at 298 K up to 9 bar. For methane experiment only 6 hours interaction time was used to get thermodynamic equilibrium at 298 K up to 9 bar without using PC control asymptotic uptake value. For demineralization experiments, first samples were stirred with 6N HCl for 24 h under nitrogen atmosphere, then filtered and washed with distilled water until the filtrate became neutral, immediately followed by %40 HF addition to HCl washed coal and stirred for 24 h under nitrogen gas in a nalgene beaker then filtered and washed with water [19]. Demineralized samples will be used for methane adsorption experiments to understand mineral matter effect. For biogasification experiment, anaerobic medium, that contains sodium carbonate and phosphate as Lewis bases for degradation of the complex structure of the coal matrix, was prepared. Then methanogenic bacteria will be added to the coal-anaerobic media to understand bioavailability of Soma lignite. 3. Results and Discussion Coal samples extracted from Soma basin are considered as low rank (lignite to subbituminuous) with high content of vitrinite /huminte maserals, high organic carbon
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Oviedo ICCS&T 2011. Extended Abstract
amount and low mineral matter. Results are shown in Table-1 and Table-2, therefore they have high micro porosity and high adsorption capacity compared to the other coal samples with the same rank. Vitrinite reflectance (%Ro) values of the samples vary between 0.42 and 0.48 and Pyrolysis Tmax values of the samples are between 393 and 412oC; both values suggest that coal samples are immature with respect to thermogenic gas generation (Table 2). Table 1. Ultimate and Proximate Analysis of the Samples Sample No JK - 1122 JK-1126 JK-1135 JK-1137
Original sample C% H%
N%
67,91 66,75 59,45 65,56
0,58 1,48 1,89 0,93
5,85 4,8 5,79 6,23
S%
Dry sample C% H%
N%
S%
O%
2,18 1,01 1,23 1,88
74,94 72,07 63,62 70,53
0,64 1,6 2,02 1
2,41 1,09 1,32 2,03
12,82 14,71 12,69 10,79
Sample No JK - 1122
Original sample Volatile Moisture % Matter % 9,37 3,52
Ash %
JK-1126
7,38
JK-1135 JK-1137
5,29 4,29 5,41 5,86
33,88
Fixed Carbon % 53,23
Dry sample Volatile Ash % Matter % 3,88 37,38
Fixed Carbon % 58,74
5,76
36,01
50,85
6,21
38,87
54,92
6,56
13,93
37,63
41,88
14,91
40,27
44,82
7,04
9,1
42,3
41,56
9,78
45,5
44,72
Table 2. Rock-Eval and Maceral Analyses Results Sample No JK -1122 JK-1126 JK-1135 JK-1137
Depth (m) 793.50-793.70 826.65 725.90-726.20 736.70-736.90
Tmax (oC) 396 412 408 393
Organic Carbon % 67.55 68.31 61.3 66.37
Huminite (%) 96 87 78 82
Liptinite (%) 2 1 19 16
Inertite (%) 2 12 3 2
Ro (%) 0.46 0.48 0.42 0.44
The desorbed gas from the coal samples were collected and used for the determination of the carbon isotopic ratio of the gasses. Results are shown in Figure 1. According to the results, majority of the collected samples are in biogenic region, represented as a green dot on the Figure 2, also one of the samples is in the mixed gas region but very close to the biogenic part, which is shown by red dot on the figure. These results confirm that origin of the CBM is the result of the bacterial activity [20].
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Oviedo ICCS&T 2011. Extended Abstract
Figure 1. Differentiation of biogenic and thermogenic gas Usually coal samples have high micro porosity therefore methane is adsorbed in this micropore structure. In order to understand primary biogenic methane potential of the Soma basin, porosity and micropore surface area of the samples were determined by CO2 sorption experiments. Adsorption isotherms obtained in this study are shown in Figure 2. D-R micropore surface area, micropore volume and micropore capacity are given in Table 3. CO2 isotherms indicate that micropore area and adsorption capacity of JK-1126 is 274,73m2/g and 47,92, respectively; the highest of all samples, . This result indicates that coal samples with high organic carbon ratio and high Ro values have greater micro porosity, and thus, gas adsorption capacity. Pore size distribution of all four samples show that micro pore size of the Soma lignite is reported below 1 nm (Figure 3.). Table 3. CO2 surface characterization results at 273 K Sample No JK - 1122 JK-1126 JK-1135 JK-1137
Organik Karbon % 67,55 68,31 61,3 66,87
DR surface 2
area (m /g) 248,891 274,73 224,909 232,653
DR micropore volume 3
DR micropore capacity (m3/g)
(m /g) 0,083 0,092 0,075 0,078
43,23 47,92 39,06 40,63
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Oviedo ICCS&T 2011. Extended Abstract
Figure 2. CO2 Adsorption Isotherms
Figure 3. Pore size distribution of the samples Nitrogen sorption in coal is a diffusion limited process, uptake is very low compared to methane and CO2. Nitrogen Isotherms up to 9 bar by using asymptotic values show that JK-1126 also has higher adsorption capacity than the others (see Figure 4.). Methane adsorption is far greater than nitrogen adsorption to the coal due to the strong interaction of the C-C Van der Waals forces and easy access (i.e.: smaller dynamic radius) to the micropore structure of the coal (see Figure 5). JK-1126 has the highest organic carbon ratio, as well as the highest adsorption capacity. Therefore, micropore structure of the coal is strongly related to the organic carbon pattern, adsorbent finds available sites in these organic patterns.
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Oviedo ICCS&T 2011. Extended Abstract
Figure 5. Methane adsorption isotherms at 298 K In the literature, there are many arguments about outgas temperature of coal, therefore outgas effect on the methane adsorption capacity of coal were investigated as shown in Figure 6., One sample was outgassed in vacuum at 105oC prior to adsorption, other sample was outgassed only in vacuum. Results show that vacuum outgassed sample has higher methane adsorption capacity than the other; this is taking to suggest that temperature could cause a collapsed micropore structure of coal, and disturbed structure cannot adsorb methane efficiently.
Figure 6. JK-1137 adsorption isotherms at 298 K with and without temperature outgas prior to experiment. After demineralization, methane sorption experiment will be performed.
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Oviedo ICCS&T 2011. Extended Abstract
Also, to understand bioavalibilty of the samples, biogasification experiments with the specific organisms have still been continued. 4. Conclusions The isotherm up to 9 bar shows that methane adsorption on coal is greater than the nitrogen adsorption isotherms. Micro porosity is strongly related to organic carbon ratio of the coal, moreover, gas adsorption capacity is directly proportional to micro porosity. 105oC is too high for the outgas process of coal, since temperature most likely cause collapse of the micropore structure. Lower temperature must be applied to get maximum gas adsorption capacity. References [1] Harris S.H. et al, Microbial and chemical factors influencing methane production in laboratory incubations of low-rank subsurface coals. International Journal of Coal Geology 2008;76:46–51 [2] Rolando M. A. Roque-Malherbe. Adsorption and Diffusion in Nanoporous Materials,CRC press, (2007) [3] Lozano-Castelló D., Cazorla-Amorós D., Linares-Solano A. Usefulness of CO2 adsorption at 273 K for the characterization of porous carbons. Carbon 2004;42:1231–1236 [4] Lowell S., Shields, J. E. Thomas Martin A.; Thommes M., Characterization of Porous Solids and Powders: Surface area, Pore Size and Density, Springer, (2004) [5] Anderson R. B., Bayer J., Hofer L. J. E., Determining surface areas from carbon dioxide isotherms. Fuel 1965;44:443-452 [6] Walker P. L. Jr., Geller I. Change in surface area of anthracite on heat treatment. Nature (London) 1956;178:1001[7] Mastalerz et al. Meso- and Micropore Characteristics of Coal Lithotypes: Implications for CO2 Adsorption. Energy & Fuels 2008;22, 4049–4061 [8] Breck D.W. Zeolite Molecular Sieves Wiley, New York (1974),p. 634 [9] Kenneth S.W.Sing, Ruth T.Williams, Review: The Use of Molecular Probes for the Characterization of Nanoporous Adsorbents. Part. Part. Syst. Charact. 2004;21:71 – 79 [10] Reucroft P. J., Sethuraman A. R. Effect of pressure on carbon dioxide induced coal swelling. Energy & Fuels 1987;1:72-75 [11] Barton A. F. M. Handbook of Solubility Parameters and Other Cohesion Parameters; CRC Press: Boca Raton, FL, (1983) [12] Nandi S. P., Walker P. L. Jr. The diffusion of nitrogen and carbon dioxide from coals of various rank. Fuel 1964;43:385-93 [13] Cazorla-Amoro´s D. et al. CO2 As an Adsorptive To Characterize Carbon Molecular Sieves and Activated Carbons. Langmuir 1998;14:4589-4596 [14] Marsh H. Determination of Surface Areas of Coals - Some Physicochemical Considerations. Fuel 1965;44:253-68 [15] Marsh H., Siemieniewska T. The surface area of coals as evaluated from the adsorption isotherms of carbon dioxide using Dubinin–Polanyi equation. Fuel 1965; 44:355-67 [16] Smith D. M., Williams F.L. Coal Sience and Chemistry(A. Volborth, Ed.), Elsevier,Amsterdam 1987, p:381
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[17] Cohen M.S., Gabriele P.D. Degradation of coal by the fungi Polyporus versicolor and Poria monticola. Applied and Environmental Microbiology 1982;44:23-27 [18] Narayan R., Ho N.W.Y. Objectives of coal bioprocessing and approaches. Am. Chem. SOC., Div. Fuel Chem. Prep. 1988;33:487-495 [19] Altuntaş, N.;Yürüm, Y. Effect of Catalysts on the Pyrolysis of Turkish Zonguldak Bituminous Coal. Energy & Fuels 2000;14:820-827 [20] İnan S. Et al., Coalbed Gas Potential In the Miocene Soma Basin (Western Turkey). 27th Pittsburgh Coal Conference 2010; Abstract Booklet:pp.21
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Oviedo ICCS&T 2011. Extended Abstract
Decoupling in Thermochemical Conversion: Approach and Technologies Guangwen Xu, Juwei Zhang, Yin Wang, Shiqiu Gao State Key Laboratory of Multi-phase Complex System, Institute of Process Engineering,Chinese Academy of Sciences, P. O. Box 353, Beijing 100080, P. R. China : Tel/Fax：+86-10-62550075, E-mail:
[email protected] Abstract All thermochemical conversion processes for coal, biomass and solid wastes involve a series of chemical reactions as well as physical variations including fuel drying/pyrolysis, char gasification, tar reforming/cracking, combustible matter combustion, water gas shift and so on. These reactions are mutually interactive to form a complicated reaction network.
In the conventional thermochemical conversion
technologies (processes), these reactions are coupled together to implement the conversion in a single reaction vessel, which may result in some undesirable effects such as low efficiency, low-value/quality products, and high pollutant emission.
The
“decoupling” means to control the interactions among different reactions to facilitate the beneficial interactions or to suppress the undesirable interactions to optimize the performances of thermochemical conversion, including facilitating the conversion for higher efficiency, improving product quality/value, increasing fuel adaptability, suppressing pollution emission, realizing poly-generation and any other benefits. The implementation of “decoupling” comprises first the breakage of inter-linked reactions to separate one or more interacted reactions and then reorganization of the separated reactions. This article highlighted the decoupling principle, and both “isolating” and “synergizing” were generalized as the two different approaches of decoupling. On the basis of the decoupling approaches, a series of decoupling thermochemical conversion (DTC) technologies were highlighted to understand their process principle and realize technology superiorities. In turn, an overview on the progress of a few typical DTC technologies under development in Institute of Process Engineering (IPE), Chinese Academy of Sciences (CAS) was presented to summarize the major results in both fundamental studies and pilot or demonstration tests. The technologies referred to in the article included the dual fluidized bed gasification (DFBG), dual bed pyrolysis Submit before January 15th to
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gasification (PG), two-stage gasification (TSG), topping combustion (TC), low-NOx fluidized bed combustion (LFBC) and decoupling combustion (DC). This summarization validated that the decoupling can provide a viable idea and practical approach of technology innovation for developing new conversion technologies that overcome the inherent disadvantages of the traditional conversion technologies and thereby realize the decoupling effects of, for example, high efficiency, poly-generation, high quality/ value of products, wide fuel adaptability, and low pollutant emission etc. 1. Introduction The thermochemical conversion supplies the major technical approaches for utilizing solid carbonaceous fuels including coal, biomass and municipal wastes. It is shown generally via three different conversion patterns: pyrolysis (carbonization or coking), gasification and combustion. In the process of each of these conversion patterns, not a single but a series of reactions occur to incur the explicit chemical changes. Figure 1 highlights the chemical behaviors involved in the thermochemical conversion of solid fuels. It is noteworthy that in Fig. 1 the words or terminologies pyrolysis, gasification and combustion refer to their individual reactions in the middle rectangle box rather than to their macro processes in the last column of the figure.
Solid fuel Oxidant (O2) Reagent (H2O, CO2, H2)
(1) Drying (2) Pyrolysis (3) Cracking/ decompositon (4) Polymerization (5) Hydrogenation (6) Gasification (7) Combustion (8) Reforming (9) Water gas shift …….
Sufficient O2 Enough long t and high T
Combustion: CO2, H2O, heat
Insufficient O2 Enough long t and high T
Gasification: CO, H2, CmHn
Lean in O2 Short t and low T
Pyrolysis: C (char), CO, H2 CmHn, (C6Hm)xOy
Fig.1 Chemical behaviors of solid fuels during the thermochemical conversion process
The reactions in the rectangle box are inter-correlated or interactive, and some are even sequential in occurrence. Under heating, the solid fuel is first dried and pyrolyzed to produce char, tarry oil (tar), steam and uncondensable pyrolysis gas consisting mainly of H2, CO, CO2, and CH4. In turn, the other reactions in box start to occur and lead to a series of interactions between/among the product/products of some reactions with the other reactions. These interactions make the involved reactions in the thermochemical
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conversion process, such as those listed in the rectangle box in Fig.1, closely intercorrelated to form a complicated reaction network described in a previous publication of ours [1].
Of these interactions, some can facilitate the conversion to lead to high
efficiency, low pollution emission and high product quality, whereas the others are not. Therefore, it deserves to control the interactions for optimizing the thermochemical conversion process.
In most commercialized as well as in-developing processes of
pyrolysis, gasification and combustion, all such reactions are arranged into a single reaction space and it is thus impossible to control any individual reaction and its interactions with the other reactions. This causes the problems related to low product quality or value, low conversion efficiency, poor fuel adaptability, high pollutant emission and so on. In order to manipulate individual reaction to avoid or weaken the undesired but strengthen the desired interactive effects, separation of the related reactions and in turn rearrangement of the separated reactions is necessary and worthwhile. This idea of reaction control has been termed “decoupling” in our previous publication [1]. Two implementation approaches for decoupling, isolating and synergizing, have been also developed and through their applications to gasification it has been found that the “isolating” and “synergizing” approaches correspond rightly to the conversion technologies based on “dual bed” and “staging”, respectively. Figure 2 summarizes the preceding inter-correlations among the reactions involved in the thermochemical conversion process, two implementation approaches of the decoupling, their implicated characteristic conversion technologies and the expected technical effects from implementing the decoupling. As one may image, the implementation of the decoupling consists of two sequential actions upon the involved reactions, separation of one or more reactions from the inter-correlated reaction network (see Ref. [1]) by breaking the linkages of the reaction or reactions with the others, and rearrangement of the separated or decoupled reactions according to the needs of reaction control. If the decoupled reactions are arranged into isolated reactors to separate their products and fully suppress the interactions between their products, the resulting decoupling is the “isolating” and the technology is dual bed conversion, realizing consequently the effects of poly-generation, high efficiency, high product quality, and wide fuel adaptability.
If the decoupled
reactions are reorganized to facilitate the beneficial interactions or to suppress the undesired interactions, the utilized decoupling is the so-called “synergizing” and the corresponding conversion technologies are usually related to “staged” shown explicitly as Submit before January 15th to
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Oviedo ICCS&T 2011. Extended Abstract
two-stage processes, fuel staging and reactant-staging. technologies
These staged conversion
are most effective to lowering pollution emission, raising efficiency and
product quality, and allowing wide fuel adaptability. In practice, the decoupling may be also implemented by executing both “isolating” and “synergizing” to have the creatively new designs for thermochemical conversion technologies. Combustion Reaction network
Fuel
Char
Residua Char
Pyrolysis
Gasification Break
Decoupling approach
Isolating (reacting independently isolated product stream)
Synergizing (Re-organization one product stream)
Derived technology
Dual-bed conversion
Staged conversion
Decoupling effect
Poly-generation High efficiency High product quality Wide fuel adaptability
Low pollution High efficiency High product quality Wide fuel adaptability
Fig. 2 Two decoupling approaches applied in the reaction network of thermochemical conversion process (taking the simple reaction network of pyrolysis, gasification and combustion for example))
In succession to our previous publication [1], this article is devoted to generalizing the conception and applications of the decoupling to all types of thermochemical conversion technologies or processes. Table 1 summarizes some typical and well-known thermochemical conversion technologies based on decoupling and their realized major decoupling effects (i.e., technology advantages). These technologies mainly involve three types of thermochemical conversion process: pyrolysis, gasification and combustion, and here pyrolysis includes also coking. It can be seen that all of these technologies have been commercialized or in the progress toward commercialization. Herein, an overview on the progress of a few typical decoupling thermochemical conversion (DTC) technologies under development in Institute of Process Engineering (IPE), Chinese Academy of Sciences (CAS) was presented to summarize the major results in both fundamental studies and pilot or demonstration tests. The technologies referred to in the article included the Submit before January 15th to
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dual fluidized bed gasification (DFBG), dual bed pyrolysis gasification (PG), two-stage gasification (TSG), topping combustion (TC), low-NOx fluidized bed combustion (LFBC) and decoupling combustion (DC). Table 1 Typical thermochemical conversion technologies with different decoupling approaches and corresponding effects Process
Decoupling approach
isolating
pyrolysis isolating
isolating and synergizing
isolating
gasification
isolating
synergizing
isolating
combustion
synergizing
synergizing
Decoupling effects Using heat of exhaust gas to regulate the moisture in coal and realize utilization of poor-coking coal, high productivity, energy saving Rapid preheating of the coal charge to realize utilization of weak-caking coal, high productivity, energy saving Multi-stage pyrolysis to improve high product quality and fuel adaptability of the technology Avoiding dilution of produced gas by N2 and combustiongenerated CO2 to produce middle-caloric fuel gas using air as a gasification reagent Co-producing pyrolysis oil and fuel gas or syngas to use coal hierarchically Reforming/cracking of tar and pyrolysis gas by catalysis of char to produce fuel gas or syngas with little tar Co-producing pyrolysis oil, pyrolysis gas, steam and elecitricty to use coal hierarchically Reduction of NO by char and pyrolysis gas to lower NO emission in the flue gas Introduction of pyrolysis gas into char bed to lower NO and CO emission in the flue gas
Main decoupled reaction (s)
Typical technology
drying
CCMC
commercialization
drying and pyrolysis (low extent)
SCOPE21
commercialization
pyrolysis (low extent)
COED
commercialization
combustion
DFBG
demonstration
pyrolysis
PG
pilot
pyrolysis
TSG
pilot
pyrolysis
TC
pilot
pyrolysis
LNFC
pilot
pyrolysis
DC (grate furnace)
commercialization
Application stage
a
a
The full names of the typical DTC technologies are coking with coal moisture control (CCMC) [2], super coke oven for productivity and environmental enhancement toward the 21st century (SCOPE21) [3], char oil energy development (COED) [4], dual bed gasification (DBG) [5], pyrolysis gasification (PG) [1], two-stage gasification (TSG) [6], topping combustion (TC) [7], low-NOx fluidized bed combustion (LFBC) and decoupling combustion (DC) in grate furnace [8].
2. Application to Gasification The gasification technologies based on the decoupling have been named the decoupling gasification (DCG) technologies in our previous publication [15]. By far, the decoupling has been widely applied to the development of gasification technologies, and Fig. 3 highlights the process principles of three typical DCG technologies developed in IPE. Of them, the DFBG and PG adopted the “isolating” decoupling approach, while the
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TSG is based on “synergizing” decoupling approach. As illustrated in Fig. 3, the DFBG was created by decoupling the char combustion from all the other gasification reactions, While the PG and TSG were born from decoupling and in turn reorganization of the pyrolysis reaction. Table 1 also briefs the realized decoupling effects for all the mentioned DCG technologies, and these effects, as detailed in our previous publication [1], have been well validated through pilot or bench tests or demonstration applications. This results in the demonstration of that the DCG technologies innovated with either isolating or synergizing approaches or both of them can effectively lower emission, raise efficiency, increase product quality or/and allow wide fuel adaptability. Flue gas
Product gas
Pyrolysis gas+tar
Riser gasifier
Riser combustor
Product gas
Fuel BFB gasifier
Fuel Distributor
Pyrolyzer
Ash
Pyrolysis gas or N2
Steam+air Air
Air (a) DFBG
(b) PG
(c) TSG Fig. 3 Schematic diagram of three DCG technolgies
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3. Application to Combustion As far, there are some application examples of decoupling approaches for the combustion process, and many of them are developed in China [7-12]. Both the isolating and synergizing approaches have been used in these combustion technologies. 3.1. Principle Figure 4 (a) shows the principle diagram of one typical combustion technology using the isolating approach. This technology was named “topping combustion (TC)” which originated from the conception of “topping” in petroleum industry and it was proposed at the Institute of Process Engineering (IPE), Chinese Academy of Science (CAS). The TC is a poly-generation technology which can co-produce light oil, pyrolysis gas, heat, and electricity. As shown in Fig. 4 (a), the linkage between pyrolysis and char combustion is broken, and these two reactions proceed in the separated reactor, which indicates the isolating approach is used. The fuel (coal) is pyrolyzed quickly in a downer reactor after mixing with the hot sand from the char riser combustor, then the produced gas (volatiles) is separated from the solid and quenched quickly to extract the light oil (the conception of “topping” is manifested here) while the secondary reactions are minimized. For this technology, the key to achieve success is the word “fast” including the fast mixing of coal with the hot sand particles, fast gas-solid separation, and the fast cooling of the volatiles. In the combustion process, the effect of low NOX emission can generally be realized through the synergizing approach, as shown in Fig. 4 (b). From Fig. 4 (b), it can be seen that after the linkage between the pyrolysis and char combustion is broken, the pyrolysis gas is introduced into the char combustion zone to make use of the interaction between the pyrolysis gas and char or the products of char combustion. Perhaps the most important interaction is the NO-char reaction whose fundamentals have been investigated in many studies in fixed beds [13] and entrained flow reactors [14] for both coal char and biomass char. Besides, the NO from char combustion can also be reduced by the reductive components such as CH4, H2 and CO in pyrolysis gas [15]. The NO formed in both pyrolysis reaction and char combustion (i.e., the NO formed from char-N and volatile-N) can be obviously decreased, compared to regular combustion technologies in which the pyrolysis reaction and char combustion proceed in a single reactor.
Combustion process (isolating) Flue gas
Hot sands Break
Combustion Pyrolysis Submit before January 15 to
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Pyrolysis gas tar Fuel
th
Air Cold sands +char
7 Fluidization gas
Oviedo ICCS&T 2011. Extended Abstract
3.2. Topping combustion The process diagram of TC technology is shown in Fig. 5(a). The TC system mainly includes a solid-solid mixer, a downer pyrolyzer, a gas-solid separator, and the riser combustor. The product distribution of the produced gas from the downer pyrolyzer was investigated in a lab-scale (8 kg/h) electricity-heated TC apparatus in which the downer pyrolyzer and riser combustor have the inner diameters of 0.039 m and 0.086 m. The details of this apparatus and experimental procedures can be referred to literature 7. A Chinese lignite which had 6.0% moisture, 36.2% volatiles, 33.5% fixed carbon, and 24.3% ash by weight (dry basis) was used. The yields of liquid including tar, light tar and water were increased with the increasing of pyrolysis temperature and the yield of light tar could reach 7.5% when the pyrolysis temperature was about 930 K. The light tar mainly consisted of acid group, aliphatic, aromatic, and polar & basic group, and their yields were 4.3%, 1.0%, 1.6%, and 0.6%, respectively.
These results indicate the high-value
compounds contained in coal can be utilized with high efficiency in the TC technologies. At present, a 5 ton/h pilot-scale TC apparatus has been built in IPE and the experiments are in progress. Submit before January 15th to
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Flue gas + ash Fuel Mixer
Riser combustor
Secondary air
Pyrolyzer (downer) Product gas
Gas-solid separator
Cooler
Char + hot sands
Tar Fluidization gas
Primary air (a) TC Flue gas
Secondary air
Riser combustor
Flue gas
Pyrolysis gas
Pyrolysis
Postcombustion
zone
Fuel Combustion zone
Pyrolyzer Char + hot sands
Fluidization gas (N2+air)
Secondary air Primary gas
Primary air (b) LFBC
(c) DC
Fig. 5 Schematic diagram of three combustion technologies with decoupling approaches
3.3. Low-NOx combustion Figure 5(b) and (c) show the process diagrams of two low-NOx combustion technologies in fluidized bed and grate furnace (fixed bed) with synergizing approaches respectively, and they were both proposed in IPE. The fluidized bed low-NOx combustion system, as shown in Fig. 5 (b), mainly consists of a bubbling fluidized bed pyrolyzer and a riser char combustor. This system is actually modified from the 50 kg/h pilot-scale PG apparatus shown in Fig. 3 (b) by introducing the produced pyrolysis gas from pyrolyzer
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into the middle of the riser char combustor (i.e. char gasifier in PG), resulting in NOx reduction by char and pyrolysis gas. In order to validate the decoupling effects of this low-NOx combustion technology, the experiments of both regular CFB combustion (i.e., the pyrolysis reaction and char combustion both occur in the riser) and low-NOx combustion were performed using a Chinese lees which had 43.8% moisture, 39.7% volatiles, 9.4% fixed carbon, and 7.1% ash by weight. Figure 6 compares the volume fractions of the main flue gas components produced from regular combustion with that from low-NOx combustion.
Compared with the regular CFB combustion, the NOx
concentration can be decreased from 800 ppm to 100 ppm, indicating the remarkable NOx reduction effects of decoupling in combustion process with synergizing approach. O2
CO
25
900
20
750 600
15
450 10
300
5 0
150 0
5
10 15 Time (min)
20
30
1050
CO2 25
NOx
O2
CO
600
20
450
15
300
10
150
0
5 0
0 0
(a)
20
40 60 Time (min)
80
NOx concentration (ppm)
NOx
Gas composition (vol.%)
Gas compositon (vol.%)
CO2
NOx (ppm)
30
100
(b)
Fig. 6 Volume fractions of the main flue gas components produced from: (a) Regular CFB combustion; (b) fluidized bed low-NOX combustion with synergizing approach
The low-NOx combustion can also be implemented in grate furnace [8], as shown in Fig. 5 (c). This type of combustion technology is generally called decoupling combustion (DC) [8]. The furnace is divided into pyrolysis zone, combustion zone (above grate), and burnout zone. The fuel are pyrolyzed quickly in the pyrolysis zone, and the produced char and pyrolysis gas enter into the char combustion zone where the NOx reduction by char and pyrolysis gas occur, then the residual char and gas move into the burnout zone to burn with secondary air. Dong et al. [16] measured the concentration of gas composition in flue gas and the averaged temperature in a 4.5 kg/h DC grate furnace and a 3 kg/h regular combustion grate furnace using a blend of 50% rice husk and 50 % coal by weight, the results are presented in Fig. 7. The two combustion patterns have a same excess air ratio. It can be obviously seen that the DC has a lower NO, lower CO emission level, and higher furnace temperature than regular combustion.
The higher CO emission for regular
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combustion can be attributed to that the fuel pyrolysis occurs above the char bed and the produced volatiles are easy to escape directly with flue gas without experiencing the high temperature char bed. Via DC, the NO concentration is reduced by 19% compared to regular combustion, though the higher CO concentration and temperature facilitate the NO formation [17, 18]. All these results justify the decoupling effects of low pollution, high efficiency and stable combustion via DC. It should be noted that the DC can be easily employed by modifying the existing grate boiler. Currently, the 1 t/h (evaporation) DC grate boiler has already been commercialized and the scale-up works (2-10 t/h) are in progress. 10000
DC RC
150
1200 8000
1000
6000 90 4000 60
800 CO (ppm)
NO (ppm)
120
600 400
2000
30 0
0 NO
CO
Averaged temperature
180
200 0
Temperature
Fig. 7 Performance comparisons between DC with synergizing approach and regular combustion (RC) in grate furnace [16]
4. Conclusions Based on the DCG proposed in our previous paper, the conception of decoupling has been extended to more general process: thermochemical conversion which commonly includes pyrolysis, gasification and combustion process.
When the decoupling
approaches are applied in the typical processes of thermochemical conversion, various types of novel technologies can be formed. In order to generalize the application of decoupling in the thermalchemical conversion and justify the superiority of the decoupling approaches, some typical decoupling application examples developed in IPE are introduced and the related experimental results are reanalyzed. As for combustion technologies, the principle of TC are based on the isolating of pyrolysis reaction from other reactions in the whole combustion process to separate the
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Oviedo ICCS&T 2011. Extended Abstract
pyrolysis products to realize the decoupling effects of poly-generation and high product quality, and the principle of low-NOX combustion in both fluidized bed and grate furnace is based on the decoupling of pyrolysis reaction from the whole combustion process to take advantage of the interactions among pyrolysis gas, char and NO to realize the decoupling effects of low pollution. Based on these result analyses, the terminology of decoupling thermochemical conversion (DTC) was proposed to distinguish the conversion technologies based on decoupling from the conventional thermochemical conversion technologies. It is anticipated for this paper that more and more DTC technologies would be derived from the conception of decoupling, since the DTC technologies can realize lots of desired decoupling effects which could not be easily obtained in the conventional thermochemical conversion technologies. Acknowledgement The authors are grateful to the financial support of the Natural Science Foundation of China (contract No: 21006110,) and Key Projects in the National Science & Technology Pillar Program (contract No: 2009BAC64B05). References [1] Zhang JW, Wang Y, Dong L, Gao SQ, Xu GW. Decoupling gasification: approach principle and technology justification. Energy Fuels 2010;24:6223–6232. [2] Kenji K, Seiji N. Coal pretreating technologies for improving coke quality. Proceeding of the 5th Internationall Congerss on the Science and Technology of Ironmaking, Shanghai, China, 2009, 361–366. [3] Taketomi H, Nishioka K, Nakashima Y, Suyama S, Matsuura M. Research on coal pretreatment process of SCOPE21. 4th European Coke and Ironmaking Congress Proceedings, Paris, France, 2000, 640–645. [4] Strom AH, Eddinger RT. COED plant for coal conversion. Chemical engineering progress 1971;67:75–80. [5] Xu GW, Murakami T, Suda T, Matsuzawa Y, Tani H. Gasification of coffee grounds in dual fluidized bed performance evaluation and parameter influence. Energy Fuels 2006;20: 2695–2704. [6] Henriksen U, Ahrenfeldt J, Jensen TK, Gobel B, Bentzen JD, Hindsgaul C, Sorensen LH. The design, construction and operation of a 75 kW two-stage gasifier. Energy 2006;31:1542–1553. [7] Wang JD, Lu XS, Yao JZ, Lin WG, Cui LJ. Experimental study of coal topping process in a downer reactor. Ind. Eng. Chem. Res. 2005;44:463–470. [8] Li JH, Bai YR, Song WL. NOx-suppressed smokeless coal combustion technique. Proceedings of International Symposium on Clean Coal Technology, Xiamen, China, 1997, 344–349.
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Oviedo ICCS&T 2011. Extended Abstract [9] Fang MX, Wang QH, Yu CJ, Shi LZ, Luo ZY, Cen KF. Development of gas steam and power multi-generation system. Proceedings of the 18th International Conference on Fluidized Bed Combustion, Toronto, Ontario, Canada, 2005, 22–25. [10] Fang MX, Cen JM, Wang QH, Shi ZL, Luo ZY, Cen KF, 25 MW circulating fluidized bed heat-power-coal gas poly-generation installation. Journal of Power Engineering (in Chinese) 2007;27:665–639. [11] Chinese Patent 01218480.2. [12] Yao JZ, Wang XQ, Lin WG, Li JH, Kwauk MS. Coal topping in a fluidized bed system. 16th International Conference on Fluidized Bed Combustion, Reno, NV, 2001, 13–16. [13] Dong L, Gao SQ, Song WL, Xu GW. Experimental study of NO reduction over biomass char. Fuel Processing Technology 2007;88:707–715. [14] Sun SZ, Zhang JW, Hu XD, Wu SH, Yang JC, Wang Y, Qin YK. Studies of the NO-char reaction kinetics obtained from drop tube furnace and thermogravimetric experiments. Energy Fuels 2009;23:74–80. [15] Giral I, Alzueta MU. An augmented reduced mechanism for the reburning process. Fuel 2002;81:2263–2275. [16] Dong L, Gao SQ, Song WL, Li JH, Xu GW. NO reduction in decoupling combustion of biomass and biomass-coal blend. Energy Fuels 2009;23:224–228. [17] Zevenhoven R, Hupa M. The reactivity of chars from coal, peat and wood towards NO, with and without CO. Fuel 1998;77:1169–1176. [18] Aarna I, Suuberg EM. The Role of Carbon Monoxide in the NO-carbon Reaction. Energy Fuels 1999;13:1145–1153.
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Integrated Process of Coal Pyrolysis with CH4/CO2 Activation by Dielectric Barrier Discharge Plasma Xinfu He, Haoquan Hu*, Lijun Jin, Yunpeng Zhao State Key Laboratory of Fine Chemicals, Institute of Coal Chemical Engineering, School of Chemical Engineering, Dalian University of Technology, Dalian 116024, P. R. China, *
[email protected] Abstract An integrated process to combine coal pyrolysis with CH4/CO2 activation by dielectric barrier discharge (DBD) plasma was reported. Coal pyrolysis was carried out under seven atmospheres, including N2, H2 and different plasma of H2, CH4, CH4/H2, CO2/H2, CH4/CO2/H2 (MG) to confirm the effect of integrated process of coal pyrolysis with CH4/CO2 activation by DBD plasma on improving tar yield. The results showed that the effect of seven atmospheres on tar yield generally has the following order: MG plasma > CH4 plasma ≈ CH4/H2 plasma ≈ CO2/H2 plasma > H2 plasma > H2 > N2, especially at low temperature range. The effect of discharge power on product yields were also investigated. The results showed that the tar yield of Shenmu coal under optimum condition is about 2.0 and 1.8 times as that under N2 and H2 at 400 °C, respectively.
1. Introduction Coal tar from low temperature pyrolysis of coal could be one of the most important sources of fuel oil and chemicals. In general, the tar yield in coal pyrolysis process is low. To improve the tar yield, many methods, including changing the pyrolysis atmosphere, [1-6] pre-treatment of coal, [7] catalytic pyrolysis [4, 7] and catalytic hydropyrolysis [4, 8] were explored. Our recent studies indicated that tar yield could increase remarkably by integrating the coal pyrolysis with partial oxidation [9] or carbon dioxide reforming [10, 11] of methane (POMP or CRMP). However, POMP process is restricted for safety concern, while CRMP process has to be operated at above 800 °C because CH4 is hard to be activated, which is higher than the optimal coal pyrolysis temperature range for high tar yield (500-600 °C), resulting in the formation of large amount of water. Many studies have been carried out using dielectric barrier discharge (DBD) plasma for methane conversion [12-15] because it can be operated at ambient temperature and atmospheric pressure, and can be initiated in large scale. In this work, DBD plasma was
1
Oviedo ICCS&T 2011. Extended Abstract
used for activating CH4/CO2 to find out whether the integrated process of coal pyrolysis with CH4/CO2 activation by DBD plasma could improve the tar yield.
2. Experimental Coal Sample Two Chinese coal samples, Huolinhe (HLH) lignite and Shenmu (SM) subbituminous coal, were crushed and sieved to 40-60 mesh for pyrolysis. The proximate and ultimate analyses of the coal samples are shown in Table 1.
2.2 Apparatus and Procedures. The schematic diagram of the experimental setup and the reactor as well as the method for measuring discharge power was described elsewhere. [16] The feed gas includes N2, H2 and different plasma of H2, CH4, CH4/H2 (1:2), CO2/H2 (1:2), CH4/CO2/H2 (1:1:2, denoted as MG) at atmospheric pressure. During the experiment, the feed gas was first introduced to the reactor for several minutes, then gas discharge was initiated and the reactor was loaded into the center of the preheated furnace (400 to 650 °C). The heating time from ambient temperature to desired pyrolysis temperature was about 3 min and the reactor was held at the temperature for a desired time. The liquid products, tar plus water, were collected in a cold trap at -10 °C, and then the water in the liquid products was separated according to ASTM D95-05e1 (2005) using toluene as solvent. In this way, tar and water yield could be calculated, respectively.
3. Results and Discussion 3.1 Effect of Temperature under Different Atmospheres. The effect of pyrolysis temperature on char, tar and water yields under N2, H2 and different plasma atmospheres are shown in Figure 1. It can be seen that char yield is the highest while tar and water yields are the lowest under N2 at the investigated temperature range, and it is the same trend as those under H2 at low temperature range.
2
Oviedo ICCS&T 2011. Extended Abstract
Coal conversion as well as tar and water yields under H2 is higher compared with those under N2 when temperature exceeds 550 °C. H2 at low temperature range is an inert gas equivalent to N2 and has no effect on coal conversion, but as the temperature increases, it can be activated and has effect on stabilizing the free radicals cracked from coal, resulting in the increase of tar and water yields. When plasma discharge was introduced, H2 can be activated or dissociated to H, H+, H2+, H3+, etc., which have higher reactivity than molecular hydrogen, [17] that’s the reason for the lower char yield and higher tar yield under H2 plasma than that under H2 at low temperature range. However, H2 molecule can be activated thermodynamically at high temperature range, which can explain why char and tar yields have little difference under H2 plasma compared with that under H2 at high temperature range. Higher water yield under H2 plasma may be ascribed to the high-energy H species excited by discharge, which has higher ability to react with oxygenous groups in coal.
90
25
N2
CH4 P
H2
CH4/H2 P
H2 P
CO2/H2 P MG P
80 70 60
400
450
500
550 o
Temperature( C)
600
6
(b)
650
Water yield(wt%,daf)
Char yield(wt%,daf)
(a)
Tar yield(wt%,daf)
100
20 15 10 5 0
400
450
500 550 600 o Temperature( C)
650
5
(c)
4 3 2 1 0
400
450
500 550 600 o Temperature( C)
650
Figure 1. Effect of temperature on char (a), tar (b) and water (c) yield under different atmospheres (240 ml/min, holding time: 7 min, Pdis: 40 W) Other plasma atmospheres showed positive effect on coal conversion and displayed lower char yield than that under H2 plasma at low temperature range (400-500 °C). Higher tar yield can be achieved in the integrated process of coal pyrolysis with gas activation by DBD plasma than that in the single process of coal pyrolysis under N2 or H2. Tar yield under the studied atmospheres generally has the following order: MG plasma > CH4 plasma ≈ CH4/H2 plasma ≈ CO2/H2 plasma > H2 plasma > H2 > N2. CH4 was dissociated to ⋅CH3, ⋅CH2, ⋅CH, ⋅C and H⋅ species in CH4 plasma, [18] these activated species combined with free radicals ruptured from coal lead to the increase of tar yield at low tamperature range, as can be seen in Figure 2(b). However, CH4 is mainly dissociated to ⋅C and H⋅ when temperature exceeds 500 °C, and will cause serious carbon deposit which can further affect the discharge. So, tar yield under CH4 plasma approaches to that under N2 while water yield approaches to that under H2 plasma when the temperature is higher than 500 °C. The addition of H2 to CH4 has the
3
Oviedo ICCS&T 2011. Extended Abstract
effect on increasing tar yield and decreasing water yield compared with that under CH4 plasma at high temperature range because H2 can eliminate carbon deposit and maintain a steady DBD. [19, 20] Tar yield under CO2/H2 plasma is about the same as that under CH4/H2 plasma when temperature is below 500 °C, but approaches to that under N2 when temperature is above 500 °C. Water yield under CO2/H2 plasma is the highest in all the experiments and keeps increasing with the increase of temperature. It’s about several times more than that under N2 and H2 at low temperature range but it’s still no more than 6 %. The high tar under CO2/H2 plasma may ascribed to the combination of oxygenous radicals and CHx radicals formed in CO2/H2 plasma [21-23] and those from coal, while high water yield due to the reverse water gas shift reaction and methanation reaction which can easily occur under the experiment conditions. MG plasma has both the advantages of CH4/H2 plasma and CO2/H2 plasma. Tar yield under MG plasma is the highest compared with that under other atmospheres, and it’s about 100 % and 77 % more than that under N2 and H2 at 400 °C, respectively. The increase of tar yield under MG plasma is more evident at low temperature range, which can be seen from Figure 2(b) that tar yield under MG plasma is about 2 and 1.1 times as that under N2 at 400 °C and 600 °C, respectively. While water yield under MG plasma is lower than that under CO2/H2 plasma though it’s higher than that under any other atmospheres. HLH SM
75 70 65 30
40
50
60
Discharge power(W)
Water yield(wt.%,daf)
80
60
10
24 (a)
Tar yield(wt.%,daf)
Char yield(wt.%,daf)
85
(b)
22 20 18 16 14
30
40
50
Discharge power(W)
60
(c)
8 6 4 2 0
30
40
50
60
Discharge power(W)
Figure 2. Effect of discharge power on char, tar and water yields under MG plasma atmosphere (500 °C; 240 ml/min; 7 min) 3.2 Effect of Discharge Power. The effect of discharge power on char, tar and water yields under MG plasma is shown in Figure 2. With increasing discharge power, both coal conversion and water yield increase in varying degrees. However, tar yield of HLH coal increases first, then decreases and has a maximum at discharge power of 40 W, while that of SM coal keeps
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Oviedo ICCS&T 2011. Extended Abstract
increasing. Discharge power is one of the most important factors affects the activation level of the feed gas. The energy and the number of the electrons in the discharge zone are directly determined by the input energy which is depending on discharge power. As the discharge power increasing, the number of high-energy electrons and the collision frequency between electron and gas molecular increases, leading to formation of more activated species [24-26] and finally resulting in the increase of volatiles. However, carbon deposit which can seriously affect the discharge is unnegligible when increasing the discharge power. Tar yield could be significantly increased by increasing discharge power on condition that the discharge is uniform and stable.
4. Conclusions Experimental results showed that coal pyrolysis coupling with gas activation through DBD plasma has effects on coal conversion and product yield, especially at low temperature range. Compared with that under H2 atmosphere, tar yield has little change but water yield shows a significant increase under H2 plasma, while tar yield increases remarkable and water yield decreases under CH4 plasma at low temperature range. CH4 plasma discharge was difficult to be maintained stable and uniform at above 500 °C because of the carbon deposit. Decreasing temperature and addition of H2 and/or CO2 into CH4 can stable the discharge and lead to the increase of tar yield. Tar yield under MG plasma is about two times of that under N2 at 400 °C. CO2/H2 plasma can also increase tar yield, but has the highest water yield in all the experiments. Increasing discharge power can improve tar yield, but water yield has the same trend. Acknowledgement. This work was performed with support of the National Natural Science Foundation of China (20576019 and 20776028), and the National Basic Research Program of China (973 Program), the Ministry of Science and Technology, China (2011CB201301).
Reference [1] Cyprès R, Furfari S. Low-temperature hydropyrolysis of coal under pressure of H2-CH4 mixtures. Fuel 1982; 61:721-4. [2] Ibarra JV, Moliner R, Miranda JL, Cyprès R. Pyrolysis under H2 and He of Utrillas coal. Fuel 1988; 67:464-8. [3] Braekman-Danheux C, Cyprès R, Fontana A, Van Hoegaerden M. Coal hydromethanolysis with coke-oven gas: 2. Influence of the coke-oven gas components on pyrolysis yields. Fuel
5
Oviedo ICCS&T 2011. Extended Abstract 1995; 74:17-9. [4] Zhou Q, Hu HQ, Liu QR, Zhu SW, Zhao R. Effect of atmosphere on evolution of sulfurcontaining gases during coal pyrolysis. Energy Fuels 2005; 19:892-7. [5] Qin ZF, Maier WF. Coal Pyrolysis in the Presence of Methane. Energy Fuels 1994;8:1033-8. [6] Sakaguchi M, Laursen M, Nakagawa H, Miura K. Hydrothermal upgrading of Loy Yang Brown coal - Effect of upgrading conditions on the characteristics of the products. Fuel Process Technol 2008; 89:391-6. [7] Hu HQ, Bai JF, Wang Y, Guo SC. Catalytic Liquefaction of Coal with Highly Dispersed Fe2S3 Impregnated in-Situ. Energy Fuels 2001; 15:830-4. [8] Li BQ, Braekman-Danheux C, Cyprès R. Catalytic hydropyrolysis by impregnated sulphided Mo catalyst. Fuel 1991; 70:254-8. [9] Liu QR, Hu HQ, Zhu SW. Integrated process of coal pyrolysis with catalytic partial oxidation of methane. International Conference on Coal Science and Technology. Okinawa, Japan, 2005. [10] Liu JH, Hu HQ, Jin LJ, Wang PF, Zhu SW. Integrated coal pyrolysis with CO2 reforming of methane over Ni/MgO catalyst for improving tar yield. Fuel Process Technol 2010; 91:41923. [11] Liu JH, Hu HQ, Jin LJ, Wang PF. Effects of the Catalyst and Reaction Conditions on the Integrated Process of Coal Pyrolysis with CO2 Reforming of Methane. Energy Fuels 2009; 23:4782-6. [12] Drost H, Rutkowsy J, Mach R, Klotz HD, Schulz G. Plasma-chemical methane conversion under nonthermal and thermal conditions: an attempt toward uniform kinetic modeling. Plasma Chem Plasma Process 1985; 5:283-91. [13] Zhou LM, Xue B, Kogelschatz U, Eliasson B. Partial oxidation of methane to methanol with oxygen or air in a nonequilibrium discharge plasma. Plasma Chem Plasma Process 1998; 18:375-93. [14] Liu CJ, Xue BZ, Eliasson B, He F, Li Y, Xu GH. Methane conversion to higher hydrocarbons in the presence of carbon dioxide using dielectric-barrier discharge plasmas. Plasma Chem Plasma Process 2001; 21:301-10. [15] Rico VJ, Hueso JL, Cotrino J, González-Elipe AR. Evaluation of different dielectric barrier discharge plasma configurations as an alternative technology for green C1 chemistry in the carbon dioxide reforming of methane and the direct decomposition of methanol. J Phys Chem A 2010; 114:4009-16. [16] He XF, Hu HQ, Jin LJ, Zhao YP. Prepr Pap - Am Chem Soc, Div Fuel Chem 2010;55:17-8. [17] Zhang YW, Ding WZ, Guo SQ, Xu KD. Reduction of metal oxide in nonequilibrium hydrogen plasma. Chin J Nonferrous Met 2004; 14:317-21. [18] Kadao S, Urasaki K, Sekine Y, Fujimoto K, Nozaki T, Okazaki K. Reaction mechanism of methane activation using non-equilibrium pulsed discharge at room temperature. Fuel 2003; 82:2291-7. [19] Dai B, Zhang XL, Gong WM, He R. Effects of Hydrogen on the Methane Coupling under Non-equilibrium Plasma. Plasma Sci Technol 2001; 3:637-9. [20] Cui JH, Xu GH, Han S. Eliminating Coke Formed in CH4 Coupling under Plasma via Pure H2 Discharge in the System. Acta Phys-Chim Sin 2002; 18:276-8. [21] Liu CJ, Xu GH, Wang TM. Non-thermal plasma approaches in CO2 utilization. Fuel
6
Oviedo ICCS&T 2011. Extended Abstract Process Technol 1999; 58:119-34. [22] Maya L. Plasma-assisted reduction of carbon dioxide in the gas phase. J Vac Sci Technol A 2000; 18:285-7. [23] Eliasson B, Kogelschatz U, Xue BZ, Zhou LM. Hydrogenation of Carbon Dioxide to Methanol with a Discharge-Activated Catalyst. Ind Eng Chem Res 1998; 37:3350-7. [24] Zhu AM, Gong WM, Zhang XL, Zhou J, Shi H, Zhang BA. Investigation on pulsed corona induced plasma for coupling of methane. Nat Gas Chem Ind 1997; 2:1-5. [25] Dai B, Zhang XL, Zhang L, Gong WM, He R, Lu WQ, Deng XL. Methane coupling under hydrogen plasma. Sci China Ser B 2001; 31:174-7. [26] Wang BW, Zhang X, Liu YW, Xu GH. Conversion of CH4, steam and O2 to syngas and hydrocarbons via dielectric barrier discharge. J Nat Gas Chem 2009; 18:94-7.
7
Effect of Steam Treatment of a Sub-bituminous Coal on Its Caking and Coking Properties Shui Hengfu
Shan Chuanjun Chang Hongtao Wang Zhicai Ren Shibiao and Kang Shigang
Lei Zhiping
School of Chemistry & Chemical Engineering, Anhui Key laboratory of Coal Clean Conversion & Utilization, Anhui University of Technology Ma’anshan 243002, Anhui Province, P.R. China
[email protected] (Shui H.) Abstract A Chinese sub-bituminous coal i.e. Shenfu (SF) coal was steam treated at different temperatures and the caking and coking properties of the treated coals were evaluated by caking indexes (G indexes) and crucible coking determinations. The results show that steam treatment can obviously increase the G index of SF coal and modify the crucible coke quality made from the steam treated SF coal blends. For the steam treated coals used in the coking coal blends instead of the SF raw coal, the micro-strength index (MSI) of the coke and particle coke strength after reaction (PSR) increased, and particle coke reactivity index (PRI) decreased, which are beneficial for metallurgical coke to increase the gas permeability in blast furnace. The removal of oxygen functional groups especially hydroxyl group thus dissociating the aggregated structure of coal during steam treatment maybe responsible for the modifying results. 1. Introduction
The rapid developing of iron-making by blast furnaces promotes the developing of coke-making industry. This is leading to the increased requirement to the coking coals. More than 300 millions tons coke productivity consumes about 400 millions tons coking coals per year in China, resulting in the shortage supply of the coking coals. Therefore opening coking coal resources is becoming one of most interesting issues in coke-making industry. On the other side, the reserve of sub-bituminous coals in China is abundant. Shengfu (SF) coal is one of the sub-bituminous coals and it has low contents of sulfur and ash. Therefore it is effective for lowering the contents of sulfur and ash in coke if SF coal can be used in coal blends. In order to modify the
caking propensity and decrease volatile material content thus increasing the amount of non-coking coals used in coke-making coal blends, many of effective methods have been carried out as pretreatment such as thermal and hydrothermal treatments [1-3]. Mukherjee et al [2] found that hydro-thermal treatment promoted the formation of a coke-like mass for non-coking coal and the relative decreases of total oxygen and hydroxyl oxygen were greater in hydro-thermal treatment than in thermal treatment without water for coal. Iino et al [3] found that water treatments of three Argonne Premium coals at 600 K increased their extraction yields greatly. We have also found that hydro-thermal treatment at proper conditions can increase the extraction yields of bituminous coals [2]. Hydrothermal treatment (liquid water or steam) researches by pioneers are all carried out at higher pressure, normally more than 20 atm, and the hydrothermally treated coals are characterized by solvent extraction. It can speculate that hydrothermal treatment of coal by steam at atmosphere maybe more beneficial for removal volatile maters and easy to be realized in industry, especially for sub-bituminous coals. In this study a Chinese sub-bituminous SF coal was steam treated at atmosphere, and the caking and coking properties of the treated coals were evaluated by crucible coking determinations. The results are positive and the steam treated SF coal can be used in coal blends of coke-making. 2. Experimental 2.1. Coal sample A Chinese sub-bituminous coal i.e. SF coal was used in this study. The properties of SF coal are shown in Table 1. The coal sample was ground and sieved passing through 200 meshes, stored under a nitrogen atmosphere and dried for 12 h under vacuum at 80℃ before use. Table 1 Ultimate and proximate analyses of SF coal
SF Coal *
Proximate analysis(wt%)
Ultimate analysis（wt %, daf）
Sample
*
C
H
N
S
O
Ad
Vdaf
Mad
78.67
5.01
1.21
0.45
14.66
5.2
38.3
10.1
by difference.
G 0
2.2. Steam treatment The steam treatment of coal was performed at a fixed bed reactor at atmospheric pressure. The reactor was constructed of 2.6 cm i.d. 314 stainless-steel pipe and was 26 cm in length. In each run, 70 g of dried coal sample was charged into the reactor then saturated steam in 100 oC was induced flowing through the coal sample in the reactor with a rate of 5ml/min. The reactor was heated by a two half external furnace to the desired temperature, maintained for 1h, then cooled down to room temperature in 1-2 h by opening the two half furnace and stopped steam flowing. The treated coal was taken out from the reactor, and then dried under vacuum at 80℃ for 12 h. 2.3. Caking index measurement The caking index (G index) was used to characterize the caking property of coal. The measurement was carried out according to National Standard of China (GB5447-85), which is based on that of Roga index. Briefly, 1g of coal was mixed with 5g of standard anthracite (Ruqigou，China). The mixture was carbonized in an inert atmosphere at 850°C for 15 min. The coke obtained was subject to the drum tests for twice, which is slightly different from the Roga index testing requiring drum tests for three times. The coal sample preparation, stirring, carbonization and drum test are all the same as those of the Roga index measurement, and the caking index G was calculated as: G= 10 +
30m1 + 70m2 m
Where m is the weight of coal sample (g), m1 and m2 are the weight (g) of the coke fraction (>1mm) after the first and second drum test respectively. 2.4. Coal extraction Coal extraction was carried out at room temperature. A mixed solvent of carbon disulfide/N-methyl-2-pyrrolidinone (CS2/NMP, 1:1 by volume) was used as solvent, as described in details elsewhere [4]. The extraction yield was then determined from the weight of the residue: Extraction yield =
1 − M r / M coal × 100 % (100 − A d ) / 100
Where, Mr is the weight of dried residue (g), Mcoal is the weight of dried coal (g), and Ad is the ash content of coal (db,%). 2.5. Crucible coking determination The carbonization experiments were carried out in an electrically-heated oven using a 300ml crucible. 300g coal blends (SF raw coal or its steam treated coal 8%, gas coal 27 wt%, coking coal 35 wt%, lean coal 10 wt% and rich coal 20 wt%) with a particle size less than 1.25mm were loaded into the crucible. An iron cake of 500g was put on the coal sample to maintain the bulk density of coal feed. The filled crucible with a cover was placed in the oven in an inert atmosphere and heated at the rate of 5~7°C/min to 400 °C, at 3 °C/min to 950°C, and held at 950°C for 30min, then cooled down to room temperature in about 12 h. The coke produced was subject to further evaluations. 2.6. Coke reactivity towards CO2 The reactivity towards CO2 of coke product was measured following a procedure based on the reported method [5]. Briefly, 20 g of cokes (3–6 mm in size) were reacted at 1100°C for 2 h with CO2 at a flow rate of 150 ml/ min. The particle coke reactivity index (PRI) was calculated as the percentage of weight loss after the reaction. Replicate runs were conducted, and the error was within 3%. PRI reported in this study are the average value of the 2 runs. 2.7. Coke mechanical strength Micro-strength of coke was determined according to the Ragan and Marsh method [6]. Briefly, two charges of coke (2 g; particle size between 0.6 and 1.2 mm) were charged into two separate cylinders (25.4 mm i.d and 305 mm long) sealed by steel dust caps, and each contained 12 steel ball-bearings (8 mm diameter). While the cokes from the crucible coking experiments were subjected to 800 rotations at a speed of 25 rpm, the weight percent of coke particles (>0.2 mm in size) was used as the indicator of the micro-strength index (MSI) of the coke. The micro-strength index of the coke after reactivity measurement is defined as particle coke strength after reaction (PSR). The MSI and PSR were reported as the average value of 2 runs. 2.8. FTIR measurement
FTIR were measured on a PE-Spectrum One IR spectrometer at a resolution of 4 cm-1. Samples for the FTIR measurement were prepared by mixing the coal sample with KBr and the mixture was pressed to form a pellet. The difference spectra between raw and treated coal were also obtained using the absorption of C=C bond stretching of aromatic rings at 1600cm-1 as a standard peak.
3. Results and Discussion 3.1. Effects of steam treatment on the caking property and extraction yield of SF coal. Table 2 shows the G indexes and extraction yields of raw and steam treated coals at different temperatures. Because SF coal is a non-caking coal, its G index is 0 as shown in Table 1. In order to differentiate varieties of G index after steam treatment, SF coal or its steam treated one was mixed with an equivalent rich coal (G index is 98) and the mixed coal was used to determine G index instead of SF raw coal or its steam treated ones. Table 2 shows that steam treatment of SF coal can increase its G index and modify its caking property. The maximal G index of 42.4 was obtained at 150℃ steam treatment, which is much higher than that of raw coal 34.6. Further increasing the steam treatment temperature the G index had a decreased tendency. The change of extraction yield after steam treatment is similar with that of G index as shown in Table 2. The extraction yield of SF raw coal is low (4.8%). After steam treatment the extraction yield increased obviously and the maximal extraction yield of 16.0 % was obtained at 200℃ steam treatment. At the range of 150-250℃ the steam treated coals gave similar extraction yields but were much higher than that of raw coal. We have reported [7,8] that the amount and composition of extractible constituents in the CS2/NMP mixed solvents have a great effect on the caking property of coal. With the increase of the extraction yield in the mixed solvents, the caking index of the coal increases. A little decrease in G index after 200℃ steam treatment maybe due to the more content of light constituents in the extractible constituents compared to those of
the coal steam treated at 150℃. It can also be observed from Table 2 that with steam treatment temperature increasing, the volatile yield continuously decreases. Table 2 Effect of steam treatment on the G index, extraction yield and volatile yield of SF coal with different temperatures G
Extraction yield (daf, wt%)
V (daf, wt%)
Raw coal
34.6
4.8
38.3
100
38.4
10.3
36.5
150
42.4
15.2
31.4
200
41.2
16.0
30.2
250
39.2
15.6
28.7
Temperature (℃)
3.2. Characterization of steam treated coal. Table 3 shows the elemental compositions of SF raw coal and its steam treated ones. It can be observed from Table 3 that the changes of N and S after steam treatment are negligible. An obvious decrease in O contents for the steam treated coals compared to that of SF raw coal could be observed. With the increasing of steam treatment temperature the O content of the treated coal kept the decreased tendency. For example, O content decreased from 14.66% of SF raw coal to 13.62% of 150℃ and further to 13.10% of 250℃ steam treated ones. This suggests that steam treatment can promote the removal of the oxygen groups in coal molecules although it has little effect on the removal of S and N heteroatoms. It is very interesting that H/Cs of steam treated coals almost keep the same as that of raw coal 0.76 as shown in Table 3. It is easy to understand that H content in volatile mater in higher than that in raw coal. Steam treatment promotes to release the volatile metters from the treated coal bringing much of light constituents to be rich in hydrogen. Table 3 Elemental analyses of SF raw coal and its steam treated ones (wt %, daf) C
H
N
S
O*
H /C
Raw coal
78.67
5.01
1.21
0.45
14.66
0.76
100
79.36
5.04
1.27
0.42
13.91
0.76
150
79.65
5.08
1.26
0.38
13.62
0.76
Temperature
(℃)
200
79.80
5.06
1.29
0.41
13.44
0.76
250
80.13
5.04
1.29
0.44
13.10
0.75
* By difference. In order to probe the mechanism for the G index and extraction yield enhancements of coal by steam treatment, FTIR measurement was used to discover the structural change of coal by the steam treatment. Figure 4 shows the FTIR spectra of SF raw coal, its steam treated one at 200 ℃ and their difference spectrum. The difference spectrum shows a decrease in intensity of bands near 3400 and 1650 cm-1 for steam treated coals, which are assigned to self-associated OH hydrogen bonds and carbonyl band (C=O) stretching in coal respectively. The decreases in intensity of bands near 3400 and 1650 cm-1 in difference spectra show the decreases in the self-associated OH hydrogen bonds and carbonyl bands of steam treated coals. That is to say one of the mechanisms of steam treatment of coal is to remove OH and carbonyl groups, therefore to decrease self-associated OH hydrogen bonds resulting in G index and extraction yield enhancements of treated coal. The removal of oxygenated a
functional groups including hydroxyl
A
b
oxygen
is
beneficial for the caking property of coal because
c
oxygenated 4000
3500
3000
2500
2000
Wavenumber /cm
1500
1000
functional
500
-1
groups during coal pyrolysis consume much amount of
Fig. 1 FTIR spectra of steam treated coal at 200 (a), SF
raw coal (b)
and the difference b-a (c)
active hydrogen, resulting in the formation of cross-linked
chars.
3.3. Crucible coking determination. Normally, the use of coal blends is a common practice in the coke manufacture
industry. And also SF coal is a non-caking coal, it can not be singly used in coke-making. As such, coal blends were used in crucible coking determinations in this study instead of a single coal. The standard coal blends used in this study were gas coal 35 wt%, coking coal 35 wt%, lean coal 10 wt% and rich coal 20 wt%. In the case of SF coal used in the coal blends, the amount of SF raw coal or its steam treated one was 8 wt% instead of 8 wt% gas coal, i.e. the amount of gas coal used in the coal blends decreased to 27%. The coke micro-strength index (MSI), the particle coke reactivity index (PRI) towards CO2 and particle coke strength after reaction (PSR) of the cokes produced in crucible coking experiments were used to assess the coke properties, as shown in Table 4. Table 4 Crucible coking determination results Coal Blends
MSI,%
PRI, %
PSR,%
Standard coal blends
63.4
41.0
49.5
SF raw coal
54.6
48.6
42.9
150
hydro-thermally treated coal
61.3
42.8
45.9
200
hydro-thermally treated coal
62.0
41.4
50.0
250
hydro-thermally treated coal
61.0
42.9
49.1
It can be observed that comparing with standard coal blends, using 8 wt% SF raw coal instead of equivalent amounts of gas coal caused an obvious decrease in MSI and increase in PRI, resulting in great decrease in PSR. Currently perhaps the two most important parameters, used and worldwide to evaluate the performance of a metallurgical coke in the blast furnace and to define the quality of coke, are the reactivity to CO2 at high temperature (coke reactivity index) and the post-reaction mechanical strength (coke strength after reaction index) [9, 10] . This result suggests that as a sub-bituminous coal, SF raw coal is disadvantage in coal blends compared to gas coal because of its non-caking property and high content in volatile yield. Table 3 shows that steam treated coals give much higher quality of cokes compared to SF raw coal. The PRIs of the cokes from steam treated coals decreased obviously compared with that of SF raw coal, and the corresponding PSRs were also higher than that of SF raw coal. The coke with the largest values of MSI, PSR and lowest value of PRI was obtained from 200
steam treated coal, and it had got to the quality of coke from
standard coal blends. The results strongly demonstrate that steam treatment is a very effective method for sub-bituminous SF coal to modify its caking and coking properties. The crucible coking determination results of steam treatment are consistent with the changes of G and solvent extraction yield shown in Table 2. The steam treated coals with increased values of G and extraction yield gave higher quality of coke when they were used in coal blends. It is well known that [2] oxygen groups, especially for hydroxyl oxygen in coal molecules are responsible for the generation of cross-link and consumption of active hydrogen, thus shortening the plastic temperature range. Steam treatment can promote the removal of OH group (indicated by elemental analyses and FTIR), thus dissociating the aggregated structure of coal and making the treated coal to be with less aggregated structure and much more amount of light fragments. These are beneficial for SF coal to modify its caking and coking properties, resulting in a higher quality coke formed. 4. Conclusions Comparing with thermal treatment, steam treatment of sub-bituminous SF coal can obviously increase its G index and solvent extraction yield and decrease its volatile yield by promoting the thermolysis of SF coal and making much more volatile matters to be released. Steam can swell the coal and dissolve the mobile, thermolytic fragments, especially for oxygen-containing materials, resulting in more light fragments existed in the steam treated coal and improved transfer of internal donatable hydrogen. Steam treatment of SF coal can promote the removal of oxygen groups, especially OH group in coal molecules. This will decrease self-associated OH hydrogen bonds in macromolecular network of coal, thus making the treated coal to be with less aggregated structure, resulting in the G index and extraction yield enhancements of the treated coal. Steam treatment is an effective method to modify the caking and coking properties of SF sub-bituminous coal. Crucible coking tests suggest that steam treatment can greatly increase the MSI, PSR and decrease the PRI of the coke when the steam treated SF coal is used in the coal blends instead of SF raw coal. The quality
of the coke obtained from 8% 200℃ steam treated SF coal in coal blends gets to that of the coke obtained from the standard coal blends, which there is no SF coal addition in the coal blends.
Acknowledgment.
This work was supported by the Natural Scientific Foundation
of China (20876001, 21076001, 20936007), and National Basic Research Program of China (973 Program, 2011CB201302). Authors are also appreciative for the financial support from the Provincial Innovative Group for Processing & Clean Utilization of Coal Resource. References [1] Saikiaa BK, Boruaha RK, Gogoib PK, Baruaha BP. A thermal investigation on coals from Assam (India). Fuel Processing Technology 2009; 90:196–203. [7] Shui HF, Lin CH, Zhang M, Wang ZC, Zheng MD. Comparison of the associative structure of two different types of rich coals and their coking properties. Fuel 2010; 89:1647–1653. [2] Mukherjee DK, Sengupta AN, Choudhury DP, Sanyal PK, Rudra SR. Effect of hydro-thermal treatment on caking propensity of coal. Fuel 1996;75: 477-482. [4] Shui HF, Wang ZC, Wang GQ. Effect of hydro-thermal treatment on the extraction of coal in the CS2/NMP mixed solvent. Fuel 2006; 85:1798-1802. [3] Iino M, Takanohashi T, Li C, Kumagai H. Increase in extraction yields of coals by water treatment. Energy & Fuels 2004; 18: 1414-1418. [5] Wietlik US, Gryglewicz G, Machnikowska H, Machnikowski J, Barriocanal C, Alvarez R, Dı´ez MA. Modification of coking behaviour of coal blends by plasticizing additives. Journal of Analytical and Applied Pyrolysis 1999; 52: 15–31. [6] Ragan S, Marsh H. Carbonization and liquid-crystal (mesophase) development. 22. Micro-strength and optical textures of cokes from coal-pitch cocarbonizations. Fuel 1981; 60:522–528. [8] Shui HF, Zheng MD, Wang ZC, Li XM. Effect of coal soluble constituents on caking property of coal. Fuel 2007; 86 : 1396–1401.
[9] Koszorek A, Krzesińska M, Pusz S, Pilawa B, Kwiecińska B. Relationship between the technical parameters of cokes produced from blends of three Polish coals of different coking ability. International Journal of Coal Geology 2009; 77:363-371. [10] Pis JJ, Menendez JA, Parra JB, Alvarez R. Relation between texture and reactivity in metallurgical cokes obtained from coal using petroleum coke as additive. Fuel Processing Technology 2002; 77–78: 199– 205.
International Conference on Coal Science and Technology (ICCS&T), October 2011 ABSTRACT THE ROLE OF COAL SCIENCE IN DEVELOPMENT AND DEPLOYMENT OF HIGH EFFICIENCY ENERGY TECHNOLOGIES D. J. Harris and D. G. Roberts CSIRO Energy Technology PO Box 883 Kenmore, Queensland, 4069, Australia
[email protected] World coal consumption is projected to grow by approximately 55% between 2007 and 20351. The non-OECD Asian nations (predominantly China and India) are expected to account for 95% of this projected growth, with China to increase its coal-fired electricity generation capacity from approximately 500 GW (2007) to approximately 1250 GW by 2035. While coal use in OECD countries is expected to increase relatively slowly over this period, and coal’s share of electricity generation in these nations is expected to fall, increased generation from coal-fired power plants is still significant: in the US it is expected to account for 26 percent of the growth in total electricity generation from 2007 to 2035. Another important factor in some OECD countries (particularly the US and Australia) is the need for new plant to replace aging installations which have relatively low efficiencies and are expected to be unattractive for retrofit of CO2 capture technologies to meet likely future greenhouse-gas emissions requirements. The research challenges needed to advance the development of low emissions coal-based power generation technologies are clearly associated with increasing efficiencies and reducing greenhouse gas emissions at large scale and low cost. In the face of strongly increasing world coal use, new technologies will be needed in the future to increase the efficiency of coal fired power generation significantly above the levels of current best practice and to facilitate the capture of CO2 for long term storage. Low Emissions Coal: Efficiency is King The average efficiency of coal-fired plants globally is currently only about 28% (HHV) with the most efficient ultra-supercritical steam plants and new IGCC demonstration technologies achieving about 45%2. The current large worldwide growth in new capacity provides an important opportunity in both developing and developed nations for development and deployment of advanced, high efficiency power generation technologies which can provide a suitable technology platform for further efficiency and cost improvements to meet the requirements for increasing levels of CO2 emissions abatement. Repowering existing coal-fired plants, where possible, to improve their efficiency, coupled with installation of new and more efficient plant, will provide significant reductions in CO2 emissions. However, to achieve very high levels of CO2 emissions reduction from fossil fuel based power generation technologies, it will be increasingly important to have in place conversion technologies with the highest possible efficiencies which are capable of reducing the amount of CO2 that must eventually be captured and stored. Due to the significant energy demands and costs associated with CO2 capture and storage (CCS), deploying the most efficient plant possible is a critical prerequisite to enable these plants to be capable of being fitted with CO2 capture technologies, either from new or as a staged retrofit in the future. The presentation will discuss research needs and technology development pathways in the areas of high efficiency, low emissions coal technologies. Particular emphasis will be on coal gasification and the important downstream syngas conversion and gas separation technologies necessary to facilitate CO2 capture and hydrogen production at a scale and cost acceptable to the power industry. Based on current Australian RD&D programs, the presentation will also include an overview of current pilot and demonstration-scale research on post combustion capture (PCC) of CO2 from conventional pulverised coal power technologies. 1
US Energy Information Administration, International Energy Outlook 2010 World Coal Association, 2011: http://www.worldcoal.org/coal-the-environment/coal-use-theenvironment/improving-efficiencies/ 2
1
Ultimately, even higher efficiency systems will be required. This presentation will also introduce examples of some novel future technologies, such as direct injection coal engines using coal water fuels and direct carbon fuel cells, which have the potential to achieve step increases in power generation efficiency with consequent reductions in the amount of CO2 that would be emitted or collected for storage. The emphasis here will be on the impact of coal properties and behaviour on the performance of these technologies and the role of coal science and associated R&D in facilitating their rapid development and uptake to meet environmental, cost and performance goals during the next 5-20 years of rapid energy demand growth. Post Combustion Capture: Starting From Where We Are While next generation, high efficiency technologies are clearly needed to meet longer term needs, post combustion capture (PCC) of CO2 from conventional pulverised coal power technologies is an important transition technology and focussed research and demonstration programs are required to support early adoption of these technologies on the most appropriate existing and new pf plant. While solvent based technologies for CO2 capture are well established in the chemical and process industries, the key challenges associated with PCC technology are associated with reducing the capital cost, energy efficiency penalties and potential environmental impacts of these large scale solvent based systems. For current systems, the efficiency penalties associated with PCC on conventional plant can be up to 10 percentage points and this major loss of efficiency (and capacity) represents the most challenging aspect for retrofit and new build applications of this technology. On the most modern pf plants, which already have high levels of flue gas treatment, coal property impacts on PCC systems are relatively minor and much of the required R&D is focussed on developing improved solvents and reducing the energy requirements of CO2 recovery. However, many existing coal fired plant, with little or no flue gas treatment, face additional constraints as coal specific contaminants can interact deleteriously with the most common solvents and further work is required to develop alternative materials and processes. IGCC: a High-Efficiency Platform for CCS IGCC technology presently achieves similar efficiency to latest PC technology (~40–45% HHV basis) but at a slightly higher capital cost. However, substantial improvements in IGCC efficiency (~ + 8 percentage points) along with significant reductions in capital cost are projected through new and improved process blocks now under development internationally. As for all power technologies, the introduction of CCS decreases the overall efficiency and increases costs of power generation. For IGCC systems, efficiency losses with currently-available (precombustion) CO2 capture technologies are expected to be approximately 6–8 percentage points and capital costs are likely to increase by up to 40%. Carbon capture technologies for IGCC applications are still early in their learning curve; therefore, as with the main IGCC plant, significant improvements to the process components can be envisaged and development of these will substantially reduce the cost and efficiency penalties associated with CCS. As IGCC and IGCC-CCS technologies begin to be implemented, initially at the commercial demonstration scale, a critical factor in the success of these projects, and in the subsequent wider deployment, will be stakeholder confidence. It is important that governments, technology developers, vendors, operators and the community can see that such technologies can operate reliably at the required scale with high availability, safety and environmental performance while meeting the necessary CO2 emissions requirements. Ongoing research is therefore required to continue to support development and deployment of large-scale IGCC and IGCC-CCS systems. Many of the major technical issues associated with the success of these initial projects, and which will underpin continued improvement of the efficiency and performance of IGCC with CCS, will rely on a detailed understanding of the behaviour of coals in the gasification process and of the resultant impacts on downstream unit operations associated with syngas cleaning, processing, separation and CO2 storage or utilisation. For example: •
Knowledge of coal gasification reactivity and conversion behaviour under conditions relevant to the specific technology and operating environment is a critical factor in efficient gasifier design and operation to ensure complete and efficient coal conversion to maximise efficiency and reduce carbon in slag to acceptable levels;
2
•
Appropriate coal characterisation, selection and preparation are key factors which define the ability of the mineral matter component of coals to form suitable slags that do not compromise gasifier operation and ensure high availability. Of particular importance in this regard is the impact of poor slagging behaviour of coals on the entire system operation. Most notably the need to operate the gasifier at higher temperatures or with excessive fluxing to successfully manage the slag in the gasifier. This has direct implications on gasifier efficiency and on oxygen demand (and costs) and can directly limit plant capacity. Excessively high operating temperatures also affects plant life and maintenance requirements – both within the gasifier and in downstream gas cooling and cleaning systems
•
In any system where coal-derived syngas is used as the basis for power generation, chemicals production, or in the manufacture of liquid fuels, the syngas must be cleaned to standards acceptable by the downstream plant. In coal gasification derived systems the contaminants include fine particles of fly ash, gaseous species containing sulphur, chlorine, fluorine, alkali metals and trace elements. Coal properties and gasification behaviour under the relevant process conditions profoundly affect syngas composition which specify development criteria for improved and breakthrough technologies to reduce the costs and energy penalties associated with syngas cleaning and processing, gas separation and CO2 capture systems
Coal Impacts on Gasification Performance Conditions inside an entrained flow gasifier are extreme: pressures are high (20–40 bar or sometimes greater) and temperatures are high (flame temperatures often over 1800 K). The ratio of oxygen to fuel is significantly lower than those used in coal combustion technologies, and the mineral matter in the coal is required to melt and flow out of the gasifier continuously. Steam is sometimes included in the feed streams to the gasifier, and some gasifiers are designed to feed coal as a coal-water slurry. These aspects of gasification mean that the extensive literature and understanding of coal performance in pf boilers has little direct application to understanding and prediction of coal performance under gasification conditions. Results of ‘standard’ combustion tests do not translate to gasification performance—new approaches, facilities, techniques, and knowledge are required. A striking point to emerge from analysis of coal performance in the complex environment in these high pressure, high temperature reaction systems is the potential impact of relatively fundamental coal properties on many of the process operations comprising the IGCC system. Even seemingly simple factors such as inherent moisture, mineral matter composition, high temperature volatile yield, char reactivity and structure, grinding behaviour, slurrying characteristics, sulphur content etc can become particularly important as they may create issues that cannot be accommodated through simple changes to operating conditions. Such issues therefore become limiting factors for the fixed plant design (eg size of ASU, gasifier, syngas cooler etc). Managing these and other coal related issues can incur significant costs and/or operating boundaries that can seriously affect plant capacity, efficiency and performance. To allow practical and reliable application of a sound fundamental understanding of gasification science to the solving of real industrial problems, knowledge of coal pyrolysis, char formation, char reactivity, slag formation and flow, and coal gasification behaviour needs to be integrated in a form that is applicable to a range of gasification technologies and, eventually, gasification-based energy systems. Fundamental, experimental gasification research needs to be undertaken in parallel with the development of detailed coal reaction and conversion models designed to allow more widespread application of the outcomes through, for example, relevant gasifer and integrated process models of the entire coal conversion, slag handling, syngas processing and gas separation systems. This can only be done effectively through close collaboration of researchers, industrial technology developers, vendors and operators. From Refineries to Power Generation: Advanced Syngas Processing for High Efficiency CCS In the chemicals and refinery industries, coal gasification, and capture of CO2 from syngas is commercially mature. A research strategy to support the rapid development and application of advanced syngas processing and gas separation technologies in the power sector will require a combination of fundamental materials development and testing programs, laboratory scale experiments, modelling projects, larger ‘research gasifier’ scale measurements and screening tests. This work would be complemented with appropriately targeted pilot plant and slipstream tests utilising syngas slipstreams such as those available from a number of international IGCC commercial, 3
demonstration and research projects. An example of a commercial facility that currently operates in this way is the Puertollano demonstration IGCC project in Spain which has a slipstream of up to approximately 2% of the syngas available for advanced technology development projects such as gas cleaning, shift and gas separation concept development, materials testing etc. Improved technology components and concepts fitting within the IGCC process flow sheet that have been identified to date, and in some cases tested using simulated syngas environments at laboratory scale, include: •
new water gas shift catalysts optimised for coal syngas and suitable for use with membrane reactor systems at higher temperatures (up to 600°C);
•
integrated high temperature (~600°C) dry syngas cleaning systems;
•
trace element capture integrated with high temperature syngas cleaning;
•
metal and ceramic membrane based Hydrogen/CO2 separation technologies;
•
integrated water gas-shift/metal membrane catalytic reactor concepts capable of enhancing hydrogen production and separation at high temperatures;
•
ion transport membrane air separation technologies have been in development for almost two decades and are now are nearing commercial availability at tonnage scales.
Further opportunities to significantly increase efficiencies are expected as syngas and hydrogen based fuel cells reach commercial availability. While these are unlikely to be available at the scale and reliability required for large scale power generation within the next 15-20 years they provide an attractive development pathway for the core, high efficiency technology platforms that are being developed and demonstrated today.
4
Gas adsorption capacity of coaly shales from Japan and USA - Implications for CO2 storage in coal-bearing formation S. SHIMADA1, Y. NISHIIRI1, N. SAKIMOTO1, K. OHGA2 and Y-S. JUN3
1
Graduate School of Frontier Sciences, 5-1-5, Kashiwanoha, Kashiwa, Chiba 277-8563 Japan,
[email protected]
2
Graduate School of Engineering, Hokkaido University, Kita 13, Nishi 3, Kitaku, Sapporo 060-8628, Japan 3
Department of Energy, Environmental & Chemical Engineering, Washington University in St. Louis
One Brookings Drive, Campus Box 1180, St. Louis, MO 63130-4899, USA Abstract The adsorption capacity of CO2 and CH4 with coaly shales was measured by the volumetric method at 35°C and 50°C, with a pressure range up to 9 MPa. The coaly shale samples were obtained from Yubari (Japan), Kusiro (Japan), Bibai (Japan), and Illinois(USA) and a gas shale sample came from Pennsylvania (USA). Adsorption capacity of the shale samples varied widely. In particular, Yubari coaly shale exhibited the largest CO2 adsorption capacity of 14 cc/g at 35°C and 5 MPa. This value is equivalent to the adsorption amount of general coal with medium CO2 adsorption capacity. The above result suggests that CO2 adsorption in shale is not negligible during CO2 storage in coal bearing formation. This study provides information about the potential challenges on low injectivity (low permeability) of coaly shale CO2 storage operation and can aid developing effective and safer CO2 storage in ECBMR. 1. Introduction Geological CO2 storage is a promising method to reduce CO2 emissions to the atmosphere in abating global warming. ECBMR (Enhanced Coalbed Methane Recovery) is a unique technology to enhance methane recovery from coal seams while storing CO2. However, the coal seam as a target formation of geological CO2 storage has a disadvantage that the coal seams are thinner compared to aquifers, while thicker formation is preferred to ensure larger CO2 storage capacity. The coal-bearing formation consists of coal seams and aquifer components (shale, sandstone and brine etc.). Therefore, from the view point of CO2 storage, the coal-bearing formation can be a very attractive target formation. The storable CO2 amount in coal seams is calculated by the sum of adsorbed gas in the 1
coal matrix and free gas in cleats. For the aquifer component, the storage amount has been calculated by the sum of free CO2 gas in pore space and CO2 dissolution in aquifer brine. The adsorption mechanism is usually not considered in aquifer components. Shales and coaly shales composing the coal-bearing formation have very fine pores and might exhibit the adsorption phenomena with CO2. Therefore, when coal-bearing formation is considered for potential CO2 storage, the CO2 adsorption contribution in formation may not be negligible in the calculation of storable CO2. Moreover, shale has sealing capacity in geological CO2 storage. As a storage site, either for depleted oil and gas field or an aquifer, one of the most important tasks for operation is to verify the sealing integrity of CO2 by cap rock of the reservoir. In Sleipner monitoring project, when CO2 is injected into the aquifer formation, the buoyancy increases and the majority of CO2 has accumulated at the bottom of the cap rock. In this case, the adsorption phenomenon is believed to occur between CO2 and the shale that located in the cap rock. Nevertheless, this kind of trapping mechanism for adsorption has not been seen from any reservoir simulator so far. However, if the CO2 adsorption can be fixed stably, it is not just contribute to the increment of capacity of storage potential, but also can be anticipated for decreasing the CO2 leakage rate as well [1, 2]. In addition, recent increasing shale gas development in the United States also promotes the importance of research regarding CH4 adsorption on shale, which previously not really given much attention. Although it is known that shale gas exists as free gas or adsorbed gas inside shale, there are not many research related to experiment of adsorbed gas, and also there is still many unclear things related to adsorption mechanism using possible adsorption model equation. The significance of research regarding special characteristic of CO2 and CH4 adsorption is also based on the idea of that similar advanced CH4 capture method for CO2 injection into coal seam might be applicable for shale gas production as well. Therefore, in this study, an experiment of CO2 and CH4 adsorption in shale has been carried out, verifying the possible adoption of the adsorption model equation for four coaly shale and a gas shale samples. 2. Experimental Method 2.1 Outline A series of experiments to measure the amount of CO2 and CH4 adsorption by using volumetric method have been carried out in this study. The method is to calculate the molar amount of gas from gas pressure injected to the cell and determines the adsorbed amount from the difference of the molar amount in gas before and after adsorption. The measuring device of volumetric method is shown in Figure 1. In order to verify the supercritical adsorption properties, experimental condition was set to be 35°C or 50°C, and up to about 9 MPa. 2
Figure 1 Schematic diagram of experimental apparatus. (a) Reference cell, (b) Sample cell, (c) Water bath for temperature control, (d) Pressure transducer, (e) Gas cylinder, (f) Vacuum pump, (v-1) and (v-2) valves By using the adsorbed amount obtained from the experiment, the amount of excess adsorption is calculated by the nth of differential excess adsorption ∆n (nexc) of the following Eq. 1,
∆n
(n) exc
Pi ( n )VR Pf( n −1)VV Pf( n ) (VR + VV ) − = + RTZ ( n ) RTZ ( n −1) RTZ (f n ) i f
(1)
where i and f are the initial and final conditions of adsorption equilibrium, respectively, at each step, VR is the reference cell volume, VV is the void volume of sample cell, Z is the compressibility factor. Regarding the compressibility factor of Z in Eq. 1, the value for CO2 was obtained from Span and Wagner equation [3] and CH4 from Setzmann and Wagner equation [4]. The nth excessive adsorbed amount was calculated by using Eq. 2 below n
(n) exc
n
(k ) = ∑ ∆nexc k =1
(2)
In this calculation, the initial value of the void volume of sample cell is calculated as constant since it is deviated due to the larger volume of adsorbed surface in the sample, which increases the amount of adsorption when pressure is high. Therefore, the value of the absolute amount of adsorption, which was taken from the volume of adsorption layer, is converted to Eq. 3. In this study, the adsorbed phase density is calculated as equal to the liquid density at normal boiling temperature for CH4 and liquid density of triple point for CO2. The value for CO2 and CH4 is 1.178[g/ml] and 0.422[g/ml], respectively.
ρg nexc = nabs 1 − ρ ad
(3)
where nabs is the absolute adsorbed amount, ρg is the gas density, and ρad is the adsorbed phase density. 3
2.2 Shale Samples The following five shale samples were used in this study: PA-GS (gas shale taken from shale gas development area in Pennsylvania, USA) Illinois (coaly shale collected from the coal-bearing formation of a coal mine in Illinois, USA) Yubari (carbonaceous shale of Yubari area, Japan) Bibai (coaly shale from Bibai coal mine, Japan) Kushiro (coaly shale from Kushiro coal mine, taken from under sea formation) In order to reduce the time to reach the equilibrium, the shale sample was crushed to 75~150 μm size. For the pre-treatment before the experiment, the sample is vacuumed for 24 hours at 60°C. 2.3 Adsorption Model Equation In this study, the measured values are fitted with four adsorption formulas shown below. [5, 6, 7, 8] Langmuir Equation (Surface adsorption model)
n=
n0 KP 1 + KP
(4)
Freundlich Equation (Surface adsorption model)
n = KP1/ t
(5)
Modified Dubinin-Radushkevich (DR) Equation (Pore-filling model)
ρ n = n0 exp − D ln ad ρg
2
(6)
Modified Dubinin-Astakhov (DA) Equation (Pore-filling model)
ρ n = n0 exp − D ln ad ρg
m
(7)
where n is the absolute adsorbed amount, n0 is the maximum adsorbed amount, and K, t, D are constants. The Langmuir and Freundlich equations are surface adsorption models, and modified DR and DA equations are pore-filling models. In the modified DA equation, the values of m are coefficients related to the pore size of the sample. Also m=2 in the modified DR equation, is mainly considered to be suited for a microporous (< 2nm) sample, and for 1 < m < 2 is for the mesoporous (2~50nm) sample in case of CO2 and CH4 adsorption. 4
3. Results 3.1 CO2 and CH4 Adsorption Measurement Figure 2 (a)-(e) shows measurement results on absolute adsorption amount for each sample. The ratio of absolute adsorption for CO2:CH4 is about 2-3 for all samples. This number is similar to the adsorption of bituminous coals. Especially for PA-GS, which was collected from shale gas (CH4) exploration fields, it might be suggested that CH4 enhanced recovery by CO2 injection is possible and similar to the enhanced coalbed methane recovery. The difference of adsorption amount at 35°C and 50°C is smaller for CH4 than CO2.
(a) PA-GS
(b) Illinois
(c) Yubari
(d) Bibai
(e) Kushiro Figure 2 Absolute adsorption amount (mmol/g) on various shale samples. 5
Figures 3 and 4 describe the absolute adsorption amounts of CO2 and CH4 at 35°C, respectively. The reversible adsorption reaction was observed (not mentioned in details here), so adsorption on shale is considered to be a physical adsorption. The adsorption behavior tended to increase linearly with increasing pressure. In addition the rapid increase of adsorption amount like coal sample in low-pressure as often described by Langmuir adsorption model was rarely seen. The increment of adsorption amount has been attributed to the micropore filling. Thus, it was suggested that the development of microporous structure similar to coal sample cannot be seen in shale sample. The result also showed that for both CO2 and CH4, Yubari exhibited the highest adsorption amount and Illinois exhibited the lowest.
Figure 3 Absolute adsorption amount
Figure 4 Absolute adsorption amount
(CO2, 35°C)
(CH4, 35°C)
3.2 Verification of Adsorption Model Equations The following Eq. 8 is used as an indicator to verify the actual value and the degree of expression for four adsorption models described in section 2.3. Figures 5 and 6 show the error comparison of CO2 and CH4 adsorption measurement results for each sample and temperature.
Error =
1 N
nexp ,i − ncal ,i ∑ nexp ,i i =1 N
2
(8)
where N is the number of data, nexp is the measured value, and ncal is the calculated value. According to the results of Figures 5 and 6, the Langmuir equation showed huge errors for CO2, while the Freundlich equation expresses the adsorption isotherms well. On the other hand, it has been observed that the Langmuir equation could fit better in CH4 than CO2. The Langmuir equation can be used for assuming that there is no interaction between adsorbed molecules, in fact, there is an intermolecular force such as dispersion force especially in CO2. Thus, it was assumed that the error in curve fitting of CO2 was higher 6
than CH4 due to this force. Furthermore, both modified DR and DA equations, which are pore-filling models, have described adsorption behavior of each sample of CO2 and CH4 accurately as opposed to than surface adsorption models. The m value that is related to the pore size in the DA equation has also been calculated. Each value falls within the range of 1.5 < m < 2. From here, it has been inferred that the microporous structure in the shale has not developed too much, but rather as a relatively small pore; that is to say, a mesoporous structure has developed.
Figure 5 Error comparison (CO2)
Figure 6 Error comparison (CH4)
4. Discussion A calculation was performed to estimate how much CO2 can be stored in the shale by adsorption. In this estimation, Illinois sample, which has the least adsorption amount, was used. In this calculation, the assumed conditions of shale formation (1m3) have been filled with 100% CO2. The amount of storage CO2, G1 is calculated using the traditional model for fluid stored in the pore spaces of shale without respect to adsorption. G2 is calculated as the adsorption amount from this experimental result. In addition comparison examination has been performed between G1 and G2. In various calculations, the following values were used; CO2 density ρCO2 = 286.914 [kg/m3], adsorbed amount nCO2 = 4.629 [kg/ton], shale porosity measured by mercury porosimeter φ shalee= 3.608%, pressure 9MPa and temperature 50°C. As the result, G1 = Vshale × φshalee × ρCO2 = 10.352 [kg/m3] and, G2 = Vshale × (1- φshale) × ρshale × nCO2 = 12.146 [kg/m3] were obtained. As a consequence, the amount of storage in coaly shale by adsorption is more than that stored in pores. Thus, it is worthwhile to consider the adsorption in shale as a storage mechanism, which has not been considered until now. 5. Conclusion 7
Through the adsorption amount measurement for five types of shales, the physical adsorption of CO2 and CH4 in shale has been accurately described by the pore-filling model equations than surface adsorption model. The absolute adsorption amount in coaly shale is relatively large enough to be included in the storage amount in the coal bearing formation. Acknowledgement This study is partly supported by the CCCU (Consortium for Clean Coal Utilization) of Washington University St. Louis, USA. The authors would like to express deeply our thanks for the research fund. References [1] Busch, A., et al.; “Effects of physical sorption and chemical reactions of CO2 in shaly caprocks”, Energy Procedia 1, 2009, 3229-3235. [2] Busch, A. et al.; “Carbon dioxide storage potential of shales”, INITERNATIONAL JOUNAL OF GREENHOUSE GAS CONTROL 2, 2008, 297-308 [3] Span, R., Wagner, W.; “A New Equation of State for Carbon Dioxide Covering the Fluid Region from the Triple-Point Temperature to 1100K at Pressures up to 800MPa”, J. Phys. Chem. Ref., 1996, Vol. 25, No. 6. [4] Setzmann, U., Wagner, W.; “A New Equation of State and Tables of Thermodynamic Properties for Methane Covering the Range from the Melting Line to 625 K at Pressures up to 1000 MPa”, J. Phys. Chem. Ref. Data, 1991, Vol. 20, No. 6. [5] Cipolla, C.L., et al.; “Modeling Well Performance in Shale-Gas Reservoirs”, 2009, SPE125532 [6] Nuttall, B.C., et al., “Analysis of Devonian Black Shales in Kentucky for Potential Carbon Dioxide Sequestration and Enhanced Natural Gas Production, Final Report”, 2005 [7] Fujii, T. et al.; “Evaluation of CO2 sorption capacity of rocks using gravimetric method for CO2 geological sequestration”, Energy Procedia 1, 2009, 3723-3730. [8] Sakurovs, R., Day, S., Weir, S., and Duffy, G.; “Application of a Modified Dubinin-Radushkevich Equation to Adsorption of Gases by Coals under Supercritical Conditions”, Energy & Fuels, 2007, 21, 992-997
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Relationships between the sorption capacity of methane, carbon dioxide, nitrogen and ethane on bituminous coals. R. Sakurovs, S. Day and S. Weir CSIRO Energy Technology PO Box 330 Newcastle 2300 Australia
[email protected]
Abstract Sequestration of carbon dioxide in coal seams can reduce atmospheric emissions of carbon dioxide. If such sequestration simultaneously results in enhanced coal bed methane (ECBM) production, some of the sequestration costs can be recovered by the value of the methane produced. This requires knowledge of both the carbon dioxide and methane sorption behaviour of coal at high pressures. In order to elucidate the connection between them, we compared the sorption of carbon dioxide, methane, ethane and nitrogen at 55 °C at pressures up to 20 MPa for a number of coals. Sorption isotherms were fitted by a modified DubininRadushkevich model. The relationship between methane and nitrogen maximum sorption capacity was particularly close: on a volume basis, the maximum sorption capacity of all coals examined for methane was twice that of nitrogen. We confirm that the ratio of maximum sorption capacity of carbon dioxide and methane decreased linearly with increasing carbon content. However, this does not necessarily indicate that low rank coals had a specific interaction with carbon dioxide not observed by methane, since similar trends were observed with the ethane/methane sorption ratio.
1. Introduction Unmineable coal seams are one option for sequestration of carbon dioxide, because they can store 6-12 % by weight of carbon dioxide [1]. Often coal seams contain methane. If carbon dioxide can be sequestered in such coal seams and the process simultaneously results in enhanced coal bed methane (ECBM) production, some of the sequestration costs can be recovered in the value of the methane produced [2].
It has long been known that coals can sorb more carbon dioxide than methane [3] although published values for the ratio of molar sorption capacity for coals vary from 2:1 to 10:1 [4]. This variation is partly because the ratios were measured at pressures below saturation and
1
carbon dioxide is adsorbed more strongly than methane, which would increase the ratio especially at very low pressures. However, of more fundamental interest is the maximum adsorption capacity of the two gases, which has not been so intensively investigated.
From basic monolayer sorption models it would be expected that, on a volume basis, the maximum adsorption capacities of gases should be roughly the same since the surface area and pore volume of the coal should be constant. Simple pore-filling models would also reach a similar conclusion. While it is true that larger molecules may not penetrate some pores that are accessible to smaller molecules [5] and that coals swell to some extent when exposed to gases that are strongly sorbed [6], which could change surface area and micropore volume, the difference in the maximum sorption capacity between methane and carbon dioxide is too great to be explained by either hypothesis. Others have suggested that there could be a specific interaction between carbon dioxide and coal that does not exist between methane and coal [2].
Sakurovs [7] found an approximately proportional relationship between the maximum sorption capacity of a coal for gases and their critical temperature if the gases accessed the coal structure equally well. This was used to explain why the maximum (volumetric) adsorption capacity for carbon dioxide is greater than that of methane by about a factor of two: the critical temperature of carbon dioxide is greater than that of methane by about a factor of two. We extend this study here to a range of bituminous coals to determine if the relationships found previously are general.
2. Experimental The excess sorption measurements of twenty-three bituminous and subbituminous coals using methane, carbon dioxide and nitrogen were determined using a gravimetric system [8]. Ethane measurements were performed on ten coals. The coals (analytical data in Table 1) were crushed with minimum fines to less than 1 mm, and the 0.5 mm - 1mm size fraction was used for all characterisation. Prepared samples were dried overnight under vacuum at 60 °C.
Sorption measurements were made using a nominal sample mass of 200 g at pressures up to 20 MPa (accurate to 0.01 MPa) and at a temperature of 55(±1) °C. Samples were maintained at each pressure step for sufficient time to allow equilibrium to establish. Gas densities at each pressure were measured using a reference cell in the isotherm apparatus. In this study, 2
the coal samples were exposed to the various gases one after the other, with carbon dioxide being the first in the series in all cases.
The densities of the coals were measured using a Quantachrome Ultrapycnometer 1000 helium pycnometer. Corrections to cell volume as a consequence of coal swelling were not applied; we assumed that the coal volume inaccessible to the gas remains constant on swelling [9, 10].
Excess sorption (Wads) was fitted to the modified Dubinin-Radushkevich equation [11]
Wads = W0 (1 − ρG / ρ L )e
−[ln ( ρ L / ρG ) RT / E ]2
+ kρG
Equation 1
where W0 is the maximum sorption capacity of the coal, ρG is the density of the gas at the temperature and pressure, ρL is the condensed gas density (assumed to be the van der Waals density of the gas [7]), R is the gas constant, T the temperature, E the apparent heat of sorption and k is the volume accessibility of the coal by the gas compared to that accessed by helium [10]. The term (1 – ρG/ρL) is the correction for the volume of gas displaced by the concentrated surface phase.
3
Table 1 analytical data for coals CH4
CO2 coal
1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23
ash density C H VM Rv,max db daf daf daf % g/cc % % % % 17.6 1.519 83.86 4.55 28.4 0.81 18.7 1.505 83.52 4.77 31.1 0.80 9.7 1.449 72.76 4.83 49.7 0.25 7.7 1.422 82.99 4.66 31.7 0.69 20.3 1.552 80.70 3.90 31.2 0.62 8.3 1.387 85.60 4.90 30.0 0.90 9.1 1.368 89.22 5.37 27.7 1.21 5.6 1.313 84.11 5.73 36.1 0.95 15.48 1.446 77.67 5.00 47.3 0.46 7.4 1.331 83.59 5.40 37.3 0.89 5.3 1.324 84.37 5.55 38.5 0.90 16.9 1.481 88.93 4.55 21.7 1.40 11.4 1.365 83.30 5.25 35.6 0.80 25.5 1.471 82.68 5.11 34.9 0.80 4.9 1.319 84.12 5.29 35.9 0.80 5.8 1.329 83.76 5.22 35.6 0.80 8.8 1.367 88.82 5.01 23.4 1.43 4.8 1.363 91.07 4.44 19.5 1.68 24.8 1.542 81.78 5.32 38.8 0.81 3.6 1.353 67.95 4.72 47.3 2.2 1.354 74.30 5.00 44.3 3.5 1.395 71.40 4.60 41.4 3.7 1.340 72.60 5.40 51.4
Vit mmf % 10.6 20.2
Lipt mmf % 5.0 4.9
29.7 23.9 33.9 85.4 82.7 85.0 88.7 89.4 28.1 64.0 40.3 64.3 59.5 76.7 89.0 28.7
3.9 1.6 2.3 0.3 4.1 5.0 2.2 3.5 0.0 1.6 1.3 3.0 1.3 0.0 1.0 6.7
W0 db kg/t 81.0 77.7 89.0 110.1 103.3 74.1 70.5 67.0 96.3 68.2 58.5 53.8 73.6 65.0 79.9 78.5 63.7 67.8 45.4 96.8 144.4 144.1 126.5
W0 rms vol db % kg/t 12.0 0.5 11.4 0.5 12.5 1.1 15.2 0.9 15.6 0.7 10.0 0.6 9.4 0.5 8.6 0.2 13.5 1.0 8.8 0.3 7.5 0.7 7.8 0.8 9.8 0.8 9.3 0.3 10.2 0.9 10.2 0.6 8.5 0.4 9.0 0.6 6.8 0.5 12.7 0.8 19.0 1.7 19.6 1.8 16.5 2.0
W0 vol % 7.1 7.1 4.8 7.1 8.6 6.6 6.3 5.3 7.4 5.4 5.2 5.4 6.0 5.6 6.9 5.7 5.7 6.2 4.0 6.3 8.7 9.1 6.6
N2 Ethane W0 vol % 4.0 3.7 2.7 4.7 3.4 3.3 2.6 3.4 2.7 2.8 3.3 2.5 3.4 2.8 3.3 3.4 2.0 4.2 4.7 3.5
W0 vol % 11.0
13.6 9.7 9.4 8.4 8.5
8.7
17.8 16.7 12.7
4
3. Results and Discussion
Figure 1. Relationship between volume sorption capacities of coals for methane versus volume sorption capacity for carbon dioxide
Figure 1 shows that the sorption capacity of coals for methane increases with the sorption capacity for carbon dioxide, though the relationship is not a proportional one.
Figure 2 Relationship between volume sorption capacities of coals for nitrogen versus volume sorption capacity for methane
5
Figure 2 shows that the sorption capacity for nitrogen is proportional to that of methane; all coals examined can sorb about twice as much methane as nitrogen (by volume).
a)
b) Figure 3. Relationship between volume sorption ratio and rank for a)CO2/CH4 and b)Ethane/CH4.
6
Figure 3 shows that the ratio of maximum sorption capacity between carbon dioxide and methane decreases, with increasing carbon content, from 2.5 to 1.5 over the range investigated here. This shows that proportionally more carbon dioxide, compared to methane, can be sorbed into lower rank coals than higher rank coals. A similar trend is evident for the ethane/methane sorption ratio. These results suggest that the relationship between sorption capacity for a gas and critical temperature observed previously[7] is only approximate, and other factors specific to coals can affect this result. It also shows that the effect is not specific to carbon dioxide, but occurs with ethane as well. This means that the reason the ratio of maximum sorption between carbon dioxide and methane decreases with increasing rank is not due to a specific interaction of the coal with carbon dioxide.
4. Conclusions Investigation of the sorption of a range of bituminous coals with different gases under supercritical conditions has shown the following: 1. There is a good correlation between the maximum sorption capacities of different gases on coals. The maximum sorption capacity of coal for nitrogen is proportional to its sorption capacity for methane. 2. The ratio of sorption of carbon dioxide to methane increases with decreasing rank and increasing sorption capacity. This variation is not due to a specific interaction of the coal with carbon dioxide since ethane, which on a volume basis is sorbed by coal about as well as carbon dioxide, has a similar effect. References [1] S. Day, G. Duffy, R. Sakurovs, S. Weir, Effect of coal properties on CO2 sorption capacity under supercritical conditions, International Journal of Greenhouse Gas Control, 2 (2008) 342-352. [2] C.M. White, D.H. Smith, K.L. Jones, A.L. Goodman, S.A. Jikich, R.B. LaCount, S.B. DuBose, E. Ozdemir, B.I. Morsi, K.T. Schroeder, Sequestration of carbon dioxide in coal with enhanced coalbed methane recovery - A review, Energy & Fuels, 19 (2005) 659-724. [3] I. Ettinger, Chaplins.A, E. Lamba, V. Adamov, Natural factors influencing coal sorption properties .3. Comparative sorptionof carbon dioxide and methane on coals, Fuel, 45 (1966) 351-&. [4] S. Harpalani, B.K. Prusty, P. Dutta, Methane/CO2 Sorption Modeling for Coalbed Methane Production and CO2 Sequestration, Energy & Fuels, 20 (2006) 1591-1599. [5] O.P. Mahajan, CO2 surface area of coals - the 25-year paradox, Carbon, 29 (1991) 735742.
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[6] S. Day, R. Fry, R. Sakurovs, Swelling of Australian coals in supercritical CO2, International Journal of Coal Geology, 74 (2008) 41-52. [7] R. Sakurovs, S. Day, S. Weir, Relationships between the Critical Properties of Gases and Their High Pressure Sorption Behavior on Coals, Energy & Fuels, 24 (2010) 1781-1787. [8] R. Sakurovs, S. Day, S. Weir, G. Duffy, Application of a modified Dubinin-Radushkevich equation to adsorption of gases by coals under supercritical conditions, Energy & Fuels, 21 (2007) 992-997. [9] S.A. Mohammad, J.S. Chen, J.E. Fitzgerald, R.L. Robinson, K.A.M. Gasem, Adsorption of Pure Carbon Dioxide on Wet Argonne Coals at 328.2 K and Pressures up to 13.8 MPa, Energy & Fuels, Accepted (2009). [10] R. Sakurovs, S. Day, S. Weir, Causes and consequences of errors in determining sorption capacity of coals for carbon dioxide at high pressure, International Journal of Coal Geology, 77 (2009) 16-22. [11] R. Sakurovs, S. Day, S. Weir, G. Duffy, Temperature dependence of sorption of gases by coals and charcoals, International Journal of Coal Geology, 73 (2008) 250-258.
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Oviedo ICCS&T 2011. Extended Abstract
CIUDEN CO2 Transport Test Rig: Technical Description and Experimental Plan B. Navarrete 1; P. Otero; I. Llavona; M.A. Delgado; Fundación Ciudad de la Energía CIUDEN, II Avenida de Compostilla nº2 Ponferrada, Spain.
[email protected]: Abstract Fundación Ciudad de la Energía (CIUDEN), a public Foundation created by the Spanish Government, is currently designing an experimental installation to study the main issues related to CO2 pipeline transport, within CIUDEN’s Technology Development Centre for CO2 Capture es.CO2. This paper describes CIUDEN´s CO2 Transport Test Rig, currently under development of the detailed engineering design and R&D program, prior to the immediate start of the construction of the Installation. The main technical characteristics of the process units are the following: (a) Pumping system to transport CO2. (b) High pressure vessel. (c) Recirculation pump and heat exchanger system in order to set operation pressures and temperatures within the range of 80 -110 barg and 10 - 30 ºC respectively. (d) Dosing equipment to injecting impurities in the CO2 flow. (e) Tube coils with variable lengths and different materials. (f) Test zones. CIUDEN has designed a complete test campaign, mainly focused on the evaluation of the effect of impurities and contaminants on the mechanical behaviour of different steels and other materials, equipment and instrumentation. It will also be studied the influence of different compositions of CO2 on the flow assurance, as well as CO2 sudden depressurization and its effects on the pipeline, equipment and instrumentation. Due to the semi-industrial size of the installation and the flexibility of its design, the results are expected to be particularly valuable for the design and construction of commercial CO2 transport pipelines and auxiliary equipments. Additionally, the facility could be used as a test facility for manufacturers of industrial transport and storage injection equipments and auxiliaries and training–room for operators in charge of the operability/maintenance of the industrial CO2 pipelines.
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Oviedo ICCS&T 2011. Extended Abstract
1. Introduction
Earlier work by Svensson et al. [1] identified pipeline transport as the most practical method to move large volumes of CO2 overland and other studies have affirmed this conclusion [2]. There is considerable experience in the transport of CO2 by pipeline primarily for use in enhanced oil recovery (EOR) operations [2, 3]. An important number of studies have been carried out in order to define sizing of CO2 pipelines and estimate the capital cost of pipeline transport. Also, models have been developed for pipelines engineering-economic evaluations [4]. However further investigations are necessary in order to solve the uncertainties related to the effect of impurities on the behavior of the fluid and their consequences upon the pipeline design and operation [5]. Fundación Ciudad de la Energía, a public Foundation created by the Spanish Government, is involved on the construction of a Technology Development Centre for CO2 Capture, hereinafter, the Centre or “es.CO2”. The main objectives of the Centre are the research, development and demonstration of efficient, cost effective and reliable carbon capture and storage technologies (CCS) [6]. The main units of the es.CO2 Centre, are the following: - Fuel Preparation unit. - Pulverized coal boiler (20 MWth) capable of operating from air-mode to full oxycombustion-mode. - Circulating fluidized bed boiler (15 MWth in air-mode, 30 MWth in oxymode). - Comburent preparation system. - Flue gas cleaning train to remove dust, NOx and SOx. - CO2 compression and purification unit (oxymode). - CO2 transport experimental installation. In addition, es.CO2 includes a biomass gasification unit of 3 MWth, as part of another initiative of the Spanish Government. Figure 1 shows a diagram of the installation [7, 8, 9]. Focused on the CO2 Transport Experimental Facility activity, the main goals will be the evaluation of the effect of impurities and contaminants and the study of the thermodynamic conditions of the CO2 on the mechanical behavior and corrosion of different steels and other materials, as well as on the flow assurance. It will also be used for industrial equipment and instrumentation testing and is suitable to study the CO2 depressurization and its effects on the pipeline and equipment.
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Oviedo ICCS&T 2011. Extended Abstract
Figure 1.- Schematic process diagram of es.CO2
2. Experimental section The CO2 Transport Test Rig at the es.CO2 Centre, , includes the following main units, as shown in Figure 2: (a) Pumping system to transport CO2 either from storage vessels (commercial quality) or the CPU (Compression and Purification Unit) of the Centre. (b) High pressure vessel to avoid fluctuations in the flow. (c) Recirculation pump and heat exchanger systems in order to set operation pressures and temperatures within the range of 80 -110 barg and 10 - 30 ºC respectively. In order to operate the test rig in thermal conditions similar to those expected in CO2 transport pipelines (mainly buried), the facility will be located inside a highly thermal isolated building with thermal control. Dimensions of the industrial building are 23x19x10 m3. It is designed considering that has to be capable of maintain isothermal conditions in the interval of 15 to 18 ºC, as well as to deal with possible enlargements and safety issues as a consequence of the use of the CO2 flow. To achieve these requirements, the building will include a heater/cooler system that will make the most of the steam produced in the es.CO2- The building will be made of pre-fabricated concrete and will have an effective ventilation system to avoid sub-oxygenated atmospheres.
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Oviedo ICCS&T 2011. Extended Abstract
Figure 2.- Block diagram of CO2 Transport Test Rig (d) Dosing equipment to add impurities and contaminants to simulate different CO2 streams composition expected to be captured in the industry: H2O, NOx, SOx, N2, O2, Ar, CO, H2, H2S, CH4. It is important to point out that besides the es.CO2 was conceived considering the development of oxycombustion processes, the CO2 transport experimental facility will test CO2 streams including typical contaminants of precombustion processes, such as CO, H2, H2S and CH4. (e) Tube coils with variable length and different materials: each coil has an equivalent length of approximately 500 m and a nominal diameter of 2 inches. Considering the number of tube coils, the length of the whole test rig will exceed 5,000 meters. It is also possible to by-pass one or several tube coils in order to be adapted to specific conditions. (f) Test zones with pipes of different diameters in order to install new equipment or instrumentation to be tested in real conditions of CO2 transport. The number of test zones is designed considering the different tests that will be carried out (see Table 1)
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Oviedo ICCS&T 2011. Extended Abstract
Figure 3. 3D simulation of the Test Rig where it can be seen the tube coils and some of the test zones [10] 3. Test campaign and test matrix To achieve the aforesaid objectives, a set of specific testing campaigns has been designed focused on the data acquisition for scaling-up the system, operator training and CO2 safety operation.Table 1 shows the type of test that will be carried out and the independent variables that will be modified during the tests duration. Table 1. Summary of the test campaign. ID
Type of test
1
Corrosion rates in different materials.
2
Flow assurance (depressurization in the line).
3
Installation of industrial instrumentation or equipment.
4
Release studies.
Independent variable Pressure, Temperature. CO2 quality. CO2 velocity. Pressure, Temperature. CO2 quality. CO2 velocity. Diameter. Pressure, Temperature. CO2 quality. TBD.
4. Conclusions
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Oviedo ICCS&T 2011. Extended Abstract
The CIUDEN´s CO2 Transport Test Rig that is installed in the CIUDEN´s Technology Development Centre for CO2 Capture will provide real basis for the design, maintenance and operation of industrial CO2 pipelines. The designed test campaigns will generate valuable information to material selection, impure CO2 behavior, depressurization and CO2 safety operation; besides this and considering that the installation is located inside a building, it will be possible to test CO2 small releases in order to study or validate dispersion models. Acknowledgement Part of the research work presented in this paper is co-financed by the European Union´s European Energy Programme for Recovery programme. The sole responsibility of this publication lies with the author. The European Union is not responsible for any use that may be made of the information contained therein. References [1] Svensson, R., et al., Transportation systems for CO2-application to carbon capture and storage. Energy Conversion & Management, 2004. 45: p. 2343-2353. [2] Doctor, R., et al., Transport of CO2, in IPCC Special Report on Carbon Dioxide Capture and Storage, B. Metz, et al., Editors. 2005, Cambridge University Press: Cambridge, U.K. [3] Gale, J. and J. Davidson, Transmission of CO2- Safety and Economic Considerations. Energy, 2004. 29: p. 1319-1328. [4] Bock, B., et al., Economic Evaluation of CO2 Storage and Sink Enhancement Options. 2003, TVA Public Power Institute: Muscle Shoals, AL. [5] IEA Greenhouse R&D Programme, Impact of impurities on CO2 Capture, Transport and Storage, PH4/32, august 2004.
[6] Cortes V J. State of development and results of oxy-coal combustion research initiative by CIEMAT in Spain. 2nd IEAGHG International Oxy-Combustion Network Meeting. Windsor, USA. January 2007. [7] Cortes V J, Navarrete B. Test Facilities for Advanced Technologies for CO2 Abatement and Capture in Coal Power Generation. 3rd IEAGHG International Oxy-Combustion Network Meeting. Yokohama, Japan. March 2008. [8] Lupion M, Navarrete B; Otero P; Cortes V J. CIUDEN CCS Technological Development Plant on oxycombustion in Coal Power Generation. 1st International Oxyfuel Combustion Conference. Cottbus, Germany. September 2009 [9] Lupion M. CIUDEN Carbon Capture Technology Development Plant. Second APP Oxy-fuel Capacity Building Course. Beijing, China. March 2010 [10] ISOLUX CORSÁN. Technical proposal for CIUDEN´s bidding process LT-2010/02
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Oviedo ICCS&T 2011. Extended Abstract
Viscosity behaviour of slags from coal-petroleum coke blends
Alexander Ilyushechkina and Marc Duchesneb a
CSIRO Energy Technology, Queensland Centre for Advanced Technologies,
Technology Court, Pullenvale QLD 4069, Australia b
CanmetENERGY, 1 Haanel Drive, Ottawa ON K1A1M1, Canada
Abstract The slagging behaviour of petroleum coke and coal blends must be known to determine suitable blending requirements for entrained-flow slagging gasification. In the present study, the viscosities of petcoke ash blended with Australian and Canadian coal ashes were measured in the temperature range of 1200-1550ºC. The effect of compositional changes in the blends on viscosity behaviour was investigated in terms of solids precipitation. At high temperatures, increasing the amount of petcoke ash (up to 100% of the coal ash weight) gradually reduces the viscosity of the blends used in this study. This reduction in viscosity is likely associated with changes in slag composition, particularly for the major components CaO, SiO2, Al2O3, and FeOx. These changes decrease the liquidus temperature, increase the amount of CaO, and decrease the Si/Al ratio. However, increasing the petcoke content in the blends may increase the temperature of critical viscosity. Slag quenching experiments indicate that the presence of vanadium in petcoke containing slags stimulates solids precipitation, which increases viscosity dramatically at lower temperatures.
1. Introduction Production of petroleum coke (petcoke), a by-product of the oil refining industry, has been and is expected to continue increasing [1]. Major factors driving this increase include the rising demand for transport fuels, the use of heavier crude oils and new environmental regulations pushing for reduced waste and highly refined fuels. Due to petcoke’s high heating value and low cost, there is much interest in its use as a primary fuel or in a coal-petcoke blend. Oil sands petcokes have been found to be particularly unreactive, with reactivities comparable to high rank coals such as anthracites [2]. In
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1
Oviedo ICCS&T 2011. Extended Abstract
addition, the high sulphur (5–6%) and metals (particularly V and Ni) content of the petcokes cause concern for downstream treatment and residue disposal. One method of mitigating undesirable aspects of the petcoke as a fuel is to blend it with other fuels, such as lignite, sub-bituminous coals and bituminous coals. The higher volatile content of these fuels could potentially make gasification of the petcoke more reliable by lowering ignition and ash fusion temperatures (for slagging gasifiers). The majority of gasifiers in operation today are of the entrained-flow type. In this type of gasifier, most of the inorganic component of the fuel (ash) is partially or fully melted, sticks to the reactor wall and flows to the bottom as slag. To determine the suitability of a fuel and appropriate operating conditions for gasification, it is important to know its slagging properties. A slag which is too viscous may accumulate on the reactor wall till the reactor is plugged, ceasing operation. As a rule of thumb, slag viscosity should not exceed 25 Pa s at the slag tapping temperature [3, 4]. The petroleum coke ash can contain high concentrations of vanadium, nickel, and iron. Vanadium oxide is a major impurity in the petcoke slags which is not found in coal ash. The oxidation state of vanadium depends upon the temperature and partial pressure of oxygen [5]. Because of the low oxygen partial pressure existing in a gasifier, a VO x phase can precipitate under reducing conditions during gasification. Previous studies have identified karelianite (V2O3) in slag which can increase the viscosity of a coal/petcoke mixture [6]. In order to prevent V2O3 from accumulating inside the gasifier, the slagging behaviour of petroleum coke and coal ash blends has to be understood. In the present study, the viscosities of Australian and Canadian coal ashes blended with petcoke ashes were measured in the temperature range 1200-1550ºC. Microstructural analysis of quenched slag was conducted to investigate a link between solid phase formation and viscosity behaviour.
2. Experimental 2.1 Coal and petcoke selection Several coal ashes and petcoke ashes were used in this study. Compositions are given in Table 1. C1 is an artificial coal ash, for which the composition is based on the major and minor oxides analysis of Genesee coal, a western Canadian sub-bituminous coal. C2 is a slag obtained from the gasification of a low-rank (sub-bituminous) coal from Western Australia. C3 is a slag obtained from gasification of an Australian high volatile bituminous coal from the Hunter Valley in NSW. Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
P1 is real ash produced from Suncor petcoke by ashing at 700-750 ºC for 20-30h. P2 is artificial Suncor petcoke ash; its composition is based on the major and minor oxides analysis of Suncor petcoke, excluding sulphur. Artificial coal and petcoke ashes were prepared by mixing laboratory or analytical grade Al2O3, CaO, Fe2O3, K2CO3, MgO, Na2CO3, NiO, S, SiO2, TiO2 and V2O5 powders. Slag samples were collected from the quench vessel of Siemens pilot plant gasifier [7].
Table 1. Composition of ashes and slags (wt.%) used in this study Samples C1* C2 C3 P1 P2* SiO2
58.02
49.62
47.73
37.77
40.66
Al2O3
23.47
18.22
24.07
16.6
17.87
Fe2O3
4.55
24.36
15.41
7.35
7.91
TiO2 CaO MgO
0.53 5.63 1.57
1.17 3.27 1.62
1.49 8.25 0.97
1.1 15.82 3.87
1.19 17.03 4.17
Na2O
2.34
0.55
0.28
1.67
1.8
K2O
0.58
0.45
0.3
0.96
1.03
V2O5 NiO
<0.2 <0.2
<0.2 <0.2
<0.2 <0.2
6.67 1.09
7.18 1.17
P2O5
<0.2
0.73
1.49
<0.2
<0.2
SO3 3.32 <0.2 <0.2 7.11 <0.2 * composition of ash used for sample preparation
2.2 Viscosity measurements Slag viscosity measurements were made using a Haake high temperature viscometer using a rotational bob [8]. Slags were placed into a molybdenum crucible and heated to 1600ºC in a neutral atmosphere (N2). To ensure the absence of oxygen, the crucible was placed in a sacrificial graphite liner. Viscosity measurements were performed using a molybdenum rotating bob connected to a viscometer head. Measurements were conducted at incremental temperature steps (20-30ºC) within the temperature range 1200-1600oC. On the cooling stage samples were held at each temperature for at least 30-min intervals to equilibrate compositions inside the crucible. The overall uncertainty in the viscosity measurements is within ±5 %, which includes the uncertainties of calibrating oil (2%) and viscometer accuracy (1%). The accuracy of the temperature measurements are within ±3 ºC. At each temperature point, the slag viscosity was coalSubmit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
petcoke slag mixtures simulating those encountered in entrained bed slagging gasifiers. Energy Fuels 2009;23:4723–33. measured at least three times to ensure the repeatability of measurements. The average of these measurements was used as the value of viscosity. The standard deviation of the measurements was below 2%, which is less than the overall uncertainty of the measurements
1.1 Analysis of ash and quenched slags Slag cold rod quenching experiments were performed in order to analyse slag phase composition at specific temperatures related to viscosity measurements. Cold rod quenching involves dipping a molybdenum rod in the molten slag after viscosity measurements, removing it and then cooling it in water. Quenched slags were investigated using Scanning Electron Microscopy (SEM) in backscattering mode. Energy Dispersive X-ray Spectroscopy (EDS) was undertaken to identify the elemental composition of selected regions of the slag. Electron Probe Microanalysis (EPMA) was used to identify the composition of solid and liquid phases in the quenched samples. Bulk chemical composition of the slag and ash were determined by Inductively Coupled Plasma - Atomic Emission Spectroscopy (ICP-AES) and confirmed by EPMA.
2. Results and discussion 2.1 Slag viscosity Slag viscosity was investigated for three different blends of coal ashes and petcoke ash: C1+xP2, C2+xP1, and C3+xP1 (x indicates petcoke/coal ash ratio). In each blend, petcoke ash content, x, varies in the range of 0.125-1. Viscosity graphs for each blend are plotted in Figure 1a-1c as a function of slag temperature. Sample C1 has a viscosity well above 25 Pa·s for the entire temperature range tested. This was not a surprise as its SiO2 content is high and there were slag plugging problems when Genesee coal was used in a pilot-scale entrained-flow gasifier [9]. Bothlimestone and dolomite are effective for viscosity reduction of artificial Genesee coal ash [10]. In the present study, adding petcoke ash also reduces the viscosity, but does not have as much of an effect as limestone or dolomite. Viscosity of the slags gradually decreases with an increasing amount of petcoke in the C1 coal ash with P1 petcoke ash blends. It is likely that the low network former (such as SiO2) and high network modifier (such as CaO and MgO) contents of the petcoke ash have a major effect on the viscosity of these blends. Submit before May 15th to
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4
Oviedo ICCS&T 2011. Extended Abstract
a)
b)
T1
c)
T2
T3
Figure 1. Viscosity of slags C1+P2 blend (a), C2+P1 blends (b), and C3+P1 blends (c), and C4+P3 blends (d). T1-T3 are the slag quenching temperatures. Viscosity of coal C2 ash is below 25 Pa·s at temperatures above 1270 ºC and addition of the petcoke ash P1 to the slag does not change its viscosity appreciably. At temperatures below 1270 ºC the viscosity of the blends becomes higher than the viscosity of the C2 slag. For slag C3, which also has a relatively low viscosity (less than 25 Pa·s at temperatures above 1320 ºC), addition of the petcoke shows the following tendencies: i) viscosity of blended samples are always lower than viscosity of C3 slag at high temperatures (above 1275 ºC); ii) for the C3+0.25P1 blend, viscosity is always lower than for C3 for the entire temperature range tested; iii) increasing petcoke ash content to x=0.5 leads to a more dramatic increase in the viscosity at low temperatures (below 1275 ºC) and viscosity of this slag becomes higher than the viscosity of the C3 slag. For all blends studied, addition of petcoke ash reduces viscosity at high temperatures where the slag is fully liquid. The effect of petcoke ash in the liquid phase is from modification of the concentration of major slag components (SiO2 Al2O3, Fe2O3
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Oviedo ICCS&T 2011. Extended Abstract
and CaO), a reduction of silica (a network former), and an increase in CaO (a network modifier). It is also possible that V2O3 in the liquid phase may play a similar role as amphoteric Al2O3: it can act as either a network former or a network modifier.
2.2 Slags microstructure In order to understand viscosity behaviour below the liquidus temperature of the investigated slags, a series of slag quenching experiments was conducted. Quenching temperatures, indicated in Figure 1, were selected where the viscosity increase is significant and where the differences between viscosities of coal slags and their blends with petcoke are clearly identified. Figure 2 shows micrographs of C1+xP2 slags quenched from 1350 ºC. Only C1 slag contains a solid phase which is identified by EPMA as mullite. Coincidentally, the viscosity of C1 is very high (see Figure 1). Both blended samples have only a liquid phase at this temperature. It is clear that for C1 coal ash, blending with P2 petcoke ash significantly changes the phase equilibria in the slag system and lowers the liquidus temperature.
a)
b) M
c) L L
L Figure 2. Microstructure of slags quenched from 1350 °C: C1 (a), C1+0.5P2 (b) and C1+P2 (c). Legend: L- former liquid phase, M- mullite c) Micrographs of C2+xP1 slags quenched from 1200 °C are shown in Figure 3. C2 slag is fully liquid, while C2+0.25P1 and C2+0.5P2 blends have spinel and spinel with feldspar, respectively. At 1200 °C, the viscosity of the C2+0.25P1 and C2+0.5P2 slags are higher than the viscosity of the C3 slag, and it is likely associated to the higher solids content in the C2+0.25P1 and C2+0.5P2 slags, which includes V-rich spinel.
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Oviedo ICCS&T 2011. Extended Abstract
a)
b)
c)
L
S L
S
F L
Figure 3. Microstructure of slags quenched from 1200 °C: C2 (a), C2+0.25P1 (b) and C2+0.5P1 (c). Legend: L- former liquid phase, F- feldspar, S- spinel The presence of vanadium spinel is also observed in C3+0.25P1 and C3+0.5P1 slags quenched from 1250 ºC (Figure 4). Pure C3 slag has only a small amount of mullite at this temperature (also Figure 4). C3+0.25P1 slag has feldspar as the primary phase field and V spinel as the secondary phase field. For C3+0.5P1 slag, the composition shifts with spinel as the primary phase field, and at 1250 ºC, spinel is already formed as large clusters. Hence sharp viscosity increase in blended slag samples (C3+xP1) may also have resulted from formation of V-rich spinels. a)
b)
c) L
S
M
L
L F
S Figure 4. Microstructure of slags quenched from 1250 °C: C3 (a), C3+0.25P1 (b), C3+0.5P1 (c). Legend: L- former liquid phase, F- feldspar, M- mullite, S- spinel A more detailed analysis of quenched slag phase compositions will be published separately.
Microstructural analysis of the quenched samples reveals that blending coal ash slags with petcoke ash has a complex effect on phase formation in slags. The petcoke ashes used in this study have the same major components as the coal ashes. Therefore, the addition of petcoke ash shifts the bulk composition for the four major components and may change the liquidus temperature in the new system, or even shift the composition to another primary phase field than that of the original coal ash. The presence of V may also shift the composition to have a spinel primary phase field and force the growth of solids at temperatures higher than the liquidus temperature of the Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
original coal ash. These microstructural changes have effects on viscosity such as: i) they change the temperature at which solids start to form; ii) below the liquidus temperature, a new solids-liquid assembly may form if the system shifts to a new primary phase field; iii) the presence of vanadium may assist spinel formation and increase the amount of solids in the slag; iv) V spinel growth consumes Fe and Al, and therefore likely changes the viscosity of the liquid phase leading it to contain less FeOx and a higher silica/alumina ratio.
3. Conclusions A series of viscosity measurements and microstructural analyses of quenched samples indicate that blending coal ashes with petcoke ash has a complex effect on the viscosity. Addition of petcoke ash in coal ashes changes the composition in terms of major components such as SiO2, CaO, Al2O3, and FeOx, as well as the Si/Al ratio. For all blends presented in this study, addition of petcoke ash lowers viscosity at temperatures above the liquidus. Since the composition of the slags changes with petcoke ash addition, the liquidus temperatures also change. The shifting of resulting compositions to a new primary phase field and formation of new solids is possible. In slags with a relatively high iron oxide content, the presence of vanadium stimulates formation of spinels at low temperatures. This may lead to a gradual viscosity increase with increasing vanadium in blended mixtures.
References [1] Roskill L, The economics of petroleum coke, 5th ed. Information Services, 2007. [2] Furimsky, E. Gasification of oil sand coke: Review. Fuel Process Technol 1998; 56: 263–70. [3] Browning GJ, Bryant GW, H.J. Hurst HJ, Lucas JA, Wall TF. An empirical method for the prediction of coal ash slag viscosity. Energy and Fuels 2003;17:731-7. [4] Folkedahl BC, Schobert HH. Effects of atmosphere on viscosity of selected bituminous and low-rank coal ash slags. Energy and Fuels 2005;19:208-15. [5] Farah HJ. V4+ in quenched calcium silicates: An electron spin resonance spectroscopic investigation. Mater. Sci. 2003;38:727–37. [6] Nakano J, Sridhar S, Moss T, Bennett J, and Kwong K-S. Crystallization of synthetic Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
[7] Roberts DG, Harris DJ, Tremel A, Ilyushechkin AY. New insights into coal conversion and slag formation during entrained flow gasification and their impacts on gasification performance. Proceedings International Freiberg Conference on IGCC&XtL Technologies, 4th, 2010. [8] Hurst HJ, Novak F, Patterson JH. Viscosity measurements and empirical predictions for some model gasifier slags. Fuel 1999;78:439-44. [9] Cousins A, McCalden DJ, Hughes RW, Lu DY, Anthony EJ. Entrained-flow gasifier fuel blending studies at pilot scale. Can J Chem Eng 1008;86:335-46. [10] Duchesne MA, Ilyushechkin AY, Macchi A, Anthony EJ, Optimization of Canadian petroleum coke, coal and fluxing agent blends via slag viscosity measurements and models. Proceedings Annual International Pittsburgh Coal Conference 27th, 2010;38-2.
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Study on Clinker generation control in coal combustion boiler ― Clinker controlling effect of Fe-based coal additive ― Nobuyuki WAKABAYASHI 1,
Hiromi SHIRAI2
Central Research Institute of Electric Power Industry (CRIEPI) 2-6-1 Nagasaka, Yokosuka-shi, Kanagawa-ken 240-0196 JAPAN E-mail:
[email protected]、
[email protected] Abstract In pulverized coal fired boilers, clinker generation due to ash adhesion on the furnace wall and superheater (SH) is a serious problem. To address this problem, the application of the coal additive that controls the generation of the clinker is expected. And the effect needs to be clarified. In this study, the effect of an Fe-based coal additive was evaluated under the condition of SH in a pulverized coal combustion test furnace (thermal input of fuel: 760 kW). As a result, it was confirmed that the ash adhesion growth rate was decreased, and the melting of the adhesion ash was controlled by applying the Fe-based coal additive. 1. Introduction In a pulverized coal fired power boiler, ash adheres to the furnace wall and superheater (SH) and generates clinker in operation. During periodic overhaul, it is necessary to ensure work safety due to the hazards of falling clinker. Therefore, clinker adhering to the furnace wall and SH must be removed. During the periodic overhaul, clinker removal work takes several days. It is an important issue to reduce the number of work days. To control clinker generation, the application of a coal additive that controls clinker generation is expected, and the effect needs to be clarified. Therefore, in this study, to clarify the effect of an Fe-based coal additive, a method of evaluating the effect of controlling clinker generation using a pulverized coal combustion test furnace was studied. Moreover, the effect of the Fe-based coal additive was evaluated using three types of coal with different melting points under SH condition (tube surface temperature: 600℃, gas temperature: 1200℃) by additive test and no-additive test. 2. Experiment 2.1 Fe-based coal additive An outline of the Fe-based additive used in this study is shown in Table 1. The additive
is a black liquid, and the Fe concentration is 10wt% as the Fe2O3 conversion concentration. In a pulverized coal fired power boiler, the recommended additive amount for coal is 1/2000 ~ 1/5000 as the coal weight ratio; this additive amount is very small. Table 1 Outline of Fe-based coal additive solution Fe2O3 conversion concentration (Additive stock solution)
10wt%
Appearance
Black liquid
Recommended amount of addition
Coal weight ratio 1/2000 ~ 1/5000 (Continuous injection)
2.2 Tested coal properties and amount of Fe-based additive added. The tested coal properties and amount of additive added are shown in Table 2. So that the effect of the additive on the ash composition will become equal for each tested coal, the amount of additive added was 0.77wt% as the ash weight ratio for coal. This amount of additive added was 1/2000 as the coal weight ratio in Coal A. The amount of change in ash composition was calculated with this additive amount. As a result, mainly Fe2O3 is increased at 0.07wt%, and it is very small. Therefore, there is no significant change in the ash composition following the additive addition. Table 2 Tested coal properties and amount of Fe-based additive added Amount of Fe-based additive added Coal properties
Coal weight ratio Ash weight ratio [wt%] Lower heating value [MJ/kg] Fuel ratio(FR) [-] Ash [wt%] Ash compositions [wt%] SiO2 Al2O3 Fe2O3 CaO MgO TiO2 P2O5 Na2O K2O SO3 Ash fusion temperatures [℃] Initial deformation (Oxidizing) Hemispherical Fluid
Coal A Coal B Coal C 1/2000 1/1222 1/1483 0.77 0.77 0.77 28.3 27.8 27.5 1.13 1.79 1.00 7.4 11.6 10.0 55.57 51.71 44.79 24.46 26.04 35.97 10.01 12.96 4.28 2.17 2.43 3.78 1.69 1.09 1.45 1.27 1.35 3.11 0.30 0.59 1.09 0.67 0.41 0.84 1.86 1.77 0.29 1.55 1.58 3.21 1,350 1,450 1,550< 1,420 1,490 1,550< 1,440 1,520 1,550<
2.3 Pulverized coal combustion test furnace The test furnace in this study is a water-cooled, horizontal cylindrical type made of steel, equipped with a single burner. The inner diameter of this furnace is 0.85m and the length is 8m. Refractory materials are coated onto the inside wall of the furnace. The thickness is 0.075m. The thermal input in the furnace is 760kW and the bituminous coal feed rate is about 100kg/h[1]. 2.4 Experimental methodology In an actual plant, the temperature of the ash adhering to the heat exchanger tube is approximately tube surface temperature at near the tube surface. On the other hand, as the ash adheres to the tube, the adhesion ash layer surface temperature comes close to the gas temperature owing to the heat thermal resistance of the adhered ash. Therefore, in this study, the characteristic of ash adhesion was focused on the difference in ash adhesion surface temperature and evaluated as follows. The characteristic of initial ash adhesion was evaluated using a cooling probe that simulates the boiler tube surface temperature. The characteristic of the adhesion ash layer that was sufficiently grown, the surface temperature of which was the gas temperature, was evaluated using a test piece without cooling. The test conditions are shown in Table 2. For the combustion conditions in this study, the O2 concentration at the furnace exit was 2%, and the coal feed rate is about 120kg/h to give the gas temperature condition that is shown below for ash adhesion evaluation (the thermal input in the furnace is approximately 930kW). For the ash adhesion evaluation conditions, the gas temperature was 1200℃ and the surface temperature of the cooling probe was 600℃. Figure 1 shows the Schematic of the test furnace, a cooling probe and a test piece without cooling that were set in the test furnace. After the temperature of the combustion air and exhaust gas became constant, a cooling probe was installed horizontally through the observation window at 3.39m from the burner exit. The gas temperature at 3.39m from the burner exit is assumed to be 1200℃. Then ash adhered to the probe for about 5 hours. During the test, the ash adhesion on the probe was recorded using a video camera that was set opposite to the probe. The probe surface temperature was controlled by adjusting the flow of coolant and was set to about 600℃ when the ash adhesion on the probe started. During the test, the coolant flow was constant. A test piece without cooling was placed at the furnace bottom before the combustion test, and was obtained after the combustion test.
Table 3 Test conditions Combustion condition
O2 concentration at furnace exit Coal input (Thermal input of fuel) ※LHV basis Ash adhesion test condition Gas temperature Probe surface temperature
φ850mm
8.00 m
2% 120kg/h (930kW) 1200℃ 600℃
Cooling probe
3.39 m Video camera
Video camera
Coolant
・・・
Primary air and puverized coal
Furnace outlet
Cooling probe Test piece without cooling
Figure 1 Pulverized coal combustion furnace and test section
3. Results and Discussion 3.1 Effect for combustion Figure 2 shows the gas temperature at the ash adhesion evaluation position on the center axis of the furnace. The gas temperature was about 1200℃ which was a test condition for each test. However, in the case of additive addition, the gas temperature was 60℃ lower than in the case of no-additive addition at Coal A and 50℃ lower than in the case of no-additive addition at Coal B. The effect of additive for combustion was evaluated using unburned fraction (U*c) that was calculated using the following equation.
𝑈𝑐∗ =
𝐴𝑠ℎ 100−𝐴𝑠ℎ
∙
𝑈𝑐 100−𝑈𝑐
× 100 (= 100 − 𝜂)
Where, Ash is the ash content in coal[%], Uc is the unburned carbon concentration in fly ash[%], and ηis the combustion efficiency[%]. Figure 3 shows the unburned fraction for each test. In the case of additive addition, the unburned fraction was lower than in the case of no additive addition for each type of coal. In the coal combustion, if the coal combustion condition is the same, the unburned fraction increases as the fuel ratio increases [1]. A similar tendency was also shown in this test. Furthermore, the
decreased amount of the unburned fraction increased as the fuel ratio increased. Thus, in Coal A and Coal B, the unburned fraction of which was relatively higher than that of Coal C because of the effect of additive, it is considered that combustion became well in the burner neighborhood by combustion promotion, the peak of gas temperature distribution moved to the burner side in the furnace, and the gas temperature decreased at the ash adhesion evaluation position.
2.0
1600 □No additive added ■Additive added
1400 1300 1220℃
1200
1160℃
1240℃ 1190℃
1185℃ 1185℃
1100 1000 900
Unburned fraction, Uc*[%]
Temperature [℃]
1500
□No additive added ■Additive added
1.8 1.6 1.4 1.2
0.9 0.4
1.0 0.8 0.6
0.4 0.1
0.2 0.0
800 Coal A
Coal B
Coal C
Figure 2 Gas temperature at the ash adhesion evaluation position on the center axis
Coal A(FR:1.13)
Coal B(FR:1.79)
Coal C(FR:1.00)
Figure 3 Unburned fraction (Uc*)
of the furnace.
3.2 Ash adhesion test using a cooling probe that simulates a boiler tube surface temperature. Figure 4 shows an example of ash adhesion on the probe. Because the combustion gas is swirling and flows toward the furnace exit in this furnace, the combustion gas collided with the probe from lower right at 45°as shown in Figure 4, and the adhesion ash grew up toward the flue gas flow. The ash adhesion characteristic was evaluated using the ash adhesion growth rate that was calculated using the following equation. 𝐴𝑠 𝑎𝑑𝑒𝑠𝑖𝑜𝑛 𝑔𝑟𝑜𝑤𝑡 𝑟𝑎𝑡𝑒 [(𝑚𝑚⁄)⁄(𝑘𝑔⁄( ∙ 𝑚2 ))] =
𝑐𝑎𝑛𝑔𝑒 𝑖𝑛 𝑎𝑑𝑒𝑠𝑖𝑜𝑛 𝑎𝑠 𝑙𝑎𝑦𝑒𝑟 𝑡𝑖𝑐𝑘𝑛𝑒𝑠𝑠 𝑝𝑒𝑟 𝑢𝑛𝑖𝑡 𝑡𝑖𝑚𝑒 [𝑚𝑚⁄] 𝑎𝑚𝑜𝑢𝑛𝑡 𝑜𝑓 𝑎𝑠 𝑡𝑟𝑜𝑢𝑔 𝑢𝑛𝑖𝑡 𝑐𝑟𝑜𝑠𝑠 − 𝑠𝑒𝑐𝑡𝑖𝑜𝑛𝑎𝑙 𝑎𝑟𝑒𝑎 𝑖𝑛 𝑓𝑢𝑟𝑛𝑎𝑐𝑒 𝑝𝑒𝑟 𝑢𝑛𝑖𝑡 𝑡𝑖𝑚𝑒 [(𝑘𝑔⁄( ∙ 𝑚2 ))]
Where, the change in adhesion ash layer thickness per unit time was given by analysis
Probe
45° ash
Figure 4 Ash that adhered to probe (5:15 after the start of the test)
Ash adhesion growth rate [(mm/h)/(kg/(h/m2)]
of ash adhesion image. Figure 5 shows the ash adhesion growth rate at each test. In the case of additive addition, the ash adhesion growth rate has decreased at Coal A and Coal B. The ash adhesion growth rate at Coal C did not change significantly. Therefore, in Coal A and Coal B, it was indicated that the ash adhesion has decreased by applying the Fe-based coal additive.
0.30 0.25
□No additive added ■Additive added 45°
0.20
0.15
Thickness of adhesion ash
0.10 0.05
0.00 Coal A
Coal B
Coal C
Figure 5 Ash adhesion growth rate
3.3 Ash adhesion test by using test piece without cooling. Figure 6 shows test pieces without cooling after the test. The ash that adhered to the test piece grew toward the gas flow as well as the adhesion ash in the cooling probe. In Coal A, the adhesion ash without the additive was hard and its appearance shows that it was melting during the test; the adhesion ash with the additive was fragile and didn’t seem to have melted from its appearance during the test. In Coal B, the appearance of the adhesion ash without the additive shows that it was melting and flowing downward during the test; the appearance of the adhesion ash with the additive shows that its melting was suppressed compared with that without the additive, and it grew toward the gas flow. In Coal C, with or without the additive, the adhesion ash was fragile, and there is no significant difference. Therefore, in Coal A and Coal B, it was indicated that the melt of the adhesion ash was controlled by applying the Fe-based coal additive.
Coal A Melted and hard ash.
No additive added
Coal B Coal C Ash was melted and flowed downward. Non melted fragile ash.
No additive added Ash melting is suppressed, and the ash grew toward the gas flow.
Fragile ash.
Additive added
No additive added Non melted fragile ash.
Additive added
Additive added
Figure 6 Ashes that adhered to test pieces without cooling. 3.4 Evaluation of the molten ash by thermodynamic equilibrium calculation It was experimentally indicated that the adhesion of the ash and the liquid phase fraction in the ash that shows the melting of the ash are correlated[2]. Moreover, it is indicated that the adhesion of the ash and the liquid phase fraction that is determined by thermodynamic equilibrium calculation based on the ash composition are correlated[3]. Thus, the liquid phase fraction in the ash for the tested coals was analyzed using a thermodynamic equilibrium calculation software FactSage Ver5.5. Then, the molten state of the ash in this test was examined. In this study, because the additive amount was 0.77wt% as the ash weight ratio for test coals, there was no significant change in the ash composition following the addition of additive. Therefore, there is also no significant change in the liquid phase fraction in the ash. Figure 7 shows the relationship between the temperature and liquid phase fraction in the ash for tested coals which was analyzed by thermodynamic equilibrium calculation. The liquid phase fraction in the ash for each tested coal changed suddenly at approximately 1200℃ which is the test condition. In this test, the changes in the gas temperature and liquid phase fraction in the ash by applying the Fe-based coal additive are as follows. In Coal A, the gas temperature decreases by 60℃ and the liquid phase fraction in the ash decreases by 13.5%. In Coal B, the gas temperature decreases by 50℃ and the liquid phase fraction in the ash decreases by 7.5%. In Coal C, the gas temperature does not change and so the liquid phase fraction in the ash also does not change. Therefore, by applying the Fe-based coal additive, in Coal A and Coal B, the causes of the ash adhesion decrease and adhesion
ash molten suppression is considered to be the gas temperature decrease and molten ash fraction decrease.
Liquid phase fraction in ash [wt%]
100 Temperature range at the test position
75
50 Coal A
25
Coal B Coal C
0 1000
1200
1400
1600
Temperature [℃]
Figure 7 Liquid phase fraction in ash by thermodynamic equilibrium calculation 4. Conclusions Under the conditions of 600℃ of tube surface temperature and 1200℃ of gas temperature, an ash adhesion test with or without additive was carried out for three types of coal with different melting points. As a result, the following differences were found by applying the Fe-based coal additive compared with the case where no additive is applied. ・ In Coal A and Coal B, the gas temperature decreased at the ash adhesion evaluation position by applying the Fe-based coal additive. This is because the peak of the gas temperature distribution moved to the burner side by applying the Fe-based coal additive. ・ In the ash adhesion test that used a cooling probe, it was indicated that the ash adhesion growth rate decreased and the ash adhesion decreased in Coal A and Coal B. ・ In the ash adhesion test that used a test piece without cooling, in Coal A, the adhesion ash with the additive was fragile and did not seem to have melted from its appearance during test. In Coal B, the appearance of the adhesion ash with the additive shows that its melting was suppressed compared with that without additive, and it grew toward the gas flow.
・ The molten state of the ash in this test was examined by thermodynamic equilibrium calculation. In Coal A and Coal B, it has been found that the gas temperature at the evaluation position decreases, and the molten state of the ash is eased as the liquid phase fraction decreases. The additive amount was very small, that is 0.77wt% as the ash weight ratio for the test coals. Therefore, there is no significant change in the ash composition following the addition of additive. References [1]Makino H, Sato M, Kimoto M, Infulence of coal properties on emission characteristics of NOx and unburned carbon in fly ash in pulverized coal combustion. JIE (Japan Institute of Energy) Journal 1994;73:906-913 [2]Ichikawa K, Oki Y, Inumaru J, Ashizawa M, Study on the mechanism of the coal ash deposit and the growth in the gasifier - The relationship between Ash melting characteristics and deposit characteristics -, CRIEPI Report,W98013 [3]Akiyama K, Pak H, Tada T, Evaluation of ash deposition behavior of upgraded brown coal (UBC) and bituminous coal, KOBE STEEL ENGINEERING REPORTS 2010, vol.60,No.1,67-70
Properties, microstructure and leaching of coal slag with additives during high temperature gasification Yoshihiko Ninomiya1*, Yajuan Wei1, Kenichi Honma2, Takao Tanosaki3 Hanxu Li4, Masato Kawaguchi5 and Norihisa Tatarazako6 1
Dept of Applied Chemistry, Chubu University, Kasugai, Aichi, 4878501, Japan
[email protected]
2
Taiheiyo Cement Corporation, Sakura, Chiba, 2858655, Japan
3
Materials Sci.& Eng., South China Univ. of Tech., Guangzhou, 510640, China
4
Dept. of Chem. Eng., Anhui Univ. of Sci. & Tech., Huainan, Anhui, 232001, China
5
Shimizu Institute of Technology, 3-4-17 Etchujima Koto-ku, Tokyo, 135-8530, Japan
6
National Institute for Environmental Studies, Onogawa, Tsukuba, 305-8506, Japan
Abstract The characteristic of gasification by-products determines the way how to make use of them. This paper was investigated the effect of MgO and CaO additives on the properties of slag. A number of mixtures of coal ash with additives were melted in a lab-scale electrically heated furnace under mildly/strongly reducing condition at 1773K. The crystallized species in the resulting slag were examined by XRD, and their microstructures were examined by SEM-EDX. The slag was also leached by water, as well as seawater; the ecotoxicity of leachant were evaluated. Furthermore, heavy metal releasing in the acid extraction also evaluated.
It was confirmed that the slag
microstructure were significantly affected by adding MgO and/or CaO to coal ash. MgO was found to be better ability than CaO on suppressing iron precipitate in the strongly reducing gas composition. The results of the leaching test of heavy metals from the slag and the ecotoxicity assay in water/seawater meet the safety regulation. However, In the case of 1M/HCl dissolution, slag generated in mildly gas composition is more stable under than that of strongly atmosphere. MgO modified slag exhibited the better ability of acid resistance than that of CaO. Keywords: additives; leaching behavior; slag utilization; gasification condition -1-
1. Introduction The integrated gasification combined cycle (IGCC) is a power generation system designed to run more efficiently than conventional pulverized coal-fired systems by combining coal gasification with a gas turbine combined-cycle system. The residues were discharged in a form of slag, which can reduced the waste volume as well as decreased emission obviously [1,2]. However, with the application of IGCC, large quality of slag is produced every year and its disposal will become a major problem， hence, efficient utilization of it is an important issue for integrated waste management around the world. Historically, slag has been used for the construction of roads and as fill material [3]. Recently, the uses of slag are expanded as cement additives, landfill cover material and for a number of agricultural applications [4]. Since slag contain a high mount of potentially toxic elements, such as Cu, Pb, Zn, As and Cr, which can be released into the environment through ageing processes and leaching, it is necessary to estimate its ecotoxcity prior to their application or use as landfill. It is well known that the properties of slag samples vary with the specific design and operating conditions in the gasifier. Studies indicate that under strongly reducing gas composition, iron oxide may be partially or fully reduced to elemental iron and may cause problems with the slag after tapping; the Shell gasifier usually generates a slag with a distinct iron-rich phase in addition to the silicate phase [5]. By contrast, in the mildly reducing atmosphere, such as air-blown gasifier, the resulting slag is glassy and granular, with no observed of iron deposit[6]. On the other hand, using entrained flow has the ability to handle practically many coals as feedstock, even the coal with a high ash melting temperature. Its fluid properties can be averted by adding flux, which improving the content of molten phase in the slag at specified temperature by forming low melting temperature eutectics [7-9]. To date, the additive used in the industrial is lime stone only. In this regard, the optimization of the properties of gasification slag by different additives is of importance. In some reported studies, the effect of additive on ash fluid temperature and viscosity of slag have been investigated [10, 11]. However, the effect of additives on the microstructure of slag, especially under different gasification condition has not been attempted.
-2-
For this study, the effect of additives on the melting behavior and properties of slag, especially under different gasification condition was investigated. The elution of metals from slag under different ecological conditions was estimated. Leaching of metals from slag was investigated using water and seawater. By using seawater, leaching behavior of heavy metals under the typical conditions encountered in the environment (especially coastal areas) can be determined. Furthermore, the acid resistance of slag was also performed, in the case of influence of acid rain.
2. Materials and Experimental procedures 2.1 Materials Two typical coals were selected in this study: one with low ash melting temperature, named as A232 here after; the other is A230. Chemical analyses of ash were conducted using X-ray florescence spectrometry (XRF, RIX 2100, Rigaku). A230 ash was mixed with the additives including reagent CaO (10%, 20%) and MgO (10%) on the mass base of ash, respectively. These two regents are of analytical grade. The mixtures were named A230-10CaO, A230-20CaO, and A230-10MgO for simplification. The chemical composition of ash and mixtures of ash with additives are given in Table 1. Table 1. Chemical composition of the mixtures of ash and additives as oxides (mass%) SiO2 Al2O3 Fe2O3 CaO MgO Na2O K2O SO3 P2O5 TiO2 MnO Total
A232 45.9 33.8 8.69 5.32 1.02 0.47 0.96 0.89 0.29 2.57 0.04 100
A230-10CaO 41.8 30.7 7.90 13.9 0.93 0.43 0.87 0.81 0.26 2.34 0.03 100
A230-10MgO 41.8 30.7 7.90 4.84 10.0 0.43 0.87 0.81 0.26 2.34 0.03 100
A230-20CaO 38.3 28.1 7.24 21.1 0.85 0.4 0.8 0.74 0.24 2.15 0.03 100
-3-
2.2 Apparatus and procedure The slag was prepared using an electrically heated furnace with a maximum temperature of 1873K, as shown in Fig. 1. In order to simulate typical gasification atmospheres, the H2/CO2 mixtures of 90/10 and 50/50 on volume fraction basis were tested in the study. The specified temperature was 1773K and the residence time of the sample in the reaction zone was 600 s. The mixture of fly ash with additive was dried at 383 K for 2 hours, and made into pellets with a 20-mm diameter and a 10-mm thickness under a pressure of 40 MPa. At the beginning, the pellet was placed in an alumina crucible at the top of the furnace. The furnace was first heated up to 1173 K with N2 atmosphere flowing in the reactor, subsequently, the gas was changed to a mixture of H2/CO2 with the previously given volume fraction. Once the specified temperature was reached, the sample-laden crucible was pushed into the high-temperature zone and was held for 10 min. Eventually, the sample was further quickly raised to the top of the furnace and instantly cooled down by N2.
Fig. 1. Schematic of the experimental rig. -4-
2.3 Characterization of slag sample The crystalline minerals present in the molten slag were identified using a powder X-ray diffract meter (XRD, RINT 2000, Rigaku) at a voltage of 40 KV and a current of 40 mA with Cu Kα radiation. The detector scanned over a 2θ range of angles from 5-80° with a step size of 0.02°. The morphology of the slag was observed using a scanning electron microscopy coupled with an energy dispersive X-ray analyzer (SEM-EDX, JEOL-6510). The slag was mounted in liquid epoxy resin, pelletized, polished, and finally sputter-coated with carbon. A computer-controlled scanning electron microscopy (CCSEM) technique was also employed to determine the visual structure, elemental composition, diameter, position, and shape factor of the precipitates in the slag [12]. The leaching procedure according to the Environmental Quality Standards for Soil Pollution of Japan [13] was applied to the slag samples in order to determine the potential leachability of their trace elements. First, the slag were pulverized to less than 100 μm and dried at 383 K for 2 hours, water and seawater were used as solution, with a liquid to solid ratio (L/S) of 10, the treatment time is 6 hours with rotate speed about 200 rpm. The concentrations of mercury (Hg), lead (Pb), chromium (Cr), cadmium (Cd), arsenic (As) and selenium (Se) in the leachates were analyzed by inductively coupled plasma atomic emission spectroscopy (ICP-AES, SPS 1700VR), Cr6+ was analyzed by colorimetry method based on the Japanese Industrial Standard Testing methods for industrial waste water K0102 (JIS K0102) [14], and pH was measured following JIS K0102 also. The toxicity of leachant was evaluated according to Microtox test by using photobacterium phosphoreum. And light measurements are made at 5 and 15 minutes intervals. The test endpoint is measured as the effective concentration of a test sample which reduces light emission by 50% after 15 minutes at 15 oC (EC50 15 MIN). The acid resistance test was performed according to the method of the Japanese Industrial Standard for chemical analysis of Portland cement (JIS R 5202-1999) [15]. The pulverized slag was digested to a L/S of 50 by 1 M HCl; this reaction was carried out in a warm water bath for 10 min. Finally, the leachant was filtered and the residue was evaluated by combusting the filter at 975±25 oC. Heavy meal concentrations in extracts were determined by ICP-AES. -5-
3. Results and discussion 3.1 Effect of additives on slag properties Fig. 2 shows the XRD patterns of the molten slag formed when CaO or MgO was added to coal fly ash, they possess a broadly amorphous diffraction pattern with a large lump centering at approximately 22-30o of 2θ
max,
signifying the presence of amorphous
species in slag. The low angle at 22-24 o of 2θ max is invariant and most likely represents a highly polymerized SiO2 net work, it is agree with other studies [15], while the high angle halo shows a strong correlation with MgO / CaO additives, representing the reduction in the polymerization of the silicate network. The halo is quite variable in intensity and 2θ max position and is also asymmetric in shape. In the case of adding 10% MgO or CaO, 2θ
max
is about 25 o. With increasing CaO addition to 20%, this
characteristic peak was shifted to a higher 2θ
max,
signifying stronger structural
modifications of a polymerization silica network.
2θ=25.8
A230-10MgO A230-10CaO A230-20CaO
A230-10MgO A230-10CaO A230-20CaO
2θ=27.5 20
40
60
80
20
2θ (?)
40
60
80
2θ (0)
Fig. 2. XRD pattern of slag
3.2 Effect of additives on the microstructure of slag According to our previous studies, Fe was partially reduced to metallic iron under strongly reducing atmosphere, which cause inhomogeneous in molten slag, hence we observed the microstructure of slag under H2/CO2=90/10 gas composition. Fig. 3 shows
-6-
the microstructures of the A230-10MgO slag, there are many bright particles with different sizes disperse in the slag, which corresponds to iron droplet. The other slag samples also show the same appearance.
Fig. 3. BSE image of slag under different gas composition To gain further insight into the effect of MgO /CaO additives on microstructure characteristic of the slag, we used CCSEM analysis to determine the diameter and position of the Fe precipitation, and the results are plotted in Fig. 4. The number of iron precipitation in the HM-10CaO, HM-20CaO and HM-10MgO slag are 3.6×105, 1.0×105, 2.9×105/ m2 and the diameters of iron particles (Dp) are of different sizes. For
40
%
30
A230-10CaO A230-10MgO A230-20CaO
20 10 0 0.0 0.5 1.0 1.5 2.0 2.5 3.0 Dp
-7-
Fig. 4. Dp distribution of iron precipitation A230-10CaO, the variation of the Dp of Fe particles is the biggest, ranging from 0.2 to 2.0 μm with the majority of the particles settling in 0.4-0.7 μm. In the case of A230-20CaO, the Fe particles formed exhibit a sharp peak around 0.2 μm. By contrast, A230-10MgO has a narrowest particle distribution, the diameter of particles are mostly less than 0.5 μm.
3.3 Effect of additives on elution properties of slag 3.3.1 Leaching test Tables 3-5 show the concentration of heavy metals in leachant from slag samples by water leaching and seawater leaching, respectively. The result indicate that, for A232 slag, regardless the slag generated from mildly reducing atmosphere or strongly reducing atmosphere, the concentrations of Hg, Cd, Pb, Cr (Cr6+), As and Se in leachant are low enough to meet the Japanese soil conservation regulations in water/seawater leaching. The same results were observed in A230-10MgO and A230-20CaO slag leaching test. The results of Microtox test of water and seawater leachant was shown in Table 5, which indicate the MgO or CaO modified slag is environment friendly under normal leaching condition. Table 2. Concentrations of the heavy metals in leachates from A232 slag samples by water leaching (mg/L) Inlet Gas:CO2/H2 =50/50 Inlet Gas:CO2/H2 =10/90 Water Sea water Water Sea water Hg <0.0005 <0.0005 <0.0005 <0.0005 Cd <0.01 <0.01 <0.01 <0.01 Pb <0.01 <0.01 <0.01 <0.01 Cr6+ <0.02 <0.02 <0.02 <0.02 Cr <0.02 <0.02 <0.02 <0.02 As <0.01 <0.01 <0.01 <0.01 Se <0.01 <0.01 <0.01 <0.01 pH 6.9 8.1 7.1 8.1
-8-
Table 3. Concentrations of the heavy metals in leachates from A230-10MgO slag samples by water leaching (mg/L) Inlet Gas:CO2/H2 =50/50 Inlet Gas:CO2/H2 =10/90 Water Sea water Water Sea water Hg <0.0005 <0.0005 <0.0005 <0.0005 Cd <0.01 <0.01 <0.01 <0.01 Pb <0.01 <0.01 <0.01 <0.01 Cr6+ <0.02 <0.02 <0.02 <0.02 Cr <0.02 <0.02 <0.02 <0.02 As <0.01 <0.01 <0.01 <0.01 Se <0.01 <0.01 <0.01 <0.01 pH 6.8 8.0 6.8 8.0 Table 4. Concentrations of the heavy metals in leachates from A230-20CaO slag samples by water leaching (mg/L) Inlet Gas:CO2/H2 =50/50 Inlet Gas:CO2/H2 =10/90 Water Sea water Water Sea water Hg <0.0005 <0.0005 <0.0005 <0.0005 Cd <0.01 <0.01 <0.01 <0.01 Pb <0.01 <0.01 <0.01 <0.01 Cr6+ <0.02 <0.02 <0.02 <0.02 Cr <0.02 <0.02 <0.02 <0.02 As <0.01 <0.01 <0.01 <0.01 Se <0.01 <0.01 <0.01 <0.01 pH 6.8 8.0 6.8 8.0 Table 5. Toxicity effect of leachates on photobacterium phosphoreum Inlet Gas: CO /H =50/50 2
A232
2
A230-10MgO A230-20CaO
water N.D N.D seawater N.D N.D (* Note: N.D means not detected.)
N.D N.D
Inlet Gas: CO /H =10/90 2
A230 A230-10MgO
N.D N.D
N.D N.D
2
A230-20CaO
N.D N.D
3.3.2 Acid resistance With deterioration of atmosphere, some places were significant impact by acid rain. The acid resistance of slag samples was investigated prior to their potentially utilization, such as landfill. Table 6 shows the dissolution ratio of slag samples in 1M/HCl. It can be seen that slag generated in mildly reducing atmosphere is more stable than that of
-9-
strongly reducing atmosphere. MgO modified slag exhibited the better acid resistance than that of CaO. In the case for A230-20CaO, the slag is depolymerized completely, the heavy metals almost dissolved in acid with slag thoroughly. Table 6. Dissolution of slag samples in 1M/HCl (%) A232
A230-10MgO
A230-10CaO
A230-20CaO
(A) Inlet Gas: CO H /=10/90
70.4
64.2
77.4
99.6
(B) Inlet Gas: CO /H =50/50
61.2
58.3
72.3
99.3
Dissolved Ratio=(A)/(B)
1.15
1.10
1.07
1.00
2
2
2
2
The relations of dissolved ratio of metals with slag were plot in Figs 5-7, they are very different from each other. Si, Al, Ca, Mg are the major component in the slag, their dissolved ratio under mildly/strongly reducing atmosphere is consisting with the dissolved ratio of slag. On the other hand, Fe and heavy metals, their dissolved ratio is different from slag. We assumed that the slag is uniform. In the case of CO2/H2=50/50 slag, the metal are mainly exist in the form of ion, they were dissolved together with glass matrix. By contrast, in CO2/H2=10/90 condition, Fe and heavy metals might be reduced to element and separated from the slag matrix, their dissolution ration in 1M/HCl is varied. Fe, Cr, As and Zn can be digested by 1M/HCl under heating condition. Their dissolved ratio is high than that of slag. While Ni, Cu, Cd, Co can’t dissolved in 1M/HCl, hence their dissolved ratio is lower that of slag. MgO exhibit better ability than CaO on suppressing metal reduction.
4. Conclusions This investigation examined the effect of additives, including CaO and MgO, on the properties of slag in several typical gasification atmospheres at 1773K. It was found that, with the addition of MgO or CaO, SiO2 network of slag is modified significantly. The results of the leaching test of heavy metals from the slag and the ecotoxicity assay in water/seawater meet the safety regulation.In the case of 1M/HCl dissolution, slag generated in mildly gas composition is more stable under than that of strongly
- 10 -
atmosphere. MgO modified slag exhibited the better ability of acid resistance than that of CaO. In strongly reducing atmosphere, iron and some heavy metals were separated
Dissolved ratio
CO2/H2=10/90 CO2/H2=50/50
from glass matrix, and their dissolution in 1M/HCl solution is different.
1.4
A232 Fe
1.2
Si Al Ca Mg
Cr As
Zn
1.15
Dissolved ratio of slag
1.0
Ni Cu Cd Co
0.8
Si Al Ca Mg Fe Cr As Zn Ni Cu Cd Co
Fig. 5. Dissolution of metal under different gas compositions
- 11 -
Dissolved ratio
A230-10caO A230-10MgO
2.0
Mildly reducing atmosphere Ca
1.5 Fe
Cr
As
Zn
Ni
1.0
Cu
Cd Co
1.20
Dissolved ratio of slag
Mg
0.5
0.0
Fe
Mg
Ca
Cr
As
Zn
Ni
Cu
Cd
Co
Fig. 6. Dissolution of heavy metal under mildly gas compositions
A230-10caO Dissolved ratio A230-10MgO
2.0
Cr Ca
Strongly reducing atmosphere As
1.5
Zn
Fe Dissolved ratio of slag
1.24
1.0 Cd Ni
Mg
0.5
Cu Co
0.0
Fe
Mg
Ca
Cr
As
Zn
Ni
Cu
Cd
Co
Fig. 7. Dissolution of heavy metal under strongly gas compositions
- 12 -
Acknowledgements The authors gratefully acknowledge financial support from the Ministry of Economy, Trade and Industry, Japan.
References [1] Clayton SJ, Stiegel GJ, Wimer JG. Gasification markets and technologies —present and future. An Industry Perspective. A U.S. department of energy report, DOE/FE-0447,2002. [2] Minchener, A.J., 2005. Coal gasification for advanced power generation. Fuel 84, 2222-2235. [3] Choudhry, V., Hadley, S.R., 1992. Utilization of coal gasification slag. Clean energy from waste and coal, (M. Rashid Khan, eds), American Chemical Society. [4] Perrry, R.T., Salter, J.A., Baker, D.C. and Potter, M.W., 1990. Environmental characterisation of the Shell coal gasification process. III. Solid by-products. 7th Annual International Pittsburgh Coal Conference September. Washington DC, pp.253-263. [5] Brink, H.M., Eenkhoorn, S., Hamburg, G. 1996. A mechanistic study of the formation of slags from iron-rich coals. Fuel 75, 952-958. [6] Hashimoto T., Sakamoto K., Kitagawa Y., Hyakutake Y., Setani N., 2009. Development of IGCC commercial plant with air-blown gasifier. Mitsubishi Heavy Industries Technical Review 46, 1-5. [7] Huggins, F.E., Kosmack, D.A., Huffman, G.P., 1981. Correlation between ash-fusion temperatures and ternary equilibrium phase diagrams. Fuel 60, 577-584. [8] Huffman G.P., Huggins F.E., Dunmyre G.R., 1981. Investigation of the high-temperature behaviour of coal ash in reducing and oxidizing atmospheres. Fuel 60, 585–597. [9] Bryant, G.W., Lucas, J.A., Gupta, S.K., Wall, T.F., 1998. Use of thermomechanical analysis to quantify the flux additions necessary for slag flow in slagging gasifiers fired with coal. Energy Fuels 12, 257–261. [10] Öhman, M., Boström D., Nordin, A., Hedman, H., 2004. Effect of kaolin and limestone addition on slag formation during combustion of wood pellet in small scale pellet appliances. Energy & Fuels 18, 1370-1376. - 13 -
[11] Song W., Tang L., Zhu X., Wu Y., Zhu Z., Koyama S., 2010. Flow properties and rheology of slag from coal gasification. Fuel 89, 1709–1715. [12] Zhang, L, Sato, A., Ninomiya, Y., 2002. CCSEM analysis of ash from combustion of coal added with limestone. Fuel 81, 1499-1508. [13] Minister of Environmental, 1991. Environmental Quality Standards for Soil Pollution. Notification NO. 46 of the Environment Agency, Japan. [14] Japanese Industrial Standards Committee. 1998. The testing methods for industrial waste water of Japanese Industrial Standard (JIS) K 0102, Japan. [15] Japanese Industrial Standard for Chemical Analysis of Portland Cement (JIS R 5202-1999). [16] Diamond, S., 1983. On the glass present in low-calcium and in high-calcium flyashes. Cement and Concrete Research 13, 459-464.
- 14 -
Oviedo ICCS&T 2011. Extended Abstract
Interpreting Coal Conversion Under Elevated H2 Pressures With FLASHCHAIN® And CBK Stephen Niksa Niksa Energy Associates LLC, 1745 Terrace Dr., Belmont, CA, USA 94002
[email protected] Abstract Rapid devolatilization at heating rates faster than 103°C/s does not provide enough time for appreciable bridge hydrogenation so tar yields are not appreciably enhanced under the H2 pressures associated with entrained coal gasification technology. Conversely, tar yields under slow heating conditions are strongly enhanced. Our proposed mechanisms for bridge hydrogenation and its associated impact on fragment recombination in FLASHCHAIN® were able to depict these tendencies within the measurement uncertainties for H2 pressures to 15 MPa. Whereas tar yields from rapid devolatilization diminish for progressively higher pressures, total weight loss passes through a minimum at some H2 pressure around 1 MPa, because char hydrogasification counteracts the lower weight loss associated with lower tar yields. A single, nth-order reaction within the CBK framework gave results within the measurement uncertainties through 15 MPa H2, and accurately interpreted a database representing 32 coals of rank from lignite to anthracite; pressures to 21 MPa; heating rates from 1 to 103°C/s; temperatures from 550 to 1100°C; reaction times to 180 s; and particle diameters to 1 mm. The assigned hydrogasification reactivities are far less variable than those for conventional char gasification, and show no consistent trend with rank. 1.
Introduction
Throughout the 1980s, coal hydropyrolysis was aggressively developed as a technological response to the OPEC oil shocks, as a means to produce chemical feedstocks from coal. The coal research community conducted extensive lab- and pilotscale testing that characterized the impact of elevated H2 pressures on the conversion chemistry of coals across the rank spectrum. Even though the interest in hydropyrolysis technology has waned, the hydropyrolysis database is directly relevant to coal gasification at moderate temperatures, particularly when gaseous hydrocarbons (GHCs) in the synthesis gas are promoted to boost calorific values. Such processes often sustain H2 pressures approaching 1 MPa, which is high enough to enhance primary tar yields, hydrogenate tar into oils and GHCs, and convert char into CH4 via hydrogasification. Submit before May 15th to
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1
Oviedo ICCS&T 2011. Extended Abstract
Yet the hydropyrolysis database has never been quantitatively interpreted to extract accurate kinetics for hydropyrolysis, secondary tar hydrogenation, and char hydrogasification and, especially, to accurately describe the rank dependences. This paper introduces interpretations for hydropyrolysis at moderate to fast heating rates based on FLASHCHAIN®[1], and for char hydrogasification based on the Carbon Burnout Kinetics (CBK) Model [2]. Following a brief review of the underlying process chemistry, we present quantitative interpretations for datasets on variations in heating rate, temperature, reaction time, H2 pressure, particle size, and coal quality. 2.
Coal Hydrogenation Chemistry
Hydrogenation chemistry during coal devolatilization enhances the yields of both tar and gas, and the enhancements are much greater under slow heating conditions than during rapid coal devolatilization.
This heating rate dependence reflects the extensive
rearrangements associated with the hydrotreatment of petroleum resids, which convert condensed aromatics into naphthene rings with subsequent scission into gaseous hydrocarbons (GHCs), especially longer aliphatics and olefins. Only very slow heating provides sufficient reaction time for such complex chemistry. For heating rates of at least 1°C/s, we propose two much less extensive roles for H2: (1) Direct hydrogenation of labile bridges in coal without any transformations of aromatic nuclei; and (2) Suppression of recombinations among the ends of mobile fragments in the condensed phase that would otherwise introduce refractory char links into the nascent char phase. These reactions process H2 absorbed into the condensed phase, either as a dissolved species in bituminous coal melts during the plastic stage or as adsorbed species on the pore surfaces of coal solids of low- or very high-rank. In either case, the levels of H2 within the condensed phase are directly related to the H2 pressure within the particle which, in turn, is determined by the transport of H2 against the outward flow of volatiles. Accordingly, a new version of FLASHCHAIN® incorporates (1) Finite-rate bridge hydrogenations that shift the selectivity of succeeding bridge conversions toward scission and away from spontaneous condensations into char links, which reduces the fragment size distribution; (2) Suppression of bimolecular recombinations that involve fragments with hydrogenated ends; (3) A transport analysis to estimate the H2 pressure within devolatilizing particles; and (4) Henrys Law to relate the internal H2 pressure and the H2 concentration in the condensed phase. Char hydrogasification is described with a special version of CBK called “CBK/G”[2]. CBK/G predicts the rate of char conversion, the char particle temperature, and the Submit before May 15th to
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2
Oviedo ICCS&T 2011. Extended Abstract
Tar Yield, daf wt. %
W eight Loss, daf w t.%
30
70
60
20
50
10 2.5
Linby HV Bit. 10s @ 700C
1 atm
2.5
25
10 20
20
40
70 15
1 atm Linby HV Bit. 10s @ 700C
10
30 1
10
100
1000
1
Heating Rate,C/s
10
100
1000
Heating Rate,C/s
Figure 1. Evaluation of (Left) weight loss and (Right) tar yields for various heating rates to 700°C with 10 s IRP under () 1 atm He and H2 pressures of (S) 2.5, (z) 10, (T) 20, and (¡) 70 atm [3].
changes in the particle diameter and bulk density as gasification proceeds, given profiles of gas temperature and radiative exchange temperature, and the pressures of H2O, CO2, CO, and H2. Since kinetics of both first- and half-order have been validated, the char hydrogasification kinetics in CBK/G are oversimplified into a single nth-order rate law in the surface concentration of H2, whose pseudo-frequency factor and apparent activation energy are specified to quantitatively interpret reported extents of char conversion over a broad domain of operating conditions. 3.
Results and Discussion
The most distinctive aspect of coal hydropyrolysis is apparent in the joint dependence on heating rate and H2 pressure in Fig. 1. The FLASHCHAIN® results correctly exhibit greater weight loss for progressively faster heating rates at the lowest pressure; neutrality at the next two intermediate pressures; and lower yields for progressively faster heating rates at both of the highest pressures.
They accurately depict tar
enhancement at both of the lower pressures; an insensitivity to pressure at the two intermediate pressures of 10 and 20 atm; and weak suppression of tar production at the highest pressure. None of the discrepancies are greater than 3 wt. %. Clearly, the proposed mechanisms accurately depict the complex joint influences of heating rate and pressure variations across the entire domain. Figure 2 presents the weight loss dynamics throughout a 10 s isothermal reaction period (IRP) at two pressures. The data were taken immediately after the heating period to Submit before May 15th to
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3
Oviedo ICCS&T 2011. Extended Abstract
50 50
2.5 atm H2
40
20
Linby HV Bit. 3 10 C/s to 700C
Weight Loss
40
40
Subbituminous, 70μm 1000 C/s, 700 C, 10 s IRP
30
30
20 20
10
10
Tar
0
0 0
2
4
6
8
10
Char Conversion, %
70 atm H2
W eight Loss & Tar Yield, daf wt.%
W eight Loss, daf wt.%
60
0 0
Isothermal Reaction Time, s
1
2
3
4
5
6
7
8
H2 Pressure, MPa
Figure 2. (Left) Evaluation of weight loss throughout a 10 s IRP after heating at 103°C/s to 700°C under () 2.5 and (z) 70 atm H2 [3]. Figure 3. (Right) Evaluation of weight loss and tar yields from subbituminous coal for heating at 103°C/s to 700°C with 10 s IRP under a range of H2 pressures [4].
700°C, then for an additional 1 and 10 s into the IRP. At both pressures, the reaction mechanisms correctly indicate that weight loss was not finished by the end of the heating period, which is due to incomplete primary devolatilization.
At both H2
pressures, only an additional 1 s was required to achieve the asymptotic volatiles yield for this temperature.
The model results are within the measurement uncertainties
throughout the 10 s IRP at 2.5 atm H2 pressure. At 70 atm H2, the results are within the measurement uncertainties for 0 and 10 s IRP, but are too low for an IRP of 1 s. Total and tar yields from a subbituminous for a broad range of H2 pressure appear in Fig. 3. The predicted tar yields and weight loss are within the experimental uncertainties throughout, except for the weight loss at the lowest pressure. However, it is hard to fathom how weight loss could diminish at the lowest test pressure in a series such as this.
Extents of char hydrogasification are appreciable at all but the lowest test
pressures, and approach 20 % for the highest H2 pressure. But this level is 6 wt. % less than the conversion of Pit. #8 under the same test conditions, and demonstrates that char hydrogasification kinetics are not necessarily faster for coals of lower rank. The temperature dependence in the char hydrogasification kinetics are clearly apparent in Fig. 4 for lignite (LY) and an hv bituminous sample (MN) for temperatures from 550 to 1000°C with 2 and 5 s IRPs after heating at 103°C/s under 7 MPa H2. The predicted
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4
Oviedo ICCS&T 2011. Extended Abstract
70
50
70
LY
40
65
MN
30
7 MPa H2, 1000C/s, 5s IRP
Pit.#8, 6.9 MPa H2 750 C/s, 1000 C, 5-20 s IRP
20
40
LY 30
20 500
MN
600
700
800
900
Char Conversion, %
50
W eight Loss, daf wt.%
W eight Loss, daf wt.%
60 60
55
10
50
0
45
1000
0
200
Temperature, C
400
600
800
1000
Particle Diameter, μm
Figure 4. (Left) Evaluation of weight loss from a lignite (LY) and hv bituminous coal (MN) for heating at 103°C/s to various temperatures with 5 s IRP under 7.0 MPa H2 [5]. Extents of char hydrogasification appear on the right ordinate. Figure 5. (Right) Evaluation of weight loss from different size cuts of Pit. #8 under 6.90 MPa H2 for heating at 750°C/s to 1000°C with IRPs from 5 – 20 s [6].
weight loss values for both coals are within the measurement uncertainties across the entire temperature range, except for the lowest temperature with the lignite. But this temperature is so low as to be unimportant in applications; consequently, the predicted extents of char hydrogasification for both coals at this temperature are insignificant. More important, both sets of predictions are based on an activation energy for char hydrogasification of 19.5 kcal/mole. This same value also accurately described the temperature dependence in the hydrogasification of a lv bituminous under the same test conditions, whose extent of char hydrogasification grew from 4 to 15 % when the temperature was increased from 650 to 1000°C. Considering the broad range of rank represented by these samples, and the associated large variation in hydrogasification rate, we fixed EH2 for all samples in this validation database and adjusted only the frequency factor to describe the varations in the hydrogasification kinetics. Primary devolatilization is unaffected by variations in particle size provided that the size is smaller than a few millimeters. But, as seen in Fig. 5, hydrogasification yields diminish for progressively larger sizes. The predictions are well within the measurement uncertainties across the entire range, although our analysis predicts that the sensitivity to size vanishes under about 100μm but this feature is not evident in the data. CBK/G
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5
Oviedo ICCS&T 2011. Extended Abstract
includes three explicit size dependences in the analysis, none of which explains the size dependence in Fig. 5. One is in the film diffusion rate, and another is in the transport rate through the internal pore system. However, since hydrogasification is relatively slow, neither transport rate ever becomes rate limiting and the effectiveness factor is set to unity. The third dependence is in an empirical expression that describes how size and density vary through a gasification history, but this factor also relaxes to a sizeindependent limit (at variable density) due to the slow hydrogasification kinetics, consistent with full accessibility to the internal surface area. We were surprised to find that the basis for the predicted size dependence in Fig. 5 is in the energy transport rates that accommodate the enthalpy requirement for hydrogasification.
Since the heat conduction rates in the analysis are inversely
proportional to the particle size, larger particles are converted at significantly cooler temperatures than smaller ones, even while the ambient conditions are the same. Consequently, the predicted extents of char hydrogasification diminish for progressively larger sizes. Our interpretation of the coal quality impacts is based on the data from Strugnell and Patrick [5, 7], who monitored weight loss, total liquids (as the sum of tar and H2O), and noncondensable gas yields from 15 candidate coals for hydropyrolysis applications, most of which were hv bituminous. For all 15 coals, weight loss was monitored after heating at 1000°C/s to 1000°C with a 2 s IRP under 7 MPa H2. Six of these coals were also tested under 7 MPa He, and three were also tested at temperatures from 550 to 1000°C with IRPs of 2 and 5 s (cf. Fig. 4). In our interpretations, the tar yields from pyrolysis were first estimated from the total liquids yields using the coal-O distributions in CO and CO2, and assuming that the tar-O fraction was 0.30, as seen in measured product distributions for numerous coals. Then the increase in the liquids yields during hydropyrolysis was attributed to tar alone to estimate the tar yields for hydropyrolysis. Parameters in FLASHCHAIN® were adjusted to fit the reported tar and total yields for the pyrolysis tests for the six coals with both pyrolysis and hydropyrolysis test data. Then the rate constant for tar hydrogenation was assigned for match the hydropyrolysis tar yield, and the rate constant for char hydrogasification was adjusted to match the measured total weight loss. These six interpretations determined a functional form for the coal quality dependence in the
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6
Oviedo ICCS&T 2011. Extended Abstract
0.40
10000 9000
7 MPa H2, 1000C/s, 1000C, 2s IRP
0.35
7 MPa H2, 1000C/s, 1000C, 2s IRP
8000 0.30 7000 6000
0.20
A8
AHY
0.25
5000 4000
0.15
3000 0.10 2000 0.05
1000
0.00
0 65
70
75
80
85
90
Carbon Content, daf wt. %
65
70
75
80
85
90
Carbon Content, daf wt.%
Figure 6. Assigned values of the frequency factors for (Left) bridge hydrogenation and (Right) hydrogasification for 15 coals tested under 7.0 MPa H2 with heating at 103°C/s to 1000°C with 2 s IRP [5].
frequency factor for bridge hydrogenation which could be applied to the other 9 coals. Finally, the frequency factor for char hydrogasification was assigned to match the predicted and measured total weight loss for the remaining coals. Since all total and tar yields from the simulations were fit to the measured values, they will not be shown. But the assigned parameter values in Fig. 6 are certainly important. The values for the bridge hydrogenation rate fall within a narrow band for ranks from lignite to hv bituminous.
For higher ranks, they surge and reach up to four times the
mean of the band for lower ranks. This abrupt transition coincides with the nearelimination of oxygen from the labile bridges in coal, which leaves them almost entirely composed of aliphatic, hydroaromatic (naphthenic), and olefinic functional groups. It seems plausible that purely hydrocarbon linking structures would be more easily hydrogenated than links that contain carboxylic acids, ethers, and esters as well. In contrast, the assigned hydrogasification reactivity values show no consistent tendencies with rank. The most remarkable feature is that the values for all but three coals lie within 25 % of a mean value. In contrast, the frequency factors for gasification in steam and CO2 vary by as much as two orders of magnitude for coals of the same rank [2], and those for char oxidation vary by up to an order of magnitude [1].
Both rates
increase for coals of progressively lower rank. However, char hydrogasification kinetics are far less variable than those for conventional char gasification, and show no consistent trend with rank. Indeed, all the values in Fig. 6 express a variation of only a
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Oviedo ICCS&T 2011. Extended Abstract
factor of four, which is relatively very modest.
The fundamentally different rank
independence for hydrogasification probably indicates that the hydrogasification surface chemistry is not coupled with the surface chemistry for gasification by steam and CO2, and that there may not be any common adsorbed intermediates or primary reaction sites. Hydrogasification rates were assigned for an additional 8 coals whose reported conversion data was significant enough to evaluate the reactivity parameter. Four gave values within the main band of values in Fig. 6 that ranged from 2450 to 3800. The remaining four assignments fell below the band, and ranged from 600 to 1865. It is conceivable that these assignments are as reliable as the other, and that the reactivity band is therefore broader than seen in 6. But we suspect that they are not, because none of these studies were reported before 1982 and these systems were subject to substantially greater measurement uncertainties than more modern systems. Hence, we tentatively expect values for the hydrogasification reactivity from 2000 to 5000 with no consistent tendency with rank. 4.
Conclusions
Elevated H2 pressures can affect both primary devolatilization and char conversion, depending on the processing conditions. However, since hydrogenation chemistry is relatively slow, its impact is governed by the time scale for particle heating. Rapid devolatilization at heating rates faster than 103°C/s does not provide enough time for appreciable bridge hydrogenation so tar yields are not appreciably enhanced under the H2 pressures associated with coal gasification technology. Conversely, tar yields under slow heating conditions are strongly enhanced. Our proposed mechanism for bridge hydrogenation and its associated impact on fragment recombination in FLASHCHAIN® were able to depict these tendencies within the measurement uncertainties for H2 pressures to 15 MPa. The predicted yield enhancements due to bridge hydrogenation were modest at even the highest test pressures for rapid heating, but became much more substantial at slower heating rates. Consequently, the proposed reaction mechanisms do correctly depict the inversion of the heating rate dependence for the greatest pressures of interest. Hydrogen fully penetrated the coal particles to achieve the same internal H2 pressure as the ambient value whenever the heating time scale was long enough to sustain bridge hydrogenation. Whereas tar yields from rapid devolatilization diminish for progressively higher pressures, weight loss passes through a minimum at some H2 pressure around 1 MPa. Higher H2 pressures promote more extensive char hydrogasification, which counteracts Submit before May 15th to
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the lower weight loss associated with lower tar yields. A single, nth-order reaction within the CBK framework gave results within the measurement uncertainties through 15 MPa H2, and accurately interpreted a database representing 32 coals of rank from lignite to anthracite; pressures to 21 MPa; heating rates from 1 to 103°C/s; temperatures from 550 to 1100°C; reaction times to 180 s; and particle diameters to 1 mm. These interpretations determined values for the bridge hydrogenation rate during devolatilization that were uniform for ranks from lignite through hv bituminous, then surged for low volatility coals.
This abrupt transition coincides with the near-
elimination of oxygen from the labile bridges in coal, which leaves them almost entirely composed of aliphatic, hydroaromatic (naphthenic), and olefinic functional groups. The assigned hydrogasification reactivities are far less variable than those for conventional char gasification, and show no consistent trend with rank. Their most remarkable feature is that the rates for all but three coals lie within 25 % of a mean value.
References [1] Niksa S, Liu GS, Hurt RH: Coal conversion submodels for design applications at elevated pressures. Part I. Devolatilization and char oxidation, Prog. Energy Combust. Sci., 2003, 29(5):425-477. [2] Liu GS, Niksa S: Coal conversion submodels for design applications at elevated pressures. Part II. Char Gasification, Prog. Energy Combust. Sci., 2004, 30(6):697-717. [3] Guell A J, Kandiyoti R. Development of a gas-sweep facility for the direct capture of pyrolysis tars in a variable heating rate high-pressure wire mesh reactor, Energy Fuels, 1993, 7:943. [4] Guell AJ, Wu R, Li C-Z, Madrali ES, Dugwell DR, Kandiyoti R. Effect of H2pressure on yields and structures of liquids from the hydropyrolysis of maceral concentrates, Fuel Process. Technol., 1993, 36, 327-32. [5] Strugnell B, Patrick JW. Hydropyrolysis yields in relation to coal properties, Fuel, 1995, 74, 481-86. [6] Anthony DB, Howard JB, Hottel HC, Meissner MP. Fifteenth Symposium (International) on Combustion, Combustion Institute, Pittsburgh, PA, 1975, p. 1303. [7] Strugnell B, Patrick JW. Rapid hydropyrolysis studies on coal and maceral concentrates, Fuel, 1996, 75, 300-06.
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Development of a Zero-emission Coal-to-Liquids Plant W. Atcheson, K. Myers, L. O’Sullivan, H. Schobert Department of Energy and Mineral Engineering and EMS Energy Institute, The Pennsylvania State University, University Park, PA 16802 USA
[email protected]
Abstract Our university has developed a process for production of a coal-based middle distillate liquid that has potential applications as a jet fuel, diesel fuel, or feed to solid oxide fuel cells. The core process involves solvent extraction of bituminous coals, followed by two-stage hydrotreating and fractionation to obtain the desired products. Our current focus is to augment the core process with operations designed to achieve a zero-emission coal-to-liquids plant. These operations include electrolytic production of hydrogen using the Cu-Cl cycle or solid oxide electrolysis cells; and algal capture of CO2 followed by cogasification of spent algae with coal.
1. Introduction From 1989 to 2009, Penn State University was involved in developing a coal-based jet fuel. The initial emphasis was on formulating, producing, and testing a fuel that could be used not only for propulsion energy but also for thermal management on board aircraft. This fuel, JP-900, should resist thermal stressing for two hours at 480°C (i.e., 900°F). This goal was met [1]. Successful testing of prototype JP-900 demonstrated its applicability to turboshaft [1] and turbojet engines. Additional tests showed its potential as a coal-based drop-in replacement for Jet A/A-1 and JP-8 [1], as well as applications in small diesel engines and as a direct feedstock to solid oxide fuel cells [2]. Therefore, JP-900 represents a coal-based material that, potentially could replace petroleum-derived middle distillate fuels in numerous applications. New, grassroots coal-to-liquids plants require substantial capital investment (well in excess of USD 100,000 per daily barrel of capacity) and long times to completion. The JP-900 project at Penn State developed the terms coalderived to designate a fuel produced entirely from coal, and coal-based to mean
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a fuel that would contain a substantial contribution from coal, but also incorporate other hydrocarbon resources—likely petroleum but also bioliquids— as part of the fuel. This concept was developed so that desirable components from coal could be blended with petroleum-derived materials for eventual processing in conventional unit operations of petroleum refining. A coal handling and processing “front end” could be retrofitted to existing refinery infrastructure. This should allow coal to contribute to total liquid fuel production, at much lower capital investment and time to completion than for a grassroots CTL plant. An additional advantage is that it is not necessary for the two major sections of a plant, i.e. obtaining primary liquids from coal and then the refining the coalpetroleum blend stream into finished products, to be co-located. Of course, the option remains to construct an integrated grassroots facility for liquids production without being tied physically or contractually to an existing refinery. Coal-to-liquids plants have been operated successfully, at least from an engineering perspective, since the 1930s. In recent years, concern has grown about their environmental impact. This is especially true with respect to CO2 emissions, but also applies to, e.g., sulfur, ash, and mercury. As an example, Fischer-Tropsch plants are considered to have CO2 emissions roughly double those of a petroleum refinery of the same daily capacity [3]. The process developed at Penn State is described in the next section. It was developed in part in an effort to reduce capital costs and time to completion for production of clean middle-distillate fuels that would contain a significant (i.e., at least 50%) contribution from coal. Currently, we are examining strategies for making this process as close to a zero-emission operation as possible. This paper reports the overall concept and some of the initial steps in process development. What is reported here is very much a “work in progress” and further results will be reported in due course.
2. Experimental and Design Section The core process. The key coal conversion step is solvent extraction of coal using petroleum-derived light cycle oil as the solvent. The preferred conditions are 360°C, 10:1 solvent:coal ratio, and 1 hour residence time. So far, our work has been limited to testing of bituminous coals of the Illinois and
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Appalachian Basins, USA. Colleagues at North-West University in South Africa are working along similar lines [4,5]. Solvent extraction is followed by solid-liquid separation, where currently the preferred approach is pressure filtration, followed by solvent recovery. Extraction uses a 10:1 solvent:coal ratio, but with recovering and recycling most of the solvent, the liquid being processed further would be a 1:1 solvent:coal extract blend. The extract yield is typically ≈50% on a dry, ash-free basis, but the best cases provide a yield of ≈70%. Conversion of the solvent–coal extract to finished fuels involves two steps of hydrotreating. The first uses a Criterion Syncat-3 catalyst at 315-370°C and 4 MPa. The primary purpose is hydrodesulfurization, accompanied by some hydrodenitrogenation and ring saturation. The second stage uses an Engelhard Redar catalyst at 175-360°C and 4 MPa to fully saturate aromatic rings. Finally, distillation separates the hydrotreated liquid into mainly jet fuel, diesel fuel, and light fuel oil, with smaller amounts of gasoline and heavy fuel oils. Using bestcase extraction results, and assuming a feedstock of 10% moisture and 10% ash yield, and final middle distillate product of 0.85 specific gravity (35° API), the yield is close to 4 barrels per as-mined ton of coal. Figure 1 provides a block-flow diagram of the major operations.
Figure 1. Core process for production of coal-based middle distillate fuels by solvent extraction and two-stage hydrotreating. As depicted in Figure 1, the process has no emission-control operations. Further,
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the source of hydrogen is not specified. In an integrated plant with abundant coal availability, a likely choice would be production of hydrogen by coal gasification followed by water-gas shift and absorption of CO2. Other sources of CO2 include fired heaters used throughout the process, and possibly some thermal decomposition of the coal during extraction. A second product of concern is the residual solid separated downstream of the extraction. This solid will include unextracted (or partially extracted) coal and mineral matter, and might be wet with process solvent and/or extract. H2S will be produced as part of the hydrodesulfurization operation downstream of the solvent recovery operation. Hydrogen production. The best way to reduce the carbon footprint is to generate less CO2 in the first place, rather than shifting the entire burden of CO2 reduction to a carbon capture and storage (CCS) system. We consider H2 production in which the majority of hydrogen needed for hydrodesulfurization and ring saturation would be generated from “non-carbon” sources. At present, we favor electrolysis of water, using “non-carbon” electricity, e.g. electricity generation from wind, solar, or nuclear sources. A second possibility is water splitting via concentrated solar power (CSP). However, there remains the issue of disposal of residual solids, where one option is to gasify it for supplemental H2 production, reducing the amount of solid waste and increasing H2 production at the same time. Further, using heat sources other than the carbon-oxygen reaction for providing the necessary energy to drive the endothermic carbon-steam reaction may reduce CO2 production from the gasification. The options in this case include CSP or use of waste heat from a co-located nuclear electric station. An alternative approach to H2 production involves the Cu-Cl thermochemical cycle [6]. This cycle decomposes water at 400-600°, with good efficiency (≈40%). This temperature range is attractive because it offers the potential for taking advantage of CSP or waste heat from nuclear reactors. Further research needs to focus on finding cell membranes that are sturdy yet have low permeability to copper. Additional parametric studies of reagent concentrations, flow rates, and temperatures should help in improving the rate of H2 production. Collateral work has explored solid oxide electrolytic cells, which also operate at high temperature [7]. Preliminary studies show that cell efficiency
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and hydrogen production appear to increase at a higher-than-linear rate with temperature. Figure 2 illustrates a block flow diagram for a possible hydrogen production section of the plant.
Figure 2. Block flow diagram for hydrogen production section of conceptual process. Residual solid. An option for dealing with residual solid is gasification for supplemental hydrogen production. With appropriate design and operating conditions, mineral matter could be converted to slag and, in turn, to a vitrified glassy material for disposal or possible commercial use (e.g. for production of fiberglass-like material for low-cost insulation). Alternatively, recovered residual solid could be used as a low-grade fuel in a facility appropriately permitted to acquire and burn waste coal. While this option provides additional revenue, it shifts the issue of CO2 production to some other, off-site location. If the solid is wet with solvent and/or extract, additional operations for washing with a light solvent, and recovering and recycling the wash solvent, would be needed. At the current stage of process design, we favor the gasification option. Process heat. Even with major changes in the H2 production section to reduce the carbon footprint, other sources of CO2 will exist. Most of the process operations take place at elevated temperature, with CO2 generation from fired
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heaters. At present we consider that this CO2 would need to be dealt with in a CCS section (discussed below). “Non-carbon” alternatives for process heat include CPS, waste heat from a nuclear reactor, or possibly the choice of hydrogen as fuel for process heaters. Sulfur capture. Hydrodesulfurization will produce H2S. Numerous commercial processes exist for H2S removal and destruction [8], the Claus process being an example. Given the availability of tested, successful processes, we presume that one of these could be licensed and used without the need for additional research or development in this area. Carbon capture and storage. At the present stage of design we consider that CO2 will inevitably produced in the plant. The current plan is to utilize algal photobioreactors for capturing CO2. The use of algae in this application is under intensive investigation in many laboratories and pilot plants around the world [e.g., 9,10]. This approach has at least three advantages. First, under optimum circumstances, some species of algae can absorb large amounts of CO2 in short times, doubling body weight in about 24 hours. Second, collecting and processing the algae provides a potential source of bio-oil, which could be used as a blend stock with the coal-petroleum products. Third, the residual algal material remaining after processing for bio-oil could be co-fed to the gasifier used for treatment of the residual coal extraction solids. Figure 3 provides a block flow diagram of the emission-control section of the process. In the longer term, the possibility of photocatalytic reduction of CO2 offers an alternative approach to CCS [11]. This technology does not seem to be as thoroughly studied nor so far advanced as algae-based CCS. However, in such reactors CO2 could be converted to useful products such as methanol.
Results and Discussion Some of the properties of prototype JP-900 fuel are provided in Table 1. Table 1. Selected characteristics of prototype JP-900 produced using refined chemical oil as a surrogate for coal extract [1]. Aromatics, % Calorific val., MJ/kg Flash point, °C Freezing point, °C
1.9 42.67 61 –67
Smoke point, mm Sulfur, % Viscosity, –20°C, cSt
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22 0.0003 7.5
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Oviedo ICCS&T 2011. Extended Abstract
This fuel was produced using two-stage hydrotreating described above. Because the solvent-extraction process was still under development at laboratory-bench scale at the time this hydrotreating was done at pilot scale, we used a surrogate material, refined chemical oil, to simulate the expected coal extract. Refined chemical oil is a distillation product from coal tar processing in metallurgical coke plants, and consists primarily of naphthalene, indane, and their alkylated derivatives.
Figure 3. Block flow diagram for CCS section of zero-emission plant.
4. Conclusions Figure 4 presents a block flow diagram for the entire zero-emission process at its present state of design and development. The inputs to the process are coal, water, and make-up solvent, as well as electricity for hydrogen production. The outputs are clean middle-distillate fuels, along with by-product sulfur and a vitrified ash or slag. Both by-products have potential market value. The prototype fuel has been tested successfully in aviation gas turbine,
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turboshaft, and small diesel engines, as well as in a solid oxide fuel cell. A current focus of our work is to determine how best to integrate the hydrogen production, CCS, and other operations
Figure 4. Layout of conceptual zero-emission plant. to work toward the zero-emission concept. Laboratory work is investigating cogasification of Kentucky bituminous coal with Chlorella vulgaris alga in an externally heated drop-tube reactor, and hydrogen production via the Cu-Cl thermochemical cycle and in solid oxide electrolysis cells. We seek to investigate the electrolysis of aqueous solutions of CO2, which could be reduced to CO and provide a useful CO-H2 gaseous product. We are beginning a study of technology assessment coupled with an economic and financial analysis of this process. These results will be published in due course.
Acknowledgements The present work was facilitated by seed funding from Professor Yaw Yeboah of the John and Willie Leone Family Department of Energy and Mineral Engineering, access to electrolysis equipment provided by Professor Serguei Lvov, coal samples from Gary Mitchell, help in numerous ways from Dr. Sharon
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Miller.
References [1] Balster LM, Corporan E, DeWitt MJ, Edwards JT, Ervin JS, Graham JL et al. Development of an advanced, thermally-stable coal-based jet fuel. Fuel Proc. Technol. 2008;89:364-78. [2] Zhou ZF, Gallo C, Pague MB, Schobert HH, Lvov SN. Direct oxidation of jet fuels and Pennsylvania crude oil in a solid oxide fuel cell. J. Power Sources 2004;133:181-7. [3] Ramage, M.P.; Tilman, G.D. Liquid transportation fuels from coal and biomass. Washington: National Academies Press; 2009; and numerous references therein. [4] Janse van Resnburg E. Solvent extraction of a South African coal using a low volatile, coal derived solvent. MS Thesis. Potchefstroom: North-West University; 2007. [5] Makgato H, Janse van Rensburg E, Neomagus HWJP, Schobert HH. Extraction of South African vitrinite-rich coal using a low-volatile, coal-derived residue oil. Proceedings South African Chemical Institute annual meeting; 2008. [6] Balashov VN, Schatiz, RS, Chalkova E, Akinfiev, NN, Fedkin MV, Lvov SN. CuCl electrolysis for hydrogen production in the Cu-Cl thermochemical cycle. J. Electrochem. Soc. 2011;266-75. [7] Stuart P, Unno T, Kilmer J, Skinner S. Solid oxide proton conducting steam electrolyzers. Sol. State Ionics 2008;179:1120-4. [8] Kohl AL, Nielsen RB. Gas purification. Amsterdam: Elsevier; 1997. [9] Packer M. Algal capture of carbon dioxide; biomass generation as a tool for greenhouse gas mitigation with reference to New Zealand energy strategy and policy. Energy Policy, 2009;37:3428-37. [10] Verma NM, Mehrotra S, Shukla A, Mishra BN. Prospective of biodiesel production utilizing microalgae as the cell factories: a comprehensive discussion. African J. Biotech. 2010;9:1402-11. [11] Indrakanti VP, Kubicki JD, Schobert HH. Photoinduced activation of CO2 on Ti-based heterogeneous catalysts: current state, chemical physics based insights, and activation. Energy Enviro. Sci. 2009;2:745-58.
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DOOSAN POWER SYSTEMS DEVELOPMENTS
OXYCOAL™ TECHNOLOGY
MD Maloney, B Dhungel, DW Sturgeon, P Holland-Lloyd Doosan
Power
Systems,
Porterfield
Road,
Renfrew,
PA4
8DJ,
UK
e-mail:
[email protected]
Abstract Doosan Power Systems (DPS) is committed to delivering a low carbon future through ongoing development, demonstration and deployment of advanced carbon abatement technologies for economically viable power generation. DPS has adopted an evolutionary approach to the design and implementation of OxyCoal™ technology: combining well proven and commercially available components with state-of-the-art technology advances. The “Demonstration of an Oxyfuel Combustion System” project, led by DPS and supported by funding from the UK Government Department of Energy and Climate Change, led to design, manufacture and testing of a 40MWt OxyCoal™ burner. The operational envelope of the 40MWt OxyCoal™ burner: start-up; shutdown; turndown and transition from air to oxyfuel firing; was safely demonstrated on DPS Clean Combustion Test Facility (CCTF). Further parametric testing investigated performance, in terms of: flame shape; flame stability; combustion and thermal performance; at realistic operating conditions. This paper will summarise the results from the 40MWt OxyCoal™ burner trials; our continuing development activities and our current OxyCoal™ commercial power generation activities in Asia and Europe. 1. Introduction Doosan Power Systems (DPS) designs, supplies and constructs advanced steam generation technology for the power industry and develops some of the cleanest, most efficient coal-fired power plant in the world. To this end, DPS is committed to delivering unique and advanced carbon capture solutions. There are three main pathways to the capture of CO2 from coal-fired power generation i.e. post combustion capture, oxyfuel combustion and pre-combustion capture. DPS offers post-combustion carbon capture technology and OxyCoal™ technology, our implementation of oxyfuel combustion technology for coal-fired utility plant. Doosan Heavy Industries (DHI) offer
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integrated gasification combined cycle technology. Oxyfuel firing involves the combustion of fuel in a medium comprising injected oxygen plus recycled flue gas and offers a means of generating CO2-rich flue gas requiring minimal treatment prior to storage or beneficial application (such as Enhanced Oil Recovery). DPS has adopted an evolutionary approach to the design and implementation of OxyCoal™ technology combining well proven and commercially available components with state-of-the-art advances in technology. DPS has completed a number of UK Government
supported
collaborative
projects
developing
oxyfuel
combustion
technology, including: Investigation of oxyfuel combustion fundamentals and underpinning technologies; Demonstration of an oxyfuel combustion system; and Impact of High Concentrations of SO2 and SO3 in Carbon Capture Applications and its Mitigation The results from the investigation of oxyfuel combustion fundamentals and underpinning technologies project have been presented previously [1, 2]. This paper presents the results of the full scale demonstration of an oxyfuel burner and pilot scale work to investigate SO2 and SO3 in carbon capture applications and their mitigation. 2. Demonstration of an Oxyfuel Combustion System Oxyfuel combustion represents one of the more promising of the technologies currently being developed for CO2 capture. The global market for CO2 capture equipment is likely to be considerable, and it is strategically important for power plant operators and equipment manufacturers to have a developed product within a timescale consistent with the market. Full scale oxyfuel burner testing on Doosan Power Systems’ Clean Combustion Test Facility (CCTF) in Renfrew, Scotland represents an important step towards commercial implementation of this technology. A key element in the development of oxyfuel combustion technology is the availability of appropriate burner technology. Therefore the objective of the OxyCoal 2 project was to design and demonstrate an OxyCoal™ burner of a type and size applicable to new build and retrofit advanced supercritical oxyfuel plant. The project had the following specific aims: •
Demonstrate successful performance of a full-scale (40MWt) oxyfuel burner firing at conditions pertinent to the application of an oxyfuel combustion process in a utility power generating plant
•
Demonstrate performance of an oxyfuel burner with respect to flame stability, NOX,
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flame shape and heat transfer characteristics •
Demonstrate the operational envelope of an oxyfuel burner with respect to flame stability, turndown, start-up, shutdown and the transition between air- and oxyfuelfiring
•
Demonstrate safe operation of an oxyfuel combustion process under realistic operating conditions
•
Generate sufficient oxyfuel combustion process performance data to inform future investment decisions
•
Demonstrate the level of technology readiness of the oxyfuel combustion process
2.1. 90MWt Clean Combustion Test Facility The 90MWt CCTF is designed primarily for the development of burners for fossil fuel firing applications and is one of the largest and most modern single burner test rigs in the world. The plant has been designed to enable burners to be developed, optimised and performance tested at full-scale prior to application in utility power plant or industrial furnaces. An overfire air system is installed which provides the ability to test two stage combustion (TSC) burners at realistic plant conditions, down to a primary zone stoichiometry of 0.75 at up to 70MWt. TSC air is supplied through any one of three injection locations, each comprising of two ports angled towards the centreline of the furnace. For OxyCoal™ burner testing, the CCTF has been upgraded with the addition of equipment and instrumentation required for safe oxyfuel firing. This included the addition of an oxygen storage facility, comprising three liquid oxygen storage tanks each with a capacity of approximately 50 tonnes and eight ambient vaporizers to supply gaseous oxygen for injection into primary and secondary flue gas recycle (FGR) streams. The primary and secondary FGR streams replace the primary air and main combustion air respectively, each having a dedicated fan. A transport FGR stream replaces the transport air stream. The transport and primary FGR streams also have additional flue gas cooling systems fitted to condense moisture, followed by in-duct heating as a means to mitigate potential PF feeding problems. A proportion of the secondary FGR stream can be redirected as overfire FGR for two-stage combustion. 2.2. 40MWt OxyCoal™ Burner The Doosan Power Systems’ 40MWt OxyCoal™ burner is an extrapolation of current low NOX air-fired burner technology which is installed in coal fired utility plants
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worldwide with many years of operating experience, to ensure compatibility with existing plant for retrofit purposes. A requirement of the 40MWt OxyCoal™ burner is the ability to operate under both conventional air firing and oxyfuel firing conditions at full burner load. The first generation 40MWt OxyCoal™ burner uses a uniform ‘simulated air’ flue gas composition, where the molar oxygen content of the primary gas is maintained as per primary air, i.e. 21%v/v. The overall stoichiometric ratio and flue gas recycle rate have been chosen so that adiabatic flame temperatures, radiant and convective heat transfer are theoretically equivalent to air firing. Consequently the 40MWt OxyCoal™ burner was designed, with the aid of CFD modelling, to best exploit a range of potential operating conditions for both air and oxyfuel firing. 2.3. Experimental Results for 40MWt OxyCoal™ Burner on the 90MWt CCTF 2.3.1. Isothermal Testing Isothermal testing was undertaken to investigate the burner aerodynamics in terms of the swirl number and pressure drop (k-factor) characteristics of the secondary and tertiary air streams. The measurement of velocity profiles supported computational fluid dynamics (CFD) modelling activities. 2.3.2. Burner Proving To prove the full-scale 40MWt OxyCoal™ burner firing bituminous coal in both air and oxyfuel firing operation in terms of flame stability and control and operability, combustion tests were undertaken and completed by January 2010. Kellingley coal was used for much of the test programme, a UK bituminous coal widely used in UK power stations and has been used on the CCTF to develop, optimize and performance test low NOX air-fired burners. El Cerrejón coal, a Columbian bituminous coal, was also used as it is a widely available world traded coal. For flame stability, ignition should be within the burner throat/quarl zone without the use of support fuel (e.g. oil). A furnace side wall camera was used to continuously monitor the flame root location and general flame shape during both air firing and oxyfuel firing operation. The observation ports arranged along one side wall of the furnace on the burner centreline were used for visual checks of flame length. The 40MWt OxyCoal™ burner at full load has the same flame length for both air and oxyfuel firing. At steady state conditions the flames were well rooted to the flameholder and similar in shape across a wide range of operating conditions for both air and oxyfuel
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Oviedo ICCS&T 2011. Extended Abstract
firing. During burner proving several methodologies for the transition between firing modes were investigated. A preferred methodology for the transition on the CCTF was identified, but all of the methods tested are practically applicable to oxyfuel boiler plant. 2.3.3. Air Ingress During initial oxyfuel trials, low CO2 flue gas concentrations of around 50%v/v dry were achieved and it was therefore necessary to investigate air ingress by means of an oxygen survey. Subsequently the sources of air ingress were identified and CO2 flue gas concentrations in excess of 75%v/v dry, and up to 85%v/v dry, were achieved by reducing the air in-leakage. Hence, consideration will need to be given to the potential sources of air ingress on a utility power plant (e.g. furnace, boiler, economiser or particulate collector) and the means of minimisation. 2.3.4. Turndown The firing load of the OxyCoal™ burner was turned down in discrete stages from 100% load to 40% load - a comparable turndown to Doosan Power Systems’ commercially available air firing low NOX axial swirl burners currently operating around the world. For each load reduction increment first the coal flow (thermal input) was reduced, followed by the primary FGR flow and finally excess oxygen level. The primary FGR turndown was typical of a normal E-mill installation (i.e. the flow was decreased linearly from 100% at mill full load to 70% of the full mill load value at 50% burner load using a linear interpolation between these two points for intermediate load settings). Steady state part load tests were performed at 32MWt, 24MWt, 20MWt and 16MWt (approximately 80%, 60%, 50% and 40% burner load, respectively). As expected, flame length reduces as load is reduced, but a stable, rooted flame was maintained for all loads down to 40%. 2.3.5. Parametric Testing To characterise and investigate the achievable performance of the full-scale 40MWt OxyCoal™ burner firing bituminous coal in both air and oxyfuel firing operation in terms of emissions and thermal performance, parametric testing was undertaken and completed over approximately 3 months. Characteristic levels of CO2, NOX, SO2, CO and carbon in ash (CIA) were determined within the stable operating range of the burner. On a volumetric basis (vppm) oxyfuel firing economiser exit NO concentration is approximately trebled compared to the air firing concentration, largely due to the dilution effects of removing N2 from the combustion process. However, on a heat input
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Oviedo ICCS&T 2011. Extended Abstract
basis (mg/MJ) oxyfuel firing economiser exit NO is reduced by approximately 50% compared to air firing at comparable stoichiometry. Both oxyfuel firing and air firing show a trend of higher economiser exit NO with increasing burner stoichiometric ratio. On a volumetric basis (vppm) oxyfuel firing economiser exit SO2 concentration is also approximately trebled compared to air firing. However, on a heat input basis (mg/MJ) oxyfuel firing economiser exit SO2 is reduced by approximately 25% compared to air firing. This reduction is largely due to dissolution of SO2 in the primary flue gas recycle (FGR) direct contact cooling system and absorption of SO2 on fly ash. On a volumetric basis (vppm) oxyfuel firing economiser exit CO concentration is similar to air firing, both being typically below 200vppm dry. For oxyfuel firing it is likely that there is more dissociation of CO2 to form CO at high temperatures within the flame due to the high CO2 content of the flue gas. However, on a heat input basis (mg/MJ) oxyfuel firing economiser exit CO is approximately equal compared to air firing, which echoes findings by Hjärtstam et al. at Chalmers University of Technology [3]. Both oxyfuel firing and air firing show the widely observed trend of rapidly increasing economiser exit CO at lower burner stoichiometric ratio. Although CIA at the economiser exit appears high, particularly at low burner zone stoichiometric ratios, when the CIA is converted to unburnt loss the value is below 1%GCV, due to the low ash content of the coal. More significantly unburnt loss is comparable for oxyfuel and air firing. The CCTF has a relatively short burnout zone residence time compared to a utility furnace and the gas temperature in the burnout zone is significantly lower, experience has shown that unburnt loss is lower on full scale utility plant compared to the CCTF. The test facility furnace and boiler thermal performance, in terms of heat release and absorption, was determined. During both air firing and oxyfuel firing testing total absorbed heat flux measurements were taken at 10 locations along the side wall of the furnace at the burner centreline elevation. The measurements show that the oxyfuel flame radiates less heat to the walls in the first half of the furnace when compared to air, though similar heat fluxes were observed in the last half of the furnace. The differences in heat flux profile can be explained by a combination of the mass flow through the burner and ash concentration. Thermal performance analysis of the CCTF boiler using Doosan Power Systems’ proprietary design programs showed that the impact of oxyfuel on boiler thermal
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Oviedo ICCS&T 2011. Extended Abstract
performance was consistent with that anticipated from the boiler modelling activities. 3. Impact of SO2 and SO3 in Oxyfuel Applications It has been shown that the concentration of sulphur containing compounds (SO2 and SO3) in oxyfuel firing increases by a factor of 2 to 5 when compared to conventional airfiring. However, the understanding of SO3, in particular, is limited in oxyfuel systems. The presence of SO3 in flue gas is associated with low temperature corrosion through the condensation of sulphuric acid in the economiser, air heater or flue gas ducts. SO3 control will therefore be required for the successful implementation of oxyfuel combustion technology. Doosan Power Systems undertook experimental investigations at the 160kWt Emission Reduction Test Facility (ERTF) to investigate the impact of oxyfuel combustion on SO3, to investigate the potential for SO3 reduction by dry sorbent injection. Tests were performed during air-firing and oxyfuel combustion using a medium sulphur bituminous coal, with and without the injection of sorbents. Four different sorbents were tested, two sorbents (A&B) were injected in the furnace (post flame region) and two sorbents (C&D) were injected in the downstream flue gas duct (lower temperature region) at a range of feeding rates. To determine the SO3 reduction potential of the sorbents, SO3 measurements were carried out at the ESP inlet using a controlled condensation measurement technique. The controlled condensation sampling train can also measure SO2 and moisture content of the flue gas. 3.2. Results 3.2.1. SO3 Reduction by Sorbent Injection Prior to investigation of sorbent injection, baseline measurements were performed for air-firing and oxyfuel combustion. The SO3 measured during air-firing was ~23ppm and during oxyfuel firing was ~32ppm. The absolute values obtained are specific to the ERTF configuration and the coal fired, however the percentage reduction obtained following the injection of sorbents should be representative of plant, generally. More than 75% of the SO3 produced was seen to be reduced by all of the sorbents tested. The SO3 concentration is seen to drop from a baseline of 32ppm to less than 5ppm by the injection of most of the sorbents, but sorbent B was seen to be slightly less effective, with a drop in SO3 concentration to 7ppm. Significantly, the SO3 reduction efficiency was seen to be higher for post combustion sorbents (Sorbents C and D) as much lower mass (molar ratio) of sorbent is required to achieve an equivalent SO3 reduction
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Oviedo ICCS&T 2011. Extended Abstract
efficiency. This may be attributed to the temperature at which SO3 conversion reaches its maximum value. Literature [4, 5] indicates that SO3 formation reaches its maximum value in the post furnace region where the temperature is below 1000°C. At the location of post combustion sorbent injection, SO3 conversion via gas phase chemistry is complete. SO3, being more reactive than SO2, will therefore react with the injected sorbents preferentially, resulting in much higher SO3 reduction efficiency. Similar results in terms of SO3 reduction and sorbent utilisation were also observed for air-firing. 4. Conclusions Following conversion to oxyfuel firing, the Clean Combustion Test Facility (CCTF) became the world’s largest demonstration of an oxyfuel combustion system, and a test programme has been successfully completed to demonstrate the full-scale Doosan Power Systems 40MWt OxyCoal™ burner on air and oxyfuel firing, achieving safe and stable operation across a wide operational envelope. Safe and smooth transitions between air firing and oxyfuel firing were demonstrated, and knowledge of fan interactions during the transitions gained from operational experience. Flame stability and flame shape under air firing and oxyfuel firing conditions were comparable, and were consistent with theoretical expectations. CO2 flue gas concentrations up to 85%v/v dry were achieved by reducing air in-leakage. Turndown of the 40MWt OxyCoal™ burner was proven from full load to 40% load, the flame remaining stable and well rooted throughout. The levels of emissions of NOX and SO2 were significantly lower for oxyfuel firing, reduced by approximately 50% and 25% respectively on a heat input basis (mg/MJ) compared to air firing, and emissions of CO and unburnt loss were comparable for oxyfuel and air firing. The variation in heat flux profiles for oxyfuel and air firing were explained by differences in operating conditions and ash concentration. Overall, the 40MWt OxyCoal™ burner performed as designed and is applicable to new build and retrofit advanced supercritical oxyfuel plant. Results from the 160kWt test facility indicate that application of dry sorbent injection technology offers a viable means of low temperature corrosion mitigation. SO3 reduction efficiency by the injection of post combustion sorbent was seen to be higher than infurnace sorbent, and is preferred to in-furnace sorbent injection, as injecting the sorbents further downstream gives rise to a reduced likelihood of negative impacts e.g. slagging and fouling.
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Oviedo ICCS&T 2011. Extended Abstract
5. Moving Forward Doosan Power Systems will undertake extended endurance testing of their OxyCoal™ burner on Vattenfall’s 30MWt Schwarze Pumpe Oxyfuel Pilot Plant, in Germany. This will enable demonstration of the full chain oxyfuel process including cryogenic air separation and CO2 compression plant. This project involves Doosan Power Systems joining Vattenfall’s Oxyfuel Technology Sharing Group. Doosan Power Systems are working together with Doosan Heavy Industries to complete a FEED study for a 125MWe Oxyfuel retrofit in South East Asia. Doosan Power Systems is pursuing further demonstrations at 250MWe to 350MWe and large scale in Europe and Asia as part of the full commercial development of the OxyCoal™ technology. Acknowledgements The authors gratefully acknowledge the grant funding provided by the Department of Energy and Climate Change and the Technology Strategy Board, and the technical and financial contributions made by the project collaborators: University of Nottingham; Imperial College London; Scottish and Southern Energy plc; Air Products plc; DONG Energy Power; Drax Power Limited; EDF Energy plc; E.ON UK plc; ScottishPower Limited; Vattenfall AB; UK Coal plc; University of Leeds; and IEAGHG. The authors also gratefully acknowledge the support of their Doosan Power Systems colleagues who contributed to the success of these projects. References [1] Sturgeon DW, Cameron ED, Fitzgerald FD, Demonstration of an oxyfuel combustion system, 9th International Conference on Greenhouse Gas Control Technologies, Washington, D.C., 16-20 November 2008 [2] White V, Torrente Murciano L, Sturgeon DW and Chadwick D, Purification of OxyfuelDerived CO2, 9th International Conference on Greenhouse Gas Control Technologies, Washington, D.C., 16-20 November 2008 [3] Hjärtstam S, Andersson K, Johnsson F, Combustion characteristics of lignite-fired oxy-fuel flames, 32nd International technical conference on coal utilization and fuel systems, Clearwater, FL, 2007. [4] Nettleton MA, Stirling R, 12th International Combustion Symposium, p.635, The Combustion Institute, Pittsburgh, 1969 [5] Senior CL, Sarofim AF, Zeng T, Helble JJ, Mamani-Paco R, Fuel Processing Technology, Vol. 63, pp. 197-213, 2000
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Oviedo ICCS&T 2011. Extended Abstract
CIUDEN CO2 Technology Development Centre on Oxycombustion M Lupion*, V J Cortes, M Gomez, A Fernandez Fundacion Ciudad de la Energia (CIUDEN) II Avenida de Compostilla nº2 24400 Ponferrada. Spain 1. Introduction Carbon dioxide has been identified as the major contributor among greenhouse gases (GHG). CCS technologies have the potential to reduce CO2 emissions into the atmosphere in order to keep climate change below 2ºC. The Europe 2020 Strategy for smart, sustainable and inclusive growth includes headline targets that set out where the EU should be in 2020, one of them relates to climate and energy. This way, Member States have committed themselves to reducing GHG emissions by 20%, increasing the share of renewables in the EU's energy mix to 20%, and achieving the 20% energy efficiency target by 2020. The European Council reconfirmed in February 2011 the EU objective of reducing GHG emissions by 80-95% by 2050 compared to 1990, in the context of necessary reductions according to the Intergovernmental Panel on Climate Change [1]. Energy projections made by the World Energy Council, the International Energy Agency (IEA) and the US Energy Information Administration give similar pictures of the dominant role of fossil fuel in the future primary energy global demand and the importance of CCS as part of the portfolio of solutions to reduce GHG emissions [2 - 4]. IEA calculation puts CCS in critical role in a least-cost pathway to reaching 450 ppm scenario where the role of CCS increases after 2030 up to 19% of needed reductions against baseline in 2050 [5]. In this context, one of the current European initiatives in terms of R&D&D on CCS is the Technology Development Centre for CO2 Capture, or es.CO2 Centre, which is supported by the Spanish Government through The Fundacion Ciudad de la Energia (CIUDEN). CIUDEN is a research and development institution created by the Spanish Administration in 2006 and fully conceived for collaborative research in CCS and CCTs thus contributing to the strengthening of the industrial and technological base in Spain and by extension in Europe. CIUDEN´s main objectives within the CO2 Capture Programme are the research, development and demonstration of efficient, cost effective and reliable CCS and advanced CCT as well as third generation flue gas cleaning through the design and operation of the es.CO2 Centre [6, 7]. The es.CO2 Centre incorporates the world’s most advanced equipment for the development of capture processes through oxycombustion, one of the three options of capturing CO2. Oxycombustion technology is based on the concept of burning the fuel in oxygen rather than air in order to eliminate nitrogen from the flue gas. The resulting flue gas stream is predominantly concentrated with CO2. The concentrated CO2 stream after processing is ready for storage. CIUDEN’s current R&D&D Programme targets on oxycombustion includes: (1) validation and scaling-up of oxyPC, oxyCFB, Flue Gas Desulphurization (FGD) and CPU technologies; (2) advanced materials for oxy-firing; (3) integration and optimization tests of the full process to produce a CO2 stream ready for transport and storage, and (4) 2nd generation oxyfuel power plants. 2. Experimental section Figure 1 shows the bird's eye view of the es.CO2 Centre with indication of the main units involved. General description of the es.CO2 Centre is found elsewhere [7-10]. This paper focuses on detailed technical characteristics of the combustion island,
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Oviedo ICCS&T 2011. Extended Abstract
particularly the PC boiler, and complementary systems. Preliminary results at the Fuel Preparation Unit as well as at the PC boiler are also included.
Figure 1. es.CO2 Centre bird´s eye view 2.1 Oxidant preparation system Independent oxidant preparation systems for each stream (Figure 2) have been provided allowing the individual study of the effects of the oxygen concentration, humidity and temperature in each stream and offering high flexibility and large capabilities. Each one of the oxidant streams is composed of: (1) air (conventional combustion mode); (2) air, oxygen and recirculated flue gases (partial oxycombustion); and (3) oxygen plus recirculated flue gases (total oxy-combustion). The upper limit of oxygen concentration compatible with the materials used is situated between 21% and 40% v. Three oxidant streams have been considered in the design for the PC boiler: primary oxidant (CB1) used to introduce the pulverised coal into the furnace through the burners and to supply part of the means of oxidation necessary for combustion, and secondary (CB2) and tertiary (CB3) oxidants introduced to keep combustion progressing. Primary oxidant (CB1) for CFB is used to fluidize the bed as well distributes the solid fuel along the bed for a proper combustion. A stream derived from CB1 provides the required fluidisation necessary for the solids return from the cyclone. Transport oxidant for limestone, fuel and sand feeds and recycling of ash is also derived from CB1. Secondary oxidant (CB2) will be responsible for supplying the rest of the oxygen required for good combustion.
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Figure 2. Oxidant preparation system Combustion gases are recirculated under oxy-mode (partial/total) in order to reach the O2 concentration required by means of two different streams: - Stream FGR1: taken downstream from the FGD unit (“wet recirculation”) - Stream FGR2: taken upstream from the FGD unit. Maximum variability is possible between the ratios of primary/secondary/tertiary streams in order to investigate the stratification effects. It is well established that secondary/tertiary oxidant streams ratio variation produces significant effects in the fluid dynamics of the combustion process, and in the development of associated phenomena, such as the generation of NOx and unburnt coal, and temperature profiles distribution. 2.2 PC boiler Combustion Chamber The boiler furnace is separated into lower area and upper area. The lower area has a square section of 4.5 m width and 5.6 m height. Its bottom is a trapezoidal section for discharge and collection of ashes. The upper area has a rectangular section with dimensions of 4.5 m depth x 5 m width and 6.7 m height. The complete furnace has a heat transfer area of 200 m2 and a volume of 187 m3. The inlets of tertiary oxidant are located: (1) two inlets above each burner in two levels; (2) in the lower area of the furnace (trapezoidal area) to improve combustion and (3) in the termination area of the furnace nose. Figure 3 shows PC boiler´s structure before and after completion.
Figure 3. PC boiler structure Burners System Four horizontal or wall burners, two on each side of the furnace facing each other, have been installed. Each one has a capacity of 5 MWth and of “turbulent” type. The burners are low level NOx burners, which capability to modifying the fluid dynamic characteristics of the various primary and secondary streams making possible the study of different flame configurations. The burner system incorporates pulverised fuel receptacles, ignition electrodes, gas valves, BMS control and monitoring trains. Each one has a high voltage electric lighter and a gas lance for the ignition of the primary fuel. The specified system for burner control includes the latest generation sensors for individual flame monitoring. The burner is made up of a central nozzle from where the pulverised coal and primary oxidant mixture are injected (CB1). Surrounding this central outlet are housed the secondary oxidant outlets (CB2) that surround the central flame and complete the initiated combustion. 3
Oviedo ICCS&T 2011. Extended Abstract
Figure 4 shows a diagram of the PC boiler indicating the radiant/convective areas and the inlets of the three oxidant streams.
Figure 4. PC boiler inlet streams Extensive instrumentation has been included in the PC system. Standard fixed process instrumentation such as flame detectors or optical pyrometers are incorporated together with specific instrumentation for research like laser-doppler or acoustic sensors. Diverse sampling points for on and off-line determinations have been arranged in the combustion chamber and sections of the boiler, uniformly distributed along the flame travel, and with accessibility to the furnace from the exterior walls of the combustor. 2.3 CFB boiler Figure 5 illustrates schematically the configuration of the CFB boiler. There are various primary oxidant nozzles above the grill to directly feed part of the primary oxidant flow (CB1) in order to maintain adequate velocity and suitable loss of load through the grill. Evaporative walls covered in a layer of abrasion resistant refractory make up the lower chamber to equip the boiler with a greater operational flexibility. The refractory surface is adjustable due to the incorporation of removable panels. A sealing non-mechanical system (Loop Seal) designed to avoid the combustion gases from the bottom of the furnace being conducted to the Solids Separator unit and located below is also provided. Solids recovered on the Solids Separator unit are conducted proportionally to one of the two outlets based on operational reasons. A furnace cooling system (INTREXTM) for solid materials returning to the furnace is located after the Loop Seal with the aim of increasing the flexibility of the system and enabling to control the average combustion temperature.
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Figure 5. CFB boiler configuration and inlet streams High-performance combustion is expected due to the turbulent mixing. However, for low-reactivity / high S content fuel, a system for reinjection of ashes has been provided. There is no need to pulverize the fuel as crushing to desired size is sufficient, consuming less power. Low SO2 emissions are anticipated due to the fact that the sulphur-limestone reaction is favourable at the range of the operating temperatures. Low NOx emissions are expected too, due to the relatively low temperature and the combustion stratification present in the furnace. In addition an ammonia injection system is available in order to study the additional emission reduction. Lower CO and CxHy emissions are also possible owing to the long residence time and turbulent mixing in the cyclone. 3. Results and Discussion As above-mentioned, preliminary results mainly regarding the Fuel Preparation Unit and the PC boiler are presented in this paper. 3.1 Fuel Preparation Unit The aim of the testing planning at this stage is double: (1) evaluation of the electricity vs. coal particle size distribution; and (2) Natural Gas (NG) consumption vs. drying capacity. Table 1 shows the characteristics of the fuel used within this testing campaign (anthracite) and the particle size distribution (PSD) achieved at both crushed (standard mesh series) and pulverized coal (sieve shaker). Anthracite Analysis Moisture content (%) Volatiles (%) Total ash (%) LCV (kcal/kg) HG Index
Crushed coal PSD % retained 6.50 14.95 15.01 4590 50
Pulverized coal PSD % retained
> 6.3 mm 0.8 >300 µm 4.0 -6.3 mm 5.8 >150 µm 2.4 -4.0 mm 26.3 >125 µm 1.0 - 2.4 mm 26.2 >90 µm 0.15 - 1.0 mm 24.3 >75 µm < 0.15 mm 16.6 >45 µm Table 1. Characteristics of crushed and pulverized anthracite
0.0 0.0 0.05 1.26 5.74 25.34
In order to grind 19.4 t of coal, 687.04 kWh were consumed by the mill, and 61 m3 of natural gas were burnt for drying purposes. This gives 35.41 kWh/t which is in line with real operating data from power plant stations for anthracite coals (36 kWh/t). Moisture of the pulverized coal is less than 1,1%. This means that more than 1047 kg of water 5
Oviedo ICCS&T 2011. Extended Abstract
were evaporated from the coal. The Hardgrove index, total moisture, input coal size, output fineness, and mill wear have direct impact on the mill output and energy consumption. 3.2 PC boiler Functionality tests are required in order to validate the PC boiler design and capabilities under oxycombustion conditions compared to air-firing conditions. Different testing campaigns have been already set including the stages: (1) start-up/shut-down; (2) steady state operation (design case); and (3) operation at minimum load (when required). The methodology for the following conditions is now available and ready for use. • Safety in oxy combustion • Switching between air and oxy modes • Functionality of auxiliary systems in oxy mode • Boiler load/firing capacity • Furnace Temperature • Control range of oxidant oxygen (in each oxidant stream) • Control range for primary, secondary and terciary stream distribution • Testing the co-firing of fuel blends • Reliability of measurements • Procedures for sampling, furnace profile, fouling/corrosion measurements Commissioning and initial tests have been carried out with low-sulphur anthracite. Local anthracite and a mixture local anthracite/petcoke are also to be tested. Characteristics of test fuels and scheduled tests are given in Tables 2 and 3 respectively. Analysis Low-S anthracite Local anthracite Petcoke Moisture % (wb) 9.9 - 12.1 6.7-11.0 4.7-9.0 Volatiles % (db) 7-15 5.7-8.5 9.6-13.0 Ash % (db) <35 31.5-38.8 0.8-6.8 Total S % (db) 0.44 1.00-1.14 4.5-5.7 LHV Kcal/kg (db) >4590 4300-5100 7100-8300 Table 2. Test fuels for PC boiler Fuel LNG Low-S anthracite Low-S anthracite Low-S anthracite Low-S anthracite Low-S anthracite Low-S anthracite Low-S anthracite Low-S anthracite
Mode
%O21
Load
FGR2 %O2 (ex)3
Air 21.00 Design 0 Air 21.00 Minimum 0 Air 21.00 Design 0 Oxy 21.00 Minimum 0 Oxy 21.00 Design 70 Oxy 25.00 Design 0 Oxy 25.00 Design 70 Oxy 27.00 Design 68 Oxy 29.00 Design 75 Table 3. Test campaigns for PC boiler
4.0 4.0 4.0 4.0 4.0 4.0 4.0 4.0 4.0
Action steady-state steady-state steady-state steady-state steady-state steady-state steady-state steady-state steady-state
1
O2 content in the oxidant stream (% db) FGR ratio (% of total flue gas wb) 3 Excess O2 in the flue gas (% db) 2
A collaborative project with the University of Santiago – Applied Maths Department aims to extend existing combustion modelling capabilities to oxy-firing conditions, applied to PC boilers. Computational Fluid Dynamics (CFD) models have been widely 6
Oviedo ICCS&T 2011. Extended Abstract
used to simulate combustion in coal-fired power stations in order to gain knowledge about certain combustion characteristics under oxy-firing conditions, such as flame aerodynamics or heat transfer [11]. Three different approaches have been conceived: (1) implementation of a model in commercial CFD-codes FLUENT; (2) development of a model using the code BTL based on physical-chemical models; and finally (3) mathematical models specifically developed within this research work such as SC3D or VFP3D using FORTRAN [12]. A commercial CFD program, ANSYS Fluent version 12.1 was used to simulate the combustion process in CIUDEN’s 20 MWth PC boiler. This code was applied to solve the appropriate transport equations for the continuous phase, whereas a Lagrangian approach was adopted to calculate the particles combustion process. Parameters considered within this study include the following: gas velocity, turbulent intensity, temperature, mass fraction of O2, CO2, CO and SO2, rate of reaction, DPM burnout, DPM mass source, DPM enthalpy and DPM evaporation/devolatilization. In order to shorten simulation processing time, the geometry of the boiler and the burners (Figure 6) have been simplified taken into account areas where boundary conditions are different.
Figure 6. Simplified geometry: boiler and burner Differences of oxycombustion compared to air-firing include among others higher gas emissivities in the furnace gases, a reduction of the volume of gases through the furnace and of flue gases (about 80%) [13]. Combustion parameters such as flame stability, ignition, radiant/convective heat transfer also changed under oxycombustion conditions. In addition the mass fraction and concentration of compounds (CO2, O2, SOx, NOx, H2O) vary in relation to the operating conditions. Figures 7 to 9 show some results of the simulation considering two different scenarios: air and oxy modes. Maximum O2 concentration considered in the oxidant stream is 30% for oxy while 25% for air-firing. Figure 7 shows that the flame area is narrower in oxy mode than in air mode at the same load due to the higher specific heat of the CO2. This is consistent with literature and expectations based on CFD modelling undertaken during the burner design stage. Figure 8 shows the high CO2 concentration in the boiler due to the recirculation of flue gas (mainly CO2). Figure 9 shows that under oxy-firing conditions combustion takes place closer to the injection zone since the O2 concentration is lower.
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OXY‐mode
AIR‐mode
Figure 7. Temperature profiles for PC boiler
OXY‐mode
AIR‐mode
Figure 8. CO2 concentration profiles for PC boiler
OXY‐mode
AIR‐mode
Figure 9. O2 profiles for PC boiler
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3. Conclusions The paper includes the description of CIUDEN Technology Development Centre for CO2 Capture located in Northwestern Spain, in particular the oxyPC 20 MWth system. The installation, first of its class, will provide real basis for the design and operation of flexible and competitive oxycombustion facilities at demonstration scale, thus accelerating the deployment of CCS technologies. Preliminary results focused on the Fuel Preparation Unit and the 20 MWth PC boiler, processing low-S anthracite as design fuel, are presented in this paper. Functionality tests have been carried out with the aim of study the capabilities of the oxy-PC system in comparison to air mode. Different testing campaigns have been set in order to develop the methodology to implement for the following stages: (1) start-up/shut-down; (2) steady state operation; and (3) operation at minimum load. A commercial CFD program, ANSYS Fluent version 12.1 was used to simulate the combustion process in the PC boiler. CFD results allow the comparison between air and oxy mode about relevant operational parameters such as temperature profiles inside the boiler, CO2 concentration and flame properties. Validation of the model is required when sufficient experimental data are available. 4. Further research work As regards forthcoming test campaigns on the PC boiler, research will focus on operating considerations include theoretical considerations such as radiant heat transfer in the furnace (combustion/heat transfer interaction), generated compounds (NOx, SO2/SO3, CO), ash properties, slagging, and fouling. For these issues there is much experience at lab scale but experience is required at larger scale. Acknowledgement. Part of the work presented is co-financed under the FP7 Programme and the European Union's European Energy Programme for Recovery programme. The sole responsibility of this publication lies with the author. The European Union is not responsible for any use that may be made of the information contained therein. References [1]
[2] [3] [4] [5] [6]
[7]
Communication from the commission to the European parliament, the council, the European economic and social committee and the committee of the regions. A Roadmap for moving to a competitive low carbon economy in 2050. COM(2011) 112 final. March 2011. International Energy Agency. Key World Energy Statistics. 2010 International Energy Agency. Technology Roadmap - Carbon Capture and Storage. October 2010. Global Carbon Capture and Storage Institute (GCCSI). Communicating for CCS deployment – Report. November 2009. J Lipponen. Carbon Capture and Storage: Potential and Challenges in the Global Context. CCT2011 Conference. Zaragoza, Spain. May 2011 V J Cortes. State of development and results of oxy-coal combustion research initiative by CIEMAT in Spain. 2nd IEAGHG International Oxy-Combustion Network Meeting. Windsor, USA. January 2007. V J Cortes. CIUDEN´s Technology Development Centre for CO2 Capture. CCT2011 Conference. Zaragoza, Spain. May 2011. 9
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[8]
[9]
[10]
[11]
[12]
[13]
M Lupion; B Navarrete; P Otero; V J Cortes. CIUDEN CCS Technological Development Plant on oxycombustion in Coal Power Generation. 1st International Oxyfuel Combustion Conference. Cottbus, Germany. September 2009. M Lupion; B Navarrete; P Otero; V J Cortes. Experimental programme in CIUDEN’s CO2 capture technology development plant for power generation. Chemical Engineering Research and Design. In press. 2010. M Lupion; R Diego; L Loubeau; B Navarrete. CIUDEN CCS Project: Status of the CO2 Capture Technology Development Plant in Power Generation. 10th International Conference on Greenhouse Gas Control Technologies, GHGT10. Amsterdam, Netherlands. September 2010. L. Álvarez, M. Gharebaghi, M. Pourkashanian, A. Williams, J. Riaza, C. Pevida, J.J. Pis and F. Rubiera. CFD modelling of oxy-coal combustion in an entrained flow reactor. Fuel Processing Technology. Volume 92, Issue 8, Pages 1489-1497. August 2011. A Bermudez, J L Ferrin, A Liñan. The modelling of the generation of volatiles, H2 and CO, and their simultaneous diffusion controlled oxidation, in pulverised coal furnaces. Combustion Theory and Modelling, 11: 949-976 (2007). Wall, T., et al., Demonstrations of coal-fired oxy-fuel technology for carbon capture and storage and issues with commercial deployment. Int. J. Greenhouse Gas Control. 2011.
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PROGRAM TOPIC: COAL GASIFICATION AND CLEAN FUELS OXY-FUEL COAL GASIFICATION IN FLUIDISED BEDS Nicolas Spiegl, Esther Lorente, Nigel Paterson, Cesar Berrueco, Marcos Millan* Department of Chemical Engineering, Imperial College London, London SW7 2AZ, UK * Corresponding Author Phone: +44(0)20 7594 1633
[email protected]
Fluidised bed gasification represents a viable technology for power generation from low value coals, biomass and wastes. These gasifiers are normally air-blown to avoid ash melting and the subsequent loss of fluidisation at the high temperatures resulting from oxygen-firing. However, this is a drawback from the point of view of carbon capture as the resulting flue gases would contain large amounts of nitrogen, making expensive N2-CO2 separation technologies necessary for CO2 capture and storage (CCS). In this work, an oxy-fuel process alternative, where the bed is fluidised with recycled flue gas (mainly CO2) and oxygen, is assessed. A continuously fed, laboratory scale fluidised bed (FB) reactor has been used to study the effect of operating conditions including pressure, temperature and gasification agent on the extent of gasification of a German lignite. Conditions necessary to avoid agglomeration of the bed material were identified. An overview on the effect of these variables on carbon conversion, gas composition and overall operability of the gasifier will be presented. Overall, this work shows that under certain operating conditions, oxy-fuel firing can enhance fuel gasification reactivities in fluidised beds and it represents a promising route to integrate gasification and CCS technology.
Introduction Fluidised bed gasifiers require oxygen to combust part of the fuel and produce the heat required for the endothermic process, a temperature moderator to keep the bed temperature well below the ash softening point and a gasification agent to convert the remaining fuel into fuel gas. FB gasifiers are normally operated between 850°C and 1000°C, depending on the ash properties, as ash softening has to be avoided. The overall process performance, e.g. carbon conversion, fuel gas heating value and composition, and energy conversion, needs to be optimised by controlling the oxidizer to fuel and gasification agent to fuel ratios.
The effect of the ratio between oxygen and carbon in the fuel (O/C ratio) on carbon conversion and on the overall operability of the gasifier has been published elsewhere [1]. In this study the influence of different gasification agents and bed temperature on the performance of CO2-blown gasification was investigated. Firstly, variations in the gasifier output due to changes in the fuel to gasification agent ratio are presented as a function of temperature. Secondly, the effect of injecting steam/CO2 mixtures as gasification agent at different temperatures is outlined.
All experiments were carried out at atmospheric pressure using a German Lignite (GL) as fuel in a bench-scale fluidised bed gasifier described in [1]. The main parameters used to judge the process performance are fuel gas composition and heating value, and carbon conversion.
Effect of CO2/C Ratio GL was gasified at 750°C, 850°C and 950°C using different CO2/C ratios. All experiments were carried out with 100% CO2 as fluidising gas. Different CO2/C ratios were achieved by changing the amount of CO2 injected into the reactor and keeping the fuel feeding rate constant. This changes the superficial velocity but keeps the pyrolysis contribution to the overall performance constant. The change in superficial velocity is seen as not significant under these conditions as discussed in more detail elsewhere [2].
950°C
850°C
750°C
Figure 1: Fuel Gas Composition in oxy-fuel gasification of German lignite at different temperatures and CO2/C ratios. (□ CO2, CO, Δ H2, × CH4)
Figure (gas composition) and Figure (carbon conversion) show the following trends with increasing CO2/C ratio: (1) at 950°C, there is a decrease in CO concentration, an increase in CO2 concentration and a constant (average 85%) carbon conversion; (2) only small changes in CO and CO2 concentrations are observed at 850°C, although there is a clear increase in carbon conversion from to 60% at CO2/C ratio of 1 to 85% at 1.8. Finally, at 750°C the increase in carbon conversion with CO2/C ratio is minor, from 32% to 37% while, similarly to the results at 850°C, there is a slight increase in CO2 and a minor decrease in CO. It seems that at a maximum possible carbon conversion is reached 950ºC. These experiments were carried out at a high superficial velocity (up to 0.27m/s) and it was observed that this results in certain amount of char particles being ejected from the bed and collected in the tar trap. Therefore the limit of 85% conversion is probably due to this loss of char particles during the experiment. At 950°C increasing CO2/C ratios does not result in increasing carbon conversion. The char-CO2 reaction at this temperature seems to be fast and not affected by an increasing amount of CO2.
Increasing the CO2/C ratio diluted the fuel gas and therefore decreased the heating value from 10 to 8.5 MJ/m3 (
Figure 1).
At 750°C the char-CO2 reaction is very slow and most of the carbon conversion is a result of pyrolysis. Therefore increasing CO2/C ratio has only a limited effect on the overall conversion. The result at 850°C is markedly different from that at 750°C. The char-CO2 reaction seems to be fast enough to be affected by the partial pressure of CO2. At a CO2/C ratio of 1.8 it reaches 85% carbon conversion, the maximum achieved in these experiments as discussed above. Both at 750 and 850°C excess CO2 dilutes the fuel gas and decreases its heating value by a similar amount (
Figure
1).
Figure 2: Oxy-fuel gasification of German lignite at different temperatures and CO2/C ratios – Carbon Conversion (950°C, □850°C, Δ750°C).
Figure 1: Oxy-fuel gasification of German lignite at different temperatures and CO2/C ratios –Fuel gas heating value LHV ( 950°C, □ 850°C, Δ 750°C).
Effect of Steam/CO2 Ratio It is well known that steam gasification is faster at lower temperatures than CO2 gasification. To investigate a potential combination of both gasification agents, CO2 was stepwise replaced with steam. Adding steam to a primary O2/CO2 blown process could be considered for two reasons. Firstly, in order to take advantage of the faster char-steam rate to increase carbon conversion and heating value of the fuel gas and
secondly, to increase the H2/CO ratio in the product gas if required by downstream applications.
Experiments were carried out at atmospheric pressure with GL. Fuel was gasified at 750°C, 850°C and 950°C using different steam/CO2 ratios (0, 0.3, 1 and 3). This results in 0, 25%, 50% and 75% of CO2 being replaced by steam (molar basis). No O2 was added to the inlet gas in this set of experiments. The total flow rate entering the reactor was kept constant at 2.64 NL/min and the superficial velocity was in the range 0.24 - 0.27 m/s. The resulting fuel gas compositions are summarised in Figure .
3 950°C
850°C
750° C
Figure 4: Oxy-fuel gasification of German lignite at different temperatures and Steam/CO2 ratios. (CO, □CO2, Δ H2, × CH4)
With increasing steam/CO2 ratio, the char-CO2 reaction is partially replaced with char-steam reaction. This change can be clearly identified in the fuel gas composition in Figure . Replacing CO2 with steam increases the H2 concentration and decreases the CO concentration in the fuel gas.
Figure shows the change in carbon conversion with increasing steam/CO2 ratio. At 950°C the maximum possible carbon conversion is achieved independently of the CO2/C and steam/CO2 ratios. At 850°C carbon conversion is increased by using 25%
steam to a level comparable to that observed in experiments carried out at 950°C. Further increasing steam/CO2 ratio did not change the carbon conversion at 850°C. As for 950°C, it is assumed that the maximum possible carbon conversion is reached due to entrainment of char particles.
3.00 Figure 5: Oxy-fuel gasification of German lignite at different temperatures and steam/CO2 ratios – Carbon Conversion (950°C, □ 850°C, Δ750°C).
Summary and Conclusions High carbon conversion and syngas heating values can be achieved by FB oxy-fuel gasification. In general, increasing temperature increases carbon conversion. However, results presented above show that gasification temperature can be decreased from 950°C to 850°C, while keeping the carbon conversion constant if (1) the CO2/C ratio is increased to 1.8 or (2) the steam/CO2 ratio is increased to 0.33. In the first case this comes at the cost of having a large amount of additional CO2 in the process, costing additional energy to heat it up and diluting the fuel gas. In the second case extra steam is need for the process and the H2/CO ratio of the fuel gas is increased. In an O2/CO2 blown gasification process CO2 is the main gasification agent. The charCO2 reaction is strongly temperature-dependent. At 750°C almost no CO2 gasification occurred, while at 950°C the rate was fast enough to achieve complete carbon conversion with a gasification agent consisting of 100% CO2 under any conditions
tested. The reaction is also strongly dependent on the CO2/C ratio, which produces an increase in conversion. On the other hand, excess CO2 dilutes the fuel gas and more energy is required to heat up the larger volume of fluidising gas.
Up to 75% of CO2 was replaced with steam to study the gasification behaviour if a mixture of these gasification agents is used. It was found that at 950°C both reactions were fast enough to achieve maximum carbon conversion. The addition of steam to an O2/CO2-blown fluidised bed gasification process could be useful to increase H2/CO2 ratio in fuel gas and to achieve higher carbon conversion at lower temperatures.
References [1] Spiegl N., Sivena, A., Lorente, E., Paterson, N., and Millan, M., “Oxy-fuel coal gasification in fluidized beds – A novel route for zero emission power generation: Equipment development and study of gasification performance” Energy & Fuels, 2010, 24, 9, 5281-5288. [2] Spiegl N., “Oxy-fuel gasification in fluidized beds” PhD Thesis, Imperial College London, 2011.
Acknowledgments: Funding from RFCS, European Commission, FLEXGAS Project, Near Zero Emission Advanced Fluidised Bed Gasification is gratefully acknowledged.
Experimental and numerical investigation of oxy-coal combustion in an entrained flow reactor
L. Álvarez1, M. Gharebaghi2, A. Williams2, M. Pourkashanian2, J. Riaza1, C. Pevida1, J.J. Pis1, F. Rubiera1 1
Instituto Nacional del Carbón, INCAR-CSIC. Apartado 73, 33080 Oviedo, Spain Centre for Computational Fluid Dynamics, School of Process, Environmental and Materials Engineering, University of Leeds, Leeds LS2 9TJ, UK
[email protected]
2
Abstract This paper presents experimental and numerical tests for coal combustion under O2/CO2 atmospheres in an entrained flow reactor (EFR) using five coals of different rank. Mixtures of O2/CO2 of various concentrations (21-35%vol O2) were used and compared with the combustion in air. The temperature profiles, burning rates, coal burnout and NO emissions in the EFR were predicted. A decrease in both experimental and numerical coal burnouts was observed when N2 was replaced by CO2 for the same oxygen concentration (21%), but an improvement in the O2/CO2 atmosphere for an oxygen concentration higher than 30%. The NO emissions during combustion in 21%O2/79%CO2 were lower than those in 21% O2/79%N2. Experimental coal burnout obtained in the EFR were used to test the accuracy of the CFD model. The numerical results obtained with the CFD model were in good agreement with the experimental results.
1. Introduction Oxy-coal combustion is considered as one of the most promising CO2 capture technologies. In addition, a significant and simultaneous reduction of NOX and CO2 can be achieved. During oxy-coal combustion, coal is burnt with a mixture of O2 and recycled flue gas (mainly CO2 and H2O). Due to the difference in gas properties of N2 and CO2, oxy-coal combustion differs greatly from air combustion in several ways, including heat transfer, flame ignition, coal burnout and pollutant formation [1]. Computational Fluid Dynamics (CFD) models have been widely used to simulate combustion in coal-fired power stations. CFD can be used to gain knowledge about some combustion characteristics like char oxidation, radiation, pollutant formation (Hg, NOX and SOX) or the impact of turbulence [2]. The aim of this paper was to develop a 1
CFD model in order to evaluate the effect of replacing N2 by CO2 on coal burnout, temperature and NO emissions during coal combustion. Experimental coal burnout results obtained in an entrained flow reactor (EFR) were used to validate the CFD model.
2. Experimental section The experimental data were obtained using an EFR, whose details have been previously reported [3]. The EFR has an internal diameter of 4 cm and a length of 200 cm, the reaction zone being assumed to extend up to a length of 142 cm. Five coals of different rank (AC, HVN, SAB, DAB and BA) were used in the combustion experiments. The proximate and ultimate analyses, and the heating values of the coals are shown in Table 1. For the combustion experiments, air (21%O2/79%N2) was taken as reference, and three binary gas mixtures of O2 and CO2 (21-35 %vol O2) were used. The experiments were performed at 1273 K; the gas flow rate was adjusted in order to ensure a particle residence time of 2.5 seconds and the oxygen excess was varied between 0 and 100%. The exhaust gases were monitored using a battery of analysers (O2, CO, CO2, SO2, NOX). Table 1 Proximate and ultimate analyses, and high heating value of the coals AC HVN DAB Sample Origin Spain Spain China Rank an sa hvb Proximate Analysis (wt.%, db) Ash 14.2 10.7 10.9 V.M. 3.6 9.2 28.8 F.C.a 82.2 80.1 60.3 Ultimate Analysis (wt.%, daf) C 94.7 91.7 81.9 H 1.6 3.5 5.0 N 1.0 1.9 0.8 S 0.7 1.6 1.2 Oa 2.0 1.3 11.1 High heating value (MJ/kg, db) 29.2 31.8 28.5 an: anthracite; sa: semi-anthracite; hvb: high-volatile bituminous coal. db: dry basis; daf: dry and ash free bases a Calculated by difference.
SAB S. hvb
BA Spain hvb
15.0 29.9 55.1
6.9 33.9 59.2
81.5 5.0 2.1 0.9 10.5 27.8
88.5 5.5 1.9 1.1 3.0 33.1
3. Modelling approach A commercial CFD program, ANSYS Fluent version 12, was used to simulate the combustion process in the EFR. This code solves the meaning transport equations for the continuous phase, and a Lagrangian approach was used to calculate particle
2
trajectories through the calculated gas field. The RNG k-ε turbulence model was used to model the dynamic of the flow [4]. The Discrete Ordinate (DO) radiation model was used and the gas mixture absorption coefficient was calculated by using the weighted sum of gray gases model (WSGGM) [5]. The devolatilisation rate of the coal was modelled using a single step first-order Arrhenius reaction. The kinetics parameters were obtained by means of the FG-DVC devolatilisation model [6]. Volatile combustion was simulated using the mixture fraction/PDF chemical equilibrium approach [7]. PDF tables for both air and oxy-fuel conditions were calculated using a pre-PDF pre-processor. Twenty species including chemical species (CO2, O2, N2, CO, H2O, H2, CH4, C2H2, and SO2) and radicals and intermediate species (C, H, O, N, S, OH, CS, …) were included. The Smith intrinsic model was employed for char combustion [8]. Devolatilisation and combustion parameters employed in this work are summarized in Table 2. Table 2 Devolatilisation and combustion reactivity data inputs for Fluent model Parameter/case
AC
HVN
DAB
SAB
BA
Devolatilisation model
Single rate
Single rate
Single rate
Single rate
Single rate
Pre-exponential factor (1/s)
3.07E4
3.6E14
4.68E11
4.68E11
2.01E11
Activation energy (kJ/mol)
228.6
229.7
155.9
155.9
148.6
Combustion model
Intrinsic
Intrinsic
Intrinsic
Intrinsic
Intrinsic
Kinetic-limited rate preexponential factor (g/cm s)
0.030198
0.030198
0.030198
0.030198
0.030198
Kinetic-limited rate activation energy (kJ/mol)
155±10
155±10
155±10
155±10
155±10
Specific surface area (m2/g)
40
40
100
100
100
Burning mode, alpha
0.25
0.25
0.25
0.25
0.25
Once validated the CFD, NO predictions were carried out as a post-processor. The successful prediction of NO emissions requires the correct representation of the fluid flow, heat transfer, combustion process, and NO chemistry. Fuel nitrogen was assumed to evolve as HCN and NH3 during devolatilisation of the coal particles. The HCN and NH3 evolved from the volatiles were competitively oxidized and reduced to produce NO and N2, respectively, via the mechanism proposed by De Soete [9]. The char nitrogen was heterogeneously oxidized to NO according to the mechanism proposed by Lockwood [10].
3
4. Results and Discussion Experimental results Coals were burnt under different levels of oxygen excess for each atmosphere studied. In this work, the fuel ratio was used to assess the oxygen excess during combustion. This parameter is defined as the ratio between the coal mass flow rate used and the stoichometric value. Fig. 1 shows the burnouts for HVN under the different experimental conditions. 700
100
HVN
HVN
600
90 NO (ppm)
Burnout (%)
500
80 70 21%O2-79%N2 21%O2-79%CO2 30%O2-70%CO2 35%O2-65%CO2
60
400 300 21%O2/79%N2 200
21%O2/79%CO2
100
30%O2/70%CO2 35%O2/65%CO2
50
0
0.2
0.4
0.6
0.8
1.0
1.2
1.4
Fuel Ratio
0.2
0.4
0.6
0.8 1 Fuel ratio
1.2
1.4
Figure 1 Burnout and NO emissions during combustion of coal HVN under different atmospheres as a function of the fuel ratio. As expected, coal burnout decreases as the fuel ratio increases due to the lower concentration of oxygen. The burnout obtained with the mixture of 21%O2/79%CO2 is lower than that of air. When N2 is replaced by CO2, the heat capacity of the gases increases, leading to lower particle temperatures [11]. As a consequence, the combustion rate of the particles decreases and the coal burnout becomes worse. Also a reduction on NO is observed due to the higher CO concentrations in oxy-fuel environments [12]. For the mixtures of 30%O2/70%CO2 and 35%O2/65%CO2 the burnout and NO emissions are higher than that in air, since higher oxygen concentration produces an enhancement in the char combustion rate.
Computed results For the CFD model the calculation domain is the reaction zone of the EFR. A threedimensional structured grid consisting of ~75,000 cells was employed to describe a quarter of the total volume. The boundary conditions, mass flow inlets and wall temperatures were established using measurements made during the experimental tests. 4
Figure 2 (a) Predicted temperature (K), (b) burning rate (kg/s) and (c) NO concentration (%) inside the EFR during HVN and BA combustion for: (I) 21%O2/79%N2, (II) 21%O2/79%CO2, (III) 30%O2/70%CO2 and (IV) 35%O2/65%CO2.
5
As can be seen in Fig. 2 (a) significant differences in the temperature contours appear when N2 is replaced by CO2 for the same oxygen concentration (cases I and II), also the temperature decreases due to the higher specific heat of CO2. When the oxygen concentration increases (cases III and IV), the temperature also increases. The predicted burning rates shown in Fig. 2 (b) suggest that devolatilisation in O2/CO2 starts earlier than in air, i.e., it is enhanced by the higher CO2 concentrations. But, after devolatilisation takes place, the burnout rate is lower in 21%O2/79%CO2 than in 21%O2/79%N2, indicating that no improvement occurs due to char-CO2 gasification. It can also be seen in Fig. 2 that for the rest of the oxy-firing cases (III and IV), when the oxygen concentration increases, there is a concomitant increase in char burning rate. In the NO concentration profiles shown in Fig. 2 (c), a decrease on NO emissions is observed for the test with 21%O2/79%CO2 (case II) in comparison with 21%O2/79%N2 (case I). As shown in Fig. 2, significant differences were found between HVN (semianthracite) and BA (bituminous), as it could be expected due to their different rank. BA reached higher temperatures and, as a consequence, its burning rate was higher. HVN and BA have similar nitrogen content but their NO emissions profiles are different, due to the different amount of nitrogen intermediates (HCN and NH3) evolved with the volatiles. To validate the CFD model, the experimental burnouts in the EFR were employed. Table 3 shows a comparison between the experimental and computed burnouts Table 3 Comparison between experimental and numerical coal burnouts for the five coals studied in air and O2/CO2 conditions. Atmosphere
21%O2/79%N2
21%O2/79%CO2
30%O2/70%CO2
35%O2/65%CO2
Coal
Exp.
Predict.
Exp.
Predict.
Exp.
Predict.
Exp.
Predict.
AC
76.8
78.5
69.0
72.1
79.7
87.3
81.6
91.0
HVN
79.5
82.3
77.2
79.7
81.7
87.9
82.4
90.9
DAB
96.3
94.9
95.8
94.6
97.3
97.8
-
-
SAB
92.5
95.9
90.2
95.1
94.4
98.6
94.7
99.8
BA
95.9
94.3
96.5
94.1
96.5
98.8
96.1
99.4
6
In all cases the predicted coal burnouts follow the same trend as the experimental results. Burnout in a mixture of 21%O2/79%CO2 is lower than that in air, but an improvement is achieved in both experimental and numerical cases when the oxygen concentration is 30% or higher. In general, a good agreement between experimental and numerical results was obtained. Although for the anthracitic coals for high oxygen concentrations (35%) the burnout is slightly overpredicted. Table 4 shows a comparison between experimental and predicted NO emissions. For most of the predicted values are similar to the experimental ones. In air firing conditions, NO emissions are slightly overpredicted specially for bituminous coals. This is due to inaccuracies when predicting temperature profiles that lead to thermal NO overpredicted values, that are highly dependant on temperature. Table 4 Comparison between experimental and numerical NO emissions (ppm) for the five coals studied in air and O2/CO2 conditions Atmosphere
21%O2/79%N2
21%O2/79%CO2
30%O2/70%CO2
35%O2/65%CO2
Coal
Exp.
Predict.
Exp.
Predict.
Exp.
Predict.
Exp.
Predict.
AC
247
226
223
198
287
330
372
418
HVN
368
435
328
319
578
555
575
610
DAB
283
333
279
263
399
384
-
-
SAB
400
480
365
365
498
519
474
521
BA
350
496
306
322
484
486
477
506
5. Conclusions •
The computed results showed changes in the temperature profiles when N2 is replaced by CO2, these changes are associated with differences in density, heat capacity and radiative properties of N2 and CO2. These differences lead to changes in coal combustion behaviour (coal burning rates and coal burnout for example).
•
A decrease on coal burnout values and NO emissions is observed in both experimental and predicted values in 21%O2/79%CO2 if compared with those in air conditions. When the oxygen concentration is increased to 30%, the burnout and NO emissions are higher than that in air.
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•
Good agreement between predicted and experimental coal burnout and NO emissions was achieved.
Acknowledgements Work carried out with financial support from the Spanish MICINN (Project PS-1200002005-2) co-financed by the European Regional Development Fund. L.A. and J.R. acknowledge funding from the CSIC JAE program, which was co-financed by the European Social Fund, and the Asturias Regional Government (Severo Ochoa program), respectively. MG acknowledges financial support from E.On UK, EPSRC, Dorothy Hodgkin Postgraduate Awards
References [1] Buhre BJP, Elliot LJ, Sheng CD, Gupta RP, Wall TF. Oxy-fuel combustion technology for coal-fired power generation. Progress in Energy and Combustion Science 2005; 31: 285-307. [2] Edge P, Gharebaghi M, Irons R, Porter R, Porter RTJ, Pourkashanian M, Smith D, Stephenson P, Williams A. Combustion modelling opportunities and challenges for oxy-coal carbon capture technology. Chemical Engineering Research and Design 2011, doi:10.1016/j.cherd.2010.11.010 [3] Arias B, Pevida C, Rubiera F, Pis JJ. Effect of biomass blending on coal ignition and burnout during oxy-fuel combustion. Fuel 2008; 87: 2753-2759. [4] Jones WP, Launder BE. The prediction of laminarization with a two-equation model of turbulence. International Journal of Heat and Mass Transfer 1972; 15: 301-314. [5] Viskanta R, Menguc MP. Radiation heat transfer in combustion systems. Progress in Energy Combustion Science 1987; 13: 97-160. [6] Solomon PR, Fletcher TH. Impact of coal pyrolysis on combustion. Symposium (International) on Combustion/The Combustion Institute 1994; 15: 463-474. [7] Sivathanu YR, Faeth GM. Generalized state relationships for scalar properties in non-premixed hydrocarbons/air flames. Combustion and Flame 1990; 82: 211-230. [8] Smith IW. The combustion rates of coal chars: a review. Symposium (International) on Combustion/The Combustion Institute 1982; 19: 1045-1065. [9] De Soete GG. Overak reaction rates of NO and N2 formation from fuel nitrogen. Symposium (International) on Combustion/The Combustion Institute 1975; 15: 10931102. 8
[10] Lockwood FC, Romo-Millares CS. Mathematical modelling of fuel-NO emissions from PF burners. Journal of the Institute of Energy 1992; 65: 144-152. [11] Berejano PA, Levendis Y. Single-coal particle combustion in O2/N2 and O2/CO2 enviroments. Combustion and Flame 2008; 153: 270-287. [12] Álvarez L, Riaza J, Gil MV, Pevida C, Pis JJ, Rubiera F. NO emissions in oxy-coal combustión with the addition of steam in an entrained flow reactor. Greenhouse Gas Science Technology 2011; 1: 1-11.
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Oviedo ICCS&T 2011. Extended Abstract
What is Reactive Surface Area in Coal Chars?
Eric M. Suuberg, Indrek Aarna1 and Indrek Külaots School of Engineering, Brown University, Providence, RI 02912 USA
[email protected] 1 Present Address: Eesti Energia, Tallinn, Estonia
Abstract Coal char reactivity towards oxidizing and reducing gases has been studied for a century, but there still remain unresolved aspects. There has been a debate regarding the best way to represent the intrinsic reactivity of carbonaceous chars, some maintaining that the reactivity should be expressed on a surface area basis, since the reaction takes place at an interface. This has led to many efforts to characterize the appropriate surface area using various techniques. In this study, the reactions of coal chars produced at different heat treatment temperatures are explored in oxidizing gas (carbon dioxide). These results point to an interesting dilemma that the “correct” way to characterize reactive surface appears to depend upon the reaction rate regime, and the role of catalytic minerals in the char. In a chemisorption rate controlled system, use of a characterization based upon a measure of microporosity (e.g., BET surface area) appears to work well, whereas this does not work in a regime in which oxide surface desorption is the principal controlling step. Hence, the different effects of thermal treatment (annealing) are also pointed out in these different regimes. 1. Introduction The reactions of coal, and other chars, with reactive gases have been the subject of study for a very long time. The question of what constitutes effective surface area for reaction (or reactive surface area) is one that has received some significant attention in such studies [e.g., see 1,2], though in many cases the question has been swept aside in empirical descriptions of the phenomena. In this paper, new results on two model reaction systems are presented, and shed new light on some aspects of the problem. 2. Experimental section Controlled experiments have been performed on chars derived from a Wyodak coal sample (particle size <150 µm) obtained from the Argonne Premium Coal Sample Bank [3]. Samples were pyrolysed in a tube furnace at selected temperatures for two hours, in flowing helium. The resulting chars were gasified in an Online Instruments TG-plus TGA system. Gasification was performed in flowing mixtures of helium and reactant gas (CO2). Samples of 30-50 mg were dispersed on circular platinum pans with a large flat surface, resulting in particle beds of about 1 mm thickness. Temperatures were selected to assure that gasification took place under reaction rate controlled (so-called "Zone I") conditions. Adsorption isotherms were determined for product chars using an automated volumetric gas adsorption apparatus (Autosorb 1, Quantachrome Co.). Adsorption of N2 and CO2 were performed at 77 K and 195 K, respectively. We feel that nitrogen offers
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1
Oviedo ICCS&T 2011. Extended Abstract
the more complete and accurate picture of porosity[1], despite the well-known problem of activated diffusional limitations at near-zero burn-off [4]. Before measurements, samples were outgassed for several hours in vacuum at temperatures in the range 573673 K. 3. Results and Discussion Figure 1 shows the results of reactivity experiments involving a series of Wyodak chars, heat treated at different temperatures. These results are displayed on a commonly employed basis of mass loss rate per unit of mass of sample. In this instance, the samples were all reacted at a partial pressure of 4.8 kPa of CO2 in mixture with helium at atmospheric pressure. The rates are low, given the low temperatures and low partial pressure of reactant. The results are taken in the Zone I intrinsic reaction rate control regime in which the reaction rate actually displays two kinetic regimes [5]. The lower temperature regime has been characterized as involving a rate controlled by desorption of oxides from the char surface, whereas the higher temperature regime is the more usually studied regime of practical interest, in which chemisorption plays a main role in determining the overall rate.
Figure 1. Wyodak char reactivity towards 4.8 kPa of CO2. The legend indicates the temperatures of char preparation. The usual type of “thermal annealing” behavior is observed in the high temperature reaction rate results; the reactivity of the chars decrease with increasing severity of heat treatment. In some instances, this behavior has been ascribed to the loss of active sites per unit of surface area. In other cases, it is a general loss of surface area that is involved. Neither mechanism can generally be ruled out.
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Oviedo ICCS&T 2011. Extended Abstract
What makes the results of Figure 1 interesting is that if they are converted to another commonly-employed way of representing reactivity, based on the BET surface area of the char samples, a very different picture emerges, as seen in Figure 2.
Figure 2. The results of Figure 1, but expressed on a surface area, as opposed to mass basis. In this instance, it is the results in the high temperature kinetic regime that are pulled into some general alignment while the results from the low temperature regime are much more spread than in Figure 1. The implication from comparison of Figures 1 and 2 is that what changes during heat treatment or annealing is the surface area available for chemisorption of the reactive gas. That is, the mass reactivity of the samples changes with heat treatment in a manner that can be largely explained by the variation of sample surface area. It is understood that BET surface area is only a crude surrogate for some measure of micropore area. It should also be noted that these samples were sufficiently burned off such that the diffusional restrictions sometimes cited as precluding the use of nitrogen for surface area measurements were not of concern (i.e., there was no need to use CO2 as the adsorbate, as is sometimes recommended for samples with very small microporosity). Figure 3 leaves little doubt as to the nature of the reaction in the high temperature regime. In this instance, one set of chars was partially demineralized by acid washing, prior to subjecting the samples to reaction. The results quite clearly demonstrate that the acid washing process practically eliminates the high temperature behavior (at least in the range of temperatures examined). What this means is that the chemisorption controlled regime is catalyzed by the presence of mineral- derived species in the coal. There is
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3
Oviedo ICCS&T 2011. Extended Abstract
nothing particularly new in this aspect of the results- the rates of CO2 gasification of coal chars are known to be strongly influenced by the presence of mineral catalysts (particularly the alkali and alkaline earth elements). Hence the surface area influence in the high temperature regime is clearly associated with the accessibility of catalytic mineral species.
Figure 3. A comparison of surface area reactivity of 1273 K Wyodak chars towards CO2 for both unmodified char, and char demineralized by acid washing. In the low temperature regime, the presence of catalysts has no apparent influence on the reaction rate. This means that the action of catalysts in this system is associated with the chemisorption process and not the desorption process, which governs the low temperature regime. The implication is also that in the low temperature regime, it is something that correlates better with total carbon mass, as opposed to micropore area or volume, which determines rate. It should be kept in mind that in all cases what are being reported are pseudo steady-state rates of carbon loss (on the timescale of these particular experiments). Hence the desorption rate associated with surface oxides does not depend upon the measured micropore surface area of the char. It is rather unlikely that the oxides are being desorbed from carbon species in the bulk of the sample (e.g. from the middle of carbon crystallites or stacks of aromatic units). Consequently, since there is a clearly established variation of micropore surface area with thermal treatment of these samples, this means that the desorbing oxides must be associated with surfaces that do not correlate with micropore area measured by gas adsorption methods. These latter methods are typically applied by us to samples that have “cleaned” surfaces, which are not heavily oxide laden, as they would be during the gasification reactions.
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Oviedo ICCS&T 2011. Extended Abstract
A possible explanation for this apparently different dependence on surface area is that when oxide desorption is limiting, the pores become filled with oxides, and there is a much less distinct “surface” involved in the phenomenon. The population of oxides is characterized by a distribution of activation energies of desorption, which is what gives rise to the pseudo activation energy observed in the low temperature regime. The exact nature of the desorption process is not known- there is quite possibly some transformation of surface oxides during the desorption, and then of course there is a restoring of the surface population through chemisorption. Overall, the rate of gasification in this regime is close to zero order in reactant gas concentration. 4. Conclusions This work addresses the issue of reactive surface area of chars from the perspective of one particular oxidizing reactant- carbon dioxide. Other work in this laboratory has focused on nitric oxide and oxygen. Each reactant offers its own unique picture, since the basic reaction mechanisms of the reactants are somewhat different from one another. Nonetheless, the results from the intrinsic reaction control regime of carbon dioxide provide an insight suggesting that the “correct” surface area depends upon the particular reaction system and regime. For carbon dioxide, chemisorption associated with catalytic minerals, seems to correlate with traditional BET (i.e., micropore) area. On the other hand, the rate of desorption of oxides from the char surface does not seem to correlate with micropore area, and instead, seems to be related to the total amount of char. In this instance, thermal annealing does not influence the rate of reaction in the desorption controlled regime. References [1] Aarna, I., Suuberg, E.M., The Role of Surface area in Char Gasification and
Oxidation, 27th Symposium (Int.) on Comb. 1998 The Combustion Institute, Pittsburgh, pp.2933-2939. [2] Külaots, I. Hsu, A., Suuberg, E.M., The Role of Porosity in Char Combustion”, Proc. Comb. Inst., 2007:31:1897-1903. [3] Vorres, K.S., The Argonne Premium Coal Sample Program, Energy Fuels, 1990: 4: 420-426 . [4] Rodríguez-Reinoso, F., Linares-Solano, A., in Chemistry and Physics of Carbon, Vol. 21, P. A. Thrower, Ed., Dekker, New York, 1989. [5] Aarna, I., Suuberg, E.M., Two Kinetic Regime Behavior in Carbon Dioxide Gasification of Carbon. Carbon, 1999: 37: 152-155,.
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5
The properties of chars derived from inertinite-rich, high ash coals and CO2 gasification: Major properties affecting reactivity. Gregory N. Okolo*, Raymond C. Everson, Hein W.J.P. Neomagus Clean Coal Technology Research Group, Unit for Energy Systems, School of Chemical and Minerals Engineering, North-West University, Potchefstroom Campus, Private Bag X6001, Potchefstroom, 2520, South Africa
Abstract With the increasing global energy demand and the decreasing availability of good quality coals, a better understanding of the important properties that control the behaviour of low-grade coals and the subsequent chars in various utilisation processes, becomes pertinent. An investigation was undertaken to determine the effects of chemical and physical properties imparted on chars during pyrolysis on the subsequent gasification reaction kinetics of typical South African inertinite-rich, high ash Highveld coals. Attempt was made at following these changes from coals to chars by a detailed characterisation of both the parent coals and the subsequent chars. Characterisation results show that, although sharp variations in properties were not observed for the original coals, the subsequent chars exhibited substantial differences both amongst themselves and from the parent coals. The increasing orders of magnitude of micropore surface area, microporosity, fraction of amorphous carbon and structural disorderliness (determined from XRD carbon structural analysis), were found to change in the transition; an indication that the chars properties are different to the precursor coals’ properties. Isothermal CO2 gasification experiments of the chars were carried out in a Thermax 500 thermogravimetric analyser in the temperature range of 900-950°C with varying concentrations of CO2 (25-100 mol.%) in the CO2-N2 reaction gas mixture at ambient pressure. Major factors affecting the gasification reactivity of the chars as it pertains to this investigation are: parent coal vitrinite reflectance, and: aromaticity, fraction of amorphous carbon, degree of disorder and alkali indices, micropore surface area, microporosity and average micropore diameter of the chars. The random pore model (chemical reaction control) was found to adequately describe the gasification reaction experimental data. Keywords: Inertinite-rich coal, Coal and char properties, Char carbon forms, Carbon crystallite analysis, Carbon dioxide gasification, kinetic modelling. 1
1.0
Introduction
Coal gasification generally involves three essential steps [1-3]: (i) The devolatilisation of the organic matter and mineral matter leading to the formation of char. (ii) The homogeneous reaction of the volatile species from step1 with the reactant gases. (iii) The heterogeneous reaction of the resultant char with the reactant gases leading to the formation of ash. The relative slowness of the third step- the heterogeneous char-gas reaction, often means that it is the rate-determining step during coal conversion. Consequently this step plays a major role in the planning for the design of a gasifier, and in the assessment of coal for use in a particular gasification technology. It is therefore critically important that the char reactions kinetics are understood. Thus, the role played by various properties and parameters as it relates to the parent coal and the subsequent chars were investigation. A detailed analysis of the effect of well-known chemical and physical properties as well as carbon crystallite properties on the gasification reactivity was carried in order to determine the most important char properties for predicting reactivity. 2.0
Experimental
Four South African low grade coals designated as coals: B, C, C2 and D2; were selected for this study. The selection of these samples was based on their approximately equal ash contents and varying inertinite-vitrinite ratios. The char production was conducted in a Packed Bed Balance Reactor (PBBR) at 900 °C with a holding time of 70 minutes. 2.1
Coal and Char Characterisation
Various characterisation techniques were used to follow the changes that occur during pyrolysis, from the parent coals to the subsequent chars. Chemical analyses of the samples comprising of proximate and ultimate analyses, calorific value and total sulphur content determination were conducted at the Laboratory of Advanced Coal Technology (ACT), Pretoria [4,5]. The petrographic analyses (reflectance properties; maceral components; char carbon forms and general condition) of the coal and char samples were done at Petrographics SA, Pretoria [4-6]. XRD was used to study the changes in carbon crystallite properties in the transition from coals to chars and was carried out at the Laboratory of XRD Analytical and Consulting, Pretoria (Courtesy of Dr Sabine Verryn). 2
The average carbon crystallite lattice parameters (Interlayer spacing, d002; crystallite height, Lc; and average crystallite diameter, La) of both coal and char samples were determined, using the Braggs equation and Scherrer equation [7-10]. The average number of aromatic layers per carbon crystallite, Nave, was calculated following the method used by Trejo and co-workers [10]. Aromaticity, fa, of the samples was determined from the area under the (002) band the γ sideband (A002 and Aγ) [7,10-13]. This was achieved using the Origin 6.1 software and the HighScore Plus curve fitting functions. The fraction of amorphous carbon, XA, was evaluated from the maximum reduced intensity of the d002 band of the samples following the method proposed by Franklin [14] and used by Ergun and Tiensuu [15] and Lu et al. [7]. The degree of disorder index, DOI of the samples was calculated from the aromaticity and the fraction of amorphous carbon, using the method proposed by Lu and co-workers [12]. Micropore surface area of the parent coal and subsequent chars was measured by CO2 adsorption using the Dubinin- Radushkevich method on a Micromeritics ASAP 2020 System. The average micropore diameter and micropore volume was determined using the Horvart-Kawazoe (H-K) method. Microporosity of both coals and subsequent chars (Dp ≤ 5Å) was determined using the CO2 adsorption data [16]. Skeletal density of the samples was determined using helium gas on a Quantachrome Stereopycnometer (Model SPY-4). 3.0
Char Reactivity
A Thermax 500 TGA supplied by Thermo Fischer Scientific, Karlsruhe, Germany was used for the char-CO2 gasification experiments within the temperature range of 900 - 950 °C. For a quantitative study of the conversion of the fixed carbon in the char only, the TGA raw data (instantaneous time, weight and temperature) were normalised on an ash free basis on the assumption that negligible product gas is produced from mineral transformations or mineral reactions with CO2 [17-26]. Reactivity of the chars in this study was confined to the initial reactivity [21,25,26,27-30]. 3.1
Kinetic Modelling using the Random Pore Model (RPM)
The experimental conditions, and the very long experimental burnout times, are such that the char-CO2 gasification reactions were considered as under chemical reaction kinetics 3
controlled regime (Regime I). Furthermore, these conditions are similar to conditions used by various investigators for Regime I char-CO2 gasification kinetics [17,18,21,24,29,31-36]. Thus, random pore model (RPM) evaluations and determination of the kinetic parameters, were based on Regime I, as modified by Everson et al. [25]. 4.0
Results and Discussion
4.1
Characterisation Results
From the results of proximate, ultimate, gross calorific value and the total sulphur content of the coal and char samples, the parent coals were characterised as high ash coals with ash content ranging from 25.6% to 29.0 wt. %, adb, and calorific values less than 21.5 MJ·kg-1. Fixed carbon and ash contents increase from coals to chars, while the volatile matter and inherent moisture contents decrease in the transition. Ultimate analysis results show that fixed carbon and total sulphur content increase, while hydrogen, nitrogen and oxygen contents decrease from coals to chars. The H/C and O/C atomic ratios were found to be drastically reduced in the transition from coals to chars. The vitrinite random reflectance results showed that according to ISO-11760 (2005) Classification of coals, three of the coal samples: coal B, C and C2 were characterised as Bituminous Medium Rank C (0.65 - 0.78 Rr %); while coal D2 was classified as Bituminous Medium Rank D (0.56 - 0.58 Rr %) [4]. Results from the maceral analysis reveals that all four coals were inertinite-rich with the ratio of inertinite to vitrinite macerals in the parent coals, ranging from 1.93 and 26.3. Total reactives, which refer to the maceral constituents that possess the propensity to react to heating ranged from 35 to 44 vol. %, mmb for all the coal samples. Char carbon forms analysis result shows that, there is a larger abundance of char carbon forms from coal inertinites in all the chars. Attempt was made to classify the maceral components of the coals and the char carbon forms of the chars into the total reactive and inert components, TRC and TIC. The coal TRC experienced percentage losses of 38.6 to 48.6% in the transition to the subsequent chars. Unlike the TRC, the samples generally exhibited gross gains in TIC between 12.2 and 24.4%. The results of the determined physical structural properties of the coal and char samples show that, the micropore surface areas of the coals did not exhibit much variations with 4
values ranging from 93.42 to 122.7 m2·g-1. Unlike the parent coals, sharp differences were observed in the micropore surface areas of the chars.
4.1.1
X-ray Diffraction (XRD) Carbon Crystallite Analysis
The diffractograms for the demineralised coal- and char- C, are shown in Figure 1. The diffractograms of the samples investigated in this study possess the same graphitic features as those reported in literature [7,11-14,26,39]. The summary of results from the carbon crystallite analysis of the coal and char samples is presented in Table 1.
Figure 1: Comparison of the diffractograms of coal and char for sample C.
Table 1: Result on carbon crystallite analysis using XRD Sample ID Coal B Char B Coal C Char C Coal C2 Char C2 Coal D2 Char D2 d002 (Å)
3.49
3.58
3.48
3.53
3.46
3.49
3.46
3.50
Lc (Å)
15.5
10.5
15.4
11.4
16.9
11.4
14.8
11.9
Nave (-)
5.45
3.92
5.41
4.23
5.90
4.26
5.28
4.40
La (Å)
9.32
7.52
12.2
10.1
12.6
11.2
12.0
10.7
fa (-)
0.80
0.88
0.81
0.92
0.85
0.95
0.75
0.85
XA (-)
0.49
0.28
0.51
0.24
0.61
0.23
0.66
0.48
DOI (-)
0.59
0.37
0.60
0.30
0.67
0.27
0.75
0.56
5
4.2
Reactivity Results
4.2.1
Experimental Results
It can be seen from the conversion curves given in Figures 2 (i), (ii) and (iv), that the char-CO2 reactions follow the Arrhenius type kinetics with increase in temperature resulting in an increase in char conversion rate. It was also observed that increase in CO2 concentration in the reaction gas composition also affected the reaction positively. Similar trends were exhibited by all four chars and is comparable to widely published results [17,18,21,23-26,29,31-35,37]. Characteristic plots of dX/dt versus fractional conversion were used to evaluate the initial reactivity of the chars. These results are shown in Figure 2 (iii). Generally, reactivity of the chars increases in the order: char C2 < C < B < D2. The reactivity of char D2 was found to be higher than the reactivity of the other chars by a factor > 4.
Figure 2: Char reactivity results: (i) Effect of reaction temperature for char C2 at 100% CO2 concentration; (ii) Effect of CO2 concentration in the reaction gas mixture for char C at 900 °C; (iii) Determination of initial reactivity for char D2 at 25% CO2 concentration; (iv) Comparison CO2 reactivity of the four chars at 900 °C, 100% CO2 6
4.2.2
Effect of Coal and Char Properties on CO2 Reactivity of the Chars
It was found from the qualitative correlation of the coal and char properties with the initial reactivity of the subsequent chars that, the properties affecting the reactivity of the chars are: the parent coal rank (vitrinite reflectance) [38]; and the aromaticity, fraction of amorphous carbon, degree of disorder index [12,13,26], alkali index [21,39], micropore surface area, average micropore diameter and microporosity [1,17,21,22,28,40] of the chars. These results are presented in Figures 3 and 4.
Figure 3: Relationship between the initial reactivity of the chars and the coals’ and chars’ properties at 100% and 50% CO2, 0.875 bar. Initial reactivity of the chars versus (i) Parent coal vitrinite reflectance; (ii) Aromaticity of chars; (iii) Fraction of amorphous carbon in chars; (iv) Degree of disorder index of chars.
7
Figure 4: Relationship between the initial reactivity of the chars and chars’ properties at different CO2 concentrations, 0.875 bar. Initial reactivity of the chars versus (i) Char’s Alkali index of chars; (ii) D-R micropore surface area of chars; (iii) Average micropore diameter of chars; (iv) char’ microporosity.
4.2.3
Char-CO2 Gasification Reaction Modelling
Kinetic parameters determined for the respective chars using the random pore model are presented in Table 2, while, the RPM fitting of the char conversion and the char-CO2 reaction rate curves are presented in Figure 3. Table 2: Summary of structural and kinetic parameters for chars B, C, C2 and D2 Kinetic Parameter ψ (-) m (-) Ea, (kJ·mol-1) k so' , (min-1·bar-1)
Char B 1.11 ± 0.10 0.52 ± 0.11 235.7 ± 32.2
Char C 1.75 ± 0.18 0.67 ± 0.03 231.7 ± 2.2
Char C2 2.58 ± 0.32 0.61 ± 0.07 227.3 ± 17.6
Char D2 1.57 ± 0.20 0.63 ± 0.05 163.3 ± 5.5
5.22 ± 2.0 ·108
4.36 ± 1.9 ·107
1.12 ± 1.0·108
2.69 ±1 ·105 8
The results show that for chars B, C and D2, (ψ < 2), the maximum reaction rate and maximum reaction surface area occurred at X = 0; while char C2 exhibited a maximum at X > 0 [36,37,40]. Pore coalescence was therefore the dominant structural mechanism during the initial stage of gasification for chars B, C and C2. On the other hand, pore growth was more significant during the initial period of char C2 gasification reaction.
(i)
(ii)
Figure 3: RPM fitting to experimental results: (i) Char C conversion at 75% CO2, 0.875 bar, (ii) Char C2 conversion rate at 25% CO2 concentration, 0.875 bar
5.0
Conclusion
The four parent coals were characterised as high ash and inertinite-rich. The transition from coal to char led to the formation of various char carbon forms in the chars. The total reactive components (TRC) generally decreased from coals to the chars, while the total inert components (TIC) exhibited gains in the transition. Except for the maceral contents and inertinite-vitrinite ratios, significant changes in properties were not observed in the four original coals. The subsequent chars, however exhibited differences, both amongst the chars and from the parent coals. Results from the carbon crystallite analysis revealed that the chars are more structurally ordered, more compact and condensed (smaller in size) than the original coals. The increasing orders of magnitude of both the fraction of amorphous carbon and structural disorderliness was found to change from coals to chars. CO2 gasification reactivity of the four chars was found to increase with increasing temperature as well as CO2 composition in the reaction gas mixture. Comparison of the 9
reactivity of the four chars shows that, the reactivity of the chars generally increases in the order: char C2 < char C < char B < char D2. The reactivity of char D2 was found to be higher than the reactivity of the other three chars by a factor > 4. Correlation of the parent coal petrographic properties to the reactivity of the respective chars gave insignificant trends except for the rank parameter, the vitrinite reflectance. Correlation of char properties with char gasification reactivity show systematic and significant trends. Kinetic parameters were determined using the RPM. The activation energy obtained for the char-CO2 gasification reactions were between 163.3 kJ·mol-1 to 235.7 kJ·mol-1; while the order of reaction with respect to CO2 concentration ranged from 0.52 to 0.67. The investigated char-CO2 gasification reactions within the specified operating conditions were found to be kinetically chemical-reaction controlled and were satisfactorily described by the RPM (Regime I).
References [1] [2] [3] [4] [5] [6] [7] [8] [9] [10]
[11] [12] [13] [14] [15]
Laurendeau NM. Heterogeneous kinetics of coal char gasification and combustion. Progress in Energy & Combustion Science 4: 221-270; 1978. Watkinson AP, Lucas JP and Lim CJ. Fuel 1991; 70: 519-527. Yu J-L, Lucas J, Wall T, Liu G, and Sheng C. Modelling the development of char structure during rapid heating of pulverised coal. Combustion and Flame 136: 519-532; 2004. Du Cann VM. Test Report- PSA 2007-016, Petrographics SA, Pretoria, South Africa; 2007. Du Cann VM. Test Report- PSA 2008-040, Petrographics SA, Pretoria, South Africa; 2008. Everson RC, Neomagus HWJP, Kaitano R, Falcon R, Van Alphen C. and du Cann VM. Fuel 2008; 87: 3082-3090. Lu L, Sahajwalla V, Kong C and Harris D. Quantitative X-ray diffraction analysis and its application to various coals. Carbon 39: 1821-1833; 2001. Davis KA, Hurt RH., Yang NYC and Headley TJ. Evolution of char chemistry, crystallinity, and ultrafine structure during pulverized-coal combustion. Combustion & Flame 100:31-40; 1995. Takagi H, Maruyama K, Yoshizawa N, Yamada Y and Sato Y. Fuel 2004; 83: 2427-2433. Trejo F, Ancheyta J, Morgan TJ, Herod AA. and Kandiyoti R. Characterization of asphaltenes from hydrotreated products by SEC, LDMS, MALDI, NMR, and XRD. Energy & Fuels 21:2121-2128; 2007. Maity S and Mukherjee P. X-ray structural parameters of some Indian coals. Current Science 91: 337-340; 2006. Lu L, Kong C, Sahajwalla V and Harris D. Fuel 2002; 81: 1215-1225. Lu L, Sahajwalla V, Kong C and Mclean A. Char structural ordering during pyrolysis and combustion and its influence on char reactivity. ISIJ International 42: 816-825; 2002. Franklin RE. The interpretation of diffuse X-ray diagrams of carbon. Acta Crystallographica 3: 107121; 1950. Ergun S and Tiensuu VH. Fuel 1959; 38: 64-78.
10
[16]
[17] [18] [19] [20] [21] [22]
[23]
[24] [25] [26]
[27] [28] [29] [30] [31]
[32] [33] [34] [35]
[36] [37] [38] [39] [40]
Webb, P. A. (2001). Volume and density determination for particle technologist. Micromeritics Instrument Corp., Norcross, Georgia, USA. http://www.micromeritics.com/Repository/files/density-determination.pdf. (Accessed 22-08-2010). Dutta S, Wen CY. and Belt RJ. Reactivity of coal and char: I. In carbon dioxide atmosphere. Ind. Eng. Chem. Proc. Des. Dev. 16: 20-30; 1977. Hampartsoumian E, Murdoch PL, Pourkashanian M and Trangmar DT. The reactivity of coal chars gasified in a carbon dioxide environment. Combustion Science and Technology 92: 105-121; 1993. Kyotani T, Kubota K, Cao J, Yamashita H and Tomita A. Combustion and CO2 gasification of coals in a wide temperature range. Fuel Processing Technology 36: 209-217; 1993. Ye DP, Agnew JB. and Zhang DK. Fuel 1998; 77: 1209-1219. Zhang L, Huang J, Fang Y and Wang Y. Gasification Reactivity and Kinetics of Typical Chinese Anthracite Chars with Steam and CO2. Energy & Fuels 20: 1201-1210; 2006. Çakal GÖ, Yücel H and Gürüz AG. Physical and chemical properties of selected Turkish lignites and their pyrolysis and gasification rates determined by thermogravimetric analysis. J. Anal. Appl. Pyrolysis 80: 262-268; 2007. Zhang J-W, Zong Z-M, Wang T-X, Xie R-L, Ding M-J, Cai K-Y, Huang Y-G, Gao J-S, Wu Y-Q and Wei X-Y. Reactivities of Shenfu chars towards gasification with carbon dioxide. Journal of China University of Mining & Technology 17: 197-200; 2007. Everson RC, Neomagus HWJP, Kasaini H and Njapha D. Fuel 2006; 85:1076-1082. Everson RC, Neomagus HWJP, Kaitano R, Falcon R and du Cann V M. Fuel 2008; 87: 3403-3408. Wu S, Gu J, Zhang X, Wu Y and Gao J. Variation of Carbon Crystalline Structures and CO2 Gasification reactivity of Shenfu coal chars at elevated temperatures. Energy & Fuels 22: 199-206; 2008. Czechowski F and Kidawa H. Reactivity and susceptibility to porosity development of coal maceral chars on steam and carbon dioxide gasification Fuel Processing Technology 29: 57-73; 1991. Senneca O, Salatino P and Masi S. Fuel 1998; 77: 1483-1493. Kajitani S, Suzuki N, Ashizawa M and Hara S. Fuel 2006; 85: 163-169. Zhang Y, Hara S, Kajitani S and Ashizawa M. Fuel 2010; 89: 152-157. Radovic LR, Steczko K, Walker PL Jr. and Jenkins RG. Combined effects of inorganic constituents and pyrolysis conditions on the gasification reactivity of coal chars. Fuel Processing Technology 10: 311-326; 1985. Matsui I, Kunii D and Furusawa T. Study of char gasification by carbon dioxide: I. Kinetic study by thermogravimetric analysis. Ind. Eng. Chem. Res. 26: 91-95. Fu W-B and Wang Q-H. A general relationship between the kinetic parameters for the gasification of coal chars with CO2 and coal type. Fuel Processing Technology 72: 63-77; 2001. Ochoa J, Cassanello MC, Bonelli PR and Cukierman AL. CO2 gasification of Argentinean coal chars: a kinetic characterization. Fuel Processing Technology, 74: 161-176; 2001. Sinağ A, Sinek K, Tekeş AT, Misirlioğlu Z, Canel M and Wang L. Study on CO2 gasification reactivity of chars obtained from Soma-Isıklar lignite (Turkey) at various coking temperatures. Chemical Engineering and Processing, 42: 1027-1031; 2003. Bhatia SK, and Perlmutter DD. A random pore model for fluid-solid reactions: I. isothermal, kinetic control, AIChE Journal, 26: 379-386; 1980. Murillo R, Navarro MV, López JM, García T, Callén MS, Aylón E and Mastral AM. Activation of pyrolytic lignite char with CO2- kinetic study. Energy & Fuels, 20: 11-16; 2006. Cloke M and Lester E. Fuel 1994; 73: 315-320. Sakawa M, Sakurai Y and Hara Y. Fuel 1982; 61: 717-720. Liu G, Benyon P, Benfell KE, Bryant GW, Tate AG, Boyd RK, Harris DJ and Wall TF. Fuel 2000; 79: 617-626.
11
CHARACTERISATION AND CARBON DIOXIDE GASIFICATION KINETICS OF HIGH ASH INERTINITE RICH SOUTH AFRICAN COALS Rufaro Kaitano*1, Raymond C Everson1 and Hein W J P Neomagus1, 1
Energy Systems Research Group, School of Chemical and Minerals Engineering, North-West University, Potchefstroom Campus, Private Bag X6001,Potchefstroom 2520, South Africa * Corresponding author: Tel: + 27 18 299 1664 Fax: + 27 18 299 1535 E-mail:
[email protected]
Keywords: High ash Coal; Inertinite, Charaterisation. Dense Chars and Random Pore Model ______________________________________________________________________________ Abstract Characterisation and carbon dioxide-char reaction kinetics of a typical South African low grade coal with an ash content of an average of 35 wt % studies were conducted. Fundamental knowledge of the reaction kinetics for char conversion at reactions conditions typical of fluidised bed gasification and combustion was obtained.The random pore model was used to describe the experimental results and the reaction rate is chemical reaction controlled. Introduction South Africa has several millions of tons of low grade coal discards which are thought to play a significant role in the country’s energy mix. However, it has been suggested that to meet the current ever tightening fossil fuel usage conditions, the route to be followed in utilisation these coal deposits is gasification. A number of gasification experiments have been identified for investigation in the laboratory, (steam, and carbon dioxide); this work investigates the behaviour of these coals in the presents of different concentrations of carbon dioxide/nitrogen. To have an insight into the characteristics of the coal traditional and advanced charaterisation investigations were carried out. Experimental The experimental apparatus used for the determination of the reactivity of the chars in this investigation was a TGA, model Bergbau-Forshung GMBH7, supplied by Deutsche Montan Technologie (DMT), Germany. This apparatus has also been used with success by many other investigators for coal/char conversion studies (Johnson J.L., 1981, Calo and Suuberg, 1999; Kajitani, 2006). The gas mixtures were varied between 100% and 20% CO2, balance N2. Advanced characterisation was carried out by various accredited laboratories in South Africa such as Secunda Coal Laboratory of South African Bureau of Standards Laboratories (SABS) and petrographic analysis was carried out by Coal and Mineral Technologies (Pty) Ltd,
Data acquisition
Microbalance
Purge
Sample lock
Gas mixer
Mass Flow Controllers
F
F
F
Pressure control
Reactor
valve
O2
CO 2
N2
Figure 1.1 Schematic presentation of the experimental set-up Results and discussion A typical experimental result obtained from the Thermogravimetric analyser at 900°C, 87.5 kPa and 100% carbon dioxide is given in Figure 1.2, where both the mass of the char sample and the reaction temperature are given as a function of time. 25 900
Mass (mg ) Temp.
Carbon
600
o
15
Temp ( C)
Char Mass (mg)
800
20
400
10 Ash
5
200
0 0
50
100 Time, t (min)
150
200
Figure 1.2: Isothermal gasification of coal char at 900˚C in 100% CO2 at 87.5kPa Model equations The overall reaction rate is: (Bhatia and Perlmutter, 1980)
dX rs (1 − X ) S o 1 − ψ ln(1 − X ) = (1 − ε O ) dt
(1.1)
ψ being the structural parameter characteristic of the initial char structure and defined as:
ψ=
4π L 0 (1 − ε ) ο 2 S ο
(1.2)
By introducing a dimensionless parameter τ , defined by Kaitano, (2007) and Everson et al., (2008a) as: r S t τ= s o 1- ε o
(1.3)
Equation (1.1) can be rewritten in the dimensionless form as follows: dX = (1 − X) 1 − ψln(1 − X) dτ
(1.4)
Integration of Equation 1.4 gives the carbon conversion X, as a function of time t, or the dimensionless time τ (implicit and explicit).Relationships for carbon conversion X in terms of time t or dimensionless time τ (implicit and explicit) obtained by integration of the above equations are as follows: In terms of time t: t=
2(1 − ε o ) ( 1 − ψln(1 − X) − 1) rs S o ψ
(1.5)
In terms of dimensionless time τ τ=
2
ψ
( 1 − ψ ln(1 − X ) − 1)
(1.6)
and explicitly as X = 1 − exp[−τ (1 +
Defining the time factor as:
ψτ 4
)]
(1.7)
tf =
rs S O (1 − ε O )
(1.8)
Equation (1.6) in terms of time t becomes X = 1 − exp[ −t f t (1 +
ψt f t 4
)]
(1.9)
A reduced time t / t X with t X the time for a fractional conversion of X, being the upper limit for reliable experimental results, can be defined, which is independent of the parameters appearing before the square root term (Equation (1.5)) and only dependant on ψ as shown in equation (1.10) where t 0.9 is the time for 90% conversion. Thus, all results for a particular coal-char gasified or combusted should be the same, which enables this property to be used for determination of the structural parameter. The numerical value of the structural parameter can be evaluated by regression of the experimental data using equation 1.10. Solver was used to iterate the value of ψ . t t 0.9
=
It should be noted that
1 − ψ ln(1 − X − 1 1 − ψ ln(1 − 0.9) − 1 t t
0.9
=
(1.10)
τ τ
0.9
Intrinsic reaction rate equation rs given by Equation (1.11) for both gasification and combustion based on an nth order power rate relationship as discussed in Section 1.2 was used. In this investigation the dependence for gasification and combustion were based on the partial pressure of carbon dioxide and oxygen concentration respectively. The equations are: Gasification:
rs = k SO exp(− E / RT ) p n
Combustion:
rs = k SO exp(− E / RT )C n
(1.11)
Figure 1.4 shows a comparison between experimental results and model predictions, a plot of conversion, X vs t/t0.9 shows that the structural parameter value is valid for description of the experimental data.
1
Conversion, X (-)
0.8
0.6
0.4 0.2 0 0
0.2 M odel 87.5kPa 87.5kPa 87.5kPa 87.5kPa 87.5kPa 87.5kPa
0.4
t/t0.9
100% 863C 100% 875C 60% 875C 80% 888C 80% 900C 40% 900C
0.6 87.5kPa 87.5kPa 87.5kPa 87.5kPa 87.5kPa 87.5kPa
0.8
1
80% 850C 80% 863C 80% 875C 100% 888C 100% 900C 60% 900C
Figure 1.4: Comparison of gasification experimental and model results at 87.5 kPa
Table 1.1 gives the evaluated parameters which are comparable with literature on gasification available, Kajitani, 2006. Table 1.1: Intrinsic reaction rate parameters of carbon dioxide gasification Kinetic Parameter Calculated Value Activation energy, E
229 (±37) kJ mol-1
Pre-exponential factor (lumped), kso
9.6·108(±2) s-1kPa-1
Partial pressure dependency, n
0.50(±0.05)
Structural parameter, ψ
1.04(±0.39)
Conclusion The results are consistent with carbon dioxide gasification literature. It was found that the reaction rate increases with increasing temperature, increasing pressure and increasing carbon dioxide concentration. The char conversion kinetics could well be described with the random pore model in the absence of diffusion limitation with a structural parameter, ψ equal to 1.04. The structural parameter was
determined by a regression technique using a reduced time parameter, which eliminates the effect of intrinsic reaction rate kinetics. The intrinsic kinetics could be accurately described with a nth order power law in combination with the Arrhenius equation and an order in carbon dioxide of 0.5, an activation energy of 229 kJmol-1, and a pre-exponential factor of 9.6·108 s-1kPa-1 were determined. The activation energy is relatively high which could be attributed to the high inertinite content of the coal. Although the ash content was high, no significant catalytic effect of the ash could be observed.
Acknowledgements Eskom and the National Research Foundation (NRF - financial support for this project. Mr. Jan Kroeze, Mr. Hennie van Zyl and Mr Adrian Brock - experimental apparatus maintenance. Ms Vivien du Cann and Dr Chris van Alphen - characterisation work and the interpretation thereof.
References: BHATIA, S.K. AND PERLMUTTER, D.D. (1981). A random pore model for fluid solid reactions: II. Diffusion
and Transport Effects, AIChE Journal 27(2):247. CALO, J.M. AND SUURBERG, E.M. (1999). High pressure/temperature thermogravimetric apparatus report. Brown University. U.S.A. EVERSON, R.C., NEOMAGUS, H.W.J.P., KAITANO, R., FALCON, R. AND du CANN, V. M. (2008a). Properties of high ash coal-char particles derived from inertinite-rich coal: II. Gasification kinetics with carbon dioxide. Fuel, 87: 3403-3408. JOHNSON J.L. (1981). Fundamentals of coal gasification, In: Chemistry of Coal Utilisation, Second Sup. Volume, Wiley Inter-science, New York: 1591-1598. KAITANO, R. (2007). Characterisation and reaction kinetics of high ash chars derived from inertinite–rich coal discards. Doctoral thesis. North-West University, South Africa. KAJITANI, S. SUZUKI, N. ASHIZAWA, M. AND HARA, S. (2006). CO2 gasification rate analysis of coal char in entrained flow coal gasifier. Fuel 85:539
Program topic: Coal gasification Gasification kinetics of coal char using direct measurement of particle temperature Ryanggyoon Kim1, Ho Lim2, Cheoloong Kim3, Juhun Song4, and Chunghwan Jeon*
1. Presenter, Doctoral Course, School of Mechanical Engineering, Pusan National University, Pusan Clean Coal Center, Republic of Korea, Ph: 82-51-510-3035, Fax: 82-51-582-9818, Email:
[email protected] 2. Doctoral Course, School of Mechanical Engineering, Pusan National University, Pusan Clean Coal Center, Republic of Korea, Ph: 82-51-510-3035, Fax: 82-51-582-9818, Email:
[email protected] 3. Senior Researcher, Energy Technology Team, GS Engineering & Construction CO., Ltd., Republic of Korea, Ph: 82-2-728-3276, Fax: 82-2-728-3544, Email:
[email protected] 4. Assistant Professor, School of Mechanical Engineering, Pusan National University, Pusan Clean Coal Center, Republic of Korea, Ph: 82-51-510-7330, Fax: 82-51-512-5236, Email:
[email protected] * Corresponding author, Associate Professor, School of Mechanical Engineering, Pusan National University, Pusan Clean Coal Center, Republic of Korea, Ph: 82-51-510-7324, Fax: 82-51-5829818, Email:
[email protected]
Gasification kinetics of coal char is obtained by using the semenov’s thermal spontaneous theory. The critical condition of thermal ignition occurs when the rates of heat generation and heat loss are equal and, in addition, when the derivatives of both these rates with respect to temperature are also equal. Especially, rate of heat generation and rate of heat loss are increased at the same time under elevated ambient pressure condition due to high density of ambient gas around coal char. Thus, in order to describe effect of elevated total system n pressure at rate of generation, the nth-order rate equation( R = kPCO 2 ) was modified to be n m R = kxCO 2 Ptotal where the correlation exponent, m, to modify the effect of the pressure was determined by linear fitting. Convection and radiation were used to explain the rate of heat Gr = ( P∞ ρ D 2 / μ 2 )1/ 2 loss which was considered by using the modified Grashof number( ,m ) under natural convection which is developed to describe the effect of elevated ambient pressure on heat loss. The ignition temperature of coal char particle is obtained by a direct measurement of the particle temperature with photo detector as well as by means of a solid thermocouple which is used as both a heating and a measuring element. The ignition temperatures for subbituminous coal wira of 0.8mm diameter have been measured at 7 different ambient pressures in the ranging from 1 to 15bar in constant volume chamber. The result shows that the coal char ignition temperature deceases with increasing ambient pressure up to the critical pressure condition(15bar) due to an enhanced rate of heat generation. This kinetics is in much closer agreement with the results of other investigators.
Oviedo ICCS&T 2011. Extended Abstract
DIRECT CTL: Innovative Analyses for High Quality Distillates A. Quignard, N. Caillol, N. Charon, M. Courtiade, D. Dendroulakis IFP Energies nouvelles, Rond-point de l'échangeur de Solaize, BP 3, 69360 Solaize, France Contact Information:
[email protected]
Abstract Distillate liquid yields from high hydrogen pressure catalytic conversion of coal processes, called Direct Coal Liquefaction (DCL), are typically high at 4 to 5 bbl/T coal on a dry ash free basis for the best available DCL processes, making them an attractive option to produce transportation fuels from coal. These yields are significantly higher than using the so called Indirect Coal to Liquid (ICL) route, i.e. gasification plus Fisher Tropsch (FT) synthesis. Nevertheless, DCL products are often considered as relatively low quality products and their chemical structure is not well known.
This work focuses on the physical / chemical standardized analyses and innovative detailed characterization of the properties and the unique composition of jet fuel and Diesel cuts obtained by DCL before and after hydroprocessing. It shows that 100% high quality fully desulphurized Jet A, Jet A-1 or JP-8 aviation fuels could be obtained when using the appropriate hydrocracking conditions. It also shows that the Diesel cut obtained from the same upgrading process can be used as a high quality base for transportation fuels with less than 5 ppm sulfur, excellent cold flow properties and good combustion characteristics, with a very specific chemical structure. This innovative detailed characterization of hydroprocessed DCL jet fuel and Diesel cuts was provided using a GCxGC method developed within IFP Energies nouvelles (IFPEN) laboratories.
1. Introduction The worldwide demand of fuels has been intensified in recent years and is expected to continue growing. To satisfy these energy requirements and diversify the source of fuels, the energy industry has to face the challenge of using alternative resources in order to produce transportation fuels while ensuring an increased predicted demand of aviation fuels and Diesel. Direct Coal Liquefaction (DCL) process, such as Axens H-CoalTS,
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Oviedo ICCS&T 2011. Extended Abstract
enables higher liquid yields than the Indirect Liquefaction (ICL) process (syngas production via coal gasification followed by Fisher Tropsch synthesis). Nevertheless, DCL products are often considered as relatively low quality products and their chemical structure is not well known. DCL main simplified process is presented in Figure 1. Usually in a typical mouth mine DCL unit, the hydrogen required for the coal high pressure catalytic hydroconversion unit and the downstream hydroprocessing of the DCL products is produced from a coal gasification unit. Figure 2 reminds a typical DCL product upgrading route with a high pressure catalytic hydroprocessing (typically hydrotreatment plus hydrocracking) of the full DCL product followed by reforming and isomerization of the hydroprocessed naphtha cuts. This work focuses on the characterization of physical and chemical properties and composition of jet fuel and Diesel cuts obtained by DCL before and after hydroprocessing using high pressure hydrotreatement (HDT) and hydrocracking (HCK), and how analytical and correlative methods are modified or completely changed by using new innovative analytical techniques. It also shows that high quality fuel components can be obtained using appropriate hydrotreating or hydrocracking conditions [1-3]. Product Treating
Coal
Sulfur Ammonia
Coal Liquefaction
Coal Preparation
Gasification/ASU
H2
Liquid Product
Product Upgrading
CO2 Capture
Figure 1. DCL main process block Raw DCL products H2
Isomerization
HDT/HCK
Catalytic Reforming
ULS Gasoline
Gasoline Blending
Butanes EtOH ULS Diesel (CN > 50) High Quality Jet Fuel
Figure 2. DCL upgrading scheme via hydroprocessing
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Oviedo ICCS&T 2011. Extended Abstract
2. Experimental section The upgrading of rough direct coal liquefaction effluents (C5-350/450°C DCL cuts) was carried out in a 1L bench unit at IFPEN facilities. The upgrading experiments were run uninterrupted under severe HDT and HCK operating conditions: total pressure of more than 120 bar, relatively moderate temperatures and quite low hourly space velocities (LHSV). The used LHSV are among the lower ones used for HDT/HCK processes in order to enhance the hydrogenation activity on such a specific feedstock with high aromatic molecules and impurities (i.e; N and O) contents. IFPEN HDT/HCK bench unit well correlates with industrial reactors, as previously demonstrated. The industrial HDT and HCK catalysts used in these experiments were base metal catalysts, in situ sulfided on the unit . After reaction, effluents were cooled, condensed and separated into a gas phase and a liquid phase. Liquid phase effluents were distilled by physical distillation (True Boiling Point - TBP) according to ASTM D2892 method and characterized using standard petroleum analyses and multidimensional gas chromatography (2D-GC or GCxGC) equipped with a flame-ionization detector (FID) and a cryogenic system [4].
3. Results and Discussion 3.1 Physical and chemical properties of DCL products. The main physical and chemical analyses of an IBP-380°C rough DCL cut, before and after high pressure HDT as well as after HCK are presented in Table 1. The rough DCL product contains a high level of nitrogen and oxygen with about 50% aromatics, almost no paraffins, resulting in a relatively low hydrogen level at 11.2 wt%. On the other hand, HDT and HCK total liquid products exhibit a much higher hydrogen content with no impurities and a much lower aromatic content. These cuts are almost made of cyclo-paraffins with 80wt% or more napthenes and a remaining low normal + iso-paraffin content. The hydrogen level is enhanced up to 13.3 wt% on the HCK product. This is a unique structure, never seen in petroleum products, with maybe the exception of hydrocracked light cycle oil (HCK LCO) from catatytic cracking (FCC).
Table 1. Physical and chemical properties of raw DCL, HDT and HCK products Cut
Unit
DCL C5380°C
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HDT DCL
HCK
3
Oviedo ICCS&T 2011. Extended Abstract
C5+
DCL C5+
0.9255
0.8805
0.8685
d154
g/cc
Kinematic viscosity at 20°C
mm2/s
8.05
5.2
4.05
Carbon
wt %
87.26
87.3
86.72
Hydrogen
wt %
11.24
12.7
13.28
Oxygen
wt %
1.28
Nitrogen
wt ppm
2100
2
<0.3
Basic Nitrogen
wt ppm
1112
Sulfur
wt ppm
102
9
1.5
Aromatic Carbon NMR
wt %
25.5
6.0
4.1
Paraffins (nP + iP)
wt %
4
5
11.4
Naphthenes
wt %
47
79.5
83.8
15.5
4.8
Aromatics
wt % 49 TBP ASTM D2892 Distillation
Naphtha (IBP-150°C)
wt %
4
6
9
Jet (150-220°C)
wt %
16
16
18
AGO (220-380°C) wt % (1)
80
78
73
3.2 Cetane number prediction by Near InfraRed (NIR). Cetane numbers were determined on Direct Coal Liquefaction (DCL) mid-distillates before and after upgrading using the standard method ASTM D 613 which is based on CFR engine measurements. These mid-distillates exhibit CFR cetane numbers higher than 30, going up to 50 for high pressure hydrocracked (HCK) DCL distillates. However, the calculated cetane indices (ASTM D 4737A/B) of DCL products are very low when compared to the CFR cetane numbers, as illustrated in Table 2. Moreover, all the high pressure hydrotreated (HDT) and HCK DCL mid-distillates were out of the current NIR model's prediction database used at IFPEN, with a strong underestimation of the CFR cetane number of about 10 points. Figure 3 depicts the NIR initial ("old") IFPEN data base with a standard calibration (blue point) based on petroleum products from a wide array of origin, in the three main components axes (CP1, CP2 and CP3) representing the NIR spectra. Although borderline, the aforementioned hydrocracked light cycle oil (HCK LCO) from catalytic cracking (FCC) are within the domain. Red points are corresponding to DCL rough
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Oviedo ICCS&T 2011. Extended Abstract
products, green points to the HDT DCL products and black points the HCK DCL ones. Red circle points are representing DCL cuts, either from rough DCL or HDT / HCK origin with measured CFR, that were used to calibrate the new basis. It clearly appears from Figure 3 that all DCL products are fully outside of the petroleum distillate region because they exhibit very specific chemical structures in comparison to conventional petroleum products used for the initial NIR calibration database. So cetane values estimated by the ASTM D 4737 method and by NIR analysis are largely underestimated because both correlative methods and the analytical protocol that were used to calibrate them are not adapted to this type of products. Thus a new model for better estimation for DCL-derived products by NIR was developed whithin IFPEN using the corresponding CFR cetane number measurements. As it can be seen in Table 2, the results of the new model are clearly improved, with a satisfactory prediction on most of the points. This study points out the strong need for detailed and well-adapted molecular analysis of upgraded DCL mid-distillates in order to interpret their unexpectedly good combustion qualities. Standard Calibration LC in calibration intial LC HDT HCK
-3
x 10 4
CP3 varexp=3
3 2 1 0 -1 -2
-8 -6
-4
-4
-2
-2
-3
0
0
x 10
-3
x 10
2
CP1 varexp=84
CP2 varexp=8
Figure 3. ACP representation of the NIR standard calibration database, DCL middistillates, HDT DCL and HCK DCL mid-distillates
Table 2. Cetane prediction on DCL products before and after upgrading CN CFR
CCI
(ASTM D 613)
(ASTM D 4737)
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CCI NIR (IFPEN
(IFPEN
old model)
new model)
5
Oviedo ICCS&T 2011. Extended Abstract
DCL
31
25.5
27
32.8
HP HDT DCL
35
27.5
26
36
MP HCK DCL
39
32.4
27
36.5
HP HCK DCL
50
38.4
35.5
45.3
3.3 Comprehensive analysis of DCL products before and after upgrading. The chemical composition of DCL kerosene and atmospheric gasoil cuts, before and after upgrading, was investigated. Since the DCL mid-distillates contain high contents of naphteno-aromatic hydrocarbons, the in-house Fitzgerald based mass spectrometry method, based on petroleum cuts and currently used in IFPEN for the analysis of petroleum atmospheric distillates, cannot be successfully applied to DCL products. As a consequence, misleading results are obtained, particularly with a significant underestimation of paraffinic compounds and an overestimation of aromatics contents. Owing to higher resolution power and enhanced sensitivity when compared to onedimensional GC, comprehensive two-dimensional gas chromatography (GC × GC) is a powerful tool for improving characterisation of petroleum samples (naphtas [5], middledistillates [4, 6-8], vacuum gasoil cuts [9, 10]). In this work, a GCxGC-FID hyphenated technique was used to provide more information about the chemical composition of DCL distillates and upgraded products (Figures 4 to 6). It should be mentioned that due to the specific composition of the products, naphtenes, naphteno-aromatics and monaromatics may be co-eluted in GCxGC analysis, making necessary a preliminary separation between saturated and non saturated molecules by off line Liquid Chromatography (LC) or in line Super Critical Chromatography (SFC). While the DCL kerosene exhibits a naphtheno-aromatic structure with almost no paraffins (Table 3), the HCK DCL kerosene almost entirely consists of saturated hydrocarbons, i.e. normal, iso-, and cyclo-paraffins, most of the saturates being condensed cyclo-paraffins (50-60%). It presents a unique and very specific chemical structure with almost no aromatics, below 5%. As shown in Table 3, the high naphtenes and poly-naphthenes content exhibits good combustion properties provided that aromatics content is low. DCL HCK complies with all the military JP-8 and civil Jet A1 aviation fuel specifications, including a good compatibility with polymer O-ring materials with a good swelling behaviour comparable to petroleum jet fuel. This is not the case for pure Synthetic Paraffinic Kerosene (SPK) from indirect coal gasification and FT synthesis. DCK HCK jet also has a high volumetric heating value, excellent cold flow properties and a very good thermal Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
stability and lubricity value. Similarly to the kerosene cuts, the gasoil cuts after hydrotreatment or hydrocracking are also very polynaphthenic and contain low aromatics contents. They meet the commercial diesel fuel specifications (Table 5) with a pretty goodcetane number, excellent cold flow properties and no sulfur. The only non-satisfied specification concerns the density for the European market because of the extremely naphthenic structure of the product. More information about the detailed chemical composition of upgraded DCL jet fuel and Diesel cuts are provided thanks to comprehensive GCxGC technique, making it mandatory to explain the good combustion qualities of HDT and HCK DCL middistillates. When compared to the usual Mass Spectrometry method, GCxGC analysis gives more accurate data and much more detail on the naphthenes and poly-naphthenes families, probably responsible for the good / very good combustion behaviour of these distillates which was not initially predicted, as well as their excellent cold flow properties. More work related to this detailed molecular analysis is needed in order to try to understand why these naphthenes exhibit such good properties, maybe using molecular computing techniques since there is no experimental data on these molecules ever obtained by chemical synthesis as pure components.
Figure 4. GCxGC chromatogram of rough DCL jet fuel (180-250°C) cut
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Oviedo ICCS&T 2011. Extended Abstract
naphtenes
Di aromatic and more
naphteno aromatics
paraffins
Figure 5. GCxGC of HDT DCL 180-250°C cut (left) and 250-380°C cut (right)
naphteno aromatics paraffins naphtens
Di aromatic and more
Figure 6. GCxGC of HCK DCL 180-250°C cut (left) and 250-380°C cut (right)
Table 3. GCxGC hydrocarbon composition of DCL products before and after upgrading DCL
DCL
HDT
HDT
HCK
HCK
KERO
AGO
KERO
AGO
KERO
AGO
3.8/2.2
4.8/1.4
3.7/1.8
4.3/1.5
2.9/5.2
5.1/7.8
Naphtenes
56.3
52.4
72.9
64.2
88.9
82.8
Aromatics
37.7
23.6
21.6
29.6
3.0
4.3
Jet A1
JP-8
775-840 > 38 < -47 > 42.8
775-840 > 38 < -47 > 42.8
n/i Paraffins
Table 4. Physical and chemical properties of HCK DCL kerosene Specifications Specific gravity Flash Point (°C) Freezing Point (°C) Heat of combustion (MJ/Kg)
DCL + HCK KERO 835- 840 50-55 < - 65 43.2
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Average Jet A /A1, JP-8 [11] 800-810 47-52 -50 43.1
8
Oviedo ICCS&T 2011. Extended Abstract
Heat of combustion (MJ/L) Sulphur (wt %) Viscosity @ -20°C (cSt) Aromatics (vol %) Smoke Point (mm) + Naphthalenes (vol %) Hydrogen Content (wt %) BOCLE (mm) JFTOT ΔP @260°C (mmHg) Tube deposit rating (visual)
36.3 < 1 ppm <8 <5 22 <1 13.5 – 14.0 0.6 <1 1
34.5 0.02 4.0-5.0 17-18 1.5 14.1-14.3 0.6-0.65
< 0.3 <8 < 25 > 19 <3 < 0.85 < 25 <3
< 0.3 < 25 > 19 <3 > 13.4
Table 5. Physical and chemical properties of HCK DCL diesel Specifications
DCL Diesel
Density (kg/m3) Sulfur (ppm) Aromatics (wt%)
855 - 885 <5 2 - 10
Diaros+ (wt%) Pour Point (°C) Cloud Filterability Plugging Point (CFPP) Cetane Number Cetane Index ASTM D4737
USA Spec. (Fuel N°1) < 15 < 35
EU Spec. > 820 & < 845 < 10 -
0-2 < - 48 <- 40
-
< 11 (< 8) < - 15 < - 18(Winter)
45 - 54 not relevant
> 40 > 40
> 51 > 46
4. Conclusions This work, based on the physical / chemical standardized analyses and innovative detailed characterization, demonstrates the unique properties and composition of jet fuel and Diesel cuts obtained by DCL followed by an adapted high-pressure hydroprocessing. It shows that 100% high quality civil Jet A / Jet A-1 or military JP-8 aviation fuel can be obtained when using the appropriate hydrocracking conditions. It also shows that the Diesel cut can be used as a high quality component for transportation fuels. In this work a powerful detailed characterization of these mid-distillate cuts was also developed, based on an innovative hyphenated LC or SFC-GCxGC method developed within IFPEN laboratories. More information about the detailed chemical composition of upgraded DCL jet fuel and Diesel cuts are provided thanks to comprehensive GCxGC technique, making it mandatory to explain the good qualities of hydroprocessed DCL mid-distillates. More work related to this detailed molecular analysis is needed in order to try to understand why these naphthenes exhibit such good properties, maybe by using molecular computing.
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Oviedo ICCS&T 2011. Extended Abstract
Acknowledgement: S. Hénon, C. Dartiguelongue, W. Weiss
References [1] Dulot H., Quignard A., Charon N., Dutriez T., Courtiade M., ISCRE 21, 2010, Philadelphia, US [2] A. Quignard et al, World CTL, 2010, Beijing, China [3] W. Weiss, H. Dulot, A. Quignard, N. Charon, M. Courtiade, N. Caillol , International Pittsburgh Coal Conference, 2010, Istanbul, Turkey [4] Vendeuvre C., Ruiz-Guerrero R., Bertoncini F., Duval L., Thiebaut D., Hennion M.C., J.Chromatogr. A, 2005;1086:21-28 [5]. Adam F., Vendeuvre C., Bertoncini F., Thiebaut D., Espinat D., Hennion M.C., J. Chromatogr. A, 2008;1178:171-177 [6] Adam F., Thiebaut D., Bertoncini F., Courtiade M., Hennion M.C., J. Chromatogr. A, 2010;1217:1386-1394 [7] Adam F., Bertoncini F., Dartiguelongue C., Marchand K., Thiebaut D., Hennion M.C., Fuel, 2009;88:938-946 [8] Omaïs B., Courtiade M., Charon N., Thiébaut D., Quignard A., Hennion M.C., J. Chrom. A, 2011; 1218:3233-3240 [9] Dutriez T., Courtiade M., Borras J., Charon N., Quignard A., Dulot H. et al, Spectra Analyse, 2010; 276:34-37 [10] Dutriez T., Courtiade M., Thiebaut D., Dulot H., Bertoncini F., Vial J. et al, J. Chromatogr. A, 2009;1216,2905-2912 [11] Hadaller O.J., Johnson J. M., World Fuel Sampling Program, CRC Report No. 647, 2006
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Oviedo ICCS&T 2011. Extended Abstract
Direct coal-liquid hydrogenation over NiMoNX /γ-Al2O3 catalysts Qi Chu, Lin Guo, Jie Feng1, Wenying Li, Kechang Xie Key Laboratory of Coal Science and Technology (Taiyuan University of Technology), Ministry of Education and Shanxi province, Taiyuan 030024, China 1. Introduction The coal-liquid from direct coal liquefaction (DCL) enriched roughly 60 vol.% aromatic compounds, which not only declined its fuel quality as jet fuel and vehicle fuel, but also caused the environmental pollution. Therefore, the coal-liquid refining is the essential step in the pursuit of higher oil quality. However, for the hydrogenation of the coal-liquid from DCL, the problem is if the hydrogenated aromatics make the cycle-breaking to be the saturated hydrocarbon, the cetane number and density of coal-liquid will be depressed. How to effectively control the aromatic hydrogenation extent is the key point of DCL refining. Many kinds of catalysts are used in the process of the coal-liquid refining, such as metal sulfur catalyst[1], noble metal catalyst[2, 3]. But all these have some problems in hydrogenation process. For instance, metal sulfur catalysts not only perform high activity under harsh operating conditions, but need presulfiding before the catalysts are used. And although noble metal catalysts possess high hydrogenation activity under mild conditions, they are easy to suffer poisoning from sulfur and nitride sources in either petroleum or coal-derived oil. In recent studies, the metallic nitride can balance the relationship between higher activity and the tolerance on sulfur and nitride sources under mild conditions, and metallic nitrides of transition metals show the similar catalytic activity as noble metal and better sulfur and nitride
resistances
in [4]
hydrodesulfurization
the
reactions
such
as
Fischer-Tropsch
reaction,
[5]
and hydrodenitrogenation , especially bimetallic nitrides of
transition metals. In this paper, the hydrogenation of direct coal-liquid over NiMoNX/γ-Al2O3 catalysts is investigated and structure properties of NiMoNX/γ-Al2O3 catalysts are discussed in order to obtain some structure information of controlling the aromatic hydrogenation. NiMoNX/γ-Al2O3
catalyst
was
prepared
hexamethylenetetramine (HMT). 2. Experimental section 2.1 Catalyst preparation 1
Corresponding author Email:
[email protected]
in
a
one-step
synthesis
using
Oviedo ICCS&T 2011. Extended Abstract
NiMoNX/γ-Al2O3 catalyst (Cat-1) was prepared by 25wt% NiMoNX loading on γ-Al2O3 using volumetric impregnation. Nickel nitrate hexahydrate (Ni(NO3)2·6H2O), ammonium heptamolybdate ((NH)4Mo7O24·4H2O), hexamethylenetetramine (HMT) with mole ratio of 14:3:20 in 15% NH3·H2O. To get well-distributed surface, the mixture salt solution and supporter was vibrated 4h by ultrasonic apparatus, then natural drying. After drying at 383 K overnight, the solid was calcined in a tube furnace under N2 flow rate of 120 sccm. The furnace temperature was increased at a rate of 5 K/min and kept at 923 K for 2 h. After natural cooling to room temperature, the product was passivated for 6h under O2/N2 (1% v/v) at a flow rate of 100 sccm. NiMoNX/γ-Al2O3 catalyst (Cat-2, Cat-3) with designed loading of 20wt% NiMoNX was obtained in order to investigate the effect of calcined atmosphere for the structure of catalysts. The difference was diverse calcined atmosphere between Cat-2 and Cat-3, air atmosphere and nitrogen atmosphere, respectively. Meanwhile, Cat-2 lacked the passivated step. Other steps were similar as Cat-1. 2.2 Activity evaluation The catalytic activities were measured in a fix-bed reactor. The length and i.d. of the reactor are 480mm and 10mm, respectively. About 3 mL catalyst (0.25-0.38 mm) was packed into the reactor. Before the experiment, the passivated nitride catalysts were reduced in the reactor by H2 at 450°C for 2hrs to remove passivation surface. The reaction was operated at 550°C under 5.0 MPa, China Shenhua DCL was pumped at 3ml/min under H2 circumstance (≥99.999, 30 mL/min). The DCL liquid and hydrogenated products were analyzed by GC-MS (GC6890/MS5973) equipmented with a HP5-MS capillary column. The temperature of injector was 250°C, column temperature was temperature programmed from 80 to 285°C at a rate of 5°C/min. The scan range of MS was 10~550amu. 2.3 Catalyst characterization BET surface area and pore size distribution BET surface area, pore volume and size distribution measurement were obtained using JW-B1212W apparatus. Surface area was calculated according to BET equation, and the BJH method was used in the measure of pore volume and pore size distribution. X-ray diffraction (XRD) The X-ray diffraction was acquired on SHIMADZU diffractometer XRD-6000 with CuKα (λ=1.5406Å) radiation (40 kV, 30 mA). Scan range is from 10° to 80° with the scan rate of 10°/min and sampling pitch is 0.02°.
Oviedo ICCS&T 2011. Extended Abstract
H2-temperature-programmed reduction (H2-TPR) To investigated active site, H2-TPR of samples was carried out by using a Micromeritics Autochem Ⅱ 2920 Chemisorption Analyzer. TPR was carried out in 10% H2/Ar mixture gas with a flow rate of 50 mL/min from 60°C to 800°C and the temperature rate of 10°C/min. TPR profile were recorded by thermal conductivity detector (TCD). 3. Results and Discussion 3.1 Analysis of the coal liquid sample: The analysis results of 180~200 °C fraction are list in ¡Error! No se encuentra el origen de la referencia.. The matching index of the compounds is over 90%.
Table 1 The main compounds of 180~200ºC fraction before and after reaction Compounds
Area% Before
After
Aromatic hydrocarbon
17.6
34.8
Naphthenic hydrocarbon
20.4
9.6
Phenol derivative
34.7
0
Alkane
4.3
4.1
Notes:* contains mono-, di-, aromatic hydrocarbon derivative 3.2 Activity evaluation: Table 1 lists the main component distribution before and after reaction. Comparing with the experimental results before and after reaction, phenol derivative disappear and the content of aromatic hydrocarbon increases from 17.6% to 34.8%. The absence of phenol derivative to benzene derivative is the reason that the content of aromatic hydrocarbon increase after hydrogenation reaction. By the results of main compounds distribution of 180~200°C that is no listed before and after reaction, the diaromatic hydrocarbon change into monoaromatic hydrocarbon after hydrogenating. Meanwhile, methyl cyclopentane, methyl cyclobutane and 3,5-Dimethylcyclopentene are detected. And the evidence of detecting three compounds as mentioned above imply that NiMoNX/γ-Al2O3 not only possesses the hydrogenation capability, but can make the compounds isomerization. 3.3 BET surface area and pore size distribution:
Oviedo ICCS&T 2011. Extended Abstract
Table 2 Structural parameters of samples Sample
γ-Al2O3
Cat-1
Cat-2
Cat-3
ABET/(m2/g)
157.6
132.6
111.7
130.3
Pore volume(cm2/g)
0.637
0.512
0.382
0.521
Average pore diameter (nm)
11.307
11.368
9.412
11.503
The structural parameters of catalysts and supporter are shown in ¡Error! No se encuentra el origen de la referencia.. With the loading increasing, there is no obvious
change on BET surface area and average pore diameter from Cat-1 to Cat-3. 3.4 XRD analysis:
♠ Ni3Mo3N ♦ NiO • NiMoO4 ♥ γ-Al2O3
Intensity(a.u.)
♠ ♦
Cat-3
♠ ♠ ♣ ♦ ♠
• •
Cat-2 •
10
♠♠
Cat-1
20
30
♥
♣
♥
♣
♦
40
♣ Mo N 2
♥
50
60
70
80
Degree(2θ)
Figure 1 X-ray diffraction patterns for Cat-1, Cat-2 and Cat-3 To identify nature of bimetallic nitrides, the XRD pattern for the catalysts are shown in Figure 1. In ¡Error! No se encuentra el origen de la referencia., the Cat-2 that is calcined under air atmosphere
indicates that the existence of α-NiMoO4 (card no.33-0948) which is in 20° to 30°. The peak at 2θ=37.5° and 43° indicate the presence of γ-Mo2N (card no. 25-1366) and NiO (card no. 65-2901). For the Cat-1 and Cat-3 which are calcined under nitrogen atmosphere, the XRD patterns show the peak of Ni3Mo3N at 2θ=40.8°, 43° and 45.2° (card no. 49-1336). However, no Ni3Mo3N peak occurs in Cat-2 and the Ni3Mo3N crystallinity of Cat-1 precedes Cat-3. The formation of Ni3Mo3N phase indicate, during the calcined process under inert atmosphere, the solid solid reaction occurs between Ni and Mo2N, producing Ni3Mo3N phase. Figure 1 reveals that the phase of Ni3Mo3N is formed by this method we used and the crystallinity is correlation with loadings. The formation of γ-Mo2N phase in the Cat-2 and Cat-3 implies that nitride phase
Oviedo ICCS&T 2011. Extended Abstract
was contained in the complex precursor from HMT rather than N2. Although Cat-3 calcined under nitrogen atmosphere, N2 did not act as a nitriding agent at setting temperature. Therefore, we can infer that NiMoNX/γ-Al2O3 catalyst was prepared from the decomposition of a molybdate HMT, and HMT acted as a nitrogen source. 3.5 H2-TPR analysis:
Hydrogen consumption(a.u.)
408
Cat-2
726 Cat-3
Cat-1 665
408 257 270 100
200
300
400
500
600
700
800
Temperature (°C)
Figure 2 H2-TPR of bimetallic nitrides H2-TPR (¡Error! No se encuentra el origen de la referencia.) of bimetallic nitrides are used to study the active sites of catalysts surface. The peak value of Cat-1 and Cat-3 which were prepared under N2 atmosphere are lower than Cat-2. For Cat-2, the peak beginning at 300°C and maximizing at 408°C is due to the reduction of surface oxygen on passivated nitride Mo[6]. However, due to the existence of nickel, the reduction peak of surface oxygen on passivated nitride Mo shifts to the lower temperature. It is obvious that Ni promotes the reduction of passivation layer of nitride. And other peak at 726°C could be contributed to the further reduction of bulk molybdenum nitride. According to Figure 1, Cat-2 possesses the NiMoO4 phase. The reduction peak of NiMoO4 always appears at 545°C and 745°C[7]. But there is no the reduction peak in Figure 2. It could be related with poor crystallinity of NiMoO4. Sundaramurthy et al.[8] observed the reduction peak of NiMo nitride around 285°C. Combining with Figure 1, in this paper, however, the reduction peak of NiMo nitride are observed at 257°C and 270°C in Cat-3 and Cat-1. And it leads to a strong interaction of NiMo nitride with γ-Al2O3. Acknowledgment Authors gratefully acknowledge the financial support from the National Basic
Oviedo ICCS&T 2011. Extended Abstract
Research Program of China (Grant No. 2011CB201300 ) References 1. Ferraz, S.G.A., et al., Influence of support acidity of NiMoS catalysts in the activity for hydrogenation and hydrocracking of tetralin. Applied Catalysis A: General, 2010. 384(1-2): p. 51-7. 2. Pedro, C., et al., Kinetic and deactivation modelling of biphenyl liquid-phase hydrogenation over bimetallic Pt-Pd catalyst. Applied Catalysis B: Environmental, 2009. 88(1-2): p. 213-223. 3. Calemma, V., R. Giardino, and M. Ferrari, Upgrading of LCO by partial hydrogenation of aromatics and ring opening of naphthenes over bi-functional catalysts. Fuel Processing Technology, 2010. 91(7): p. 770-6. 4. Santillán-Vallejo, L.A., et al., Supported (NiMo,CoMo)-carbide, -nitride phases: Effect of atomic ratios and phosphorus concentration on the HDS of thiophene and dibenzothiophene. Catalysis Today, 2005. 109(1-4): p. 33-41. 5. Yuhong, W., et al., Characterization and catalytic properties of supported nickel molybdenum nitrides for hydrodenitrogenation. Applied Catalysis A: General, 2001. 215(1-2): p. 39-45. 6. Chu, Y., et al., NiMoNx/γ-Al2O3 catalyst for HDN of pyridine. Applied Catalysis A: General, 1999. 176(1): p. 17-26. 7. Xu, B., C. Zhu, and Q. Li, Surface Species of Molybdenum-Nickel Loaded Catalysts and Their Catalytic Activity. Acta Physico-Chimica Sinica, 1994. 10(6): p. 543-8. 8. Sundaramurthy, V., A.K. Dalai, and J. Adjaye, The effect of phosphorus on hydrotreating property of NiMo/γ-Al2O3 nitride catalyst. Applied Catalysis A: General, 2008. 335(2): p. 204-210.
Oviedo ICCS&T 2011. Extended Abstract
Preparation of activated carbons from direct coal liquefaction residue Jianbo Zhang, Lijun Jin, Bo Qiu, Haoquan Hu* State Key Laboratory of Fine Chemicals, Institute of Coal Chemical Engineering, School of Chemical Engineering, Dalian University of Technology, Dalian 116024, China. * E-mail:
[email protected]. Abstract Activated carbons (ACs) are widely used in many areas, and the surface and structural properties are often necessary to be controlled. In this work, ACs were prepared from direct coal liquefaction residue by KOH activation. SiO2, SBA-15, Al2O3 and Na2SiO3 were used as additives by mixing with the carbon precursor before carbonization, so as to adjust the porous structure. The result confirmed that the porous structure of the carbon was easy to adjust by introducing some additives. And the additive affected the porous structure mainly by forming salts as nanostructured particles with KOH. Nanopores will be created by washing off the space occupied by the mineral matter or the salts in the carbonized carbon. The catalytic performance for catalytic methane decomposition depends on the porous structure of the carbon material. 1. Introduction Activated carbons (ACs) are widely used as adsorbents [1-3], catalysts [4] or catalytic supports [5] because of their highly porous texture and large adsorption capacity. In a specific use, however, the surface and structural properties are often necessary to be controlled. These properties are closely related with both the activation process and the nature of the precursor [6]. Basically, ACs can be prepared by physical and chemical activation [1-3,6] or a mix of both [7]. Due to a lower temperature, higher yield and easier to obtain ACs with a large surface area, chemical activation often arouses more attention. Coal liquefaction residue (CLR), similar as both coal and pitches, seems to be a promising precursor for ACs. It is low cost and easy availability in China. Preparing ACs from CLR can not only effectively utilize CLR, about 30% of raw coal, but provide marketable carbons. In this work, Shenhua CLR was used as the carbon precursor to prepare ACs by KOH activation. It was mixed with certain additives before carbonization, so as to adjust the porous structure of the carbon. Owing to nanostructured silicate can be easily produced by the reaction of KOH and SiO2 [8], new nanopores are expected after washing off the silicate occupied the inner space of ACs.
1
Oviedo ICCS&T 2011. Extended Abstract
The effect of the amount and type of the additives, washing solvents used after carbonization on the surface and structural properties was discussed. And the catalytic activity of the ACs for methane decomposition was also investigated. 2. Experimental section 2.1 Materials. The CLR sample was obtained from a pilot plant for direct liquefaction of Shenhua coal in China. It was crushed and sieved to a particle size range of 150-250 µm before use. Proximate and ultimate analyses of the CLR are shown in Table 1, and the ash composition in the CLR is presented in Table 2. Table 1. Proximate and ultimate analyses of Shenhua CLR. Proximate analysis (wt.%) Ultimate analysis (wt.%)[a] [b] M A V FC C H N S 0.0 21.7 28.7 49.6 86.7 5.2 0.9 3.1 [a] dry ash-free basis. [b] by difference.
Table 2. Chemical compositions of the ash in Shenhua CLR. Composition Fe2O3 SiO2 CaO SO3 Al2O3 Percentage (wt.%) 28.7 27.9 15.0 14.1 11.6
MgO 0.9
O[b] 4.1
Others 1.8
2.2 Preparation of ACs. KOH was used as the activating agent. SiO2 with average pore diameter Dp of 10.0 nm, BET surface area SBET of 725 m2 g-1, total pore volume Vt of 0.95 cm3 g-1, SBA-15 with dp of 5.0 nm, SBET of 950 m2/g, Vt of 1.10 cm3/g, Al2O3 with a particle size of 75-150 µm, and Na2SiO3•9H2O were used as additives, respectively. KOH activation was carried out using an impregnation method. About 5 g of the CLR together with additives was mixed, by stirring for 24 h with a solution that contained 10 g KOH, 50 ml deionized water and 10 ml ethanol at ambient temperature. 1, 3, 4, 5 or 7 g of SiO2 depending on the ratio of SiO2/CLR (Si/R) used (1:5, 3:5, 4:5, 5:5 or 7:5) was used as an additive for the carbon precursor. The resulting slurry was dried at 120 oC overnight in vacuum before carbonization. Then the carbonization was conducted in a horizontal tube furnace under nitrogen atmosphere with a flow rate of 200 ml/min: Heating continuously from room temperature up to 400 oC with a rate of 2 oC/min, to 600 oC with 1 oC/min and keeping for 90 min, then to 900 oC with 2 oC/min and holding for 240 min before cooling. For comparison, some other samples were prepared with SBA-15, Al2O3 or Na2SiO3•9H2O, respectively, as an additive. And another one was
2
Oviedo ICCS&T 2011. Extended Abstract
also prepared without any additive. To evaluate the effect of the washing process, the carbonized samples were washed by different solvents: (1) NW-SiRC, without washing. (2) W-SiRC, washing with deionized water. (3) HCl-SiRC, washing with 2 M HCl followed with deionized water. (4) SiRC, washing with 2 M HF, 2 M HCl and deionized water by sequence to be nearly free mineral. The washing processes of all the other samples were similar to SiRC. All the samples, together with the experimental conditions, are listed in Table 3. Table 3. Experimental conditions for the production of carbon samples. Additive Sample Washing agent Type Mass (g) RC [a] 1SiRC SiO2 1 [a] SiRC SiO2 3 [a] 4 [a] 4SiRC SiO2 5SiRC SiO2 5 [a] 7 [a] 7SiRC SiO2 NW-SiRC SiO2 3 W-SiRC SiO2 3 [b] 3 [c] HCl-SiRC SiO2 SBA-RC SBA-15 3 [a] Al2O3-RC Al2O3 3 [a] Na2SiO3 3 [a] Na2SiO3-RC [a] 2 M HF, 2 M HCl and deionized water. [b] deionized water. [c] 2 M HCl and deionized water.
2.3 Characterization. N2 adsorption/desorption isotherms of the carbons were measured at -196 oC in an Autosorb 2420 physical adsorption apparatus and the information on pore structure was obtained by using BET and BJH methods. SEM (KYKY-2800B) and TEM (Philips Tecnai G2 20) was applied to record the morphology of the samples. The ash content of the samples was determined by burning the carbons at 900 oC in a TG analyzer (Mettler Toledo TGA/SDTA851e). And the ash composition in the carbon was obtained by X-ray fluorescence spectral analysis (SRS-3400, Germany).
2.4 Catalytic methane decomposition. The activity of the carbon samples towards methane decomposition was evaluated in a vertical stainless steel fixed-bed reactor (8 mm i.d.) at 850 oC under atmospheric
3
Oviedo ICCS&T 2011. Extended Abstract
pressure. Isothermal experiments were carried out with 50 ml/min of the total feed gases (20%CH4/N2). 0.20 g catalyst was used along with the total volumetric hourly space velocity of 15,000 ml/(h·gcat). The gas products were analyzed by an online gas chromatography (GC7890Ⅱ) with a thermal conductivity detector (packed with 5A molecular sieve) and a flame ionization detector (GDX502 packed column). 3. Results and Discussion Figure 1 shows the nitrogen adsorption/desorption isotherms and the corresponding pore size distributions for the ACs prepared with silica or silicate as an additive. The isotherms are typical type IV obviously exhibiting hysteresis loops, indicative of certain contribution of mesopores to the total porosity. A wide hysteresis loop is observed in RC, which is consistent with mesoporous carbons [9]. In contrast, SiRC exhibit a narrower width in its hysteresis loop, and the adsorptions below P/P0 being 0.1 obviously increase, suggesting the increase of microporosity with the addition of SiO2 to the carbon precursor CLR. But the average mesopore width for the samples studied still concentrates at about 3.5 nm. Compared the surface and structural properties in Table 4, the samples prepared with SiO2 has a smaller SBET and Vt than before. It is a result from the weakened activation of the carbon matrix, owing to the reaction of SiO2 with the same initial amount of KOH. But the contributions of micropore surface (Smic) and volume (Vmic) to SBET (Smic/SBET) and Vt (Vmic/Vt) distinctly increase. In another word, the microporosity of the carbons increases while the mesoporosity is simultaneously preserved, which may be a result caused by the formed silicate from the reaction of KOH and SiO2. When using Na2SiO3 instead of SiO2, the similar increase in microporosity of the sample can be achieved, as shown in Figure 1 and Table 4. It confirms that the additional SiO2 affects the porous structure through forming silicate. b
a
dV/dlogDP
400
200 0.0
RC SiRC Na2SiO3-RC
1.2
3 -1
Vads/ cm gSTP
600
1.6
0.2
0.4
0.6
P/P0
0.8
1.0
0.8 0.4 0.0 1
10
100
DP/ nm
Figure 1. N2 adsorption/desorption isotherms (a) and the corresponding BJH pore size distributions (b) of the samples prepared with silica or silicate as an additive.
4
Oviedo ICCS&T 2011. Extended Abstract
When increasing the addition amount of SiO2, the hysteresis loop becomes narrower and the adsorption drops gradually (Figure 2a). As shown in Table 4, the smaller drops in Smic and Vmic, comparing with those in SBET and Vt respectively, cause an increasing microporosity. The weakened KOH activation may be one of the reasons. Using SBA-15 as an additive instead of SiO2, the result is similar as that with a high dosage of SiO2. However, the carbon Al2O3-RC prepared with Al2O3 as an additive has an increased adsorption and a wider hysteresis loop (Figure 2b). The surface and structural properties suggest that it has an improved mesoporosity compared with RC. Despite that all the mesopore distribution were kept nearly the same, these results make sure that the porous structure of carbon materials is easy to adjust by introducing some additives to KOH activation of the carbon precursor.
3 -1
Vads/ cm gSTP
a
b
600
c 400
600
300
400
400
200 200 0.0
0.2
0.4
0.6
0.8
1.0
200 0.0
0.2
0.4
0.6
0.8
100 0.0
1.0
0.2
0.4
0.6
0.8
1.0
P/P0
d
e1.6
1.0 SiO2 : CLR 1:5 3:5 4:5
dV/dlogDP
0.8 0.6
f SiRC Al2O3-RC SBA-15-RC
1.2
0.6 W-SiRC HCl-SiRC SiRC
0.4
0.8
0.4
0.2 0.4
0.2
0.0
0.0
0.0 1
10
100
1
10
DP/ nm
100
1
10
100
Figure 2. N2 adsorption/desorption isotherms (a,b,c) and the corresponding BJH pore size distributions (d,e,f): The ratio of addition of SiO2 to CLR (a, b), the samples prepared with different additives (c, d), and effect of the washing solvent on the sample prepared with silica (c, d). To examine the effect of mineral matter from the initial CLR and the formed salts on the surface and structure of the carbons, the carbonized samples (NW-SiRC, W-SiRC, HClSiRC and SiRC) were washed by different solvents. With reduction of the mineral matter and the salts (in Table 5), the adsorptions increase gradually (Figure 2c), indicative of the increased SBET and Vt (in Table 4). As to NW-SiRC, the surface area and pore volume are negligible. Obviously, the pores that were filled before washing were created by vacating the inner space occupied by the mineral matter and the salts in
5
Oviedo ICCS&T 2011. Extended Abstract
ACs. When the washing intensity is not enough, some mineral matter and the formed salts will be removed incompletely (in Table 5), so that some mesopores of the samples prepared with SiO2 would be embellished as micropores. Despite the ash composition of the obtained sample SiRC is similar to RC when the mineral matter and salts are removed completely; the surface and structure properties are distinctly different. The change is attributed to the salt formed by the additive and KOH in the activation process, similar as the report [10]. Larger size of particles of the formed salt will create larger pores in the carbon after washing off.
Table 4. Surface and textural properties of the carbons. SBET Smic Vtot Sample Smic/SBET 2 2 (m /g) (m /g) (cm3/g) RC 1661 631 0.38 0.89 1SiRC 1515 870 0.57 0.78 SiRC 1175 706 0.60 0.59 4SiRC 820 667 0.81 0.37 1331 874 0.66 0.66 Na2SiO3-RC SBA-RC 1081 868 0.80 0.51 Al2O3-RC 1507 78 0.05 0.97 NW-SiRC 0.1 W-SiRC 737 298 0.40 0.41 1049 824 0.79 0.49 HCl-SiRC
Vmic (cm3/g) 0.27 0.35 0.29 0.27 0.36 0.35 0.05 0.12 0.33
Table 5. Ash composition of the carbon samples. Compositions (wt.%) Sample [a] Sum SiO2 K2O Fe2O3 SO3 CaO RC 0.5 0.1 0.0 0.1 0.1 0.0 NW-SiRC 73.7 21.0 34.6 5.5 1.2 6.0 W-SiRC 43.8 11.4 10.8 6.9 1.3 7.3 HCl-SiRC 10.8 9.6 0.2 0.1 0.3 0.2 SiRC 0.4 0.1 0.0 0.0 0.1 0.1
Vmic/Vtot 0.30 0.45 0.49 0.72 0.55 0.68 0.05 0.29 0.67
Al2O3 0.0 4.1 4.4 0.1 0.0
DAV (nm) 3.1 3.2 3.4 4.0 3.2 3.9 3.1 3.4 3.7
Others 0.2 1.3 1.7 0.3 0.1
[a] determined by TG analysis in air at 900 oC.
SEM and TEM images of NW-SiRC and SiRC are shown in Figure 3. A lot of aggregates were formed after carbonization (Figure 3a) and no pore is present in NWSiRC (Figure 3c). However, in spite of the amorphous structure of SiRC (Figure 3b), many disordered nanopores with sizes around 3.5 nm at various places can be seen (Figure 3d), which confirms that the nanoparticles of mineral matter (or related salts) are responsible for creating the porosity of the carbon.
6
Oviedo ICCS&T 2011. Extended Abstract
a
b
c
d
Figure 3. SEM (a,b) and TEM (c,d) micrographs of the samples: NW-SiRC (a, c), and SiRC (b, d). Generally, the catalytic activity and stability of carbon materials are closely related with their structural and surface chemistry [11,12]. As shown in Figure 4a, methane conversion in CMD at 850 oC on RC has a high stability in spite of a low initial conversion of about 26%, which is consistent with the catalytic performance of mesopore carbons [13]. In contrast, the carbons prepared with additives have a higher initial catalytic activity but a poor stability, owing to the increased microporosity [11,12,14]. Because the pores are filled, NW-SiRC has a very low catalytic activity for CMD. With the mineral matter or the salts that occupied the pores are removed gradually, the initial activity is obviously improved with the increased porosity, with methane conversion up to about 30% (Figure 4b); but down to 25% due to the reduction of the active metal; then back to 31% because of the improved porous structure. When more additives are used, more KOH will be consumed by the additive, so that the activation of the carbon matrix is weaker, because KOH was kept with the same initial
Conversion of CH4 (%)
amount. As a consequence, methane conversion drops more remarkable (Figure 4c). a 35
b
30 25
c
35
35 SiO2 : CLR
RC SiRC Al2O3-RC SBA-15-RC
30
NW-SiRC W-SiRC HCl-SiRC SiRC
25
25
20
20
15
15
15
10
10
10
5
20
0
60
120
180
1:5 3:5 4:5 5:5 7:5
30
5 0
60
120
180
0
60
120
180
Time (min)
Figure 4. Catalytic methane decomposition at 850 oC on the carbons prepared with: different additives (a), different washing solvents (b), and varying ratios of SiO2 to CLR (c).
7
Oviedo ICCS&T 2011. Extended Abstract
4. Conclusions It is confirmed that the porous structure of the carbon prepared from CLR is easy to adjust by introducing some additives to KOH activation. The additive affects the porous structure mainly by forming salts as nanostructured particles with KOH. Nanopores will be created by washing off the space occupied by the nanoparticles of mineral matter or the salts in the carbonized carbon. Using SiO2, SBA-15 or Na2SiO3 as an additive, the microporosity of the carbon prepared from CLR will increase. In contrast, the mesoporosity will be improved with Al2O3 as an additive. And the catalytic performance for CMD depends on the porous structure of the carbon material.
Acknowledgements We thank the National Natural Science Foundation of China for the financial support (No. 20906009), the Fundamental Research Funds for the Central Universities (No. 893367), and the National Basic Research Program of China (973 Program), the Ministry of Science and Technology, China (No. 2011CB201301).
References [1] Namasivayam C, Kavitha D. Removal of Congo Red from water by adsorption onto activated carbon prepared from coir pith, an agricultural solid waste. Dyes Pigment 2002; 1:47-58. [2]
Mall ID, Srivastava VC, Agarwal NK, Mishra IM. Removal of congo red from aqueous solution by bagasse fly ash and activated carbon: Kinetic study and equilibrium isotherm analyses. Chemosphere 2005;4:492-501.
[3] Kobya M, Demirbas E, Senturk E, Ince M. Adsorption of heavy metal ions from aqueous solutions by activated carbon prepared from apricot stone. Bioresour Technol 2005;13:1518-21. [4] Rivera-Utrilla J, Sanchez-Polo M. Ozonation of 1,3,6-naphthalenetrisulphonic acid catalysed by activated carbon in aqueous phase. Appl Catal B-Environ 2002;4:319-29. [5] Corma A, Garcia H, Leyva A. Catalytic activity of palladium supported on single wall carbon nanotubes compared to palladium supported on activated carbon Study of the Heck and Suzuki couplings, aerobic alcohol oxidation and selective hydrogenation. J Mol Catal A-Chem 2005;12:97-105. [6] Lozano-Castello D, Lillo-Rodenas MA, Cazorla-Amoros D, Linares-Solano A. Preparation of activated carbons from Spanish anthracite I. Activation by KOH. Carbon 2001;39:741-9. [7] Hu Z, Srinivasan MP, Ni Y. Preparation of mesoporous high-surface-area activated carbon. Adv Mater 2000;12:62-5. [8]
Mori H. Extraction of silicon dioxide from waste colored glasses by alkali fusion using potassium hydroxide. J Mater Sci 2003;38:3461-8.
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Oviedo ICCS&T 2011. Extended Abstract [9] Kruk M, Jaroniec M. Determination of mesopore size distributions from argon adsorption data at 77 K. J Phys Chem B 2002;106:4732-9. [10] Lee HI, Stucky GD, Kim JH, Pak C, Chang H, Kim JM. Spontaneous phase separation mediated synthesis of 3D mesoporous carbon with controllable cage and window size. Adv Mater 2011, DOI: 10.1002/adma.201003599. [11] Suelves I, Pinilla JL, Lázaro MJ, Moliner R. Carbonaceous materials as catalysts for decomposition of methane. Chem Eng J 2008;140:432-8. [12] Lee SY, Ryu BH, Han GY, Lee TJ, Yoon KJ. Catalytic characteristics of specialty carbon blacks in decomposition of methane for hydrogen production. Carbon 2008;46:1978-86. [13] Serrano DP, Botas JA, Pizarro P, Guil-López R, Gómez G. Ordered mesoporous carbons as highly active catalysts for hydrogen production by CH4 decomposition. Chem Commum 2008;48:6585-7. [14] Muradov N, Smith F, T-Raissi A. Catalytic activity of carbons for methane decomposition reaction. Catal Today 2005;102-103:225-33.
9
Char characterisation from Oxyfuel combustion Nuamah, A.,1,2. Lester, E.,1 Drage T.,1 Riley, G2. 1
Faculty of Engineering, University of Nottingham, Nottingham, NG7 2RD, U.K.
2
Fuels and Combustion, RWEnpower, Windmill Hill Business Park, Whitehill Way, Swindon,
Wiltshire, SN5 6PB
[email protected]
Abstract
Oxyfuel combustion is one of the promising carbon capture technologies with a significantly lower cost to the environment in terms of carbon emissions. Whilst coal combustion in air is widely understood, there is less available literature on combustion in a high CO2 atmosphere, particularly in terms combustion reactivity and the impact on char morphology. Changes to the intrinsic reactivity of the char and the development of active sites or internal porosity of the char particle could all result in a different burnout profile.
This paper analyses chars produced under air and oxyfuel conditions at the RWE nPower 0.5 MWt combustion test facility (CTF) at Didcot. Chars were collected at the back end of the rig using a water-cooled probe. Initially, the exercise was conducted in air with different oxygen contents. The same exercise was replicated with oxyfuel conditions, with a 72% recycle ratio and varying oxygen contents in the furnace. These oxyfuel conditions were carried in both wet and dry recycle state and their results compared
1.
Introduction
Meeting the increasing demand for energy, whilst avoiding any adverse effects to the environment, is a serious concern for many power generators. The IEA has stated that energy supplies, to support the global economy over the next 25 years, are vulnerable as a result of under-investment, the possibility of environmental catastrophe or a sudden supply interruption (IEA, 2006).
Carbon dioxide (CO2) is generally accepted to be the primary green house gas responsible for global warming and subsequent climate change. Most CO2 is emitted from the burning of fossil fuels in power generation plants. However, the use of fossil fuel in power generation will continue to dominate the energy sector for the foreseeable future. Technological advances are needed to address the issue of greenhouse gas emissions from power plants. These technologies will need to be adopted to reduce emissions from fossil fuel-fired plant whilst supporting governmental policy initiatives (law, regulations etc). There are strong arguments for urgent action in reducing carbon emissions. China, the World’s biggest emitter of carbon, has recently announced its commitment to reduce their emissions by 40-45% per GDP by 2020 based on 2005 level (Uwasu M. 2010). The UK demonstrated a similar commitment by setting a stringent emission target of carbon emissions reduction of 34% by 2020 and 80% by 2050 all based on 1990 levels (DECC,2011). The strategies to achieve these targets require carbon abatement technologies that include the development of large scale fossil fuel power plants. Carbon capture and storage (CCS) has been identified as a means to reduce emissions from power plants by as much as 90% with the potential to reduce global carbon dioxide emissions by 28% by 2050 (Sturgeon, Cameron et al. 2009).
Oxyfuel combustion is a CCS technology that uses pure oxygen, rather than air as the oxidising agent in the combustion process. This produces a highly concentrated CO2 gas stream as the flue gas, which can be easily captured and sequestered. Oxyfuel combustion can also be adapted to retrofit old plant and it has significant benefits in terms of emissions of carbon, NOx and SOx (Smart, Patel et al. 2010). However, the replacement of nitrogen with carbon dioxide causes a dramatic increase in the radiative heat transfer in the furnace due to the increase in CO2 and water vapour in the flue gas and a corresponding increase in emissivity of the furnace gas. In addition, to achieve a similar adiabatic temperature as in air, a number of conditions have to be changed, such as the mass flow rate and velocities of the secondary and primary flow, which would also interfere with the burner aerodynamics, resulting in changes in fuel ignition properties, flame propagation, shape and residence time (Smart, Patel et al. 2010). These differences between combustion in air and in oxyfuel environment which have been established from both pilot scale and laboratory scale studies (Wall et al., 2009), can be explained by differences in gas properties between CO2 and N2.
Density: CO2 has a molecular weight of 44, compared to 28 for N2, making the density of the flue gas in oxyfuel combustion higher. Heat capacity: CO2 has a higher heat capacity than N2 Radiative properties of the furnace gases: oxyfuel combustion has higher CO2 and H2O content, both having high emitting power. Diffusivity: the oxygen diffusion rate in CO2 is 0.8 times that of N2. There have been a lot of studies on oxyfuel combustion as a potential candidate for CCS (Sturgeon et al 2009, Tigges et al 2008). Most of the research have been centred on demonstration (Sturgeon et al 2009, Holtl et al 2008, Li et al 2008, Covino et al 2008, Wall et al 2009, Jia et al 2008), retrofit ability (Tigges et al 2008), and flame properties (Seepana et al 2008, Von Scheele et al 2006, Blasiak et al 2007, Narayanan et al 2006). One of the most interesting studies was undertaken by Wall and his co-workers (Wall et al 2009), where a comprehensive review of oxyfuel coal combustion was undertaken. Studies on oxyfuel combustion on a pilot-scale furnace and a laboratory scale drop tube furnace were presented and compared with computational fluid dynamic calculations. The work includes a comprehensive assessment on oxyfuel combustion in a pilot-scale oxyfuel furnace, modifying the design criterion for an oxy retrofit by matching heat transfer, the development of a new 4grey gas model (which accurately predict emissivity of the gases in oxy-fired furnace for furnace modelling), initial measurements of coal reactivity for air and oxyfuel tests at laboratory and pilot scale; and predictions of observed delays in flame ignition with oxyfiring.
There are still aspects of the technology that need to be investigated to make the technology commercially viable including; O2 handling in boiler design, large scale oxygen production, questions over CO2 disposal that is contaminated with SOX and NOX, where is the recycled flue gas taken from, where O2 should be injected etc. This paper shows some preliminary characterisation results from pilot scale trials (0.5 MWt) for a single coal operating air fired and oxyfuel fired conditions with a variable oxygen content and water content. The purpose of the work is to show if differences exist between combustion techniques, in terms of reactivity and morphology, and whether these differences have implications in terms of burnout reactivity and kinetics.
2.
Methodology
2.1
Flue gas recycle
One of the important parameters in oxyfuel combustion is the recycle ratio. The recycle ratio establishes a similar heat flux in the furnace as conventional air fired and also ensures that there is enough gas stream in the chamber to carry the heat through in the absence of nitrogen. Oxyfuel combustion can be classified as either wet or dry depending on where the flue gas is tapped from at the downstream of the chamber (Smart, Patel et al. 2010). Recycle ratio is defined as;
100
In this paper, the recycle ratio was simulated by using CO2 in a once through oxyfuel system for flexibility, and the wet oxyfuel conditions was also simulated by injecting steam at a rate of 42kg/h into the furnace.
2.2
Combustion test facility
RWE npower’s 0.5MWth combustion test facility (CTF) is a refractory lined combustion chamber with an inner cross section of 0.8m x 0.8m and about 4m in length (Figure 1). Outside of the chamber, a water jacket layer is fitted to remove the input energy. 0.5MWth International Flame research Foundation (IFRF) aerodynamically air staged burner is fitted and operated under the baseline operating conditions. Sampling and measurement are undertaken through viewing ports on the centre line of the sidewall. Although, the rig was not originally designed to operate with oxyfuel conditions, a once through oxyfuel system has been adopted which maintains flexibility in simulating different recycle configuration and varying oxygen levels in various gas streams (Figure 2).
Figure 1 – the overview of the 0.5MWt CTF rig at Didcot.
Figure 2 – the recirculation and fuel injection system for the oxyfuel trials on the CTF rig.
Samples were collected with the aid of a water cooled probe which is inserted through the port after the narrowing of the furnace, approximately 7 metres from the injection point (Figure 1), where the samples are extracted through the probe and collected in a clean container.
2.4 Experimental Conditions
The following conditions were used to obtain samples of flyash from the rig;
Air test 2% Oxygen 3% Oxygen 4% Oxygen Oxyfuel (dry) 72% recycle rate 1% Oxygen 2% Oxygen 3% Oxygen 4% Oxygen Oxyfuel (Wet) 72% recycle rate, 42kg/hr steam rate 1% Oxygen 2% Oxygen 3% Oxygen 4% Oxygen The latter experiments used a steam recycle rate of 42kg/hr to simulate wet conditions from recycled flue gases.
2.4 Morphological Analysis Powdered samples were analysed using an FEI SEM microscope and sectioned chars were analysed using oil immersion microscopy with a Ortholux Pol II BK Leica Microscope and a 32x objective.
2.5 Thermal Analysis The intrinsic reactivity of the chars was assessed using a TA instruments Thermogravimetric Analyser. Air flow was maintained at 100cm3/min whilst heating the samples from ambient to 900oC at a ramp rate of 10oC/min. The peak temperatures (where the combustion reaches a maximum) and burnout temperatures (where the combustion rate falls to 10% of the peak temperature) were recorded. The carbon content, or loss on ignition (LOI) of the flyash samples was also calculated from this test.
3.
Results
3.1 Thermal Characterisation Results Figure 3 shows the results for the LOI for each sample. There are similar trends for all three conditions (air, oxyfuel, oxyfuel with moisture) with increasing oxygen content resulting in a lower residual carbon content. The only exception is the 4% oxygen sample with oxyfuel/moisture.
4.5 4
LOI (Carbon content %)
3.5 3
Air test
2.5 oxyfuel (dry) 72% recycle rate
2 1.5
oxyfuel (Wet) 72% recycle rate, 42kg/hr steam rate
1 0.5 0 1% Oxygen 2% Oxygen 3% Oxygen 4% Oxygen
Figure 3 – Carbon content vs CTF firing conditions
Figures 4a, b and c show the intrinsic reactivity profiles for the 3 firing conditions. In each case, as the total carbon content decreases, the intensity of the peak also decreases. Interestingly, the 2% oxygen air fired condition shows a more reactive peak or shelf around 500oC which is gone with the 3% and 4% conditions. It is unfortunate that a 1% condition was not obtained for the air fired runs, since this sample would possibly confirm the existence of a more reactive peak at higher LOI levels. The peak temperatures for all 3 profiles in Figure 4a are around 600oC, whereas the oxyfuel samples show a 10oC increase. The oxyfuel/moisture samples are more intriguing. As the oxygen content increases, the peak temperatures increase. The 1% oxygen sample shows some evidence of the double peak seen with the 2% oxygen air sample and has the most reactive char (573oC), followed by 590oC,
595oC and 630oC for the 2, 3 and 4% oxygen runs. The 4% oxygen sample also shows some signs of the generation of an unreactive carbon form at 900oC. This could be a result of high temperature annealing (Hurt et al., 1998) but more work is needed to confirm what is causing this extra peak. 0.25
dW/dt
0.2 0.15 Air test 2% Oxygen
0.1
Air test 3% Oxygen
0.05
Air test 4% Oxygen
0 0
500
1000
(a) Temperature (oC) 0.25 oxyfuel (dry) 72% recycle rate 1% Oxygen
dW/dt
0.2 0.15
oxyfuel (dry) 72% recycle rate 2% Oxygen
0.1 0.05 0 0
500
1000
(b) Temperature (oC)
oxyfuel (dry) 72% recycle rate 3% Oxygen
0.3
dW/dt
0.25 oxyfuel (Wet) 72% recycle rate, 42kg/hr steam rate 1% Oxygen
0.2 0.15 0.1 0.05 0 0
500
1000
oxyfuel (Wet) 72% recycle rate, 42kg/hr steam rate 2% Oxygen
(c) Temperature (oC)
Figure 4a-c – The TGA profiles for the 3 different firing tests.
3.2 – Morphology Results Figure 5 shows the three basic char structures that tend to result from partial combustion of coal particles. Thin walled chars are relatively quick to burnout, with walls that are predominantly less than 3 microns. Thick walled chars still have significant porosity but have walls with a thickness over 3 microns and this increase in thickness results in an increase in burnout time. The solids chars are generally from inertinite type macerals in the coal (sclerotinite, macrinite, fusinite etc) and represent the most significant source of carbon (per particle) in the flyash. These solid structures take the longest time to burnout out.
Figure 5 – the three basic char structures from coal
Table 1 shows the data for the three firing conditions and Figure 6 shows example structures from various samples.
Thin Thick Solids
Oxyfuel (Wet) Oxyfuel (dry) 72% recycle rate, Air test 72% recycle rate 42kg/hr steam rate 2% 3% 4% 1% 2% 3% 4% 1% 2% 3% 4% 69 58 72 74 80 80 78 84 77 73 76 23 27 18 26 14 16 14 14 20 19 18 8 15 10 0 6 4 8 2 3 8 6 Table 1 – the char structures found in the flyash samples.
Figure 6 - Examples of the chars found in the flyash samples using oil immersion microscopy
The clearest trend is the higher levels of solids in the air fired tests. Oxygen levels appear to make no difference to the total quantity of each type, but the solids content is approximately twice that seen in the other samples. The 1% oxygen oxyfuel sample appears to produce a quantity of thick walled chars but this is the only real difference between the two oxyfuel fired tests.
Figure 7 - Examples of the chars found in the flyash samples using SEM analysis
Examples of char structures from SEM analysis are shown in Figure 7. There are cases where the chars are intimately mixed with the mineral particles and others were the char is discrete, with little or no ash inclusion. Whilst this is significant in terms of ash collection in the precipitators, the relation to combustion burnout or burnout potential is less clear. Overall SEM is more of a qualitative tool, and it so was difficult to find significant differences between the three firing conditions.
4.
Discussion
With all tests (apart from 4% oxygen with wet recycle) increasing the O2, results in a decrease in residual carbon in the flyash. The reason for the anomalous result is not clear and more work is needed to clarify. A 1% oxygen test (in air) would also help to identify further trends including the presence of a higher reactive carbon fraction at 500oC, which is present in the 2% oxygen/air run and the 1% oxyfuel run.
With most conditions the reactivity of the carbon does not change – i.e. the peak temperature and peak position appears to remain relatively constant, showing a similar burnout rate with all conditions. Increasing the oxygen content does not preferentially burnout the more reactive fractions or result in thermal annealing to create a more unreactive fraction.
Oxyfuel conditions appear to be more efficient than air based runs when dry. Dry oxyfuel runs result in approximately 50% less carbon in the flyash when compared to wet oxyfuel or air based runs.
More work is needed using the test facility with a broader suite of coals. Obtaining samples closer to the injection point would be particularly useful in order to observe intermediate char structures rather than residual structures in the flyash. This would allow direct comparison between char types during the oxyfuel process and generate comparative data to better explain burnout kinetics.
5.
References
Blasiak, W., Yang, W.H., Narayanan, K., Von Schéele, J. Flameless oxyfuel combustion for fuel consumption and nitrogen oxides emissions reductions and productivity increase. Journal of the Energy Institute. 2007, 80(1) 3-11.
Covino, B.S., Matthes, S.A., Bullard, S.J. Effect of oxyfuel combustion on superheater corrosion. NACE - International Corrosion Conference Series. 2008, 084561-0845610.
DECC (UK Department of energy and climate change) (2011), A low carbon UK. Available at; www.decc.gov.uk. Accesed on [30/05/2011].
Höltl, Werner et al. Oxyfuel combustion for low-calorific fuels. Proceedings of the 2008 Global Symposium on Recycling, Waste Treatment and Clean Technology, REWAS. 2008,151-157.
Hurt R.H., Sun J-K, Lunden M. A kinetic model of carbon burnout in pulverized coal combustion. Combust Flame 1998, p113:
181 –9 Jia, L., Tan, Y., Anthony, E.J. Oxyfuel combustion using CFBC - Some early lessons. 25th Annual International Pittsburgh Coal Conference, PCC – Proceedings. Pittsburgh. 2008.
Li, Zhen-Shan, Zhang, Teng; Cai, Ning-Sheng. Experimental study of O2-CO2 production for the oxyfuel combustion using a Co-based oxygen carrier. Industrial and Engineering Chemistry Research. 2008, 47(19), 7147-
Narayanan, K. Wang, W.; Blasiak, W.; Ekman, T. Flameless oxyfuel combustion: Technology, modeling and benefits. Revue de Metallurgie. Cahiers D'Informations Techniques. 2006, 103(5) 210-217.
Seepana, S., Jayanti, S. A new stable operating regime for oxyfuel combustion. ASME International Mechanical Engineering Congress and Exposition. 2009, 3, 435-444.
Smart, J. P., R. Patel, et al. (2010). "Oxy-fuel combustion of coal and biomass, the effect on radiative and convective heat transfer and burnout." Combustion and Flame 157(12): 22302240. Sturgeon, D. W., E. D. Cameron, et al. (2009). "Demonstration of an oxyfuel combustion system." Energy Procedia 1(1): 471-478.
Tigges, K.-D., Klauke, F., Bergins, C., Busekrus, K., Niesbach, J., Ehmann, M., Vollmer, B., Buddenberg, T., Wu, S. Kukoski, A.
Oxyfuel combustion retrofits for existing power
stations. Air and Waste Management Association - 7th Power Plant Air Pollutant Control 'Mega' Symposium. 2008,1, 08-122.
Uwasu M., J. Y., Saijo T. (2010). "On the Chinese Carbon Reduction Target." Sustainability 2: 1553-1557
Von Schéele, J.
Oxyfuel technology: Applications for and operational experience with
flameless combustion. Gaswaerme International. 2006, 55(3) 186-189.
Wall, T. Liu, Y. Spero, C. Elliot, L. Khare, S. Rathman, R. Zeenathal, F. Moghtaderi, B. Buhre, B. Sheng, C. Gupta, R. Yamada, T. Makina, K., and Yu, J. An overview on oxyfuel coal combustion-state of the art research and technology development. Chemical engineering research and design. 2009, 87, 1003-1016.
Oviedo ICCS&T 2011. Extended Abstract
Avoiding Mercury Emissions From Coal Combustion: Post Combustion Research Strategies M. R. Martínez-Tarazona, M. A. López-Antón, R. Ochoa-González, A. Fuente-Cuesta, P. Abad- Valle, J. Rodríguez-Pérez, F. Inguanzo-Fernández, M. Díaz-Somoano and R. García, Instituto Nacional del Carbón (INCAR), CSIC, Apartado 73, 33080-Oviedo. Spain.
[email protected]
Abstract
This work summarizes results of several research projects carried out with the aim of avoiding mercury emissions from coal combustion power plants. In all the cases, the strategy for mercury removal focused on post- combustion processes. Four approaches are discussed: i) the use of carbon sorbents before the particle control devices, ii) the retention and oxidation of mercury on fly ashes, iii) the use of regenerable sorbents at the end of the process and iv) the retention and stabilization of mercury in wet scrubbers. The objective is to compare and evaluate different options for Hg control.
1. Introduction Of the toxic trace elements present in coal and evaporated during combustion, Hg is undoubtedly the element of greatest concern from the environmental point of view. Energy production from coal is the main anthropogenic source of this element. Recent estimations suggest that fossil fuel combustion produces 45% of the Hg emitted throughout the world from human activities [1]. The problem of Hg emission and capture in the coal combustion process as well as the role that pollution control systems designed for other contaminants can play in this capture are the subject of study around the world. The measures proposed to reduce Hg emissions can be divided into two groups: i) actions before combustion and ii) postcombustion processes (gas cleaning). Pre-combustion Hg reduction is usually tackled either by preparing coal blends to obtain a low Hg fuel or by coal cleaning to reduce the amount of Hg present in the combustible [2] Mixing additives with the combustible in order to modify Hg performance and retaining Hg in the sub-products have also been considered as possible options [3]. Post-combustion technologies are mainly based on the use of sorbents injected into the gaseous stream and subsequently retained in particulate control devices [4]. The gas
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Oviedo ICCS&T 2011. Extended Abstract
cleaning systems already installed, such as scrubbers used to control SO2 emissions, may also be useful for the simultaneous capture of oxidized Hg [5]. An alternative postcombustion option would be to use specific purposed designed sorbents of Hg. This work summarizes the results of projects developed to reduce emissions of Hg into the air from coal combustion. All the results included in this paper were obtained at laboratory scale. However, the approach used for the projects was based on preliminary studies of Hg behaviour and distribution in coal and coal combustion by-products (CCBs) [6].
2. Experimental section Various materials were considered as solid sorbents for use for use in different conditions : 2.1.-Carbon sorbents to be used before particle control devices: Biomass gasification chars from i) sunflower husks (SH), ii) poultry litter (PL), iii) wood pellets (CW), iv) wood waste (WW1, WW2) and v) mixtures of paper and plastic waste (PW1, PW2, PW3) all of which were obtained from a pilot gasification plant of 500 kW equipped with a circulated fluidized bed (CFB) BIVKIN gasifier and commercial activated carbons (Filtracarb D47/7+S and Norit RB3 and RBHG3) were evaluated for their ability to capture and oxidize mercury. 2.2-Four fly ashes from pulverized coal combustion power plants (CTA-O, CTLO, CTSR-O, CTE-O) and one from Fluidized Bed Combustion (CTP-O) were evaluated for their ability to capture and oxidize Hg. All the ashes were obtained from the combustion of coals of different rank and characteristics. Fractions of the fly ashes enriched in unburned carbon were also studied. 2.3.-Au-loaded regenerable sorbents were prepared by two methods based on the formation of Au colloids supported on activated carbons. In one method a solution of HAuCl43H2O mixed with polyvinyl alcohol (PVA) is reduced with NaBH4 to form the Au sol that is used to impregnate the support. In the second method HAuCl4 3H2O is added to a solution of tetrakis-(hydroxymethyl)phosphonium chloride (THPC) in NaOH where the Au sol is formed. Sorbents with different Au contents and prepared by the two methods were evaluated. The experimental lab scale devices used in the study consisted of glass reactors into which each sorbent was placed to form a fixed bed. The temperature was kept
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Oviedo ICCS&T 2011. Extended Abstract
between 40 to 150 ºC. Elemental mercury (Hg0) in gas phase was obtained from permeation tubes, and oxidized mercury (Hg2+) from a solution in a commercial evaporator. The concentration of Hg in gas phase was variable but of the order of 100 μg m-3. Synthetic gas mixtures containing the species found in coal combustion (CO2, O2, SO2, NO2, NO, HCl, N2 and H2O), were used to transport the mercury species through the sorbent at different flow rates. Breakdown curves were recorded using continuous mercury monitors (UT 3000 and VM-3000), depending on the mercury concentrations. The mercury content of the sorbent after the retention experiments was also determined by means of an automatic mercury analyzer (AMA). 2.4-Retention in the limestone slurries was performed at lab scale to simulate the liquid and slurry compositions typical of wet scrubbers. These systems consist of three parts: i) a mercury and flue gas generation unit, ii) a glass reactor containing the scrubbing slurry or absorption solution and iii) a continuous mercury emission monitor to measure the Hg0. The Hg0 concentration was recorded as a function of time. The pH and redox potential in the reactor were measured continuously. A HgNO3 solution, which was stabilized in HCl medium, was evaporated continuously at 200ºC to generate Hg2+(g) species.
3. Results and discussion When comparing the capacities of solid sorbents to capture mercury species in gas phase, it is important to bear in mind that each one of the materials studied will be used in different conditions. Biomass chars were evaluated for use, by injection, in the flue gases and were compared with activated carbons. The evaluation of fly ashes was performed with the aim of optimizing their performance in electrostatic precipitators or baghouses. Regenerable sorbents were developed to be used as fixed sorbents at the end of the cycle. Some chars resulting from biomass gasification, principally those obtained from plastic-paper waste, showed high mercury retention capacities in simulated coal combustion flue gases. These low-cost sorbents were found to be an attractive alternative to activated carbons for direct injection into power plants. Apart from having the advantages of high carbon contents and good textural characteristics, the char samples were found to contain a large amount of chloride. In the case of the fly ashes, our laboratory scale studies confirmed their different
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Oviedo ICCS&T 2011. Extended Abstract
behaviors in terms of Hg capture and oxidation depending on their characteristics. However, a general pattern of behavior was ascertained. The retention of Hg in fly ashes is mainly produced by the reaction between their components and the gases in the atmosphere. Moreover, Hg0 may oxidize on the surface of the fly ash. The ashes with the highest Hg retention capacity give rise to the highest Hg oxidation. The species present in the gas may bind to the surface of the ashes altering their Hg retention capacity. The retention capacity may increase, as in the case of HCl(g), or decrease, as in the case of SO2 or H2O(g). In fact, HCl (g) and SO2 (g) (the latter in the presence of O2), have a great influence on Hg retention and oxidation in fly ashes. The homogeneous oxidation of Hg0 at 120°C was observed to occur in an atmosphere of O2 + SO2. If the atmosphere also contains CO2 and H2O(g) the proportion of Hg2+(g) increases significantly in the presence of fly ash. In an atmosphere of HCl(g), Hg retention increases and promotes oxidation in the presence of fly ash. Carbonaceous matter is also involved in most of the mechanisms between Hg and fly ash. These carbon particles not only play an important role in Hg retention but they may be the medium via which the oxidation of Hg0occurs. The development of regenerable sorbents was evaluated by using an activated carbon as support (Norit RB3), impregnated with different amounts of Au by means of two impregnation methods [7]. The chief factors that determine the retention of Hg by Auloaded activated carbons are the homogeneity and deposition of the Au particles on the carbonaceous support rather than the amount of Au. Although the quantities of Hg retained during a given period may be similar, the efficiency of Hg capture mainly depends upon the distribution and accessibility of the Au particles. Different Hg adsorption rates were observed depending on the Au impregnation method used (PVA or THPC), with higher efficiencies generally corresponding to THPC, which acts both as a reducing agent and as an inhibitor of large particle formation. The performance of the wet scrubbers is the key to Hg capture and re-emission. Since Hg2+ is the Hg species retained in the scrubbers, it is essential to avoid any reaction that would favor a reduction of Hg within the system. The results of the experiments performed at laboratory scale indicate that sulfite ions are able to convert the Hg2+ captured in the scrubber into Hg0. In the presence of gypsum slurry, retention increases with higher percentages of O2, since O2, inhibits the formation of sulfite ions. Lower concentrations of SO2(g) prevent the formation of Hg0 because the formation of sulfur-reducing species in the suspension decreases. Water recirculation in the scrubber
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Oviedo ICCS&T 2011. Extended Abstract
produces a gradual increase in the concentration of dissolved Hg, favoring Hg2+reduction reactions, and a decrease in the efficiency of retention. Another variable that affects Hg retention is the pH. In a suspension containing limestone at a pH of around 7 the reduction of Hg2+ was detected, whereas at a pH below 4, the Hg was more stable. An increase in gas velocity also significantly affects the efficiency of Hg2+ retention. Also responsible for reducing Hg are the metals present in the limestone or fly ash reaching the scrubber. These metals are capable of transforming the Hg2+dissolved in the suspension into Hg0. In so far as halides are concerned, these do not inhibit the reduction of mercury in the presence of gypsum slurry. Leaving aside technical or economical considerations, it may be said that the main processes for Hg capture involve the oxidation of Hg. In order to capture any amount of Hg that, even in the optimum conditions of Hg capture, may still be present in gas phase as Hg0 after desulfuration, the use of regenerable sorbents will be needed.
Acknowledgements The financial support for these works was obtained from several projects. RFCRCT-2007-00007, CTQ2008-06860-C02-01, PIF-06-050, RFCR-CT-2006-00006. The authors thank the Energy Research Centre of the Netherlands for preparing the chars employed in this study.
References [1] Pacyna, EG, et al, 2010. Global emission of mercury to the atmosphere from anthropogenic sources in 2005 and projections to 2020. Atmos Environ 44, 2487-99. [2] López-Antón, MA, Diaz-Somoano, M, García, AB, Martínez-Tarazona, MR, Evaluation of mercury associations in two coals of different rank using physical separation procedures (2006) Fuel 85, 1389-95. [3] Vosteen, B, et al, Process for removing mercury from flue gases, United States Patent Nº US 6878358B2 (apr. 12, 2005). [4] Lee, S., Se, Y., Jurng, J., Lee, T., Effects of Reaction Conditions on Elemental Mercury Oxidation in Simulated Flue Gas by DC Nonthermal Plasma Atmos. Environ. (2004) 38, 488793. [5] Diaz-Somoano, M., Unterberger, S., Hein, K.R.G., Using Wet-FGD systems for mercury removal J. Environ. Monitor. (2005) 7, 906-9.
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Oviedo ICCS&T 2011. Extended Abstract [6] López-Antón, MA, Diaz-Somoano, M, Diaz, L, , Martínez-Tarazona, MR, Avoiding mercury emissions by combustion in a Spanish CFBC plant, Energy and Fuels (2011) 25, 3002-08. [7] Rodríguez-Pérez, J, López-Antón, MA, Díaz-Somoano, M, García, R. Martínez-Tarazona, MR, Development of Gold Nanoparticle-Doped Activated Carbon Sorbent for Elemental Mercury Energy and Fuels (2011) 25, 2022-27.
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Oviedo ICCS&T 2011. Extended Abstract
Interpreting the Re-Emission of Elemental Mercury During Wet FGD Scrubbing Balaji Krishnakumar1, Stephen Niksa1, and Naoki Fujiwara2 1 2
Niksa Energy Associates, LLC; Belmont, CA 94002 USA;
[email protected] Coal and Environment Research Laboratory, Idemitsu Kosan Co., Ltd., Chiba 299-0267 Japan
Abstract Whereas the conventional FGD analysis works well in relating gross FGD operating conditions to SO2 absorption efficiencies, it is unsuitable for trace metal transformations, especially for Hg re-emission. The reason is that conventional formulations implicitly assume that all oxygen is consumed in sulfite oxidation at the liquid interface on a slurry droplet, and evaluate the rate of sulfite oxidation as an average based on bulk liquid concentrations. The requisite analysis for Hg2+ chemistry must automatically shift the redox potential of bulk liquid in the spray from oxidizing to reducing through a balance among the finite-rate reagent fluxes that participate in sulfite oxidation. Our SO2 capture analysis allows oxygen to penetrate through the liquid film and accumulate in the bulk liquid if its concentration exceeds the stoichiometric requirement for bisulfate oxidation. We also propose a finite-rate reaction for Hg(II) reduction that depends on temperature, pH, and S(IV) species which, in the absence of sufficient oxygen, promote Hg(II) reduction to Hg0 vapor. This approach predicts reasonable extents of Hg removals and Hg reemission under realistic FGD operating conditions, without any parameter adjustments. Hg reemission can occur along the entire absorber at a rate that accelerates slightly along the upper elevations. Smaller slurry droplets re-emit more Hg because they sustain faster absorption of all the reagent gases involved in Hg(II) reduction. The total S(IV) species concentration in the slurry solution promotes re-emission; conversely, the analysis correctly predicts less re-emission for greater Cl ion concentrations in the slurry, in accord with a well-established tendency. Simulations with added S(II) species, as released from commercial additives to suppress Hg reemission, correctly gave complete precipitation of Hg(II) as HgS(s). These predictions are consistent with full-scale field tests where the addition of NaHS almost completely suppressed Hg re-emission at even the lowest additive levels.
1.
Introduction
The Hg field-testing literature shows that wet FGDs typically capture at least 90 % of the oxidized Hg (Hg2+) in flue gas, but virtually none of the elemental Hg vapor (Hg0), because Hg0
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Oviedo ICCS&T 2011. Extended Abstract
is insoluble in aqueous solutions. However, several full- and bench-scale tests have documented Hg re-emission whereby the Hg0 concentration at the FGD exit is greater than that at the inlet, suggesting that some of the absorbed Hg2+ must be reduced in the solution and re-emitted as Hg0 vapor. Hg re-emissions have also been observed in natural waters and atmospheric cloud and rain-water cycles [1-3], where S(IV) species promote the reduction of Hg2+. Bench-scale studies under well controlled scrubbing conditions showed that Cl- and O2 suppress re-emission, whereas Ca2+ and Mg2+ species promote it. Our goal is to develop a quantitative analysis that identifies and rank-orders the factors involved in Hg re-emission during wet FGD scrubbing, and can support efforts to mitigate Hg re-emission from full-scale, commercial wet FGDs.
2.
FGD Analysis
To simulate Hg transformations in wet FGD scrubbers, one must first model SO2 absorption to describe the major solution species (S, Ca, Cl) and their variations with FGD operating conditions (L/G, droplet diameter, T).
Given the absorber temperature and the inlet
compositions of flue gas and slurry, a conventional FGD analysis predicts the correct pH and reasonable SO2 absorption rates along the absorber. It also depicts realistic tendencies for variations in the inlet SO2 level, and identifies distinctive conditions where elevated HCl levels will reduce SO2 removal efficiencies by 5 to 10 %, and greater SO2 removal for smaller slurry droplets, all else the same. Whereas the conventional FGD analysis relates gross FGD operating conditions to SO2 absorption efficiencies, it is unsuitable for trace metal transformations, especially for Hg reemission. The reason is that conventional formulations implicitly assume that all oxygen is consumed in sulfite oxidation at the liquid interface on a slurry droplet, and evaluate the rate of sulfite oxidation as an average based on bulk liquid concentrations. The requisite analysis for Hg2+ chemistry and, presumably, other trace metal transformations must automatically shift the redox potential of bulk liquid in the spray from oxidizing to reducing through a balance among the finite-rate reagent fluxes that participate in sulfite oxidation. Our SO2 capture analysis allows oxygen to penetrate through the liquid film and accumulate in the bulk liquid if its concentration exceeds the stoichiometric requirement for bisulfate oxidation. We also propose a finite-rate reaction for Hg(II) reduction that depends on temperature, pH, and S(IV) species which, in the absence of sufficient oxygen, promote Hg(II) reduction to Hg0 vapor. The computerized implementation covers the process chemistry occurring in both a counterflow absorber and slurry holding tank. The tank analysis determines the required limestone feedrate Submit before May 15th to
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for a given tank RTD and the solids PSD into the spray nozzles within the absorber. It also determines how much liquid must be extracted to maintain a target Cl concentration in the tank. Among these results, only the limestone PSD affects the behavior in the absorber. The absorber analysis determines the SO2 capture and the compositions of liquid and solids that return into the holding tank. Given the absorber temperature and the inlet compositions of flue gas and slurry, the analysis predicts the SO2 capture efficiency; complete slurry composition and flue gas composition; slurry pH; the quantitative enhancement of mass transfer by acid dissociations in the slurry; and the relative contributions of liquid and gas resistances to the overall mass transfer rate. All quantities are resolved as functions of distance along the absorber axis.
3.
Results and Discussion
The base test case process parameters in the simulations correspond to NYSEG’s Kintigh station, which was tested under a U.S DoE program to evaluate additives for high SO2 removal [4]. In the simulations we assume that a single spray header delivered the entire liquid flow into an absorber section of 10 m height with a diameter of approximately 12 m. The calculated value of the liquid pH is 5.65, in good agreement with the measured value of 5.67. All rate constants in the analysis were assigned values from the characterization literature and were not adjusted further in the course of this work. The rate constant for Hg2+ reduction was based on a reported value for conditions that mimicked Hg transformations in cloud water [3], in spite of the fact that HgSO3 is unstable under FGD operating conditions. The % Hg re-emission (based on Hg2+ at the FGD inlet) and total Hg removal (% Hg-T) for the base case are presented in Fig. 1 along with the SO2 removal. In this and all succeeding figures, the absorber inlet is on the right and the slurry spray manifold at the absorber exit is on the left. The Hg2+ absorption approached 99 % at the exit of the absorber and is not shown. Whereas SO2 was continuously removed along the length of the absorber, most of the Hg was removed in the lower section of the countercurrent FGD at a much greater efficiency than SO2 removal. We attribute this high level of Hg capture near the entrance to the high aqueous solubility of HgCl2. As the flue gas moves towards the top of the absorber, there is a slight decrease in the Hg-T removal. This is because at the higher Hg2+ concentrations in solution, the rate of reduction and re-emission increase, as seen in Fig. 1. For the baseline case with 20% Hg0, the re-emission gradually increases to approximately 20 % and the Hg-T removal reaches 65 %. The effect of variations in inlet SO2 partial pressures was examined by altering the bulk liquid S(IV) concentrations. Since the baseline case S(IV) concentration corresponds to an SO2 partial
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Oviedo ICCS&T 2011. Extended Abstract
Figure 1. SO2 and total Hg removals (left y-axis) and percentage Hg re-emission (right y-axis) along the length of the absorber. pressure of 3275 ppm, a lower partial pressure must result in a lower S(IV) concentration in a closed loop absorber. A change in inlet SO2 partial pressure alone is not sufficient to alter the liquid phase chemistry because the aqueous S(IV) concentrations are high relative to SO2 absorption in a single absorber flow cycle. To examine the model response to variations in the inlet SO2 partial pressures, four bulk liquid S(IV) concentrations between 3.5 and 6.5 mM were examined. All other simulation conditions were kept the same as the base case and only the alkalinity was adjusted for each S(IV) concentration to maintain the solution pH. The Hg-T removal and Hg re-emission for the different S(IV) concentrations are shown in Fig. 2. As expected, the SO2 removal was unaffected by the change in liquid S(IV) concentration. In all cases, the absorption of Hg2+ exceeded 99 % and is not shown in the figure. As the S(IV) concentration was increased from approximately 3.5 to 6.5 mM, the Hg-T capture efficiency decreased from 72 to 62 % with an increase in Hg re-emission by a similar percentage. The increase in re-emission is directly related to an increase in S(IV) concentration and the consequent increase in Hg-S(IV) species levels in the bulk liquid. For a two-fold increase in S(IV) concentrations, the Hg-S(IV) fraction in the solution at the absorber exit increased from 14 to 36 % with an associated decrease in Hg-Cl levels. The Hg re-emission therefore appears to linearly track the Hg-S(IV) levels in the solution. Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
Figure 2. (Left Panel) Effect of bulk liquid S(IV) concentration on total Hg removal (left y-axis) and percentage Hg re-emission (right y-axis); (Right Panel) Effect of slurry Cl− concentration on total Hg removal (left y-axis) and Hg re-emission (right y-axis). The Cl− ion plays a crucial role in the capture of Hg in scrubber solutions. In closed loop FGDs, the recirculating slurry quickly becomes concentrated in species like HCl that are absorbed from the flue gas and are not removed by precipitation reactions such as those of calcium sulfite and sulfate. The Cl− ion concentration is regulated to control corrosion problems by continuously purging a portion of the liquid stream from the reaction tank. The concentrations of Cl-species are therefore much greater than they would be in open-loop systems, as seen in the present case. Similar to examination with varying S(IV) concentration, we varied the slurry Cl− ion concentration between 0.55 and 0.68 kmol/m3 and adjusted the alkalinity to maintain the slurry pH. The Hg-T removal and Hg re-emission for the different Cl− ion concentrations appears in Fig. 2, which clearly shows that for progressively greater Cl− ion in the solution, the Hg(II) species preferentially partition into Hg-Cl complexes that reduce re-emission, which is consistent with the observations in bench- and pilot-scale experiments reported in the literature. In the present simulations, as the Cl− ion concentration in the liquid was increased from 0.5 to 0.68 kmol/m3, the aqueous Hg(II) ions associated with Cl increased from 66 to 84 % with an associated decrease in Hg-S(IV) complexes. The change in Cl− did not have an effect on Hg2+ absorption since it was in excess of 99 % in all cases. The increased partitioning of Hg(II) in Hg-Cl complexes decreased the re-emissions by 11 %. For the case under simulation with an inlet flue gas Hg2+ of 80 %, the Hg-T removal increased from 62 to 72 % as the Cl− ion concentration in the liquid was increased from 0.5 to 0.68 kmol/m3. Most additives to control Hg re-emission from scrubbers, such as Na2S4, TMT-15™ and NaHS, Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
Figure 3. Hg-T removal along the length of the absorber with and without an S(II) concentration of 1 µM. release S(II) ions in the slurry and precipitate Hg as HgS (s). We examined the addition of small amounts of S(II) into the base case scrubber solution and simulated Hg capture. Even at 1 µM concentrations of the additive, all the Hg(II) precipitated as HgS(s), which establishes the maximum Hg2+ concentration in the solution. For 1 µM of S(II) added to the solution, the maximum total dissolved Hg(II) was only 10–33. The Hg-T removal with and without S(II) addition is shown in Fig. 3. With 80 % HgCl2 in the flue gas, essentially all Hg2+ was retained in the solid phase. These predictions are consistent with the B&W tests at Endicott station [5] where the addition of NaHS almost completely suppressed Hg re-emission and the extent of reemission did not bear any relation to the additive feed rate. We also simulated the addition of S(II) under oxidizing conditions where we increased the pH to 6.5 and reduced the S(IV) species concentration such that oxygen was in excess.
Such
conditions arise in lime-based FGDs that operate at a higher pH and partition most of the S(IV) species into SO32– and other forms instead of HSO3−, under a positive redox potential due to excess O2. Under such conditions, equilibrium shifts the added S(II) species to their oxidized
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Oviedo ICCS&T 2011. Extended Abstract
forms which prevents any precipitation of HgS solids. This is consistent with B&W tests at the Mg/lime FGD at Zimmer station [5] where an NaHS additive did not improve the baseline Hg-T capture efficiency even at high additive feed rates. Our model does not predict any re-emission under excess O2 conditions, so the Hg capture profile is similar to the one shown in Fig. 3 depicting almost complete capture of Hg2+. For lower S(IV) concentrations and excess O2, our analysis explains the ineffectiveness of S(II)-based additives, although it also gives no reemission for such conditions.
It is conceivable that both the oxidation of S(II) and the
precipitation of HgS(s) may be kinetically limited, in which case the relative rates will determine whether any HgS solids would be precipitated before the oxidation of S(II) is complete. 4.
Conclusions
Our FGD scrubbing analysis predicts reasonable extents of Hg removals and Hg re-emission under realistic FGD operating conditions, without any parameter adjustments. Whereas most Hg2+ in flue gas is captured near the flue gas inlet, at the bottom of the absorber, Hg re-emission can occur along the entire absorber at a rate that accelerates slightly along the upper elevations. Whereas predicted Hg removals are insensitive to droplet diameter, smaller slurry droplets reemit more Hg because they sustain faster absorption of all the reagent gases involved in Hg(II) reduction. The total S(IV) species concentration in the slurry solution promotes re-emission; conversely, the analysis correctly predicts less re-emission for greater Cl ion concentrations in the slurry, in accord with a well-established tendency. Simulations with added S(II) species, as released from commercial additives to suppress Hg re-emission, correctly gave complete precipitation of Hg(II) as HgS(s). These predictions are consistent with full-scale field tests where the addition of NaHS almost completely suppressed Hg re-emission and the extent of reemission did not bear any relation to the additive feed rate. Acknowledgement. Financial support for this work from Japan’s New Energy Development Organization, with administration through the Coal and Environment Research Laboratory of Idemitsu Kosan Co., Ltd., is gratefully acknowledged.
References [1] Brosset C. The behavior of mercury in the physical environment. Water Air & Soil Pollution 1987, 34, 145–166. [2] Munthe J, Xiao ZF, Lindqvist O. The aqueous reduction of divalent mercury by sulfite. Water Air & Soil Pollution 1991, 56, 621–630. [3] Van-Loon L, Mader E, Scott SL. Reduction of the aqueous mercuric ion by sulfite: UV Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
spectrum of HgSO3 and its intramolecular redox reaction. J. Phys. Chem. A 2000, 104, 1621– 1626. [4] Radian Corp. High SO2 removal testing. Topical report to U.S. DoE under Contract # DEAC22-92PC91338, Results of sodium formate additive tests at New York State Eleectric & Gas Corporation’s Kintigh station. Feb 1997. [5] Amrhein GT, Bailey RT, Downs W, Holmes MJ, Kudlac GA, Madden DA. Advanced Emissions Control Development Program- Phase III. Final Report to U.S. Department of Energy under contract# DE-FC22-94PC94251, July, 1999.
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Oviedo ICCS&T 2011. Extended Abstract
Performance Simulations for Co-Gasification of Coal and Methane Stephen Niksa1, J.-P. Lim2, D. del Rio Diaz Jara2, D. Steele2, D. Eckstrom2, R. Malhotra2, and R. B. Wilson2 1 2
Niksa Energy Associates LLC, Belmont, CA, 94002 USA;
[email protected] Chemistry and Chemical Engr Dept, SRI International, Menlo Park, CA 94025 USA
Abstract In the process under development, coal suspended in mixtures of CH4, H2, and steam is rapidly heated to temperatures above 1400°C under 5 - 7 MPa for at least one second. The coal first decomposes into volatiles and char while CH4 is converted into CO/H2 mixtures. Then the char is converted into CO/H2 mixtures via steam gasification on longer time scales, and into CH4 via hydrogasification.
Throughout all stages,
homogeneous chemistry reforms all intermediate fuel components into the syngas feedstock for methanol synthesis.
Fully validated reaction mechanisms for each
chemical process were used to quantitatively interpret a co-gasification test series in SRI’s lab-scale gasification facility. Methane conversion in the gas phase increases for progressively hotter temperatures, in accord with the data. But the strong predicted dependence on steam concentration was not evident in the measured CH4 conversions, even when steam concentration was the subject test variable. Char hydrogasification adds CH4 to the product gas stream, but probably converts no more than 15 to 20 % of the char in the tests. The correlation coefficient between predicted and measured char conversions exceeded 0.8 and the std. dev. was 3.4 %, which is comparable to the measurement uncertainties. The evaluation of the predicted CH4 conversions gave a std. dev. greater than 20 %. Introduction The simulation work in this paper supports process modeling work to extrapolate gasifier performance from lab-scale datasets to the behavior in commercial gasifiers, via quantitative interpretations of the SRI test data and simulations based on validated reaction mechanisms. Experimental and Simulation Strategy The test coal used in this study was Powder River Basin (DECS-26) subbituminous with 43.6 as rec’d wt. % volatile matter; 5.0 % ash; and 0.1 % moisture and with 73.9 daf wt. % C; 5.6 % H; 19.0 % O; 1.1 % N; and 0.6 % S. The coal feed was classified to a mean size of 90 µm. The reactor was configured with a radiant igniter section and three Kanthal extensions for a total furnace length of 120 cm. The residence time for most runs was 1.40 s. The gas mixtures contained CH4, steam, H2, and CO (but no O2) in Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
proportions to achieve target ratios of the reactants. Argon was added to balance the flow. In most tests, the CH4:coal ratio (by carbon) was maintained at 40:60. Eighteen runs were qualified for the interpretations in this paper. In all tests, the pressure was 3.0 MPa and the gas flowrates were specified to deliver the coal suspension into the reactor from below at a nominal gas velocity of 30 cm/s, which is sufficient to prevent coal loss due to sedimentation. The coal flowrates were assigned from the reported suspension loadings, which varied on a weight basis (g-coal/g-gas) from 3.1 to 9 wt. %. Two runs had no coal flow. The carrier gas contained from 9 to 28 % steam; from 2.5 to 8 % CH4; and from 20 to 30 % H2; with a balance of Ar. These H2 levels were sufficient to produce negligible levels of soot in the recovered solid products when expressed as a fraction of the total carbon fed into the reactor. Temperatures for the extended section of the reactor were fixed at 1400°C in all runs but three, in which they were raised to 1450 and 1500°C. The igniter section was kept at 1500°C for all runs except three, in which it was raised to 1550°C. Hence, the primary variations in this test series were the coal loadings and steam and CH4 levels, with less variation in temperature and H2 level. Thermal histories for mean gas temperature and mean particle temperature were assigned from previously developed CFD simulations for the same pressure and gas flowrate and slightly lower coal loadings and slightly different carrier gas compositions. Along the transit time coordinate in the thermal histories, the inlet to the igniter is at 200 ms and the outlet is at 500 ms. The flow heats upstream of the igniter inlet due to radiation leakage into the upstream flow channel. Beyond the inlet, the gas temperature surges at almost 6000°C/s, and closely approaches the nominal reactor temperature at the inlet to the extended reactor section. The particle temperature considerably lags the gas temperature, and particles heat at 3350°C/s. Validated reaction mechanisms for each chemical process were used to quantitatively interpret the conversions for coal, char, and CH4 in individual tests. Primary devolatilization was described with NEA’s FLASHCHAIN® [1]. The distributions of primary volatiles were first processed through a secondary pyrolysis mechanism that converted tar into polynuclear aromatic hydrocarbons (PAH) and additional noncondensible gases, and then the PAH was instantaneously hydrogenated into additional gaseous hydrocarbons (GHCs), according to a global process [2]. The predicted distributions of products from primary devolatilization, secondary volatiles pyrolysis, and instantaneous PAH hydrogenation are collected in Table 1. Primary products are the products that leave the coal phase, before they undergo any secondary conversion. These products contain tar, albeit at the relatively low yield of about 10 daf wt. % because of the elevated pressure and high O-content of this particular coal. The latter feature explains why this distribution is dominated by the oxygenated gases, which comprise about 60 % of the ultimate volatiles yield. The minor amounts of GHCs and heteroatom species round out the distribution, and about half the coal reactant is converted into char. Both of the distributions that account for secondary volatiles conversion have the same yields of total volatiles and char, because secondary conversion does not deposit material on the parent particles. The hydrogenation products contain neither tar nor soot. Whereas the associated yield of PAH appears in the table, this yield is in parentheses to denote that PAH is ultimately hydrogenated into oils, C2H4, and CH4 [2]. Consequently, this distribution contains no product heavier than oils. Its levels of oxygenated gases are comparable to the others because all tar-O is released as these species (since PAH contains no oxygen, by definition). But it has the greatest contribution from GHCs, by far. The CH4 and C2H4 levels are especially elevated. In all succeeding simulations of the process chemistry, the volatiles were
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Table 1: Predicted volatile product distributions.
dry, ash-free wt. % Volatiles Tar Soot PAH Oils Char H2 O CO2 CO H2 CH4 C 2 H6 C 2 H4 C 2 H2 C3H6 HCN H2 S
Table 2: Comparison of equilibrium and kinetically limited product mole fractions
Primary
20 Pyrolysis
20 Hydrogenation
51.6 9.8 48.4 3.6 10.4 14.6 2.5 4.5 0.7 1.9 1.6 1.0 0.6
51.6 8.8 48.4 3.6 10.4 16.3 4.2 0.6 5.6 1.1 0.6
51.6 (7.2) 0.7 48.4 3.8 10.8 15.4 1.2 7.3 0.8 6.6 1.7 1.3 0.6
for uniform feed compositions at three reactor temperatures.
CH4 H 2O H2 CO CO2
1400°C Equil Reform Chem 59 ppm 0.035 0.118 0.165 0.489 0.414 0.090 0.058 0.006 0.007
1450°C Equil Reform Chem 11 ppm 0.011 0.161 0.176 0.380 0.354 0.072 0.062 0.009 0.009
1500°C Equil Reform Chem 7 ppm 0.003 0.161 0.164 0.379 0.373 0.072 0.070 0.008 0.008
assigned as the secondary hydrogenation products, because the yields of recovered soot were always negligible fractions of the total carbon fed into the reactor. Char gasification by steam with inhibition by H2 and CO was described with CBK/G, a mechanism previously validated for applications at elevated pressures in complex syngas mixtures [3]. Simultaneous char conversion via CO2 gasification was formally included, but found to be negligible for the very low CO2 pressures in the tests. Hydrogasification of char by H2 was described by a version of CBK that was recently validated with the literature database on hydropyrolysis. Reforming chemistry in the gas phase was described with 566 reactions among 154 species as heavy as phenol that were assembled from literature submechanisms which were extensively validated for applications at reducing conditions under elevated pressures. Due to the emphasis on comprehensive chemistry, these mechanisms were implemented in simulations that suppressed all aspects of turbulent fluid mechanics and convective transport phenomena. Series of CSTRs were used to coarsely resolve the time evolution of the reforming chemistry in nominal plug flow along the test reactor [4]. Finite-rate kinetics for both devolatilization and char gasification determined the incremental additions of gaseous fuel products to each CSTR in the series. Results and Discussion A first series of simulations estimated the relaxation time to equilibrate the gas compositions in the available transit time of 1.4 s. Table 2 compares the predicted product compositions from the kinetic simulations to the equilibrium compositions at three temperatures. All three cases had identical feedgas compositions. The product gas composition is not equilibrated at 1400°C, but does more closely approach the equilibrium composition at 1450°C, and nearly achieves it at 1500°C. We therefore expect that products recorded at 1400°C, in the bulk of this test series, will contain appreciable levels of intermediates and major products in proportions that differ significantly from the equilibrium products. Conversely, the product compositions from tests at 1500°C may be closely approximated by the equilibrium compositions. Our primary focus in this evaluation was to determine whether or not the simulation results can accurately depict the extents of CH4 reforming and coal conversion over the broad range of steam concentrations represented in our test matrix. To address this issue, we developed two interpretations based on disparate contributions from steam gasification and hydrogasification. Our most plausible analysis is based on default gasification reactivity parameters for the subject coal from NEA’s reactivity database, without further adjustment. This case gives char conversion mostly from
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100
100
Runs 77 - 94 SG Series r = 0.81 σ = 3.4 %
XCH w/ H2O & H2 Gasification
Predicted XCOAL
90
80
80
40
20
0
Runs 77 - 94 SG Series r = 0.60 σ = 23.5
4
70
60
-20
-40
60 60
70
80
90
100
Measured XCOAL
-40
-20
0
20
40
60
80
100
Measured XCH , % 4
Fig. 1: (Left Panel) Evaluation of the predicted extents of coal conversion via devolatilization and steam and H2 gasification based on default reactivities. Fig. 2: (Right Panel) Evaluation of the predicted extents of CH4 conversion via devolatilization and steam and H2 gasification based on default reactivities. steam gasification, with char conversions from hydrogasification of 15 to 18 %. It will be denoted by “SG.” The second case was deliberately formulated to have hydrogasification predominate, which required an increase in the hydrogasification reactivity by an order of magnitude. This adjustment is especially troublesome because the hydrogasification reactivities of a very diverse assortment of coal types do not vary by more than a factor of five, unlike the very strong dependence on coal quality in reported steam gasification reactivities. Char conversions from steam gasification ranged from only 4 to 9 % under this scenario. This case will be denoted by “HG.” Extents of coal conversion from the SG scenario are evaluated in Fig. 1. The predicted coal conversions from the SG analysis is accurate across the entire range of test conditions, in so far as the parity line in the figure is nearly the same as the regression line through the plotted points. The correlation coefficient exceeds 0.8 and the std. dev. is 3.4 %, which is comparable to the measurement uncertainties. But the char conversions from the HG analysis (not shown) were badly skewed off the parity line, so that conversions from 70 to 80 % were badly over-predicted, whereas those from 80 to 90 % were badly under-predicted. Clearly, a predominance of hydrogasification is inconsistent with the sensitivity to steam concentration in the reported extents of coal conversion. The regression line through the points from the HG scenario gave a correlation coefficient below 0.7 and a std. dev. of 1.9 %. The evaluation of the predicted CH4 conversions from the SG scenario appears in Fig. 2; that for the HG scenario was comparable. Both simulation scenarios seriously over-predict the lowest measured conversions, but have reasonably accurate predictions for intermediate CH4 conversions, and over-predictions for the greatest CH4 conversions. The case with excessive hydrogasification gave two tests with negative conversions (for which CH4 production via hydrogasification overcompensates for CH4 reforming), and one of these matched the measured conversion. Whereas the SG results are correlated more strongly with the measured values, neither scenario explains even most of the dispersion in the data. The std. devs. are comparable and greater than 20 % with both scenarios. It is difficult to select one scenario on the basis of the predicted CH4 conversions because both give weak performance, and because of another ambiguity in the dataset pertaining to the dependence on the steam concentration. For the five tests in which
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only the steam concentration was increased in succeeding tests from 10 to 30 %, the chemistry simulations predict that the CH4 conversions increase monotonically from 15 to 55 %. But the strong sensitivity to steam concentration in the simulated reforming chemistry is not clearly seen in the measured CH4 conversions, which show little, if any, monotonic tendency. Even among the measured CH4 conversions from all the tests at 1400°C with 30 % H2 but variable amounts of CH4, the measured values are uncorrelated with the steam concentrations across the entire range. Moreover, regressions of these CH4 conversions with the CH4 concentrations and with the coal loadings were just as bad. One explanation for this apparent inconsistency is that the unaccounted for source of CH4 that diminished the measured CH4 conversions is independent of the concentrations of both steam and CH4. Such is the case for the char hydrogasification mechanism. But the HG scenario performed markedly worse in interpreting the predicted extents of char conversion, and no better than the SG scenario in interpreting the CH4 conversions. Hence, our simulations to clarify various scaling issues were based on the SG scenario, for which steam gasification is the predominant mechanism of coal conversion, although char hydrogasification contributes appreciable amounts of CH4 to the product stream. Conclusions To summarize the main findings from our interpretation of the lab-scale test data: (1) Homogeneous reforming chemistry generates equilibrium gas compositions at 1500°C in the proposed transit time, but not at any of the lower temperatures in the test matrix. (2) Methane conversion in the gas phase increases for progressively hotter temperatures, so the temperature must be actively controlled to meet the conversion targets. But the strong predicted dependence on steam concentration was not evident in the measured CH4 conversions, even when steam concentration was the subject test variable. (3) Char hydrogasification adds CH4 to the product gas stream, but this process probably converts no more than 15 to 20 % of the char. The correlation coefficient between predicted and measured values exceeded 0.8 and the std. dev. was 3.4 %, which is comparable to the measurement uncertainties. (4) The evaluation of the predicted CH4 conversions is less satisfactory. The analysis seriously over-predicts the lowest measured conversions; becomes reasonably accurate for intermediate CH4 conversions; and over-predicts the greatest CH4 conversions. The std. dev. is greater than 20 %. Acknowledgment This work was sponsored by the Stategic Technology Office of the U. S. Defense Advanced Research Projects Agency under contract HR0011-10-C-0049 entitled, “JP-8 From Coal Via Methanol.” The views expressed are those of the author and do not reflect the official policy or position of the Deperatment of Defense or the U.S. Government. Distribution Statement “A” (Approved for Public Release, Distribution Unlimited). References [1] Niksa S, Liu GS, Hurt RH: Coal conversion submodels for design applications at elevated pressures. Part I. Devolatilization and char oxidation, Prog. Energy Combust.
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Oviedo ICCS&T 2011. Extended Abstract
Sci., 2003, 29(5):425-477. [2] Nelson PF, Huttinger KJ: The effect of hydrogen pressure and aromatic structure on methane yields from the hydropyrolysis of aromatics, Fuel,1986, 65: 354. [3] Liu GS, Niksa S: Coal conversion submodels for design applications at elevated pressures. Part II. Char Gasification, Prog. Energy Combust. Sci., 2004, 30(6):697-717. [4] Niksa S, Liu S: Advanced CFD post-processing for p. f. flame structure and emissions, 28th Int. Technical Conf. on Coal Utilization and Fuel Systems, Coal Technology Assoc., Clearwater, Fl, March, 2003.
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Biomarker and Petrographic Evidence for the Origin and Maturity of OilProne Arctic Coal and Associated Bitumen C. Marshall1, D.J. Large1, C.E. Snape1, W. Meredith1, B. Spiro2, I Mokogwu1 1
Department of Chemical and Environmental Engineering, University of Nottingham, University Park, Nottingham, United Kingdom, NG7 2RD Tel: +44 115 951 4166; Fax: +44 115 951 4115
[email protected] 2 Dept of Mineralogy, Natural History Museum, Cromwell Road, London, SW7 5BD, UK Abundant oil prone coal (Type III Kerogen) deposits are preserved within the high latitude, middle Palaeocene, Todalen Member of the Central Tertiary Basin, Spitsbergen. These coals provide an opportunity to understand the processes which control Arctic oil-prone coal formation and maturation. This paper uses high resolution sampling through a 1.5 m section of Longyear Seam from Mine 7, Longyearbyen, Svalbard. . Bitumen is unevenly distributed in the Longyear Seam with total extractable bitumen greatest at around 1.3-1.4 m. Bitumen yields are lower and more variable in the lower Longyear Seam with a transition to higher, more stable bitumen yields at around 0.8m above seam base. Increased bitumen yields towards the top of the seam are reflected by increased suppression of vitrinite reflectance (Ro) with values ranging from Ro 0.78 to 0.48%. The maceral composition of the Lower Longyear is characterised by high maceral variability and generally increased inertinite content. The transition seen in bitumen yields is replicated by the macerals with a rapid decline in inertinite (in particular fusinites) and increasing detrital and oil-prone macerals representing transition from oligotrophic to rheotrophic conditions during peatland deposition.
The
apparent link between maceral composition and bitumen yield within the Longyear coal is attributed
to
differences
in
oil
proneness
of
maceral
subgroups
and
their
augmentation/reduction of bitumen migration pathways. 1. Introduction
Oil prone coals are economically important as oil source rocks and as targets for coal liquefaction, however detailed research regarding the formation and characterisation of these coals has been limited to relatively few basins situated mostly within the southern hemisphere [1-5]. The factors and conditions required for oil prone coal formation are generally not well known, even within well studied basins such as the Gippsland (Australia) and Taranaki (New Zealand) Basins. However, two main groups of oil-prone maceral have been identified, comprising of liptinites and perhydrous vitrinites. Liptinites are hydrogen rich and are
generally derived from decayed leaf matter, plant waxes and resins, whereas the origins of perhydrous vitrinites are less well understood,
with
amorphous
vitrinite
apparently impregnated with hydrogen rich material.
Work regarding oil prone coals
within
the
northern
hemisphere,
particularly in the high latitudes is rare. The extensive coal deposits found within the Central Tertiary Basin, Svalbard have been the subject of only limited published research [6-8] with a focus upon the coals bulk characteristics, with only Orheim et al. 2007 [6] noting their oil proneness. This study will focus upon a 1.5m section cut from the Longyear seam from Mine 7, Figure 1 – Location map and regional geological map of Svalbard and major settlements.
close to the settlement of Longyearbyen, Svalbard (Figure 1).
High resolution sampling of the Longyear coal allows detailed study of the distribution of bitumen in the coals, highlighting any in-seam migration. Comparison of bitumen yields with maceral data will allow the identification of the hydrological and peatland environments required to form Arctic oil-prone coals.
2. Methodology
A 1.5m section of Longyear seam was taken from Store Norske Grubekompanie AS (SNSK) Mine 7 close to Longyearbyen, in the Arctic archipelago of Svalbard during a field visit in May 2010. The sample was cut into subsamples of average thickness 2.5 cm and crushed into <1mm and <100 µm fractions for maceral analysis and Soxhlet and Accelerated (ASE) solvent extraction respectively. The Tissue Preservation Index (TPI) and Gelification Index (GI) by Diessel [9] were calculated using the following formulae;
The coals were extracted using both techniques described above with 93:7 DCM/Methanol, however due to time constraints the less efficient ASE extraction method was used with an extraction period of 45 minutes/sample. The ASE consistently yielded less bitumen than the Soxhlet technique, however the long run times (2-3 weeks/sample) were considered too long for the number of samples required for extraction. Consequently, the Soxhlet technique was considered to represent the Total Extractable Bitumen and the ASE samples were calibrated using Figure 2. The maceral composition and vitrinite reflectance of the Longyear samples were determined using reflected light microscopy with point counts of 500 and 100 for maceral and vitrinite reflectance R0, respectively.
Figure 2 Calibration Graph between Accelerated Solvent Extraction and Soxhlet Solvent Extraction for selected Longyear coals 3. Results and Discussion
3.1 Maceral Analysis
Observation of the detailed maceral composition of the Longyear seam seen in Figure 3 shows that the Longyear seam contains significant amounts of fluorescent amorphous vitrinite which is recognised as a key indicator of oil-prone nature [9-10]. The seam exhibits
two distinct characters with a more variable, inertinite rich lower seam (0-0.75 m above seam base) and a more stable, low inertinite, high detrovitrinite, sporinite and fluorescent vitrinite upper seam (0.75-1.5 m).
The lower Longyear seam has a maceral composition of 70-80% vitrinite, 8-25% inertinite, 5-10% liptinite and around 3% mineral matter. The bulk vitrinite content of the lower seam remains relatively constant throughout the lower Longyear however the detrovitrinite and fluorescent perhydrous contents are generally lower than observed in the upper seam. The perhydrous vitrinites within the lower seam also fluoresce less strongly (dull brown) indicating that they are less oil-prone than those observed within the upper seam.
The TPI is thought reflective of both the preservation potential of the coal source material and the degree of preservation through burning [9]. The GI index is thought to reflect the hydrological conditions of peat during formation with higher values indicating stable hydrological conditions whereas lower values indicate more variable hydrology leading to oxidation of the peatland and preservation. Inertinite peaks seen within the lower seam are predominantly fusinite which is indicative of burning [12] preceded by decreasing semifusinite away from the oxidation front.
The lower seam exhibits a variable TPI which reflects the inertinite evidence of periodic oxidation and combustion of the peatland. Variability in organic material preservation may also reflect a less stable hydrological regime implying that during this period the peatland was more likely to be an ombrotrophic raised bog depending on local precipitation to maintain the water table. Consequently, this led to drier conditions within the peatland, reducing the degree of decomposition and increasing the risk of peatland fires.
These
conditions appear to reduce the oil potential of the peatland leading to lower amounts and fluorescence of perhydrous vitrinite.
The lower-upper seam transition is extremely rapid and is seen in many maceral profiles. It is preceded by a fusinite peak, indicating intense burning may have reduced the peatland to the level of the water table. However unlike earlier events (40cm above seam base), the raised bog did not recover and instead shifted from an ombrotrophic to rheotrophic hydrological regime. This transition from raised bog to fen is evidenced by a rapid decrease in inertinite, an increase in detrital macerals (sporinite and detrovitrinite) and a change to low
maceral variability in the upper seam indicating a switch from a less stable precipitation to a more stable groundwater fed regime.
In the upper seam the bulk vitrinite content is similar to that of the lower seam at 70-80%. However, the inertinite content exhibits low variability (~10%), large quantities of liptinites (~15%) appear in pulses and mineral matter is seen to increase gradually up seam to a maximum of ~10%.
The vitrinite composition of the upper seam shows a rapid increase in the number and fluorescence (Orange brown) of perhydrous vitrinites compared to the lower. This provides evidence that the fen depositional environment is more oil-prone than that that of the earlier raised bog. The fluorescence of vitrinite also appears to increase from 80 cm above seam base to the top of the seam providing evidence of possible migration within this section. This may be explained by upper perhydrous vitrinites absorbing migrated bitumen over time or may be solely due to differences in original bitumen content. The amounts of detrital macerals (detrovitrinite, sporinite, liptodetrinite) are also observed to increase above the lower-upper seam transition possibly providing a migration pathway for bitumen migration.
The increase in detrital macerals may also reflect the wetter conditions experienced in the later fen, allowing organic matter to be more easily transported and decomposed. This would is reflected by the decrease and stability of TPI in the upper seam and increase in GI post raised bog-fen transition.
Evidence from other perhydrous coals [4] indicates effective
decomposition is a requirement for the formation of perhydrous vitrinites and the rise of detrital macerals would support this.
Evidence from maceral analysis of the Longyear seam indicates that it can be divided into two peatland environments, an earlier ombrotrophic raised bog and a later rheotrophic fen. The earlier raised bog was hydrologically variable with drier and wetter periods, with periodic fire events. This variable environment enhanced preservation, however it appears this limited the oil potential.
The later fen was more hydrologically stable, providing
conditions for efficient decomposition and appears to have created an environment with greater oil-potential. Increased vitrinite fluorescence and perhydrous vitrinite content upseam may provide evidence of migration in this section.
Height above sean base (cm)
Figure 3 – Seam Profile of Longyear Coal (a) Bitumen Yield (ASE) vs. Bitumen Yield Soxhlet. (dashed line) (b) Vitrinite Reflectance (Ro) (c) % Inertinite (fusinite, semifusinite, funginite and inertodetrinite) (d) Perhydrous Vitrinite % (Fluorescent amorphous vitrinite) (e) TPI (see methods for calculation) (f) Log GI (see methods) (g) Mineral % (h) Liptinites (sporinite and cutinite) (i) Detrovitrinite (Vitrinite <10µm)
Height above sean base (cm)
3.2 Suppression of Vitrinite Reflectance (Ro) The vitrinite reflectance (Ro %) of the Longyear seam decreases from Ro 0.7 to 0.48% from the seam base to seam top (Figure 3). The large apparent change in maturity has previously been unreported due to a lack of high resolution sampling, however Orheim et al. 2007 [6] reports a significant maturity gap between the bulk Longyear seam and Svea seam (another seam within the Central Tertiary Basin) attributing this to the insulating effect of the lower coal seam.
However the diversity of Ro values and systematic change throughout the
Longyear seam indicates that this cannot be the case. The gradient seen within a single 1.5m coal seam represents a difference in burial depth in excess of 1km when calibrated to the depth vs. Ro graph used by Throndsen (1982) [13]. The oil-prone nature of the Longyear seam and the generally increased quantities of fluorescent perhydrous vitrinites with increasing height above seam base would indicate that this is caused by suppression of Ro by bitumen. This is supported by the distribution of bitumen within the Longyear coal with generally lower quantities of bitumen at the base of the seam and hence a lower degree of Ro suppression. Consequently it is assumed that the true rank of the Longyear coal is represented by the vitrinite reflectance seen at the base of the seam at Ro ~0.78%, similar to that reported by Orheim et al. 2007 [6] for the nearby less oil-prone Svea Seam.
3.3 Distribution of bitumen The measured and calibrated total bitumen extract profile of the Longyear seam is shown in Figure 3. The lower Longyear seam (0-0.75m above seam base) is characterised by variable bitumen contents ranging from 5% to 8% with the basal contact of organic shales yielding ~2% bitumen. The upper Longyear seam (0.75-1.5m above seam base) displays different characteristics with a trend of increasing total extractable bitumen from 4% to a maximum of 15% close to the top of the seam. At the top of the seam there is a rapid decline in the amount of bitumen extracted despite the bitumen impregnated appearance of this coal in hand specimen. The distribution of bitumen through the Longyear seam provides support for the evidence from coal macerals indicating that the earlier raised bog environment was less oilprone than the later fen environment.
The bitumen distribution within the upper Longyear seam resembles a typical migration profile; however migration within coals is considered difficult [10,14,15]. Comparison with various maceral parameters within the Longyear coal indicates that the increase in bitumen content may be linked to the stabilisation and reduction of inertinite content at around 75cm above seam base. At this time there is also a general increase in detrovitrinite and sporinite contents which have been highlighted as potential pathways for coal migration [14]. The change in bitumen yield could be attributed in differences in the amount of oil pronemaceral, however this is thought unlikely due to the relatively minor increases in oil-prone maceral contents in the upper seam compared to the large changes in bitumen yield seen within the upper seam. Detrital macerals are increasingly fluorescent above the raised bogfen transition and adjacent to perhydrous vitrinite boundaries indicating expulsion of oil into the surrounding matrix. Where detrovitrinites are absent from grain boundaries low fluorescence is observed indicating that the majority of migration is through the fine grained matrix. Observation of some perhydrous vitrinite boundaries grading into detrovitrinite may also indicate that expulsion from vitrinite may lead to extra detrovitrinite formation, potentially increasing porosity throughout the oil window. Unlike many bitumen migration studies [10], cleats within the Longyear coals do not appear to function as conduits for migration, with many cleats in-filled by mineral matter prior to generation. The evidence from the coal bitumen yields indicates that the later fen provided conditions which were more favourable for oil-prone coal formation and subsequent migration through detrital maceral migration pathways. The evidence from the Longyear coals indicates that migration within perhydrous coals is dependent upon the textural relationship between different maceral groups and the amounts of bitumen available for migration. 4
Conclusions
1. The maceral composition of the Svalbard coals shows that the seam can be divided into two hydrological regimes with an earlier ombrotrophic raised bog environment and a later rheotrophic fenland environment. The earlier raised bog appears to have been susceptible to desiccation and fire which appears to have limited the oil-potential of this coal.
The later fenland environment appears to have allowed greater
decomposition of organic matter, characterised by greater detrital material and
increased perhydrous vitrinite content and indicating this environment has greater oil potential 2. The maturity of the Longyear seam is thought to be around Ro 0.78%, however the large variation in vitrinite reflectance between the seam base and seam top makes this maturity parameter unreliable.
This variation is attributed to the increasingly
perhydrous nature of vitrinites towards the top of the seam suppressing Ro. 3. The bitumen yield profile of the upper Longyear seam appears to suggest migration. Migration is thought to be facilitated by the greater detrovitrinite contents of the upper seam providing migration pathways towards the seam top.
The enhanced
fluorescence observed towards the seam top is thought to be due to the adsorption of migrated hydrogen rich material into the perhydrous vitrinite structure.
Acknowledgements Financial Support by NERC Studentship is gratefully acknowledged. Many Thanks to Alv Orheim (GeoArktis AS), Malte Jochmann and Marte Oiesvold from Store Norske Grubekompanie AS (SNSK) who made sampling at the Longyearbyen Mine 7 possible.
References [1] Killops S.D, Raine J.I, Woolhouse A.D, Weston R.J. Chemostratigraphic evidence of higher-plant evolution in the Taranaki Basin, New Zealand, Org. Geochem 1995:23:5:429-45 Z [2] Killops S.D, Funnell R.H, Suggate R.P, Sykes R, Peters K.H et al. Predicting generation and expulsion of paraffinic oil from vitrinite rich coals, Org Geochem 1998:29:1:1-21 [3] Killops S.D, Mills N, Johansen P.E. Pyrolytic assessment of oil generation and expulsion from a suite of vitrinite-rich New Zealand coals Org Geochem 2008: 39:1113-8 [4] Sandison C.M, Alexander R, Kagi, R.I, Boreham, CJ. Sulfurisation of lipids in a marine influenced lignite. Org. Geochem 2002:33:1053-77 [5] Curry, D. J., Emmett, J. K. Hunt, J. W. Geochemistry of aliphatic-rich coals in the Cooper Basin, Australia, and Taranaki Basin, New Zealand: implications for the occurrence of potentially oilgenerative coals. In: Fleet, A. J. & Scott, A. C. Coal and coal-bearing strata as oil-prone source rocks: an overview, GSL Special Publications 77; 1994 [6] Orheim A, Bieg G, Brekke T, Horseide V, Stenvold J. Petrography and geochemical affinities of Spitsbergen Paleocene coals, Norway [7] Cmiel S.R, Fabianska M.J. Geochemical and petrographic properties of some Spitsbergen coals and dispersed organic matter, ijcoalgeo 2004:57:77-97
[8] Lewinska-Preis L, Fabianska MJ, Cmiel S, Kita S, Geochemical distribution of trace elements in Kaffioyra and Longyearbyen coals, Spitsbergen, Norway. Ijcoalgeo 2009:80:211-23 [9] Diessel C.F.K On the correlation between coal facies and depositional environments. In Proceedings of the 20th Newcastle Symposium ‘Advances in the Study of the Sydney Basin’ University of Newcastle, Newcastle, Australia p.19-22 [10] Wilkins RWT, George S.C, Coal as a source rock for oil: a review. Ijcoalgeo 2002: 50:317-61 [11] Diessel C.F.K, Gammidge L. Isometamorphic variations in the reflectance and fluorescence of vitrinite-a key to the depositional environment [12]
Collinson
Fire
and
M.E.,
episodic
Steart, runoff
D.S., and
Scott,
A.C.,
deposition
at
Glasspool, the
I.J.
and
Hooker,
Paleocene-Eocene
J.J..
boundary.
Journal of the Geological Society, London 2007:164: 87-97 [13] Throndsen T, Vitrinite reflectance studies of coals and dispersed organic matter in Tertiary deposits in the Adventdalen area, Svalbard. Polar Research 1982:2:77-91 [14] Stout, S. A. Chemical heterogeneity among adjacent coal microlithotypes implications for oil generation and primary migration from humic coal: In: Fleet, A. J. & Scott, A. C. Coal and coalbearing strata as oil-prone source rocks: an overview, GSL Special Publications 77; 1994 [15] Huang D. Advances in hydrocarbon generation theory II. Oils from coal and their primary migration model Journal of Petroleum Science and Engineering 1999:22:131-9
COAL QUALITY OF A MEGA COAL BASIN FROM XINJIANG, NORTHWEST CHINA
JING LI1,2, XINGUO ZHUANG1, XAVIER QUEROL2, ORIOL FONT2, PATRICIA CÓRDOBA2 1
Institute of Sedimentary Basin and Mineral, Faculty of Earth Resources, China University of Geosciences, Hubei, 430074, People's Republic of China, 2Institute of Environmental Assessment and Water Research, CSIC, C/ LLuis Solé Sabarís s/n, 08028 Barcelona, Spain
[email protected]
ABSTRACT This study focuses on the coal quality analysis of a mega coal basin in Xinjiang, northwest China-Junggar coal basin, as well as the leaching features of coal combustution by-products from power plants fed with Junggar coal. Based on the mineralogical, and geochemical analyses and leaching test, Junggar coal has a high coal quality: low ash yield, very low sulphur contents, low mineral and most trace element contents. Moreover, according with coal composition, fly ash and slag are characterized by low concentrations of most trace elements except B and Mo, the fly ash leachates have low trace element concentrations as well, which indicates a high potential utlilization of fly ash and slags arising from Junggar coal, with very low threat to environment. KEY WORDS: geochemistry, leaching, fly ash and slag, Junggar coal. INTRODUCTION The Junggar coal basin in Xinjiang, Northwest China contains super coal deposits and it is currently one of the most important coal mining resources in Western China (Figure 1). The Junggar coal basin is mainly composed of Northern Junggar coalfiled, Eastern Junggar coalfield, and Southern Junggar coalfield. The coal reserves of Junggar coal basin amount to 40% of China’s coal resources. Moreover, only in Eastern Junggar coalfield, the estimated up to date reserve is 164Gt[1]. Junggar coal basin is an typical intracontinental foerland basin[2]. According to the exploration data, the Eastern Junggar coalfield contains up to 13 workable coal seams of Middle Jurassic Xishanyao Formation, with the accumulated coal thickness of up to 103m, with maximum thickness of a single workable coal seam reaching 100m.
On the basis of borehole core data, this study mainly aims at dealing with the coal quality characteristic of this mega Junggar coal basin through geochemical, and mineralogical theories and methods. In addition, with the aim of evaluating the quality of the combustion by-products generated from coal-fire power plants in Xinjiang, which were mainly fed with the Junggar coal, the geochemical features of these fly ash, and slags were also investigated. Special emphasis on leaching potential and partitioning of trace elements among fly ash and slags were analyzed for the subsequent determination of potential applications of coal combustion by-products.
Figure 1: The location of Junggar coal basin and two power plants. METHODOLOGY Several coal samples were collected from borehole cores of six coal mines in different part of Junggar coal basin, respectively (Table 1). Three fly ash and two slag were also sampled from Hongyanchi (HONG) and Weihuliang (WEI) coal-fire power plants, that were mainly fed with Junggar coal (Table 1). Proximate analysis of coal, slag and fly ash samples were performed following the ISO-589, 1171, and 5623 recommendations. The mineralogical characteristics and particle morphology of coal, fly ash and slag samples were investigated by Powder X-Ray Diffraction and Scanning Electron Microscope with Energy Dispersive X-ray analyzer (SEM-EDX).
Table 1: Sampling list of coals, fly ashes and slags Types
Coal
Fly ash Slag
Sample Names ZNX 01-20 ZDZ 01-10 ZDS 01-06 ZDL 01-10 ZDD 01-27 ZDJ 01-26 HA WAc, WAf HS WS
Numbers. 20 10 6 10 27 26 1 2 1 1
Sample Locations Southern Junggar
Eastern Junggar
HONG power plant WEI power plant HONG power plant WEI power plant
Coal, fly ash and slag samples were acid-digested by using a two-step digestion method devised by Querol et al[2]. Then most trace element contents were analysed on the resulting solution by Inductively-Coupled Plasma Mass Spectrometry (ICP-MS), and the contents of major and selected trace elements were determined by Inductively-Coupled Plasma Atomic-Emission Spectrometry (ICP-AES). Specifically, Mercury analysis were directly analysed on feed coal, fly ash and slag samples, using a LECO AMA 254 gold amalgam atomic absorption spectrometer (GA-AAS). The European Standard leaching test EN-12457[3] was applied to fly ash and slag samples to determine the leaching potential of major, and trace elements. The pH and ionic conductivity of the leachates were determined by conventional methods. The content of major and trace elements of the leachates were determined by ICP-AES and ICP-MS. The content of Hg was determined directly on leachates by the same procedure as for the solid coal samples using GA-AAS. COAL CHARACTERISTICS Based on data collected from exploration drilling of six coal mines, as well as the proximate analysis of samping coals, the Junggar coal is characterized by middle moisture content (12%), low ash yields (8.2%, dry basis), middle-high volatile matter (31%, dry, ash-free basis), and very low sulfur content (mostly less than 0.2%, dry basis) (Table 2). Table 2: Proximate analysis and sulfur content ( %) of the studied coals M-moisture, A-ash yield, V-volatile matter, S-sulfur, ad- air dry basis, db- dry basis, daf- dry a, ash-free basis.
% ZNX ZDZ ZDS ZDL ZDD ZDJ Average M ad 3.9 12 15 13 15 13 12 A db 9.5 3.8 14 7.4 4.5 12 8.2 V daf 31 28 41 34 33 27 31
S db
0.2
0.1
0.6
0.3
<0.1
0.2
<0.2
According to the coal rank parametres, the coal rank of Junggar coal falls within sub-bituminous to bituminous coal.
Mineralogical Characteristic The minerals present in very low contents in the all Junggar coal mines were mainly quartz, kaolinite, with traces of siderite, calcite and dolomite (Table 3). Table 3: Mineral contents ( %) in the studied coals from six coal mines K-kaolinite, Q-quartz, Cal-calcite, Dol-dolomite, Sid-siderite, Illi-illite, Py-pyrite, Gy-gypsum, nd-not detected. %
K
Q
Cal Dol
Sid
ZNX ZDZ ZDS ZDL ZDD ZDJ
2.4 1.2 6.7 3.5 2.0 4.1
2.6 0.9 3.3 2.7 2.0 5.5
1.2 0.2 2.4 0.2 0.2 0.1
1.1 nd 0.1 nd 0.1 0.1 0.4 0.1 0.4 <0.1 0.6 nd 0.6 0.1 0.1 0.1 <0.1 <0.1 0.2 <0.1 0.2 <0.1 nd nd
1.9 0.2 0.4 0.2 0.4 0.1
Illi
Py
Gy
others Sum 0.2 0.7 0.2 0.3 0.8 1.3
9.5 3.8 14 7.8 5.6 11
Illite, pyrite, and gypsum were also present in trace amount in most Junggar coal mines (Table 3), and ankerite, microcline, anhydrite, anorthite, clinochlore, and albite were in some coal mines. Meanwhile, illite-montmorillonite, saponite, aragonite, and palygorskite were only detected in ZDJ coal mine, and rhodochrosite was detected in ZDZ coal mine. Based on the borehole exploration data, despite the low contents in whole junggar coal basin, mineral contents were relatively lower in coal-bearing areas with thick and few coal seams (ZDD, ZDZ, and ZDL). Where the coal seams split and thin, the mineral contents were relatively high. Geochemical Characteristic The concentrations of most major and trace elements in Junggar coal were very low compared with their concentration ranges in worldwide and Chinese coals[4-8]. Mn, P, and Ta concentrations in some Junggar coal mines, were still in their typical worldwide range, but slightly higher than their concentration ranges in Chinese coals. Moreover, the contents of Sr and Ba in some Junggar coal mines were even higher than their maximum values reported for worldwide coals. The low ash yield, very low sulfur content, low mineral contents, and low trace element concentrations indicate a high coal quality of Junggar coal. The low content of environmental relevance trace and rare earth elements is an advantage for industrial application of Junggar coal.
CHRACTERISTICS OF COMBUSTION BY-PRODUCTS Compared with European fly ashes, except B and Mo, concentrations of most trace elements in the sampling fly ashes fell in the lower part of the European concentration ranges[9]. Moreover, with the exception of Hf, Ni, Co, and Li, most of the trace elements showed a trend of enriching in fly ashes rather than in slags, and fine fractions (WAf) of fly ash had higher concentration of most trace elements than coarse fraction (WAc), which probably was attributed to the larger surface area in finer particles of fly ash than coarse ones, allowing for absorption of volatile elements on the surface of the fly ash[2,10]. Leaching Characteristic According to the leaching test, the pH values measured for leachates reach 11-12 for fly ash and 9-10 for slag. The conductivity of the leachates reached from 380 to 8000 μS/cm2. The pH and conductivity of studied fly ashes all fell within the higher ranges reported for European fly ashes. With respect to the element concentrations, most of the trace elements, including toxic elements have very low leachable proportions, compared with European fly ashes. In addition, according to the reference for characterization of landfill materials of European Council Decision 2003/33/EC (EULFD)[11], the leachable concentrations of some toxic metals from fly ashes are all taken as non-hazardous landfill materials, and except Ba, Cr, Mo, and Se, most of them even belong to inert landfill materials. CONCLUSION The present mineralogical, geochemical analyses of Junggar coals and the leaching test for fly ash and slags indicate the following properties. The very large coal reserves explored at Junggar coal basin are made up of very high quality coal, with low ash yield, very low sulphur contents, low mineral contents, as well as low concentrations of trace element. Due to the very low leaching potential of most trace elements, the fly ash and slag arising from the combustion by-products in power plants fed with Junggar coal have a very high potential for utillization, with very low environmental limitations. ACKNOWLEDGEMENT The author would like to show great gratitude to the Spanish Ministry of Foreign Affairs for supporting this study, and to the Nineth Geological Team of Xinjiang Bureau of Prospecting and Development of Geology and Mineral Resources, Urumqi, Xinjiang, China, for the coal and fly ash samples collecting. This study had financial support from the National Natural Science Foundation of China (Nos. 40572089 and 40972104). REFERENCES
(1) Zhou J., Zhuang X., Alastuey A., Querol X., and Li J.. International Journal of Coal Geology, 2010, 82, 51-67. (2) Querol, X., Whateley, M.K.G., Fernandez-Turiel, J.L., and Tuncali, E.. International Journal of Coal Geology, 1997, 33, 255-271. (3) British Standards Institution. BSI, London, 2002. (4) Swaine D.J.. Butterworths, London, 1990. (5) Ketris M.P., and Yudovich Y.E.. International Journal of Coal Geology, 2009, 78, 135-148. (6) Dai, S., Zhou, Y., Ren, D., Wang, X., Li, D., and Zhao, L.. Science in China Series D: Earth Science, 2007, 50, 678-688. (7) Zhuang, X., Querol, X., Zeng, R., Xu, W., Alastuey, A., Lopez-Soler, A., and Plana, F.. International Journal of Coal Geology, 2000, 45, 21-37. (8) Dai, S., Li, D., Chou, C., Zhao L.., Zhang, Y., Ren, D., Ma, Y., and Sun, Y.. International Journal of Coal Geology, 2008, 74, 185-202. (9) Morenoa N., Querol X., Andrés J.M., Stanton K., Towler M., Nugteren H., Janssen-Jurkovicová M., and Jonese R.. Fuel, 2005, 84, 1351-1363. (10) Dai, S., Zhao L., Peng S., Chou C., Wang X., Zhang Y., Li D.http://www.sciencedirect.com/science?_ob=ArticleURL&_udi=B6V8C‐4VXMPHW‐1&_user=422227 2&_coverDate=04%2F01%2F2010&_rdoc=1&_fmt=high&_orig=gateway&_origin=gateway&_sort=d& _docanchor=&view=c&_searchStrId=1745953589&_rerunOrigin=google&_acct=C000048559&_versio n=1&_urlVersion=0&_userid=4222272&md5=2670498fe337a23d81181ee4096e96ec&searchtype=a ‐ aff2 and Sun Y.. International Journal of Coal Geology, 2010, 81, 320-332.
(11)European Council Decision 2003/33/EC, Official Journal of the European Communities, 2003. 16, 27-49.
Oviedo ICCS&T 2011. Extended Abstract
Attempted production of blast furnace coke from Victorian brown coal Alan L. Chaffee1,2, M. Mamun Mollah1, Ralph S. Higgins3, Marc Marshall1,2 and W. Roy Jackson1,2 1
School of Chemistry, Monash University, Victoria 3800, Australia
2
Centre for Green Chemistry, Monash University, Victoria 3800, Australia
3
Consultant, 60A Arkaringa Cr, Black Rock, Victoria 3193, Australia
Corresponding Author Name: Alan L. Chaffee Email:
[email protected]
Abstract
This project aims to make coke suitable for use in blast furnaces from Victorian brown coal. The ever-increasing cost of coking coals makes this prospect attractive because Victorian brown coal is cheap, easily accessible and has very low concentrations of mineral impurities. However, in its as-mined condition, Victorian brown coal does not form coke. For example, earlier work [1] produced a hard char which was too fragile and too reactive to be used in blast furnaces.
In the current project the effect of adding carbon-rich materials derived from brown coals is being evaluated as a means of obtaining coke precursors. These precursors are then subjected to various thermal treatments, including low-temperature air-curing and heating in a furnace at high temperatures, to obtain products which can be tested to ascertain whether they have the physical properties required for blast furnace coke. In addition, the effect of inorganic cementing agents, either alone or in conjunction with the brown-coal-derived materials, is under investigation.
Oviedo ICCS&T 2011. Extended Abstract
1. Introduction Coking coals which are generally of bituminous rank, when heated, usually as small pieces, melt and agglomerate, obliterating their original shape. The coke is produced in large, fused strong pieces of low reactivity which are suitable for the production of iron in a blast furnace. In contrast, brown coal does not melt on heating and the solid residue from the carbonization is not massive, fused or inherently strong. It is characterized by relatively high reactivity, resembles charcoal more than coke and is usually referred to as a char [1, 2]. Attempts to overcome these problems have involved starting with large pieces of brown coal, briquettes obtained by agglomerating finely ground brown coal under high pressure and subjecting them to a carefully defined heating regime. Preliminary investigations showed that a hard reactive char could be obtained by modulating the heating rate during carbonization[3], but without any understanding of the processes involved, it was difficult to scale up the experiment or even to allow for variation in the total heating time or the size of the briquettes used. Rummel [4], in earlier work on German brown coals, noted that cracking and weakening of the briquettes occurred because of the differential shrinkage stress caused by the temperature difference between the surface and the core of the briquettes. He suggested that the heating rate should be controlled so that during carbonization the temperature difference between the core and surface of the briquette was constant. Experiments in which the heating rate was modified using this criterion were carried out by Megler and Kennedy [5]. The modified heating rate did not closely resemble that which was found empirically desirable, and did not give stronger char than that obtain by a constant rate of temperature increase. For this study, a different criterion for determining the optimum heating rate for a given total heating time was investigated, and a brief account is given of preliminary experiments for preparing additives which could decrease the reactivity of the char. 2. Experimental 2.1 Materials For the heating rate tests, standard industrial H-type Yallourn briquettes manufactured by the State Electricity Commission of Victoria were used. The manufacture of such briquettes is discussed in Herman [6]. The briquettes were cut in half at right angles to the longest dimension, as only individual halves would fit in the furnace. For reactivity-decreasing additives production, Loy Yang (Victoria) run-of-mine low-ash coal, ground to -3mm and dried at 1050C under N2 for 3 h, was used. 2.2 Shrinkage and temperature gradient measurements and other briquette tests The apparatus used was that described by Megler and Kennedy [5], with minor modifications [7]. A stainless steel chamber was heated electrically by an external Nichrome coil wound so as to produce a uniform temperature in the chamber. The briquette under test rested on silica spacers on a steel plate suspended by steel rods from the lid of the furnace. A steel plunger resting on top of the briquette passed through the lid via a gas-tight connection made by means of a rubber bellows such that the plunger could move freely; its movement was recorded by a dial micrometer. Control tests made by substituting a stainless steel block for the briquette made it possible to allow for movement of the plunger due to thermal expansion effects in the furnace [5]. One thermocouple was in contact with the surface of the briquette
Oviedo ICCS&T 2011. Extended Abstract
and another, enclosed in a twin-hole porcelain insulator, was inserted in to a hole drilled into the briquette so that the thermocouple junction was at the geometric centre of the briquette. Both thermocouples were of heavy gauge, for sturdiness. Thus the shrinkage and temperature at the surface and in the core of the briquette could all be measured simultaneously. The thermocouple reading at the surface of the briquette was used to regulate (manually) the electrical input to the furnace to obtain the required rate of temperature rise. A small rate of nitrogen flow through the furnace was maintained during the trial to prevent burning of the briquette. For the experiments of this study, the arrangement of the thermocouples was changed. Two very fine gauge thermocouples enclosed in fibre glass insulators both passed into holes drilled in the briquette. The hole for the surface thermocouple ended just below the surface of a polished or extrusion face, whereas the other extended to the geometric centre of the briquette. This arrangement ensured that reliable readings could be obtained with minimum disturbance of the temperature distribution in the briquette. The other change was that the briquette was heated in the furnace for at least 10 h at 200oC under the stream of nitrogen before the carbonization test began, to ensure that the briquette was dry and thus to separate shrinkage effects due to drying from those due to volatilization. After this drying procedure, the temperature was raised from 200 oC to 800 oC in 6h at rates determined by the experiment being conducted. The carbon content of briquettes raised to various temperatures was measured. 2.3 Autoclave experiments to produce additives The dried coal (3-6g) was charged into a 70mL stainless steel autoclave which was evacuated, weighed, pressurized with 3MPa (cold) N2 and weighed to determine the free space in the autoclave. It was evacuated, filled with 3-6MPa H2 and weighed. The autoclave was lowered into a preheated sand bath, and came to temperature in 5-6 min. It was held at temperature for 1 h, removed, allowed to cool and weighed. The gas was vented through a gas chromatograph and analyzed, but the analyses are not discussed here. The solid and liquid products were scraped and washed out of the autoclave into a flask with dichloromethane (DCM). The product was freed from water by Lundin distillation, and then ultrasonicated for 10 min, filtered, more DCM added to wash the filter cake back into the flask, and ultrasonication and filtration were repeated. The DCM insolubles were dried at 105 oC in flowing N2 for at least 2 h, cooled and weighed. Most of the DCM was removed from the DCM solubles by rotary evaporation and n-hexane (20:1 by weight) added. The mixture was ultrasonicated for 3 min and filtered. The insoluble material, asphaltene, was dried in a vacuum oven (0.1kPa, 55 oC), cooled and weighed. Most of the n-hexane was removed from the filtrate by rotary evaporation and the hexane-solubles (oil) stored for future use as cementing additive for the briquettes to be carbonized.
Oviedo ICCS&T 2011. Extended Abstract
3. Results and Discussion 3.1 Carbon content of heated briquettes Table 1 gives the carbon content of Yallourn brown coal heated in an inert atmosphere to various temperatures. It will be seen that at 800-900 oC the non-mineral fraction of the char is approaching 100% carbon, so that the fuel has been fully upgraded.
Table 1: Increase of carbon content of Yallourn brown coal char. Temperature (oC)
Carbon (wt% daf)
20
67.2
200
68.6
400
77.2
600
92.0
800
96.4
900
97.6
3.2 Shrinkage-temperature measurements and the ideal heating curve The shrinkage was determined as a function of core temperature of the briquette for a linear heating rate of 100oC/hour over the temperature interval 200-800oC (The core temperature rather than the surface temperature was used because the shrinkage will follow the former more closely [7]). Three trials gave similar curves, which agreed closely with each other and were similar to that deducible from Fig. 4 of [5] except near 400oC. The briquette began to shrink in the vicinity of 300oC and the shrinkage rate passed through a maximum ca 400oC before settling down to a steady rate which was maintained to ca 700oC. Above 700oC the shrinkage rate decreased, to give a total shrinkage to 800oC of about 18% in linear dimension relative to the dried state. The core-surface temperature difference as a function of surface temperature showed two peaks, centered at ca 300oC and 500oC, as found by Megler and Kennedy [5], but the first peak was much less pronounced (ca 40oC rather than the 90oC found by [5]). The second was also about 40oC, and the difference remained about 10oC right up to 800oC, as found by Megler and Kennedy. The differential shrinkage between the surface and core of the briquette was calculated from these results as a function of surface temperature. It rose to a peak of 2.2% near 400oC, then plateaued at ca 1% in 430-550oC, before falling steadily to a negligible value at 800oC (Fig.1).
Oviedo ICCS&T 2011. Extended Abstract
Ideal
3
Differential Shrinkage (%)
Linear heating N J 2
1
0 200
300
400
500
600
700
Briquette Surface Temperature (oC)
Figure 1: Differential shrinkage for linear heating (100oC/min) compared to differential shrinkage when the theoretical heating cycle was used (experiments J, N) and the ideal theoretical shrinkage. Modified from [7]. A theoretical heating cycle over the time of the experiment was calculated so that the rate of heating at any surface temperature was made inversely proportional to the differential shrinkage at this temperature. If it be assumed that the differential shrinkage is proportional to the rate of heating, this should give constant differential shrinkage during heating and hence minimize cracking stresses on the briquette. Experiments in which this calculated heating cycle was used (J, N) rather than a constant rate of heating showed that the maximum differential shrinkage was indeed reduced from 2.2% to 1.2% (Fig.1). The theoretical heating cycle was extrapolated above 800oC at the rate calculated for 750800oC. Fig.2 compares the theoretical cycle with the average of those obtained during carbonization runs in the ‘South Melbourne carbonizer’ which yielded a high proportion of hard char, using the same total heating time. The general features are similar in that, in both cases, the heating rate is lower below 600oC and higher in the range 600-900oC. It is possible that improvement in char quality would result from following the theoretical curve more closely.
Oviedo ICCS&T 2011. Extended Abstract
Figure 2: Comparison of theoretical heating curve with the successful empirical heating curve. Modified from [8]. 3.3 Additives for reducing char reactivity As noted above, char lumps obtained by adopting a suitable heating regime are hard, but still too reactive for use in blast furnaces. One suggestion for decreasing the reactivity is to add a binder to the briquettes, which may be derived from the coal itself by adding tar produced during the first part of the carbonization [9] or as a coal solvent extract [10]. In this study an attempt will made to obtain a binder by hydrogenating the coal to be coked, with as little preparation and under as mild conditions as possible, compatible with obtaining a good yield of oil or asphaltene which could be used as binders. Initially a Loy Yang low-ash coal was hydrogenated at 405oC for 1 h under 3MPa hydrogen pressure in a 70mL autoclave. Variation in the charge (3 or 6g) or the particle size of the coal (-3mm, -0.25mm) had little effect on the yields; the DCM insoluble, water and gas yield was 24.5±0.5 wt% db and the asphaltene yield was negligible (<1 wt% db) in all cases. Since the gas yield (mainly CO2) was ca 8-10 wt% db and, on the basis of previous, large scale experiments, the water yield would be expected to be ca 5-10 wt% db, the oil yield was small. More severe conditions, and iron catalyst will be trialled next to improve the yield of useful products. After a binder is produced, it will be added to the coal and carbonization attempted possibly with an optional air-curing stage which has been reported to improve the char strength [9, 11].
Oviedo ICCS&T 2011. Extended Abstract
4. Conclusion A successful heating cycle was calculated which was similar to that found empirically to give hard, but reactive brown coal char, and a beginning has been made in obtaining a suitable binder which it is hoped will reduce the reactivity of the char. Acknowledgement. This project is supported by the Energy Technology Innovation Strategy (ETIS) of the State Government of Victoria (Kyushu/Victoria Brown Coal Research and Development Initiative). References [1] [2] [3]
[4]
[5] [6] [7]
[8] [9] [10] [11]
Kennedy GL, Evans DG. Metallurgical fuel from Victorian brown coal. Journal of the Institute of Fuel 1958;31:242-247. Kennedy GL. Review: Development of the production of metallurgical fuel from brown coal. Journal of the Institute of Fuel 1960;33:598-608. Higgins RS, Kennedy GL, Evans DG. The development of brown coal char as a new metallurgical fuel. Proceedings - Australasian Institute of Mining and Metallurgy 1960;No. 195:103-116. Rummel R. Uber Die Abhangigkeit der Koksgute von der Beschaffenheit der Braunkohlenbriketts und der Fuhrung Ihrer Verschwelung; Oel und Kohle 1944;40:709-723. Megler VR, Kennedy GL. Heating of brown coal briquets to produce strong char. I. Critical assessment of existing theories. Fuel 1961;40:255-274. Herman H. Brown coal : with special reference to the State of Victoria. Melbourne. State Electricity Commission of Victoria; 1952;192-471. Ellis HJW, Kennedy GL. Further investigations of the thermal and shrinkage behaviour of Yallourn briquettes during carbonization; interim report submitted to Gas and Fuel Corporation of Victoria; 1964, March. Kennedy GL. The Production of Hard Char from Yallourn Briquettes, final report submitted to Gas and Fuel Corporation of Victoria; 1963, July. Moran RF, Joseph RT. Applicability of coals to the FMC coke process. Trans. Soc. Min. Eng. AIME. 1976;260:29-32. Hamaguchi M, Inoue A. Iron ore-containing coke and their manufacture. JP 2009180046 2011032370, Taylor JW, Coban A. Formed coke from lignite, and the critical role of air. Fuel 1987;66:141-142.
Oviedo ICCS&T 2011. Extended Abstract
Chlorine Retention during the Pyrolysis of a Western Australian Lignite in a Fluidised-bed Reactor
Jie Zhang, Cyril Kelly, Angelina Rossiter, Shan Wang and Chun-Zhu Li Fuels and Energy Technology Institute, Curtin University of Technology, GPO Box U1987, Perth, WA 6845, Australia.
[email protected] Abstract Chlorine in coal is an important consideration for coal utilisation both in terms of environmental protection and equipment corrosion. A high-chlorine (> 5 wt%) Western Australian lignite was pyrolysed in a fluidised-bed reactor to investigate the retention of chlorine in char during pyrolysis as a function of temperature and particle size. The results showed that a significant fraction of chlorine can be volatilised during pyrolysis. As the pyrolysis temperature increased from 400 to 600 °C, the chlorine retention increased significantly mainly due to the enhanced inter-particle volatile-char interactions at higher pyrolysis temperatures. As the lignite average particle size increased from 0.03 to 4.00 mm, the chlorine retention reached a maximum at the average particle size of 0.80 mm. This was again due to the changes in the extents of both intra-particle and inter-particle volatile-char interactions with particle size.
1. Introduction Chlorine is a minor element in coal that greatly affects the coal utilisation both in terms of the formation of air pollutant (HCl) and the corrosion of equipment. Pyrolysis is the initial step of many coal thermo-chemical utilisation processes [1]. Chlorine in coal can undergo various physical and chemical transformations during pyrolysis, greatly affecting the fate and distribution of chlorine in the pyrolysis products [2-4]. The retention of chlorine in char during the pyrolysis of Victorian brown coal in a fluidised-bed/fixed-bed reactor [2] increased rapidly with increasing temperature as a result of intensified inter-particle volatile-char interactions. Continuing our previous study [2], the main purpose of this study was to investigate the retention of chlorine during the pyrolysis of a high-chlorine (> 5 wt%) lignite. In addition to the inter-particle volatile-char interactions, this study also focuses on the importance of intra-particle volatile-char interactions on the retention of chlorine in char during pyrolysis.
1
Oviedo ICCS&T 2011. Extended Abstract
2. Experimental Section Two high-chlorine Western Australian lignite samples were used in this study. Sample A had a particle size range of 0.063-0.600 mm and contained 12.3 wt% of chlorine. Sample B had a chlorine content of 5.47 wt%, from which six sub-samples were prepared having average particle sizes of 0.03, 0.21, 0.48, 0.80, 2.00 and 4.00 mm. The samples were dried at 105 °C overnight in nitrogen prior to pyrolysis. The pyrolysis experiments were carried out using a lab-scale fluidised-bed reactor system [5]. Briefly, it consisted of a feedstock hopper, a feeding screw, a fluidised sand bed, two cyclones and a novel oil collection system. About 1.5 kg silica sand (0.3550.500 mm) was fluidised by pre-heated nitrogen. After the desired temperature was reached, the lignite was fed at 10 g/min into the fluidised sand bed for the lignite particles to be heated up rapidly. The retention of char in the fluidised bed changed with the char particle size. While the large char particles remained in the sand bed, the small char particles were elutriated out of the fluidised bed and then captured with two cyclones maintained at 420 and 400°C. The chlorine contents of lignite and char samples were measured according to the Eschka method (AS1038.8.1). Briefly, 2 g of the sample was oxidised in intimate contact with and covered by 5 g of the Eschka mixture in a muffle furnace at 675 °C to remove the organic material and convert all chlorine into chloride. The oxidised ash samples were dissolved in nitric acid and the chlorine contents were then determined with a modified Volhard method.
3. Results and Discussion Figure 1 shows the chlorine retention in char for Sample A as a function of pyrolysis temperature. With increasing pyrolysis temperature from 400 to 600 °C, the chlorine retention increased from 81% to 95%, in broad qualitative agreement with the previous observation on the pyrolysis of Victorian brown coal in a fluidised-bed/fixed-bed reactor [2]. The increases in the chlorine retention were mainly due to the enhanced volatilechar interactions. With increasing temperature, more active sites (e.g. reactive free radicals) in the coal/char matrix were formed to react with Cl-containing species (e.g. HCl) in the volatiles to re-capture the released chlorine. Figure 2 shows the chlorine retention in char during the pyrolysis of Sample B at 500oC. As the average particle size increased from 0.03 to 4.00 mm, the chlorine
2
Oviedo ICCS&T 2011. Extended Abstract
retention reached a maximum value of around 90% at the average particle size of 0.80 mm. This was mainly due to the changes in the extents of both intra-particle and interparticle volatile-char interactions. Firstly, with increasing average particle size, the chlorine-containing species (volatile precursors) would have to diffuse through an increasing length of distance within the pyrolysing particles to be released. This increased retention time of volatile precursors within the pyrolysing particles would encourage the intra-particle volatile-char interactions to increase the chlorine retention. Secondly, the increases in particle size caused the weight percentage of the char remaining in the fluidised-bed reactor (termed bed char) to increase rapidly from almost 0% to 99.5%. In other words, instead of being elutriated out of the bed, more char particles stayed as the top layer of the fluidised bed (char particles having lower density than the sand particles). The increased percentage of bed char meant that the volatiles formed from the lignite particles fed into the reactor at a later stage would have to interact with the char particles formed from the lignite 100
Chlorine retention, %
90
80
70
60
50 400
450
500
550
600
o
Temperature, C
Figure 1 Chlorine retention in char with increasing temperature (Sample A). 100
Chlorine retention, %
90
80
70
60
50 0.0
0.5
1.0
1.5
2.0
2.5
3.0
3.5
4.0
Average particle size, mm
Figure 2 Chlorine retention in char at 500 °C as a function of particle size (Sample B).
3
Oviedo ICCS&T 2011. Extended Abstract
particles fed into the reactor at an earlier stage, enhancing the inter-particle volatile-char interactions. The enhanced intra-particle and inter-particle volatile-char interactions thus resulted in significant increases in chlorine retention with increasing particle size. Further increases in lignite particle size (> 0.80 mm), however, caused the particles to “sink” into the fluidised bed, reducing the extent of inter-particle volatile-char interactions. The reduced extent of inter-particle volatile-char interactions was therefore the key reason for the decreases in the chlorine retention in char for large particles (Figure 2).
4. Conclusions The chlorine retention in char during the pyrolysis of a high-chlorine Western Australian lignite has been studied in a fluidised-bed reactor. As the temperature increased from 400 to 600 °C, the chlorine retention increased significantly mainly due to the enhanced volatile-char interactions. As the average particle size increased from 0.03 to 4.00 mm, the chlorine retention reached a maximum value of 90% due to the changes in the extents of inter-particle and intra-particle volatile-char interactions.
Acknowledgments This study was supported by the Western Australian State Government via the Centre for Research into Energy for Sustainable Transport (CREST) and the Spitfire Oil Pty Ltd.
References [1] Li C-Z. Advances in the Science of Victorian Brown Coal. UK: Elsevier; 2004. [2] Quyn DM, Wu HW, Li C-Z. Volatilisation and catalytic effects of alkali and alkaline earth metallic species during the pyrolysis and gasification of Victorian brown coal. Part I. Volatilisation of Na and Cl from a set of NaCl-loaded samples. Fuel 2002; 81(2): 143-9. [3] Tsubouchi N, Ohtsuka S, Nakazato Y, Ohtsuka Y. Formation of hydrogen chloride during temperature-programmed pyrolysis of coals with different ranks. Energy Fuels 2005;19: 554-60. [4] Takeda M, Ueda A, Hashimoto H, Yamada T, Suzuki N, Sato M, et al. Fate of the chlorine and fluorine in a sub-bituminous coal during pyrolysis and gasification. Fuel 2006;85:23542. [5] Garcia-Perez M, Wang XS, Shen J, Rhodes MJ, Tian FJ, Lee WJ, Wu H, Li C-Z. Fast pyrolysis of oil Mallee woody biomass: Effect of temperature on the yield and quality of pyrolysis products. Ind. Eng. Chem. Res. 2008; 47:1846-54.
4
THE FATE OF Hg AT TWO COAL POWER PLANTS EQUIPPED WITH FGD Córdoba P 1, Font O1, Izquierdo M1, Querol X1 1
Instituto de Diagnóstico Ambiental y Estudios del Agua (IDÆA-CSIC). Jordi Girona 18-26. E-08034- Barcelona. Spain. E-mail:
[email protected] ABSTRACT As result of sampling campaigns carried out in 2007 and 2008 at two Spanish power plants (PP1 and PP2) equipped with a forced oxidation wet flue gas desulphurisation (FGD) system with water re-circulation, a differential speciation of Hg in the gas OUTFGD was revealed. At PP1 and PP2 a high proportion (86-88%) of Hg escapes the electrostatic precipitator (ESP) in gaseous form, being Hg2+ (75-86%) the specie incoming the FGD at two power plants. By contrast, at PP2, Hg2+ was the prevalent Hg specie OUT-FGD (66%), whereas Hg0 (71%) was the major Hg specie OUT-FGD at PP1, reaching Hg reduction levels of 65% and 21% at PP2, and 21% at PP1.The differential speciation of Hg in the FGD system at PP1 and PP2 may attributable to: i) the high soluble salt concentrations of the FGD water streams at PP2, which reduce the gaseous solubility and probably, the gas retention efficiencies; ii) the entraining of HgCl2 droplets by the gas OUT-FGD; and iii) the operational conditions at PP2 such as, limestone purity, electrostatic precipitator gas temperature, use of additives, fluoride and/or sulphate complexes, and the S/F and S/Cl ratios in the scrubber. The potential parameters controlling Hg speciation and partitioning are currently investigated. Key words: FGD; ESP; Hg speciation; Partitioning.
INTRODUCTION Coal-fired utility boilers release around 50 tons of Hg annually or about one-third of the total anthropogenic emission [1]. On March 15, 2005, US EPA announced the Clean Air Mercury Rule (US EPA 2005) to permanently limit Hg emissions from coal-fired power plants. The first-phase cap is 38 tons annually beginning in 2010, with a final cap set at 15 tons starting in 2018, resulting in nearly 70% reductions from 1999 emission levels. In Europe, Hg emissions from power plants are not currently regulated but the Pollutant Release and Transfer Register (PRTR) of Industrial Emissions into air, water and land (Regulation (EC) No 166/2006) has established the threshold emissions for large combustion plants (10kg/year), and are used as Hg reference rate emission. Mercury in the flue gas usually exists in three forms [2]: oxidized (Hg2+), elemental (Hg0), and particle-bound (Hg). The ratio between them depends on the coal type and composition, and on combustion and flue gas conditions [3]. Owing to its high volatility, Hg0 usually occurs in vapour form, whereas Hg2+ can be removed in wet scrubbers [4] being retained in the FGD gypsum end-product and/or emitted in low proportions to the environment by the gas OUT-FGD. Therefore, it is significant for power plants to measure Hg speciation and their concentrations in the flue gas, hence that appropriate strategies can be applied. As result of sampling campaigns carried out in 2007 and 2008 at two Spanish power plants (PP1 and PP2) equipped with a forced oxidation wet flue gas desulphurisation (FGD) system with water re-circulation, a differential partitioning between output streams (gypsum and gypsum slurry water) and different speciation of Hg in the gas stream OUT-FGD was revealed. At PP1 and PP2, a high proportion (86-88%) of Hg escapes the ESP in a gaseous form, being Hg2+ (75-86%) the specie incoming FGD at the two power plants. By contrast, the speciation of the Hg species OUT-FGD is the opposite between PP1 and PP2. At PP2, Hg2+ is the prevalent Hg specie OUT-FGD in 2007 (66%) and 2008 (87%), whereas Hg0 (71%) is the prevalent Hg specie OUT-FGD at PP1, reaching high reduction levels of total Hg (71 and 65% at PP1 and PP2 2007) but very low (21%) at PP2 2008. The differential partitioning and gaseous speciation of Hg in the FGD system at PP1 and PP2 may attributable to: a) the high soluble salt concentrations of the FGD water streams at PP2, which reduce the gaseous solubility and probably, the gas retention efficiencies; b) the entraining of HgCl2 droplets by the
gas OUT-FGD; and c) the operational conditions at PP2 such as, limestone purity, use of additives, fluoride and/or sulphate complexes, and the S/F and S/Cl ratios in the scrubber. In order to corroborate the different Hg speciation in the gas OUT-FGD between PP1 and PP2 and determine the possible causes for the low retention efficiencies in PP2 2008 obtained from gaseous measurements and partitioning studies during the sampling campaigns, the aims of this work are: i) to study the speciation of Hg and other trace inorganic trace elements in FGD waters collected during the sampling campaigns, ii) to reproduce the desulphurisation process under the same operational conditions at PP1 and PP2 in lab scale; and iii) to determine the causes that may explain the different Hg speciation in the gas OUT-FGD and retention efficiencies by means of experimental process.
2. EXPERIMENTAL 2.1 Chemical analysis Water streams and trapping solutions from flue gas sampling carried out in the 2007 and 2008 sampling campaigns were directly analysed by Inductively Coupled Plasma Atomic Emission Spectrometry (ICP-AES) for major and minor elements and by Inductively-Coupled Plasma Mass Spectrometry (ICP-MS) for most trace elements. Chloride contents were measured by HPLC, and fluorides were determined by ion selective electrode.
2.2 Geochemical modelling The PHREEQC code (version 2.0) was used for calculation of the aqueous speciation of Hg and the remaining elements present in the FGD waters, and for the saturation index (SI) with respect to selected minerals and solid phases of the elements enriched in the FGD waters at PP1 and PP2. The thermodynamic database of PHREEQC was enlarged with data from other geochemical codes (MINTEQA2) for the speciation of As and U.
2.3 Design of the FGD device in lab scale A schematic of the FGD system and the internal components are presented in Figure 1. Sampling points, air, N2, HCl, SO2, and Hg 0 and Hg
2+
gas injection locations, as well
as the remaining devices are indicated in Figure 1. The PSA 10.534 series, CavKitCalc
software, was used to calculate the flow rate and temperature settings required to control of the main N2 carrier gas flow. The N2 temperature and flow rates were 40°C and 19.9 L/min, respectively. A total of 2 flowmeter, previously calibrated, were used as suppliers of the air (g) and the secondary carrier gases flow of N2 (g). A permeation device of HCl was used as source of Cl2 (g). The concentration of HCl (g) was controlled and calibrated as part of the study. A constant mass flow of HCl was provided by a diffusion device held at a constant lab temperature (25°C) and with a constant known flow of N2. The HCl device consisted in a very small diameter hole (0.5 mm) through the cap of a small (20 mL) sample vial. The concentrated HCl was placed (10 mL) in the bottom of an empty gas washing bottle. The flow of N2 designed for the constant mass flow of HCl was 0.54 L/min which allowed an HCl flow of 0.25-0.51 mg/min. Sulphur dioxide was metered into the mercury-laden gas as 100% SO2 through a calibrated flowmeter. The rig was designed to enable the ratio of Hg
0
and Hg
2+
concentrations in the gas
phase to be matched to those prevailing during the sampling campaign at the power plants under study. The Hg sources consisted in an Hg
0
and Hg
2+
permeation tubes
linked in series, each in its own temperature-controlled environment, in two water baths (Figure 1). Elemental and Hg2+ concentration were controlled by varying the temperature and the gas flow rate of the N2 carrier gas. Owing to the complexly of working with Hg2+ (g), the path of Hg2+ to the reactor vessel was heated up to 80°C with a thermo-tube controlled by temperature controller which allowed a higher release of Hg2+. The combined total flue gas was supplied by means of a porous sintered glass bubbler (forming a 154° angle with respect to the reactor vessel) into a closed, stirred vessel, which provided the FGD unit. The vessel was charged initially with a scaled volume of limestone slurry (2.5 L) and an incremental withdrawal of gypsum slurry from the vessel. The experiments were run up during 3h. The gas OUT-FGD from FGD unit, 25 L/min, passed through a tube system where the gas was split in two flow paths. A fraction of gas flow, around 500 mL/min, went to Hg speciation unit (KCl impingers), whereas the remaining fraction was vented as waste. In
the KCl impingers all Hg
2+
was trapped and reduced to Hg
0
by means of SnCl2
solution (2% SnCl2 in 10% KOH) being subsequently measured by Sir Galahad detector after experiments. The Hg
0
in the gas OUT-FGD was directly and continuously
measured by the Sir Galahad detector.
Figure 1. Diagram of the lab scale FGD. 2.4. Mercury speciation unit The Hg speciation unit was established as alternative for Hg2+ measurements owing to Sir Galahad device is specific for Hg0 measurements. The Hg speciation unit used in this experimentation consisted in a part of the Onthario hydromethod focused basically on the Hg2+ determination. The Ontario Hydro method was developed by Keith Curtis and other researchers at Ontario Hydro Technologies in late 1994. Since testing with EPA Method 29 appeared to show that some of the Hg0 was captured in the nitric acid– hydrogen peroxide (HNO3–H2O2) impingers, an attempt was made to more selectively capture the Hg2+ by substituting three aqueous 1N potassium chloride (KCl) impinge solutions for one of the HNO3–H2O2 solutions. In these experiments, the Onthario Hg speciation unit consisted in three KCl solution impingers. In each impinger, 100 mL of 1N KCl solution was placed in which after time
of sampling period all the Hg2+ in the gas IN and OUT-FGD was trapped in the sampling train. Once the sampling period was finished, the 300 mL of the KCl solution was joined in only one sample for the reduction of Hg2+ to Hg0 by means of a SnCl2 solution. At the end of this process the Hg2+ converted in Hg 0 was measured by means of the Sir Galahad.
2.3 Sir Galahad method The Sir Galahad analyzer can be used in a variety of gaseous media including combustion flue gas. The analyzer is based on the principle of atomic fluorescence (AF), which provides an inherently more sensitive signal than atomic absorption. The system uses a gold-impregnated silica support for pre-concentrating the Hg and separating it from potential interferences that degrade sensitivity. The system was calibrated using Hg 0 as the primary standard.
2.5 FGD experimental method Two different experiments were designed for the Hg speciation measurements at the two power plants: a) experiment with process water for limestone slurry preparation, and b) experiment using filtered water for limestone slurry preparation. At PP1, a fraction of filtered water is re-circulated to the scrubber after a purge process, which reduces the concentrations of some elements (especially Cl); while the remaining filtered water fraction is used for limestone slurry preparation together a fraction of process water. At PP2, all filtered water is directly re-circulated to the scrubber and limestone slurry is only prepared with process water. The a and b experiments reproduce the first operation cycle conditions of a FGD process and the FGD system conditions during the sampling campaigns, after a number of water re-circulations, respectively. We tested whether a) the experiment using filtered water corroborates the Hg speciation in the gas OUT-FGD at PP1 and PP2; and b) in the case of the prevalence of Hg2+ in the gas OUT-FGD, if it is due to the entraining and evaporation of HgCl2 droplets dissolved in the gypsum slurry by the gas OUT-FGD and/or due to the high salt content in FGD waters at PP2, which may reduce the activity of water for the gaseous retention.
The Hg IN-FGD was calibrated before starting the experiment to simulate the Hg0/Hg2+ ratio in the gas IN-FGD at the two power plants, and to check the accuracy of the Hg analytical and speciation measurements.
3. RESULTS AND DISCUSSION 3.1 Characterisation of FGD waters The limestone and gypsum slurry waters at PP1 had very high concentrations of S, Mg, Ca, and Cl, followed by B, Li, Se, Ba, Mo, and especially Ni, Zn, Co and U in the gypsum slurry water. In the case of the limestone slurry water, the high contents of the above elements are attributed to the re-circulation of a fraction of filtered water for limestone slurry preparation. Furthermore, U may form soluble uranyl complexes with carbonates that may promote presence of U bearing complexes in the gypsum slurry water [5]. At PP2, the limestone and gypsum slurry waters showed high concentrations of elements associated with soluble salts (Ca, K, Mg, Na, S, Sr, and Cl) and major and minor elements (B, Mn, Cu, Se, Cu Ba, and U), which give rise a high ionic strength in these water streams, specially the free Cl-. The high concentration of these elements in the limestone and gypsum slurry water may be ascribed to the dissolution of limestone components and to the complete and continuous re-circulation of the filtered water to the scrubber, respectively. The addition of Al to the scrubber at the PP2 modifies the partitioning of these elements, promoting the presence of soluble Al-fluoride complexes and reducing the fraction retained in gypsum. In addition, the highly acid insoluble Alfluoride ralstonite (NaMgAlF6.H2O) [6] is formed instead of fluorite (CaF2) and MgF2, probably because ralstonite forms an iso-structural solid solution between hydrated Alfluoride complexes rich in Na and Mg compounds [7].
3.2 Speciation of FGD waters The geochemical calculations performed with PHREEQC reveal that at 7.6 and 4.6 pH and 60ºC, conditions of the gypsum slurry sampled at PP1 and PP2, respectively, MgSO4 and SO42- and Cl- , as free aqueous ions, are the main species present in gypsum slurry waters at both power plants. Regards the similarities, Hg0 is the main specie of Hg in gypsum slurry water at both power plants. In spite of Hg0 is the main specie of Hg, it should be significant to pointed out the occurrence of HgCl2 in the gypsum slurry
waters at both power plants. Boric acid is the main specie of B in gypsum slurry also at both power plants. With respect to the differences, Al is present in hydroxide form (Al (OH)-4) in gypsum slurry water at PP1, being the highly soluble AlF3 complexes the main specie of Al in gypsum slurry at PP2. The F speciation is related with the aforementioned Al speciation in gypsum slurry which is controlled by the use of Aladditives. In this regard, MgF2 is the main specie of F in gypsum slurry at PP1 and also at PP2 in 2008, being AlF3 the main specie of F in gypsum slurry in 2007. Regards Se, SeO32- is the main specie in gypsum slurry at PP1, whereas at PP2, HSeO3- is the main specie of Se in gypsum slurry. Differences were also found for U speciation. At PP1, U is present in hydroxide form, whereas at PP2, U is present as sulphate form in gypsum slurry, probably to the high sulphate contents in gypsum slurry water due to the Alsulphate addition. Regards the predicted solid phases in gypsum sludge, gypsum is in equilibrium and slightly saturated in gypsum slurry water at both power plants. Fluorite (CaF2) is saturated in gypsum sludge at PP1 and also at PP2 in 2008, showing the opposite trend in 2007 where is in equilibrium together ralstonite. Related to this, MgF2 is in equilibrium in gypsum sludge at PP1 and slightly saturated in 2008 at PP2, also due to the Al-additive used at PP2. As stated previously, high Al dosages enhance the formation of water soluble AlF3 complexes leading to F-rich FGD waters, and also modifying the speciation of other elements. Thus, the injection of Al-additive promotes the AlF3 formation in gypsum slurry water and simultaneously the formation and precipitation of highly insoluble ralstonite at high (when high levels of dissolved Mg and Na are present in FGD waters); At lower dosages of Al-additives (PP1), HF reacts with Ca and Mg giving rise to the CaF2 and MgF2 (slightly and moderately water soluble species, respectively) formation reaching saturation easily, and consequently precipitate in FGD gypsum. This explains the higher SI of fluorite in the gypsum slurry in 2008 (1.31) than in 2007 (-0.04). As regards Hg, Hg2Cl2 is far below saturation in gypsum slurry at PP1, whereas at PP2, is saturated in gypsum slurry in both sampling campaigns. Mercury chlorine is far below saturation (sub-saturated) in gypsum slurry at both power plants in line with the predicted occurrence in gypsum slurry waters.
3.3 Results of Hg gaseous speciation OUT-FGD at lab scale The results obtained in experiments using filtered water corroborate the speciation of Hg in the gas OUT-FGD obtained in sampling campaigns at the two power plants (Table 1). At PP1, Hg0 (66-60%) is the prevalent specie in the gas OUT-FGD, whereas at PP2, Hg2+ (57-66%) is the predominant specie of Hg during the overall process of the reaction, reaching the highest proportion in gas OUT-FGD after 3h experiment. The gaseous Hg speciation (Hg0/Hg2+ ratio) OUT-FGD obtained in this lab scale experiments are in line with the speciation results obtained in the sampling campaigns at PP1 and PP2. Table 1. Hg gasesous speciation in lab scale experiments PP1 sampling Gas IN-FGD Gas OUT-FGD PP1 Filtered water experiment Gas IN-FGD Gas OUT-FGD (1-2h) Gas OUT-FGD (3-4h) PP2 2007 sampling Gas IN-FGD Gas OUT-FGD PP2 2008 sampling Gas IN-FGD Gas OUT-FGD PP2 Filtered water experiment Gas IN-FGD Gas OUT-FGD (2h reaction) Gas OUT-FGD (3h reaction) PP2 Process water experiment Gas IN-FGD Gas OUT-FGD (2h reaction) Gas OUT-FGD (3h reaction)
% Hg2+
% Hg0
75 29
25 71
78 34 37
22 66 60
86 66
14 34
88 87
12 13
89 57 66
11 43 34
82 37 64
18 63 36
In the experiment using process water (PP2), Hg0 (63%) is the prevalent specie of Hg in the gas OUT-FGD after 2h reaction, which is most probably due to similar conditions (low Cl content in slurry water) of most wet FGD systems (such as PP1). Probably the low Cl in slurry waters and the high SO2 levels leads to the trapping of Hg2+ as Hgsulphate species in gypsum. Alternatively, reduction processes of Hg2+ to Hg0 may also occur, being subsequently emitted by the gas OUT-FGD.
After 3h reaction the gaseous Hg speciation is the opposite being Hg2+ (64%) the prevalent specie of Hg in the gas OUT-FGD. Diverge of gaseous Hg speciation, from Hg0 to Hg2+, in the gas OUT-FGD suggest that HgCl2 is increasingly dissolved in the gypsum slurry water enhanced by the simultaneously dissolution HCl, being subsequently entrained by the gas OUT-FGD and /or not dissolved after 3h reaction due to the increase of the ionic strength promoted also by the high occurrence of free Clions in gypsum slurry water. The possibility that the unusual speciation of Hg OU-FGD was due to oxidation processes of Hg0 to Hg2+ should be omitted given the constant Hg0 flow (20.2-20.3 ng/min after 2 and 3h reaction).
3.3. Relationship between the gypsum slurry water composition and retention and speciation of gaseous Hg Since a high content of salts reduce the water solubility of gaseous species, a relationship between the ionic strength and retention of gaseous Hg was research. As previously stated, the differential gaseous speciation of Hg in the FGD system at PP1 and PP2 may attributable to: a) the high soluble salt concentrations of the FGD water streams at PP2, which reduce the gaseous solubility and probably, the gas retention efficiencies; b) the entraining of HgCl2 droplets by the gas OUT-FGD. a) Salt concentration of FGD Waters (ionic strength) The FGD waters at PP2 are characterised by a high concentration of soluble salts giving rise to a high ionic strength (especially Cl-) which may reduce the activity of the water for the gaseous retention. High ionic strength avoids the interaction between the solidwater-gaseous phases in the scrubber. Thus at PP2, the Hg2+ may enter and leave FGD system without interaction and therefore being emitted OUT-FGD as Hg2+.This could explain the differential Hg behaviour between PP1 and PP2 given the higher ionic strength (by a factor of 2-2.6) of gypsum slurry water collected during the sampling campaigns at PP2 than at PP1 (Table 2 and 3). b) Entraining of HgCl2 droplets by the gas OUT-FGD. At PP1, the geochemical modelling of the gypsum slurry water reveals that Hg0 is the primary specie of Hg in this water stream such as in the gaseous stream OUT-FGD
obtained from the gaseous speciation during the 2007 sampling campaign. As previously stated this was subsequently corroborated by the experiments at lab scale. Given the high removing efficiency for Hg2+ and the low abatement of Hg0 by wet scrubbers [8]; it could be stated that the Hg0 emission OUT-FGD in this plant is mainly due to a high gaseous retention of Hg2+ as HgCl2 and HgSO4 species adsorbed and retained in gypsum, respectively as reported by Rallo et al [9]. A partial reduction of Hg2+ to Hg0, may also account for increasing the Hg0 emission. At PP2, the geochemical modelling of the gypsum slurry water reveals that Hg0 is also the primary specie of Hg in this water stream, being the HgCl2 the second specie of Hg in the gypsum slurry water in the two sampling campaigns. By contrast Hg2Cl2 is slightly saturated in the gypsum sludge at PP2 in the two sampling campaigns (Table 3). In this regard, HgCl2 can react with Hg0 giving rise to Hg2Cl2. The formation of Hg2Cl2, in conjunction with the re-circulation of filtered water and given the low solubility of this specie (12 mg/L) [10] may support the Hg2Cl2 saturation (Table 9) in the gypsum sludge in the two PP2 sampling campaigns. However, the inverse process may also occur. It is demonstrated [10] that Hg2Cl2 decomposition at the scrubber temperatures (60-70oC) releases Hg0 and HgCl2, which may promote the dissolution of the highly soluble HgCl2 (71.5 mg/L) [10] in the gypsum slurry water and the subsequent entraining by the gas OUT-FGD. The aforementioned thermal decomposition of Hg2Cl2, in conjunction with the slight decrease of the SI of Hg2Cl2 from 2007 to 2008 sampling (Table 3) could entail a higher release and subsequent dissolution of HgCl2 gypsum slurry water in 2008 sampling with respect to 2077. High occurrence of HgCl2 in water may results in higher proportion of Hg2+ entrained by the gas OUT-FGD explaining the lower Hg retention in 2008 than in 2007. Table 2. Speciation of aqueous complexes and solid phases in gypsum slurry from PP1. Ionic strength 0.377 Activity SI Hg0
Hg(0)
5.044e-008
Hg2Cl2
-5.36
Hg2+
HgCl(OH)
2.891e-011
HgCl2
-10.28
Hg(OH)2
1.465e-011
Hg2SO4
-14.29
HgCl2
8.586e-012
HgSO4
-19.16
Table 3. Speciation of aqueous complexes and solid phases in gypsum slurry from PP2. Ionic strength 2007 0.894 2008 0.611 Activity 2007
Hg
Hg
0
2+
Activity 2008
SI 2007
SI 2008
Hg0
9.648e-006
1.015e-005
Hg2Cl2
0.59
0.51
HgCl2
4.017e-008
3.163e-008
HgCl2
-6.61
-6.72
HgCl3-
3.190e-008
2.174e-008 Hg2SO4
-8.79
-8.69
HgCl42-
8.124e-009
4.791e-009
-15.94
-15.86
HgSO4
Given that Hg2+ has been corroborated as the prevalent gaseous Hg specie OUT-FGD, and that water speciation demonstrates the occurrence of HgCl2 complexes and Hg2Cl2 as species in the gypsum slurry water and sludge, respectively the entraining of HgCl2 droplets dissolved the gypsum slurry by the gas OUT-FGD in conjunction with the high ionic strength of the FGD waters are the major causes for the Hg2+ emission by the gas OUT-FGD. The significance of each cause is under study
4. CONCLUSIONS The predominant species for the elements of environmental concern in FGD waters from PP1 and PP2: U, Se, F, Hg, Al, are present in oxide and hydroxide form at PP1 whereas at PP2, the species are associated to sulphate, chloride, and fluoride ion forms. The Hg speciation in the gas OUT-FGD as result of gaseous speciation measurements carried out in 2007 and 2008 at PP1 and PP2 is corroborated by experimental test at lab scale. At PP1, Hg0 (66%) is the prevalent specie in the gas OUT-FGD; whereas at PP2, Hg2+ (57-66%) is the predominant specie OUT-FGD. The experiment using process water at PP2 demonstrates that speciation of gaseous Hg OUT-FGD is being modified with the increase of the reaction time. Initially similar Hg speciation than commonly found in power plants, such as PP1, was obtained (prevailing Hg0 over Hg2+), while the opposite speciation was obtained at the end of experiment (3h) supporting the speciation of gaseous Hg in the gas OUT-FGD. A relationship between enrichment of FGD water in salts and gaseous Hg retention may explain the differential speciation of Hg OUT-FGD between PP1 and PP2 given the
higher ionic strength of gypsum slurry from PP2 than from PP2; which reduces the activity of water for dissolution of gaseous Hg. The comparison of the experiment using filtered and process water reveals that the entraining of HgCl2 by the gas OUT-FGD in conjunction with the high ionic strength of the FGD waters are the major causes for the Hg2+ emission by the gas OUT-FGD.
5. REFERENCES [1] U.S. Environmental Protection Agency (EPA). www.epa.gov/mercury/. [2] Galbreath KC, Zygarlicke CJ. Mercury transformation in coal combustion flue gas. Fuel Process Technol 2000; 65–66:289–310. [3] Stergaršek A, Milena H, Jože K, Janja T, Peter F, David K, Radojko J, Vesna F, Maja P, Iztok H, Jože L, Branko D, Majda Č.The role of flue gas desulphurisation in mercury speciation and distribution in a lignite burning power plant. Fuel 2008; 87: 3504-3512. [4] Lindqvist K, Johansson M, Aastrup A, Anderson L, Bringmark G, Hosvenius L, Hakanson A, Iverfeldt M. Meili B. Tim, Mercury in the Swedish environment — recent research on causes, consequences and corrective methods, Water Air Soil Pollut. 55 (1991) 1–261. [5] Barisic D, Lulic S, Miletic P.Radium and Uranium in phosphate fertilizers and their impact on the radioactivity of waters. Wat Res 1991; 26:607-611. [6] Font O, Querol X, Moreno T, Ballesteros J.C, Giménez A. Effect of aluminium sulphate addition on reducing the leachable potential of fluorine from FGD gypsum. In proceedings: 2nd International Conference on Engineering for Waste Valorisation. Patras. Greece. 2008. Paper 150. ISBN: 978-960-589 530-101-9. [7] Pauly H. Ralstonite from Ivigtut, South Greenland Amer Miner 1965; 50: 18511864. [8] Senior C.L. 2001. Behaviour of mercury in air pollution control devices on coalfired utility boilers, Proceedings of 21st Century: Impacts of Fuel Quality and Operations, Engineering Foundation Conference, 2001. [9] Rallo M, López-Anton M.A, Perry R, Maroto-Valer M.M. Mercury speciation in gypsums produced from flue gas desulfurization by temperature programmed decomposition, Fuel 2010; 89:2157-2159.
[10] Handbook of Chemistry and Physics. 91 st ed. 2010-2011. Electronic version.
Oviedo ICCS&T 2011. Extended Abstract
Aqueous chemistry of mercury in flue gas desulphurization conditions R. Ochoa-González, M. Díaz-Somoano and M. R. Martínez-Tarazona Instituto Nacional del Carbón (INCAR), CSIC, Apartado 73, 33080-Oviedo –Spain(
[email protected]) Abstract Oxidised mercury is captured in wet flue gas desulphurization systems (FGD) with a high degree of efficiency due to its solubility in water. However, once it dissolves, it can react with slurry constituents and be re-emitted to the atmosphere. This fact is mainly due to the sulfite ions formed during the absorption of SO2. In this paper, the absorption and reduction of mercury in scrubbing solutions containing sulfite ions and different concentrations of anions has been evaluated. According to our experimental results, fluorine, chlorine and bromine ions contribute to the stabilization of mercury in these systems. In contrast, nitrate and sulphate ions do not affect mercury behavior.
1. Introduction According to the Environmental Protection Agency (EPA), coal combustion is considered as the largest source of anthropogenic mercury emissions. Numerous studies have shown that some degree of mercury removal can be achieved using existing conventional air pollution control devices (APCD) commonly used for the control of NOx, SO2, and particulate matter pollutants from coal-fired flue gas [1]. For example electrostatic precipitators (ESP), fabric filters (FF) and/or Wet FGD already installed in power plants can be used to control mercury emissions at a low cost. When combustion occurs, the mercury present in coal turns into elemental mercury (Hg0). During the cooling of the gases Hg0 may be partially oxidized through different mechanisms, oxidized mercury (Hg2+) being the predominant species in the flue gas. Although the primary function of the FGD is to remove SO2, several field tests show that wet FGD can also remove oxidized mercury from the flue gas stream but it is ineffective in removing elemental mercury [2-3]. In addition, there is evidence that oxidized mercury can be reduced to elemental mercury in a wet FGD system (re-
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emission), primarily by S (IV) (sulfite and/or bisulfite) species in the scrubbing solution [4-7]. The challenge for mercury control using wet FGD systems is to investigate in depth the aqueous chemistry of mercury in order to find a way to favour the presence of oxidized mercury in the scrubber solution and to prevent the reduction of oxidized mercury in the wet FGD process. In this work, a laboratory-scale study was carried out to evaluate the influence of different anions on mercury re-emission in wet FGD systems. The objective of this work was to optimize the conditions in desulphurization plants in order to increase mercury capture.
2. Experimental section In order to investigate oxidized mercury absorption and its reduction by sulfite ions, a lab-scale FGD device was built (Figure 1). This system is composed of three parts: i) a mercury generation system, ii) a reactor containing the scrubbing solution and iii) a continuous mercury emission monitor for elemental mercury.
Activated carbon
Hg generator
Hg analyser VM 3000 HOVACAL
Gas outlet Gas inlet
KCl 0.1M Empty
Reactor Peristaltic bump
pH-meter
0.000
Water bath Balance
N2
O2
Figure 1. Scheme of the system and FGD reactor. A commercial device (HovaCAL) was used for generating mercury species in the flue gas. An aqueous mercury solution was evaporated continuously at a temperature of 200ºC. A flow of nitrogen conducted the gaseous mercury to a glass reactor. The reaction vessel was stirred and kept at constant temperature (40ºC). The scrubbing
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Oviedo ICCS&T 2011. Extended Abstract
solution contained limestone and the different ions. The pH and redox potential of this solution were measured continuously. A flue gas containing N2 or N2 + O2 was used to conduct the outgoing mercury to the measuring system. Any Hg2+ not dissolved in the scrubbing solution was captured by two impingers containing a potassium chloride solution, while Hg0 was measured continuously by the mercury analyzer (VM 3000). For each test, 100 mL of an aqueous solution containing 1 mM of sodium sulfite was placed in the reactor. 3 L/min of N2 containing 130 µg/m3 of Hg was passed through the liquid. HSC Chemistry 6.1 software was used to explain the experimental results and to predict the reaction mechanisms.
3. Results and discussion The influence of several anions on mercury capture in the scrubber liquor was investigated. The gaseous mercury species were passed through an aqueous solution containing chloride, fluoride, bromide, nitrate, sulphate and carbonate ions for 200 min. The concentration of Hg0 coming out of the reactor was measured continuously. The results (Figure 2) show that sulfite ions contribute to the re-emission of elemental mercury after 120 minutes. However, oxidized mercury was captured when the scrubber solution contained chloride (Figure 2a). When no HCl was present in the scrubbing liquor no elemental mercury emission was observed for 120 min. When the HCl concentration was increased to 0.1 and 0.5 mM the reduction started earlier, at 100 and 50 min. respectively. When HCl reached 2 mM a different behaviour was observed with a limited range in which mercury reduction occurs. For higher chloride concentrations no mercury reduction was observed for the 200 min of test duration. These results can be attributed to the formation of different mercury chloride ions (HgCl-; HgCl2, HgCl3-; HgCl42-) [7]. A second set of tests was carried out to compare the effect of the chloride but now resulting from the dissolution of KCl. The results (Figure 2b) showed a similar trend though the kinetics of the process was slightly slower. This may be a result of the different stabilities of the species formed due to the presence of other sulphur species such as bisulfite ions in solution. Similar results (Figure 2c-d) were obtained when the tests were carried out with different concentrations of fluoride or bromide ions. The reduction of mercury was delayed by the presence of halogen ions until no mercury reduction was detected at all
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during the experiment. The results show that bromide is the most efficient anion for keeping oxidised mercury in solution. 50
0 mM HCl
40
0.1 mM HCl 30
0.5 mM HCl 2 mM HCl
20
5 mM HCl 10
0 mM KCl 0.5 mM KCl
40
[Hg0] (μg/m3)
[Hg0] (μg/m3)
50
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5 mM KCl 20
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20 mM KCl 50 mM KCl
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t (min)
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1 mM KF 2 mM KF
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0 0
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20
2 mM KBr 10 mM KBr
10 0
t (min)
(c)
0
50
100
150
200
t (min)
(d)
Figure 2. Re-emission of mercury by using different scrubbing solutions in typical wet scrubber conditions. Influence of chloride (a,b), fluoride (c) and bromide ions (d). Below, several reactions with their corresponding thermodynamical constants are proposed. Whereas HgF2 decomposes in water to form HgO [1] (Table 1), chloride and bromide ions form stable complexes with Hg2+ in accordance with the reactions [2]-[5] shown in Table 1. Moreover, as mercury is able to form relative stable compounds with sulphite ions, chloride complexes may form in the solution other species [6]–[8] (Table 1). In addition, several tests were carried out to evaluate the influence of oxy-anions on mercury reduction. After the concentrations of nitrate and sulphate had increased a similar behaviour to that described for halogen ions was observed (Figure 3a-b). The probable reactions [9]-[10] (Table 1) and thermodynamic equilibrium constants favoured the reduction of HgCl2 as the dominant process instead of the formation of mercury nitrate:
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Oviedo ICCS&T 2011. Extended Abstract
Table 1. Possible reactions predicted by means of thermodynamical calculations (HSCChemistry 6.1) Reaction HgF2 (ac) + H2O → HgO (ac) + 2HF (ac) Hg2+(ac) + Cl-(ac) ↔ HgCl+(ac) HgCl+(ac) + Cl-(ac) ↔ HgCl2(ac) Hg2+(ac) + Br-(ac) ↔ HgBr+(ac) HgBr+(ac) + Br-(ac) ↔ HgBr2(ac) HgSO3(ac) + Cl-(ac) ↔ ClHgSO3-(ac) HgCl2(ac) + SO32-(ac) ↔ ClHgSO3-(ac) + Cl-(ac) ClHgSO3-(ac) + Cl-(ac) ↔ Cl2HgSO32-(ac) Hg2+ (ac) + NO3- (ac) ↔ Hg(NO3)2 (ac) HgCl2(ac)+SO32-(ac)+H2O↔Hg(g)+SO42-(ac)+2Cl-(ac)+2H+(ac)
K40ºC — 1.45 107 3.24 106 5.24 108 1.51 108 — — — 1.83 2.54 1012
[1] [2] [3] [4] [5] [6] [7] [8] [9] [10]
Finally a batch of tests was carried out in the presence of carbonate ion, the results of which are shown in Figure 3c. Similar to the effect of halogen ions, mercury reduction was slower until carbonate concentrations higher than 10 mg/L were reached. The effect of the carbonate was to increase the pH, which, in turn, made the sulfite ions more stable. This suggests that stable complexes could be formed between carbonate, sulfite and Hg2+ ions, such as HgSO3 or Hg(SO3)22-. 50
50
0 0mM mMNaNO3 NaNO3 [Hg] (μg/m3)
[Hg0] (μg/m3)
40
NaNO3 5 5mM mM NaNO3
30
1010mM mMNaNO3 NaNO3
20
2020mM mMNaNO3 NaNO3
10
5050mM mMNaNO3 NaNO3
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0 mM Na2SO4 0 mM Na2SO4
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2.5 mM (sulfato de 4sod 2.5 mM Na2SO 5 mM sulfatos (sulfato 5 mM Na2SO 4 sodio) 1010 mM sulfatos mM Na2SO(sulfat 4 sodio)
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0 mg Na 2 CO3 /L 3 mg Na 2 CO3 / L 5 mg Na 2 CO3 /L 10 mg Na 2 CO3 /L 50 mg Na 2 CO3 /L 500 mg Na 2 CO3 /L
30 20 10 0 0
100
200
300
400
t (min)
(c) Figure 3. Emission of Hg using a scrubbing solution contaning nitrate (a) sulphate (b) and carbonate ions (c). 0
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Oviedo ICCS&T 2011. Extended Abstract
4. Conclusions The re-emission of mercury with fluoride, chloride, bromide, nitrate, carbonate, sulfite and sulphate ions in the scrubbing solution was studied. Fluoride, chloride and bromide ions contribute to mercury stabilization, bromide being the most efficient anion for retaining mercury in aqueous solutions. High carbonate concentrations prevent the reduction of mercury by sulfite ions, probably due to the formation of HgCO3 or HgSO32- complexes. However, nitrate and sulfate ions do not influence the equilibrium between mercury and sulfite ions and, as a consequence, the re-emission of mercury is not promoted.
References [1] [2] [3] [4] [5] [6] [7]
Pavlish J.H., Hamre L.L., Zhuang Y., 2010, Fuel 89, 838-847. Brown T.D., Smith D.N., Hargis R.A., Jr., O’Dowd W.J., 1999, J. Air Waste Manage. Asocc. 19, 1-97. Van Loon L., Mader E., Scott S.L., 2000, J Phys Chem A 104, 1621-1626. Chang J.C.S., Ghorishi S.B., 2003, Environ. Sci. Technol. 37, 5763-5766., 2003 Chang J.C.S., Zhao Y., 2008, energy ¬ Fuels 22, 338-342. Diaz-Somoano M., Untenberger S., Hein K.R.G., 2007, Fuel Processing Technology 88, 259-263. Senior C., 2007, Environmental Engineering Science 24 (8), 1129-1134.
Acknowledgements Part of this work was supported by the project ABETRAP (Abatement of Emission of Trace Pollutants by FGD from co-combustion and environmental characteristics of byproducts) RFCR-CT-2006-00006. We also thank FICYT for a predoctoral grant.
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Oviedo ICCS&T 2011. Extended Abstract
Program topic – coal combustion
The Effect of Particle Size and Petrographic Composition on Combustion Behaviour of selected Russian Coals Hudspith. V3, Nuamah. A1,2, Scott. A.C3, Drage. T1, Powis. J2, Riley. G2, Collinson. M.E3, Lester. E1 1
Faculty of Engineering, The University of Nottingham, Nottingham, NG7 2RD, U.K.
2
Fuels and Combustion, RWEnpower, Windmill Hill Business Park, Whitehill Way,
Swindon, Wiltshire, SN5 6PB 3
Department of Earth Sciences, Royal Holloway University of London, Egham, Surrey,
TW20 0EX, U.K.
Abstract Two size fractions (53-75µm and 106-125µm) of pulverised fuel from five high volatile bituminous coal samples were selected to determine the effect of particle size and petrographic composition on combustion behaviour. Three Late Permian coal seams (numbered 78 (S1 and S2), 88, 91) and one stockpile (incorporating 13 seams) were sampled from two open cast mines in the Kuzbass region of Russia. In addition, sample 78 S1 represents a heat affected coal from a modern coal seam fire. The samples showed petrographic variation both between seams and size fractions. The burnout of the largest size fraction was measured using a drop tube furnace operating at 1300oC using a range of oxygen contents and residence times. The heat affected coal showed different chemical, petrographic and poorer burnout behaviour than the non heat affected coal samples.
1. Introduction The importance of coal in the current and future global energy mix cannot be overestimated. Even though it poses a major threat to the environment if not carefully and efficiently combusted to reduce the emission of CO2 [1], it is one of the major fuels needed to meet the energy needs of the ever increasing global population and economies [2]. The initial combustion of coal results in the generation of chars after the release of
1
Oviedo ICCS&T 2011. Extended Abstract
volatiles and moisture [3]. The combustion of this char depends on, the rank [4], maceral/ microlithotype associations [5], particle size, shape, structure and chemical composition [6], as well as the reflectance of vitrinite and inertinite [4,5]. It is important to understand the extent these factors influence char combustion, which is vital in determining the carbon loss in ash and ultimately boiler design [7]. Carbon loss in ash is an important parameter because it indicates combustion efficiency and also affects the sale and use of the pulverised fuel ash [8]. The aim of this study is to carry out a series of combustion tests on a Drop-Tube Furnace (DTF) and analyse the samples using a thermogravimetric analyser (TGA), image analysis procedures, and burnout to investigate the effect of particle size and petrographic composition on selected high volatile Russian coals.
2 Experimental section 2.1 Coal preparation The five samples were milled then sieved to 53-75µm and 106-125µm fractions and run through an airjet sieve to remove any fines. The 53-75µm fraction represents mid range commercial pulverised fuel for combustion whereas the 106-125µm fraction is more representative of the larger particles that might not have sufficient time to burnout completely. The larger size fraction may also help to highlight the impact of petrographic variations between the samples.
2.2 Drop Tube Furnance Drop-Tube Furnace refire studies were undertaken at the Fuel Characterisation Laboratory Facility (CTF) at RWE Npower in Didcot. The DTF is made up of screw feeder, feeder probe and collector probe, and a controllable gas system. Sixteen grams of each size fraction was injected into the feeder probe at a flow rate of 15g/hour. The larger samples were pyrolysed in a drop-tube furnace at 1300ºC with 1 vol% oxygen at 200ms and then refired at 5 vol% oxygen and with residence times of 200 and 500ms. Burnout (or using the ash tracer method) was determined by ashing the coal, char and refired char samples at 815°C for two hours in a muffle furnace. Burnout is calculated using equation 1 below. Volatile Yield V V or Burnout (% daf) = Whereby:
A1 =
A 2 - A1 4 10 A 2 (100 - A1)
(1)
% dry ash content of coal
2
Oviedo ICCS&T 2011. Extended Abstract A2 =
% dry ash content of char.
2.2 Maceral Analysis Pulverised fuel petrographic analysis was undertaken on polished blocks, under oil using a Leica DM 2500P reflectance microscope and a ×20 objective. Petrographic and microlithotype counts were taken using a Kötter graticule using the methodology outlined in [9,10]. Intersections on embedding resin were not counted. Maceral groups were identified using ICCP schemes [11,12].
2.3 Proximate and ultimate Analysis Proximate and ultimate characterisation analyses were undertaken on bulk pulverised fuel samples at the RWE Npower Fuel Characterisation Laboratory Facility in Swindon using British standards [13-18]. Results are shown in Table 1.
3 Results and Discussion 3.1 Characterisation properties of coal Sample
Ash
Moisture
VM
Name
(%)
(%)
(%)
78 S1
21.87
16.5
42.18
78 S2
4.56
4.06
88
5.34
91 Stockpile
FC (%)
Net CV
C
H
N
S
(MJ/Kg)
(%)
(%)
(%)
(%)
35.64
13.3
68.48
2.32
2.83
0.19
35.16
59.25
29.44
81.73
4.96
2.62
0.29
6.81
38.92
53.66
27.32
79.16
5.10
2.41
0.21
13.70
5.68
43.31
45.71
24.23
76.46
5.50
2.47
0.30
7.07
4.41
38.72
54.24
28.35
81.32
5.25
2.75
0.43
TABLE 1 Proximate data for bulk unsieved pulverised coal samples. Moisture and ash are reported to air dried basis. volatile matter, and ultimate analyses (C, H, N, S) are all reported to dry ash free basis.
3.2 Maceral content and microlithotype associations Petrographic composition varies between seams (Table 2), with vitrinite totals from 65% (78 S2) to 76.7% (88). Liptinite, 2.5% (stockpile) to 7.11% (88). Inertinite from 14% (88) to 30.24% (78 S2). Mineral matter, 0.34% (78 S2) to 3.48% (stockpile).
Petrographic and microlithotype variation is also seen between size fractions. Inertinite fractures differently to the other macerals and is more brittle than other maceral groups [20, 21, 22]. Liptinite tends to be more prevalent in the larger size fraction, since it is
3
Oviedo ICCS&T 2011. Extended Abstract
more difficult to grind than the other macerals [21]. The smaller particle size shows more liberation of the individual macerals i.e. the monomaceral vitrite is more abundant in the 53-75µm fraction in all samples (Table 3). Minerite is also more abundant in the 53-75µm fraction (78 S2, 88, 91). Conversely, liptinite is more abundant in all 106125µm fractions (Table 2). Liptite was not observed and clarite is low in total abundance up to 7.9% (91). Bi- and tri- macerals are also more resistant to grinding than monomacerals [7]. The majority of liptinite in the 106-125µm fraction must therefore come from trimacerite microlithotype associations (up to 45.98% microlithotype totals (78 S2); Table 3). Inertinite is more abundant in the 53-75µm fraction 88 and stockpile samples. Of the inertinite maceral group, fusinite and macrinite are the only macerals more abundant in the 53-75µm fraction in four samples. Semifusinite is more abundant in the 53-75µm size fraction in two of the samples (88, 91). The majority of inertinite is present in mixed maceral assemblages as vitrinertite (11% – 29.5%) or trimacerite (7.5% - 46%) only 1.2% - 15% of the total is the monomaceral inertite. This difference in microlithotype assemblages between the two size fractions may affect the reactivity of the pulverised fuel [7].
The heat affected coal (78 S1) is chemically (Table 1) and petrographically different from the other four samples (Table 2; Table 3). The vitrinite and inertinite macerals have qualitatively high reflectance (comparable to coal of anthracite rank) and there is no liptinite present. The clast shape is rounded with regular cracks around the perimeter, suggesting volatile loss. Vitrinite is the most abundant maceral (up to 35%), the majority of which is as the microlithotype vitrite (16% – 23.8%; Table 3) and vitrinertite (8.8% - 12.5%). Inertinite is a minor component (up to 4%) either as inertite (2.1%) or vitrinertite.
4
OF FUNGINITE
MACRINITE
TOTAL INERTINITE
MINERAL MATTER
TOTAL POINTS
1.37
0
0
2.64
4.01
2.95
948
1.18
0
0
0.67
2.02
2.60
1,191
11.35
0
0.68
7.26
23.04
1.59
881
9.44
5.14
0.93
4.3
30.24
0.34
1,187
5.60
0
0.44
3.24
19.88
1.62
679
8.05
0
0.1
2.72
14.32
1.88
957
4.55
0.29
0
5.58
21.29
3.08
681
3.73
0
0
1.08
23.55
3.34
1,019
3.46
1.33
0
3.11
19.01
3.02
1,126
NUMBER
SECRETINITE
78 S1 22.89 0 0 0 (53-75µm) 78 S1 34.93 0 0.17 0 (106-125µm) 78 S2 71.62 3.75 3.41 0.34 (53-75µm) 78 S2 65.37 4.04 2.11 8.34 (106-125µm) 88 74.82 3.68 4.42 6.19 (53-75µm) 88 76.7 7.11 0.42 3.03 (106-125µm) 91 72.69 2.94 8.96 1.91 (53-75µm) 91 67.52 5.59 8.54 10.21 (106-125µm) Stockpile 75.49 2.49 5.51 5.6 (53-75µm) Stockpile 75.84 4.69 1.2 5.09 (106-125µm) TABLE 2 Summary of petrographic point count
INERTODETRINITE
SEMIFUSINITE
INERTINITE (%)
FUSINITE
TOTAL LIPTINITE (%)
TOTAL VITRINITE (%)
SAMPLE NAME
Oviedo ICCS&T 2011. Extended Abstract
7.83 0 0 1.87 16 3.48 1,494 data of pulverised coal 53-75µm and 106-125µm size
fractions. Data are reported to a mineral included basis.
5
MINERITE (%)
DURITE (%)
CLARITE (%)
INERTITE (%)
3.3 6.4 4.8 0.0 6.8 5.3 7.3 5.6 6.1 9.5
LIPTIITE (%)
16.1 0.0 2.1 0.0 0.0 8.8 0.0 78 S1 (53-75µm) 23.8 0.0 0.0 0.0 0.0 12.5 0.0 78 S1 (106-125µm) 27.7 0.0 5.0 2.3 0.0 29.5 30.7 78 S2 (53-75µm) 17.3 0.0 11.8 0.9 1.0 22.9 46.0 78 S2 (106-125µm) 52.1 0.0 11.8 4.9 0.6 14.9 9.0 88 (53-75µm) 34.7 0.0 2.1 4.2 1.7 18.8 33.2 88 (106-125µm) 51.0 0.0 15.4 7.9 0.0 11.0 7.5 91 (53-75µm) 45.8 0.0 11.5 3.2 1.3 12.6 20.1 91 (106-125µm) 50.8 0.0 8.0 3.5 0.9 19.8 10.9 Stockpile (53-75µm) 30.2 0.0 1.2 7.6 1.4 21.1 29.2 Stockpile (106-125µm) TABLE 3 Summary of microlithotype associations within the five samples studied
VITRITE (%)
TRIMACERITE (%)
VITRINERTITE (%)
SAMPLE NAME
Oviedo ICCS&T 2011. Extended Abstract
a
b
c
d
e Figure 1 Mosaic Images for the 5 coal samples (106-125 micron); (a)- S1(heat affected) (b)- S2 (c)- 88 (d)- 91, (e)- Stockpile
6
Oviedo ICCS&T 2011. Extended Abstract
3.3 Burnout behaviour The burnout profiles in Figure 2 (below) include volatile release and char burnout for the larger size fraction. Samples with better burnout behaviour should have less combustible components remaining at earlier residence times [23]. Furthermore Cloke et al. [24] have shown that particle size can also affect burnout. Clearly 79 S1 shows the poorest burnout rates, with an almost linear burnout profile. Figure 3 shows that the chars structures appear to be almost solid, post pyrolysis. This must relate to the heat affected nature of the particles, slowing down the pyrolysis rates, lowering the formation of open pores within a char structure. In a diffusion controlled reaction, it is likely that burnout would be limited to the external surface, allowing slow burnout of the solid particles. This would explain the relatively slow burnout rates. 78 S2 forms predominantly thick walled spheres (crassispheres) or solids. In addition, from Figure 3, Stockpile and coal 88 form the thinnest char structures and this is reflected in the highest burnout rates in Figure 2. However, all the samples, by 700ms total refire time achieve 70-85% burnout. The remaining char structures, at these latter stages of combustion, require further investigation in order to explain which morphologies are responsable for this residual level of carbon in ash. 0.9 0.8 0.7
Burnout
0.6 91 106‐125
0.5
78 S1 106‐125
0.4
Stockpile 106‐125
0.3
78 S2 106‐125 88 106‐125
0.2 0.1 0.0 0
200
400
600
800
Residence Time (ms) Figure 2 Burnout behaviour of refired char from the larger pulverised coal fraction (106-125µm).
7
Oviedo ICCS&T 2011. Extended Abstract
From Figure 3, Coal 91 clearly has a significant number of solid chars which relate to the higher levels of inertite particles (Table 3) that are present in the original coal.
a
b
c
d
e Figure 3 Char Morphology of the pyrolysis char from the larger pulverised coal fraction (106-125µm)
(a)- S1(heat affected) (b)- S2 (c)- 88 (d)- 91, (e)- Stockpile .
3
Conclusions
i.
Petrographic variations are seen both between seams and size fractions, with vitrinite ranging from 65 – 77%, liptinite (2.5% - 7.1%) and inertinite (16 – 30%). Microlithotype associations aid in explaining these differences and are more informative than petrographic data alone.
ii.
The heat affected coal (78 S1) is chemically (Table 1) and petrographically different from the other four samples (Table 2; Table 3). Vitrinite is the most
8
Oviedo ICCS&T 2011. Extended Abstract
abundant maceral (up to 35%), the majority of which is as the microlithotype vitrite (16% – 23.8%; Table 3) and vitrinertite (8.8% - 12.5%). Inertinite is a minor component (up to 4%) either as inertite (2.1%) or vitrinertite. iii.
Burnout profiles show that the heat affected coal 78 S1 has the poorest burnout rates. From char analysis, it produces char morphologies that are denser with low porosity. The stockpile and coal 88 have the best profiles for burnout and this is reflected in the petrographic compositions where the inertinite is relatively low (compared to the other coals) and the char structures are the thinnest.
iv.
The quantity of dense chars appears to relate directly to the petographic composition of the initial coals.
Acknowledgement We would like to thank NERC, RWE Npower for sponsorship and RWE Npower for training and technical support. We would also like to thank Neil Holloway and Sharon Gibbons for technical support at Royal Holloway.
References [1] Hadjipaschalis I, Kourtis G, Poullikas A. Assessment of oxyfuel power generation technologies. Renewable and sustainable energy review 2009; 13: 2637-2644 [2] Wall T, Liu Y, Spero C, Elliot L, Khar S, Rathman R, Zeenathal F, Moghtaderi B, Buhre B, Sheng C, Gupta R, Yamada T, Makino K, and Yu J. An overview of oxyfuel coal combustion-state of the art research and technology development. Chemical engineering research and design 2009; 87: 1003-1016. [3] Cloke, M., Wu, T., Barranco, R. and Lester, E. Char characterization and its application in a coal burnout model. Fuel 2003; 82: 1989-2000. [4] Lester E, Cloke M. The characterization of coals and their respective chars formed at 1300ºC in a drop-tube furnace. Fuel 1999; 78: 1645-1658. [5] Choudhury N, Biswas S, Sarkar P, Kumar M, Ghosal S, Mitra T, Mukherjee A, Choudhury A. Influence of rank and macerals on the burnout behaviour of pulverized Indian coal. Int. J. Coal Geol. 2008; 74: 145-153. [6] Card, J.B.A., Jones, A.R. A drop tube furnace study of coal combustion and unburned carbon content using optical techniques. Combustion and Flame. 1995; 101: 539-547.
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Oviedo ICCS&T 2011. Extended Abstract
[7] Cloke M, Lester E, Belghazi A. Characterization of the properties of size fractions from ten world coals and their chars produced in a drop-tube furnace. Fuel. 2002; 81: 699-708. [8] Lester, E. Characterisation of coals for combustion. PhD thesis 1994. [9] BS 6127.4: 1990, ISO 7404-4:1988, British Standard, Petrographic analysis of bituminous coal and anthracite – Part 4: method of determining microlithotype division, carbominerite and minerite composition [10] BS 6127.3: 1995, ISO 7404-3:1994, British Standard, Petrographic analysis of bituminous coal and anthracite – Part 3: method of determining maceral group composition of bituminous coal and anthracite. [11] International Committee for Coal and Organic Petrology (ICCP). The new vitrinite classification (ICCP System 1994), Fuel 1998; 77: 349-358. [12] International Committee for Coal and Organic Petrology (ICCP). The new inertinite classification (ICCP system 1994), Fuel 2001; 80: 459-471. [13] BS 1016-105:1992, methods for analysis and testing of coal and coke, part 105: determination of gross calorific value: 1-11. [14] BS 1016-104.3:1998, ISO 562:1998, methods for analysis and testing of coal and coke, part 104: proximate analysis, section 104.3: determination of volatile matter content, second edition, 1998-02-01, 1-6. [15] BS 1016-104.4:1998, ISO 1171, Solid mineral fuels – determination of ash, third edition, 1997-12-15, 1-8. [16] BS 1016-104.1:1999, ISO 11722, solid mineral fuels – Hard coals – Determination of moisture in the general analysis test sample by drying in nitrogen, first edition 199905-01, 1-5. [17] BS ISO 17246:2005, Coal - Proximate analysis, International Standard, First Edition, 2005-05-01, 1-2. [18] BS ISO 17247:2005, Coal – Ultimate analysis, International Standard, First Edition, 2005-05-01, 1-4. [19] Cloke, M., Wu, T., Barranco, R. and Lester, E. Char characterization and its application in a coal burnout model. Fuel 2003; 82: 1989-2000. [20] Unsworth JF, Barrat DJ, Robert PT. Coal quality and combustion performance: an international perspective. Coal science and technology series 19, Amsterdam: Elsevier; 1991. P350. [21] Stach E. The microlithotypes of coal and their strength. In: Stach, E., et al., Stach’s Textbook of Coal Petrology, Gebr. Borntraeger, Berlin-Stuttgart; 1982, p.173-177.
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Oviedo ICCS&T 2011. Extended Abstract
[22] Taylor GH, Teichmüller M, Davis A, Diessel CFK, Littke R, Robert P. Organic Petrology, Gebrüder Bornstraeger, Berlin, Stuttgart; 1998. [23] Bostick NH, Betterton WJ, Gluskoter HJ, Nazrul Islam M. Petrography of Permian “Gondwana” coals from boreholes in northwestern Bangladesh, based on semiautomated reflectance scanning. Org. Geochem. 1991; 17: 399-413. [24] Cloke M, Lester E, Gibb W. Characterisation of coal with respect to carbon burnout in p.f.-fired boilers. Fuel 1997; 76: 1257-1267.
11
The effect on coal flotation of segregating maceral groups based on particle size Esmaeil Jorjani, Sina Esmaeili, Maedeh Tayebi Khorami Department of Mining Engineering, Science and Research Branch, Islamic Azad University, Tehran, Iran Email:
[email protected]
Abstract In this study, maceral groups in 15 samples from the central Alborz coal mine of Iran were segregated based on the different size fractions of -500+150, +150-75, -75+40, -40+25, and -25 µm. A total of 75 polished sections were prepared, and maceral groups were determined using microscopic studies. For the majority of the samples, vitrinite was concentrated in the fine size fractions, i.e., -40+25, and -25 µm. For the different size fractions of -500+150, +150-75, 75+40, -40+25, and -25 µm, samples with maximum and minimum assays of vitrinite (minimum and maximum assays of liptinite, respectively) were used to determine the effect of segregating maceral groups based on particle size on coal flotation. The results showed that samples with higher assays of liptinite (lower assays of vitrinite) produced a concentrate with higher combustible value and recovery. Liptinite had a positive effect and vitrinite had a negative effect on combustible value and recovery of coal flotation concentrate. It can be concluded that, in a coal flotation system, segregating the maceral groups by particle size is more important than particle size itself. Keywords: Coal, Flotation, Maceral, Grindability.
1. Introduction Grindability of coal is an important technological parameter that indicates the ease of pulverizing a coal in comparison to a reference coal that is usually determined by Hardgrove Grindability index (HGI) [1]. HGI is most directly related to the maceral and maceral group composition, but it is also dependent on rank and mineral content [2].
1
The three maceral groups are vitrinite, liptinite, and inertinite, which in turn can be subdivided into finer classifications [3]. Grindability studies [4-8], have demonstrated that, as grinding progresses, the most brittle microlithotypes, such as vitrite and inertite, will partition preferentially to the finer fractions. The remaining coal is enriched in the harder microlithotypes, i.e., the liptinite-rich and inertinite-rich bimaceral and trimaceral varieties [2]. In addition, several researchers have studied the relationship between HGI and coal petrology [9-13]. Hower and Wild demonstrated a strong correlation between an increase in liptinite content and a decrease in HGI for coals of narrow vitrinite reflectance ranges for a suite of carboniferous-age coals [10]. Trimble and Hower corroborated
this
relationship
between
an
increased
concentration
of
liptinite-rich
microlithotypes and a decrease in HGI for eastern Kentucky coals from narrow vitrinite reflectance ranges [8]. Jorjani et al. studied the relationship between petrography with grindability for Kentucky coals and found that HGI decreased as the content of liptinite in the coal increased, but HGI increased when the content of vitrinite in the coal increased [12]. Chehreh Chegani et al. also found that HGI was decreased by higher liptinite, semi-fusinite, and resinite contents in coal [13]. Froth flotation is a technique that is used for fine coal beneficiation. Since froth flotation relies on the surface properties of coal for the separation of particles, coal petrology plays an important role in the process [2]. Arnold and Aplan [14] measured contact angles for several U.S. coal samples to quantify the hydrophobicity of individual coal macerals. It was found that the order hydrophobicity of the coal macerals was liptinite > vitrinite > inertinite and that these macerals have typical ranges of contact angles of 90–130º, 60–70º and 25–40º, respectively. In addition, Arnold and Aplan reported a number of studies that examined maceral or lithotype partitioning, generally without consideration of the maceral association. They noted that the order of decreasing floatability was liptinite > vitrinite > fusinite and vitrain > clarain > durain > fusain. They found that vitrite was concentrated in the faster floating fractions and that inertite was concentrated in the slower floating fractions [15]. Hirt and Aplan examined coal from eastern Kentucky and arranged the macerals according to decreasing of floatability in order of pseudovitrinite (high Rmax) > pseudovitrinite (low Rmax) > vitrinite (high Rmax) > vitrinite (low Rmax) = micrinite = exinite = semifusinite > resinite > fusinite [16].
2
In this study, a series of coals from the central Alborz coal mine in Iran was investigated with the objective of determining any possibility of the segregating of maceral groups in the different size fractions of -500+150, +150-75, -75+40, -40+25, and -25 µm; the effects of this partitioning on the combustible value (CV) and combustible recovery (CR) of coal flotation concentrate were determined as well.
2. Experimental 2.1. Sample preparation, screening, and group maceral analysis Fifteen samples, namely Alborz Kavan (A), Alborz Kavan (B), Hamidzadeh (A), Tunnel No. 8 (A), Tunnel No. 8 (B), Fajr Mamashi (A), Fajr Mamashi (B), Kar Sang (A), Kar Sang (B), Lavij (A), Lavij (B), Mine No. 20 (A), East of flayer 4 (A), East of flayer 4 (B), and Simab (A), were prepared from different mines at the Alborz coal deposit, and the –500-μm size fraction of these samples was separated and used for further studies. The coal mines sampled in this study were those that were among the most heavily mined in the region. The –500-μm fraction of the samples was wet screened at different size fractions of -500+150, -150+75, -75+40, -40+25, and -25 μm. Subsequently, 75 subsamples were prepared and used for the preparation of polished sections according to the ISO 7404-2 standard method [17]. The maceral group analysis, i.e., assay of liptinite, vitrinite, and inertinite, was performed using microscopic studies and according to the ISO 7404-3 standard method [18]. In our study, the group macerals were counted based on an organic-only basis, i.e., liptinite + vitrinite + inertinite = 100 %.
2.2. Flotation studies The flotation tests were conducted in a 2.5-L Denver laboratory flotation cell, and CV and CR were determined for the flotation concentrate. The CV and CR obtained for the flotation experiments were calculated using the following equations: CV (%) = 100 ‐ Ash (%) (1) CR (%) = [(Wc × (100‐Ac))/(Wf × (100‐Af))] × 100, (2)
3
where Wc is weight of clean coal (%); Wf is weight of feed (%); Ac is ash content of clean coal by weight (%); and Af is ash content of feed by weight (%). The optimized conditions that were determined and used in the flotation studies were a rotation rate of 1400 rpm, kerosene (collector) with a dosage of 969 g/ton, pine oil (frother) with a dosage of 117 g/ton, a pulp density at 16 wt%, and a pH value of 8. The total froth collecting time was 120 sec [19].
3. Results and discussion 3.1. Maceral group assay in different size fractions Figs. 1 to 5 present the assay of liptinite, vitrinite, and inertinite in different size fractions, i.e., -500+150, -150+75, -75+40, -40+25, and -25 μm, for the 15 samples. In the examined samples, the assay of inertinite was low in all of size fractions, so we focused on the trends of liptinite and vitrinite.
80 70
Assay (%)
60 50 40 30 20
Liptinite Vitrinite Inertinite
10 0
Fig. 1. Assays of liptinite, vitrinite, and inertinite in the -500+150-μm size fraction of different samples
4
80 70
Assay (%)
60 50 40 30 20 10
Liptinite Vitrinite Inertinite
0
Fig. 2. Assays of liptinite, vitrinite, and inertinite in the -150+75-μm size fraction of different samples
80 70
Assay (%)
60 50 40 30 20 10
Liptinite Vitrinite Inertinite
0
Fig. 3. Assays of liptinite, vitrinite, and inertinite in the -75+40-μm size fraction of different samples 5
90 80
Assay (%)
70 60 50 40 30 20 10
Liptinite Vitrinite Inertinite
0
Fig. 4. Assays of liptinite, vitrinite, and inertinite in the -40+25-μm size fraction of different samples
90 80
Assay (%)
70 60 50 40 30 20 10
Liptinite Vitrinite Inertinite
0
Fig. 5. Assays of liptinite, vitrinite, and inertinite in the –25-μm size fraction of different samples 6
According to Figs. 1 to 5, the -500+150-μm size fraction of liptinite and vitrinite were not segregated in the examined samples due to the integration of these two group macerals. However, segregation became evident as size fractions from -150+75 μm to -25 μm were considered. For the size fraction of -75+40 μm, the assay of vitrinite was higher than the other maceral groups in all of the samples examined, except for Alborz Kavan (A) and (B) samples; while, for the size fractions of -40+25 μm and -25 μm, the assay of vitrinite was higher than liptinite in all of the samples. The results show that vitrinite was partitioned preferentially in the fine size fractions (-40+25 μm and -25 μm), while the liptinite grade decreased.
3.2. Flotation experiments The samples with maximum grade of vitrinite (minimum grade of liptinite) are East of layer 4 (A) (Fig. 1), East of layer 4 (A) (Fig. 2), Kar Sang (A) (Fig. 3), Alborz Kavan (B) (Fig. 4), Kar Sang (B) (Fig. 5) for different size fractions of -500+150, -150+75, -75+40, -40+25, and -25 μm, respectively. In contrast, the samples with minimum grade of vitrinite (maximum grade of liptinite) were Kar Sang (A) (Fig. 1), Alborz Kavan (A) (Fig. 2), Alborz Kavan (B) (Fig. 3), Tunnel No. 8 (A) (Fig. 4), East of layer 4 (B) (Fig. 5) for different size fractions of -500+150, 150+75, -75+40, -40+25, and -25 μm, respectively (Table 1). In each size fraction, two samples with the maximum and minimum grade of vitrinite (Table 1) were used to determine the effect of the quality and quantity of the maceral groups in the samples on the flotation process. The results are shown in Table 2.
Table 1. Samples with maximum and minimum assay of vitrinite in different size fractions Size fraction (μm)
Sample with maximum assay of vitrinite
Sample with minimum assay of vitrinite
-500 +150 -150 +75 -75 +40 -40 +25 -25
East of layer 4 (A) East of layer 4 (A) Kar Sang (A) Alborz Kavan (B) Kar Sang (B)
Kar Sang (A) Alborz Kavan (A) Alborz Kavan (B) Tunnel No. 8 (A) East of layer 4 (B)
7
Table 2. Results of flotation experiments
-500 +150 -150 +75 -75 +40 -40 +25 -25
Sample
Ash (%)
CV * (%)
Vitrinite (%)
Liptinite (%)
Concentrate weight (%)
Concentrate ash (%)
Kar Sang (A) East of layer 4 (A) Alborz Kavan (A) East of layer 4 (A) Alborz Kavan (B) Kar Sang (A) Tunnel No. 8 (A) Alborz Kavan (B) East of layer 4 (B) Kar Sang (B)
17.26 13.45 34.28 13.39 34.72 17.21 35.86 35.68 19.67 22.85
82.74 86.55 65.72 86.61 65.28 82.79 64.14 64.32 80.33 77.15
22.94 73.76 27.34 67.51 36.55 70.54 52.41 85.41 48.26 82.77
75.33 24.89 70.77 30.78 61.33 26.51 45.71 12.85 49.28 14.12
83.11 85.06 66.76 84.67 65.41 65.73 54.42 56.69 68.66 73.21
2.88 11.77 9.57 14.37 15.24 20.39 8.50 12.78 5.08 12.00
Concentrate CV * (%) 97.12 88.23 90.43 85.63 84.76 79.61 91.50 87.22 94.92 88.00
Concentrate CR** (%) 97.55 86.71 91.85 83.71 84.92 63.20 77.64 76.87 81.13 83.51
*CV: Combustible Value ** CR: Combustible Recovery
Figs. 6 and 7 show the combustible value (CV) and combustible recovery (CR) of the concentrate for the samples with maximum and minimum grades of vitrinite in different size fractions. With regard to Figs. 6 and 7, it can be concluded that liptinite has a positive effect on the combustible value and combustible recovery of the concentrate when the particle size is constant in a coal flotation system.
100.00
Concentrate CV (%)
Size fraction (μm)
95.00 90.00 85.00 80.00 75.00 70.00
Samples with maximum grade of Vitrinite and minimum grade of Liptinite
65.00
Samples with maximum grade of Liptinite and minimum grade of Vitrinite
60.00 ‐500 +150
‐150 +75
‐75 +40
‐40 +25
Size fraction (micron)
Fig. 6. Combustible value of the flotation concentrate 8
‐25
Concentrate CR (%)
100.00 95.00
Samples with maximum grade of Vitrinite and minimum grade of Liptinite
90.00
Samples with maximum grade of Liptinite and minimum grade of Vitrinite
85.00 80.00 75.00 70.00 65.00 60.00 ‐500 +150
‐150 +75
‐75 +40
‐40 +25
‐25
Size fraction (micron)
Fig. 7. Combustible recovery of the flotation concentrate 4. Conclusions
The results of petrographical analysis of different size fractions showed that the vitrinite content increased and liptinite content decreased as particle size decreased.
For the size fractions of -40+25 µm and -25 µm, the assay of vitrinite was higher than that of liptinite in all of the samples that were examined.
The results of this study are compatible with the Hardgrove grindability index values determined for maceral groups by previous researchers. Vitrinite has a higher grindability index than liptinite.
When particle size is constant, samples with a high grade of liptinite can produce a concentrate with a higher combustible value and a higher combustible recovery value.
It can be concluded that, in a coal flotation system, the segregation of macerals by particle size is more important than the particle size itself.
9
Acknowledgment The authors gratefully acknowledge the financial support provided by the Science and Research Branch of Islamic Azad University.
References [1] Speight GJ. Handbook of Coal Analysis. Wiley-Interscience, Chichester, West Sussex, England, 2005,161. [2] Suarez-Ruiz I, Crelling JC. Applied Coal Petrology: The Role of Petrology in Coal Utilization. Elsevier, Ltd. 2008. [3] Miller BG. Coal Energy Systems, 242. Elsevier publication, Burlington, USA. 2005. [4] Hower JC. Interrelationship of coal grinding properties and coal petrology, Minerals and Metallurgical Processing, 1998;15:1–16. [5] Hower JC, Calder JH. Maceral/microlithotype analysis of the Hardgrove grindability of lithotypes from the Phalen coal bed, Cape Breton, Nova Scotia. Minerals & Metallurgical Processing. 1997;14 (1):49–54. [6] Padgett PL, Hower JC. Hardgrove grindability study of Powder River Basin and Appalachian coal components in the blend to a Midwestern power station. Minerals & Metallurgical Processing. 1997;14 (3): 45–49. [7] Trimble AS, Hower JC. Maceral/microlithotype analysis of the progressive grinding of a central Appalachian high volatile bituminous coal blend. Minerals and Metallurgical Processing. 2000;17: 234–243. [8] Trimble AS. Hower JC. Studies of the relationship between coal petrology and grinding properties. International Journal of Coal Geology. 2003; 54:253–260. [9] Hower JC, Graese AM, Klapheke JG. Influence of microlithotype composition on Hardgrove grindability for selected Eastern Kentucky coals, International Journal of Coal Geology. 1987; 7: 227–244.
10
[10] Hower JC, Wild GD. Relationship between Hardgrove Grindability Index and petrographic composition for high-volatile bituminous coals from Kentucky. Journal of Coal Quality. 1988; 7:122–126. [11] Hower JC, Wild GD. Maceral/microlithotype analysis evaluation of coal grinding: examples from Central Appalachian high volatile bituminous coals. Journal of Coal Quality. 1994;13: 35–40. [12] Jorjani E, Hower JC, Chehreh Chelgani S, Shirazi MA, Mesroghli Sh. Studies of relationship between petrography and elemental analysis with grindability for Kentucky coals. Fuel.2008; 87:707–713. [13] Chehreh Chelgani S, Hower JC, Jorjani E, Mesroghli Sh. Bagherieh AH. Prediction of coal grindability based on petrography, proximate and ultimate analysis with multiple regression and artificial neural network models. Fuel Process. Technol. 2008;89:13–20. [14] Arnold BJ, Aplan FF. Coal froth flotation: the response of coal and mineral particles to reagent and circuit variations, Advances in Mineral Processing, Society of Mining Engineers, 1986, p. 351. [15] Arnold BJ, Aplan FF. The hydrophobicity of coal macerals. Fuel. 1989;68: 651–658. [16] Hirt WC, Aplan FF. The influence of operating factors on coal recovery and pyritic sulfur rejection during coal flotation, in: Dugan PR, Processing and Utilization of High Sulfur Coal, IV, Elsevier, Amsterdam, 1991;p. 339. [17] CSN ISO 7404-2. Methods for the petrographic analysis of bituminous coal and anthracite. Part 2. Method of preparing coal samples. Praha: Cesky Normalizanci Institut; 1995 p. 16. [18] CSN ISO 7404-3. Methods for the petrographic analysis of bituminous coal and anthracite. Part3: method of determining group composition. Praha: Cesky Normalization Institut; 1997 p. 12. [19] Alimohammadi, Sh, Optimization of operational variables in central Alborz coal flotation circuit, M.Sc. dissertation, Science and Research branch, IAU, Tehran, Iran, 2010, p. 86-116.
11
Oviedo ICCS&T 2011. Extended Abstract
Dimethyl ether production from Victorian brown coal and biomass: A comparative process modelling study Kazi Bayzid Kabir, Gabrielle Grills, James Walter and Sankar Bhattacharya Department of Chemical Engineering, Monash University, Clayton, VIC-3168, Australia. Email:
[email protected] Abstract Victorian has large reserves of brown coal with a resource life of around 500 years at current consumption rates. Forest industries in Victoria produces huge amount of forestry residue from its operation. Both brown coal and biomass can be used for synthesis of liquid fuels through gasification. Among different alternatives, dimethyl ether stands out to be product because of its environmentally benign properties and wide applicability. Simulation studies using Loy Yang brown coal and Eucalyptus saw dust were carried out to find the syngas composition and subsequent DME yield over a range of variables and coal-biomass blends. FACTSAGE and Aspen Plus were used for the simulations. The results indicate that the syngas composition from both brown coal and biomass is suitable for DME synthesis since the appropriate ratio of H2 and CO for DME synthesis can be achieved during gasification around 900°C. Steam gasification of Loy Yang coal at 900°C can produce the required syngas composition without inclusion of water gas shift reactor in the subsequent downstream sections. For biomass, partial oxidation with sub-stoichiometric oxygen only but without any added steam, in the gasifier followed by high temperature shift (HTS) at 360°C is the most suitable option for obtaining the required ratio of H2 to CO. Biomass gasification at the same temperature as that for coal can save significant amount of steam during gasification if water gas shift reactor is included in the subsequent process. The process models also reveal that biomass does not prove beneficial in terms of CO2 emission and DME yield if calculations are made on the basis of carbon fed to the process. For same amount of feedstock, DME yield is 38% higher for coal than for the biomass. Results these simulations form the basis of DME synthesis experiments as part of a wider experimental project. 1. Introduction Victorian brown coal has long been the main energy source for the social and economic development of Victoria. It is likely to continue to be the most important and cheapest energy source for Victorian economy. However, the high moisture and high reactivity of dried brown coal mean that the coal cannot be readily exported, and has to be used locally for power generation by direct combustion. With current consumption rate the brown coal reserve has a resource life of around 500 years [1].
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As far as biomass is concerned, Victoria has 8.3 million hectares of forest including 7.9 million hectares of native forest and 360,000 hectares of plantations. Current average sawlog harvest from Victorian forests is approximately 530,000 m3 per annum [2, 3]. Residues produced at each stages of wood and wood-product processing can be used as potential energy source to reduce pressure on conventional fossil fuels. Gasification is a technology that converts brown coal or biomass into a gaseous product (synthesis gas) which can be fed into a gas turbine or a fuel cell for electricity generation at much higher efficiencies than the existing pulverised-fuel combustion system [4]. The synthetic gas is also a necessary feedstock for production of liquid fuels [5], such as DME, which is an exportable product.
Dimethyl ether (DME) is a non-toxic, non-carcinogenic and non-corrosive compound. It can be used as an alternative to diesel fuel due to its high cetane number and the low emissions of CO, NOx, and particulates upon its combustion [6]. DME can be manufactured in large quantities from raw materials such as natural gas, petroleum, coal or biomass. The syngas produced from these feedstocks can then be converted to methanol which on dehydration gives DME. Till now, the majority of the DME plants worldwide use natural gas as feedstock. Both Victorian brown coal and forestry residue has significant potentials as feedstock for DME synthesis.
As part of a wider experimental project on DME synthesis from low-rank fuels, this current work involves equilibrium modelling of brown coal and biomass gasification using a proprietary package FACTSAGE which is also used to predict the amount of alkaline metals and the other impurities in the syngas. Based on this equilibrium modelling, an appropriate temperature was chosen for the steady state process modelling using Aspen Plus. The overall process model included drying, equilibrium model for gasification, gas cleaning, kinetic model for DME synthesis in single step, and product purification. Both coal and biomass, and coal-biomass blends in different ratios were considered in the study.
2. Simulation Basis and Methodology Loy Yang brown coal and Eucalyptus saw dust were considered for this simulation study. Analyses of the coal and saw dust is shown in Table 1. Equilibrium model in
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Oviedo ICCS&T 2011. Extended Abstract
FACTSAGE was used to predict the syngas composition for Loy Yang brown coal and biomass over the temperature range of 800-1500°C. Process models for one-step DME production from Victorian brown coal and biomass residues have also been developed in ASPEN PLUS. For consistency in comparison, both biomass and coal were assumed to have equal moisture content of 62% which was dried down to 12% moisture content before feeding to the gasifier.
Soave-Redlich-Kwong (SRK) equation of state was selected for the simulation, since it has been reported to give better property estimation than Redlich-Kwong (RK) or PengRobinson (PR) for methanol synthesis, water gas shift reaction (WGSR) [7] and DME synthesis[8].
Table 1: Composition of Loy Yang coal and Eucalyptus sawdust (% dry basis) Ultimate analysis C H Loy Yang
N
S
O Ash
Ash analysis Cl SiO2 Al2O3 Fe2O3 K2O MgO Na2O CaO SO3 P2O5
65 4.6 0.72 0.5 25.5 3.57 0.11 56.5
20.5
4.6 1.3
3.6
4.7 1.6
5 0.2
Eucalyptus saw 50.7 6.2 0.16 0.01 42.6 0.3 0.01 18.9
1.8
2.6 9.6
9.9
3.4 32.7
16 1.7
Coal
dust
For gasification an equilibrium model has been used, whereas calculations for DME synthesis were performed using a kinetic model. The gasifier model comprised of a yield reactor for the pyrolysis of coal followed by a Gibbs reactor for the conversion of the pyrolysis products (e.g. volatiles and char). A gasification temperature of 900°C was chosen for the study which would produce lower concentrations of alkaline species in the syngas. Gasifier pressure was 30 bar. A plug flow reactor with LangmuirHinshelwood-Hougen-Watson (LHHW) kinetic model was used as the DME synthesis reactor. Kinetic data for the syngas to DME bi-functional catalyst were taken from the work of Ni et al [9]. Pressure, temperature and space velocity for DME reactor were 60 bar, 240°C and 800 ml.gcat-1.h-1, respectively. The schematic of the steady state process model for DME synthesis from Victorian brown coal and biomass is shown in Figure 1.
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Oviedo ICCS&T 2011. Extended Abstract
3. Results and Discussion 3.1 Simulation of coal and biomass gasification using FACTSAGE The equilibrium model in FACTSAGE was used to determine thermodynamic possibility of formation of 926 species (279 gases, 157 liquids and 490 solids) during steam gasification (ATR-auto thermal reforming) of brown coal and biomass at 30 atm and the temperature range of 500-1400°C. Biomass completely gasifies at 700°C while coal requires a minimum of 900°C for complete conversion.
Since both coal and biomass have chlorine and alkaline metals, some of the alkaline metals and chlorine is expected to be released into the gas. Figure 2 shows the major chlorine and alkaline metal containing species in the coal syngas. Loy Yang and other Victorian brown coals have high sodium content and therefore NaCl concentration is high in the syngas. Both NaCl and KCl concentration increased with temperature at the expense of lower HCl concentration. HCl reacts with NaOH and KOH at the gaseous phase to give NaCl and KCl.
Figure 1: Steady state of the process model for coal/biomass/syngas to DME
Figure 3 shows the major chlorine and alkaline metal species in biomass syngas. In contrast to the coal, biomass has higher potassium content and hence has higher potassium containing species. But the amount of potassium containing species is still
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Oviedo ICCS&T 2011. Extended Abstract
low since the ash content is much lower than the brown coal. KOH is the dominant species here since chlorine content in biomass is extremely low. It is apparent from both Figure 2 and 3, alkaline species go into gas phase at higher temperature and hence would cause more problem for the downstream units. Therefore gasification at lower temperature would be beneficial to avoid initiation of corrosion in the downstream units by alkaline metal species.
Figure 2: Chlorine and alkaline metal species in Loy Yang syngas
The equilibrium predictions also showed that the solid phase (i.e. ash) starts melting at temperature close to 1300°C.
3.2 Process model for coal/biomass to DME production Equilibrium predictions using FACTSAGE was then used in the steady-state process simulation of DME synthesis from Loy Yang coal and biomass, as independent feeds and as blends of different proportions. The model includes drying, gasification, gas cleaning, syngas-to-DME synthesis and product purification. Gasification was carried out at 900°C and 30 bar while DME synthesis pressure and temperature was 60 bar and 240°C, respectively.
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Oviedo ICCS&T 2011. Extended Abstract
Our earlier simulation work [10] showed that H2 to CO molar ratio in the syngas is very important for optimal yield of DME. The simulation here showed that for both brown coal and biomass, syngas close to an appropriate ratio of 1.5 [10] can be produced. This is a definite positive aspect for brown coal (compared to black coal) since reported simulation work on black coal to DME synthesis showed that it requires additional hydrogen-rich feedstock (i.e. natural gas) because of high C/H ratio [11]. C/H ratio in brown coal is significantly less than black coal which makes it suitable for DME synthesis.
Figure 3: Chlorine and alkaline metal species in biomass syngas
An appropriate ratio of H2 to CO for DME synthesis can be obtained either by gasification followed by water gas shift, or by gasification alone. This depends on the gasification temperature. A comparative study of both the options was performed to assess the necessity of water gas shift reactor in the process. If the required gas composition can be achieved without the shift reaction, it would essentially save both the capital and operating cost associated with the shift reactor. To simulate the water gas shift reactor a plug-flow reactor with LHHW kinetic model was used [12]. With inclusion of water gas shift can only save 5-10% steam for gasification at 900°C
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Oviedo ICCS&T 2011. Extended Abstract
producing similar quantity CO2. Hence, the water gas shift can clearly be avoided if gasification is carried out using at 900°C.
For biomass, partial oxidation (with sub-stoichiometric oxygen only, without any added steam) in the gasifier followed by high temperature shift (HTS) at 360°C is the most suitable option for obtaining the required ratio of H2 to CO. Since the C/H ratio in biomass is lower than brown coal, the required syngas composition for DME syntheis can be achieved without any steam addition in the gasifier. Oxygen requirement for gasification stage and the total of steam requirement (for HTS and gasification stages) biomass gasification is significantly lower than those required for brown coal gasification.
The DME yield was calculated for coal, biomass and their blend at different ratios. Major products from the process are DME and CO2. The selectivity and yield of DME and CO2 was calculated based on the carbon fed to the process: Yield =
Selectivity =
Carbon in DME/CO 2 Total carbon fed to the process
2 × Carbon in DME/CO 2 Carbon other products other than DME/CO 2
Yields of DME and CO2 as a function of coal and biomass blends are shown in Figure 4. The yield is almost unchanged for all the cases where the calculations were performed on carbon basis and the blending of biomass with coal does not have any significant effect. However if coal is compared to biomass, the DME yield is higher since coal has higher carbon than biomass. For same amount of coal and biomass feedstock, coal results in 1.38 times the DME than that obtained for biomass alone. Similar trend is also observed for CO2 generation. Figure 5 shows selectivity of DME and CO2 for the process. The selectivity for CO2 slightly increases while the selectivity of DME decreases with higher biomass in the blend fed to the gasifier. The DME synthesis reactor model here assumes three concurrent reaction occurring in the synthesis reactor: CO + 2H2 → CH3OH CO2 + 3H2 → CH3OH + H2O 2CH3OH → C2H6O
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Oviedo ICCS&T 2011. Extended Abstract
It is clear that CO2 and DME in inversely related to each other in terms of selectivity. Increase in CO2 selectivity cause decreased DME selectivity. The summation of the selectivity of DME and CO2 are constant for all the cases.
Figure 4: DME and CO2 yield for coal-biomass blends
Figure 5: Selectivity of DME and CO2 for coal-biomass blends
4. Conclusions and practical implications
Both biomass and brown coal can be gasified with high carbon conversion at relatively low temperatures of around 900°C. FACTSAGE equilibrium models revealed that,
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Oviedo ICCS&T 2011. Extended Abstract
biomass gasification can be carried out at these temperatures without significant release of alkaline and chlorine containing species. However, for Loy Yang brown coal significant amount of alkaline species are released at these temperatures. In principle, lower temperature gasification would also be beneficial because of lower energy requirement. However, practical issues such as evolution of tar (particularly for biomass) and gaseous contaminants, associated gas cleaning, and bed agglomeration problems should be evaluated through targeted experiments.
Gasification simulation revealed that the syngas composition from both brown coal and biomass is suitable for DME synthesis since the appropriate ratio of hydrogen and carbon monoxide for DME synthesis can be achieved easily. Steam gasification of Loy Yang coal at 900°C can produce the required syngas composition without inclusion of water gas shift reactor in the subsequent downstream sections.
For biomass, partial oxidation (with sub-stoichiometric oxygen only, without any added steam) in the gasifier followed by high temperature shift (HTS) at 360°C is the most suitable option for obtaining the required ratio of H2 to CO. Biomass gasification at the same temperature as that for coal can save significant amount of steam during gasification if water gas shift reactor is included in the subsequent process.
The process models also reveal that biomass does not prove beneficial in terms of CO2 emission and DME yield if calculations are made on the basis of carbon fed to the process. For same amount of feedstock, DME yield is 38% higher for coal than for the biomass.
5. Acknowledgement Kazi Bayzid Kabir gratefully acknowledges the Department of Primary Industries (DPI) - State Government of Victoria, Brown Coal Innovation Australia, and the Department of Chemical Engineering, Monash University for his postgraduate scholarships.
References [1] ABARE. Australian Energy Resource Assessment. Canberra: Geoscience Australia and ABARE; 2010.
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9
Oviedo ICCS&T 2011. Extended Abstract
[2]
VAFI.
Forestry
in
Victoria:
Key
Issues
and
Facts.
in:
http://www.vafi.org.au/documents/forestrykeyissuescomp.pdf; 2008. [3] ABARE. Australian Forest and Wood Products Statistics, March and June Quarters. Canberra; 2007. [4] Bell DA, Towler BF, Fan M. Coal Gasification and Its Applications. 1st ed., Oxford: Elsevier; 2011. [5] Olah GA, Goeppert A, Prakash GKS. Beyond Oil and Gas: The Methanol Economy. John Wiley & Sons; 2009. [6] Arcoumanis C, Bae C, Crookes R, Kinoshita E. The potential of di-methyl ether (DME) as an alternative fuel for compression-ignition engines: A review. Fuel 2008; 87: 1014-30. [7] Graaf GH, Sijtsema PJJM, Stamhuis EJ, Joosten GEH. Chemical equilibria in methanol synthesis. Chem Eng Sci 1986; 41: 2883-90. [8] Shim HM, Lee SJ, Yoo YD, Yun YS, Kim HT. Simulation of DME synthesis from coal syngas by kinetics model. Korean J Chem Eng 2009; 26: 641-8. [9] Nie Z-G, Liu H, Liu D, Ying W, Fang D. Intrinsic kinetics of dimethyl ether synthesis from syngas. J Nat Gas Chem 2005; 14: 22-8. [10] Kabir KB, Hein K, Bhattacharya S. Dimethyl Ether production from Victorian Brown Coal: A Model integrating Drying, Gasification and DME Synthesis processes. in:
36th Internation Technical Conference on Clean Coal and Fuel Systems,
Clearwater, USA; 2011. [11] Zhou L, Hu S, Chen D, Li Y, Zhu B, Jin Y. Study on Systems Based on Coal and Natural Gas for Producing Dimethyl Ether. Ind Eng Chem Res 2009; 48: 4101-8. [12] Hla SS, Park D, Duffy GJ, Edwards JH, Roberts DG, Ilyushechkin A, Morpeth LD, Nguyen T. Kinetics of high-temperature water-gas shift reaction over two iron-based commercial catalysts using simulated coal-derived syngases. Chem Eng J 2009; 146: 148-54.
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10
Synchronous Fluorimetric Characterization of Preasphaltene and Asphaltene from Direct Liquefaction of Coal Zhicai Wang, Cheng Wei, Hengfu Shui, Zushan Wang, Chunxiu Pan, Shibiao Ren and Zhiping Lei School of Chemistry and Chemical Engineering, Anhui Key laboratory of Clean Coal Conversion & Utilization, Anhui University of Technology, 243002 Ma’anshan, China
Abstract: In order to explore the molecular structure of coal and study the mechanism of coal direct liquefaction, the heavy intermediates of coal liquefaction such as asphaltene and preasphaltene were separated into a series of sub-fractions with different solvents as eluant by column chromatography. The aromatic ring scale distribution of sub-fractions was determined by synchronous fluorescence combined with FTIR, GPC and elemental analysis. The results indicated that the preasphaltene mainly contains lots of single ring and a certain amount of multi-rings (more than 3) fused aromatic structures, but asphaltene show abundant 2-3 rings fused aromatic structures. Aromatic structures with fused rings exist mainly in the weak polar fractions eluted by toluene and tetrahydrofuran mixed solvent. The polar fractions contain a lot of single ring aromatic nucleus and a few of 2 and 3 rings aromatic structures with abundant functional groups containing oxygen. The distribution of aromatic nucleus scale and the linkage formation of aromatic nucleus are main structure differences between asphaltene and preasphaltene. Keywords: Synchronous Fluorescence, Asphaltene, Preasphaltene, Structural Characterization Introduction Preasphaltenes (PA) and asphaltenes (AS) are not only very important heavy intermediates of direct liquefaction of coal, but also remain partially in the liquefied products. As intermediates, their reactivities directly influence on the yield of liquefied oil. As heavy components in the liquefied raw oil, they pose many difficulties in hydro-refining process such as the deposition on the catalyst to result in the deactivation of catalyst. In addition, they are also easy to plugging of lines and equipment through the sedimentation and/or coking on the devices. Therefore, the investigations of PA and AS compositions and structures are an absolute essential, especially in China. Since PA and AS have very complex compositions and poor solubility in organic solvent, their identifications chemical are near impossible. In general, the characterizations of structural feature by modern analytical techniques have been extensively carried out such as the analysis of functional group, element constitution, etc. Kanda et al. [1] suggest that one molecule of original AS consists of 1 to 3 structural unit on average, in which there are 2 or 3 aromatic rings with 1 or 2 naphthenic rings, 0 or 1 alkyl carbon atoms, 0 or 1 OH groups and 0 or 1 hetero-rings.
The aromatic structural features of PA and AS such as the size of aromatic ring and the degree of condensation are important for their upgrading and stability. Ultraviolet spectrometry (UV) and fluorescence spectrometry (FL) are always used to determine the aromatic system in raw oil and coal derivates [2-4]. However, the PA and AS all show very widen peak with no structural characteristic in their UV spectra before fine separation [4]. Similarly, the application of FL spectrometry is also limited for PA and AS because they contain very complex phosphors and energy transfers [2]. In addition, 1H-NMR and 13C-NMR can determine the contents of aromatic hydrogen and carbon but they can not give the distribution of aromatic ring number. For example, Seshadri K S et al. suggested by the characterizations of 13C-NMR and FTIR that AS and PA separated from the products of coal liquefaction are ‘oligomeric’ in structure, with aromatic clusters linked by carbon bridges with different functional groups [5]. Compared to conventional fluorescence at fixed monitoring wavelength, synchronous fluorescence spectrometry (SFL) shows a much higher spectral resolution of the complex mixtures by selecting the difference of wavelength according to the sample structure [6]. Katoh et al. thought that the potential usefulness of SFL as a method of analysis of complex mixtures as coal-derived liquids despite the limitation of the method that some molecules give only weak peaks [7]. Kister J et al. determined quantitatively the polyaromatic hydrocarbons (PAHs) in extractables of coal by SFL, in which three main synchronous fluorescence spectral regions was fined according to the number of lineally condensed aromatic rings [8]. Kershaw et al. found that the tar from the pyrolysis of Loy Yang coal shows two characteristic peaks in the constant-energy synchronous spectrum, and suggested that constant-energy SFL is a valuable method for monitoring the concentration of the larger aromatic ring systems in pyrolysis tars [2]. Li et al. found that the pyridine-solubles of coal consist of 3-5 aromatic rings by the SFL at a constant wavelength difference of 14 nm [9]. In our previous works, the fluorescence spectroscopy had been used to characterize the structures and aggregations of AS and PA from the liquefaction products of coal, respectively [10, 11]. In order to determine the scale distribution of fused aromatic nucleus (FAN) in the heavy liquefied products, the sub-fractions separated from AS and PA were respectively analyzed by the SFL. Experiments Preparation and fraction of AS and PA Table 1 Liquefaction yields and results of PA and AS elemental analysis Fractions
PA
AS
Parent coal
Yield/%
SF
Wdaf /%
H/C
O/C
C
H
N
S
O*
22.0
79.16
7.84
1.41
0.56
11.03
1.19
0.10
SL
17.2
77.91
6.31
1.72
1.10
12.96
0.97
0.12
SF
14.0
80.61
6.65
1.58
0.60
10.56
0.99
0.10
SL
11.0
73.74
7.09
1.84
0.85
16.48
1.15
0.17
* by difference.
AS and PA were prepared by the soxhlet extraction of coal liquefaction products. In this work, two Chinese coals including Shenfu (SF) sub-bituminous coal and Shengli lignite (SL), were hydro-liquefied to prepare corresponding AS and PA in a batch autoclave at 400 oC, 5 MPa (initial pressure) H2 in tetralin solvent for 30 min. Liquefaction yields and results of PA and AS elemental analyses were listed in Table 1. A detailed description can be found elsewhere [12]. AS and PA were further fractionated by column chromatography, in which 200 mesh silica was used as adsorbent, and toluene (T), mixing solvents of THF/T with different ratios, THF, methanol were used as eluents, respectively. In a typical run, the mixture of 0.5 g sample and 1 g silica was placed on the top of the column, and the column was eluted sequentially with the following solvents: T (only used in the separation of AS), T/THF (4:1, volume), T/THF (1:2, volume), THF and methanol. After eluting the column with a particular solvent, the solvent was removed by rotary evaporation and drying on vacuum oven at 80 oC. The sub-fractions, which defined sequentially as A1, A2, A3, A4 and A5 for AS, and P1, P2, P3 and P4 for PA, were weighted and used to analysis. Characterization of sub-fractions The element analysis was carried out in Elementar Vario EL III, and the FTIR spectrum was determined using a Nicolet 6790 IR spectrometer at ambient temperature by KBr. SFL spectra and GPC curves of all sub-fraction samples were determined in their THF solution. In general, sub-fraction sample was dissolved in THF, and the solution was left in ultrasonic bath for 30 min and subsequently was standing overnight to assure complete dissolution. Finally, the solution was diluted with THF to 5 mg/l. The SFL spectrum was recorded on a Hitachi F-4600 spectrophotometer with 150 W Xenon lamp as the excitation source. A slit width of 5 nm was selected. The spectrum was scanned from 200 nm to 650 nm at a speed of 240 nm/min and a constant wavelength difference of 17 nm. A Shimadzu SPD-20A system, with LC-20 AT UV/Vis detector, was used to record the GPC curve at 254 nm. The GPC columns (8 mm×300 mm, 10 μm) was a Shim-pack GPC-8025. Ultra-pure THF (HPLC grade) was used as eluent at 1 ml/min flow rate, and the injection volume was 20 μL. All graphs were plotted using Origin 7.0 software. Results and discussion AS separation and its sub-fractions characterization Table 2 shows that the AS from SL lignite can be separated into four sub-fractions, including A1, A3, A4 and A5 in turn, but that from SF coal only has two sub-fractions obtained as A2 and A5. A1 and A2 are the major sub-fraction of AS from SL lignite and SF coal, respectively, and A3 and A5 all are lower than 7 %. It is suggested that the separation properties of AS varied with the parent coal, and the weak polar sub-fractions eluted by toluene or its mixed solvent are much more than the polar ones.
Meanwhile, Table 2 also shows that the Odaf% of sub-fraction distinctly increased but the Cdaf% decreased with increasing of polarity of eluant for two AS. The atomic ratios of H/C and O/C of AS sub-fractions from SL coal are significant higher than those from SF. It indicates that the element compositions of AS sub-fractions vary with their parent coals. Although the retention of sub-fraction from same AS becomes strong with the increasing O/C, but the retention of A2 from SF coal (O/C=0.09) is stronger than that of A1 from SL lignite (O/C=0.15), suggesting that the polarity is not the only factor to influence the separation property of AS. Table 2 Results of AS separation and their sub-fractions element analyses Fractions
Parent coal
Yield/%
A1
SL
A2
Wdaf /%
H/C
O/C
14.73
1.13
0.15
0.58
9.90
0.99
0.09
1.61
0.88
14.66
1.20
0.15
7.50
1.12
1.21
17.14
1.23
0.18
70.89
6.20
1.67
0.98
20.26
1.05
0.21
55.61
5.65
1.86
1.70
35.18
1.22
0.47
C
H
N
S
O*
72.1
75.47
7.09
2.03
0.68
SF
93.7
81.26
6.68
1.58
A3
SL
3.5
75.32
7.53
A4
SL
17.9
73.03
SF
6.3
SL
6.5
A5 * by difference. SF
SL A2
A1
T/%
T/%
A3
A5
A4 A5
4000
3500
3000
2500
2000
1500
Wavenumber / cm
1000
4000
500
3500
3000
2500
2000
Wavenumber / cm
-1
1500
1000
500
-1
Fig.1 FTIR spectra of AS sub-fractions from SF coal and SL lignite respectively 0.20
SF
0.20
0.15
Intensity a.u.
0.15
Intensity a.u.
SL
0.10
0.10
A1 A3
A2
0.05
0.05
A4 A5
A5 0.00
0.00 4
6
8
Retention time / min
10
12
4
6
8
10
12
Retent time / min
Fig.2 GPC curves of AS sub-fractions from SF coal and SL lignite respectively
Fig.1 compares respectively the FTIR spectra of AS sub-fractions from SL
lignite and SF coal. The area of hydroxyl absorption peak at the range of 3700~2500 cm-1 basically increases with the O/C of sub-fractions, suggesting that most of oxygen in AS consist in the hydroxyl function group. Meanwhile, the structure and content of carbonyl are different in the AS sub-fractions. A1 and A2 show mainly the carbonyl of diarylketone and quinine, but the carbonyl in sub-fractions A3~A5 consists in the formation of ketone and ester. Further, Fig.2 shows that the GPC curves of sub-fractions from different AS, respectively. In sub-fractions from SL lignite, the retention time of A1, A5, A4 and A3 decrease in turn. Although two sub-fractions from SF lignite all do not displayed normal peak, in which includes a strong solvent peak, but it can still be observed that the retention time of A2 is more than that of A5, which is similar to those sub-fraction from SL lignite. It may be speculated that the major sub-fraction A1(SL) or A2(SF) has a small effective scale, but A3 and A4 have a large effective scale in THF solvent. Meanwhile, the peak intensity of different sub-fractions in normalized GPC curves is different, suggesting that the chromophore content varies with the sub-fractions because of the ultraviolet detector (254 nm) used in the detection of GPC. A5 of two AS show the smallest content of chromophore compared to other sub-fractions. 0.12
0.12
SF
285.6
SL
285.8
0.10
Relative Intensity
Relative Intensity
0.10
0.08
0.06
362.6
0.04
A2
339.2 366.0
0.02
0.08
357.6 332.4
0.06
332.4 364.0
A5 A3
0.02
A5
A4
0.00 200
A1
0.04
0.00 300
400
Wavelength / nm
500
600
200
300
400
500
600
Wavelength / nm
Fig.3 SFL spectra of AS sub-fractions from SF coal and SL lignite respectively
In the SFL spectra of AS sub-fractions shown in Fig. 3, all sub-fractions show a multiple peaks SFL spectrum in the range between 250 nm and 550 nm. The major sub-fractions A1(SL) and A2(SF) show respectively the strongest SFL peak at 357.6 nm and 362.6 nm, and other minor sub-fractions display different SFL spectra with the major sub-fraction, in which the strongest peak all appear in 285.6 and 285.9 nm. Although A1 and A5 also show a strong peak near to 285 nm, but A1(SL) shows a strong shoulder peak at 332.4 nm and lower SFL strength than A2(SF) above 400 nm. Except the peak at 286 nm, other two peaks can be observed in A3~5 sub-fractions SFL spectra at 332.4 (339.2 SF) nm and 364.0 (366.0 SF) nm, respectively. The wavelength of corresponding SFL peaks of SF sub-fractions is slightly more than those of SL sub-fraction. According to the assignment reported by Kister et al. [8, 13], the range less than 300 nm can be attributed to the aromatic nucleus only with one aromatic ring, such as alkylbenzene. The spectral region between 300 nm and 340 nm can be assigned to alternant aromatic nucleus with two fused rings. Similarly, the spectral ranges of 340 nm-400 nm and more than 400 nm corresponded respectively to the aromatic nucleus
with three and more fused rings. Compared with the PAHs, side chains attached to aromatic rings and aromatic heterocycles containing N, S or O can produce a bathochromic effect [8]. Therefore, the fluorophor of AS sub-fractions consist mainly of 1~3 rings aromatic nucleus. The content of 3 rings FAN in the major sub-fractions A1(SL) and A2(SF) is the highest, and the content of single rings aromatic nucleus in the minor sub-fractions A3~5 is the highest, too. Sub-fraction A1(SL) contains more double rings FAN than sub-fraction A2(SF), but latter contains more 4 rings FAN than former. For minor sub-fractions, the content of double-rings FAN is more than that of 3 rings FAN. Based on above characterization results of AS sub-fractions, it is suggested that the aromatic structures of AS consist of 1~4 rings nucleus, in which single-ring aromatic nucleus mainly consist in polar sub-fractions A3~5 and FAN (2~3 rings) consist in weak polar sub-fraction A1 or A2. The AS from SF coal contains a certain amount of 4 rings FAN compared to that from SL lignite. Further, it may be the self-aggregation and/or the solvation of polar sub-fractions that A3~5 with small aromatic ring show larger effective scale than A1 or A2 with FAN structure. The influence of aggregation on the GPC of PA result has also been found in our previous work [11]. PA separation and its sub-fractions characterization Table 3 shows that all PA from SL lignite and SF coal can be separated into P1~P4 four sub-fractions. The content of sub-fraction P2 is the highest in the SL PA but that of sub-fraction P1 is the highest in the SF PA. Relative to major sub-fraction P1 and P2, the content of P3 and P4 are low, especially the content of P4 only is loss than 8 %. It suggests that the content of polar sub-fractions of SL PA is obviously more than that of SF PA. Meanwhile, Table 3 also shows that the Odaf% of PA sub-fractions increase basically with the increasing polarity of eluant. P4 shows the lowest Cdaf% and the highest Odaf% than other sub-fractions and P2 shows the lowest H/C. The H/C of PA from SF coal and its major sub-fraction were obviously higher than those from SL lignite, but their O/C are on the contrary. However, compared with the corresponding sub-fractions of AS, the sub-fractions of PA from SF coals show higher H/C and O/C, but those from SL lignite are on the contrary. Table 4 Results of PA separation and their sub-fractions element analysis Fractions
P1
P2
P3 P4
Parent coal
Yield/%
Wdaf /% C
H
N
S
O*
H/C
O/C
SF
52.2
79.28
8.03
1.36
0.65
10.68
1.22
0.10
SL
24.2
80.73
6.90
1.43
0.97
9.97
1.03
0.09
SF
31.1
79.32
6.45
1.73
0.46
12.04
0.98
0.11
SL
53.2
78.31
6.09
1.96
1.06
12.58
0.93
0.12
SF
13.7
83.27
10.45
0.76
0.36
5.16
1.51
0.05
SL
15.0
76.63
6.50
1.27
1.19
14.41
1.02
0.14
SF
3.0
56.90
7.06
1.88
1.02
33.14
1.49
0.44
SL
7.6
68.70
5.57
1.88
1.53
22.32
0.97
0.24
* by difference.
Fig.4 shows the FTIR spectra of PA sub-fractions from SL lignite and SF coal. For SF PA sub-fractions, the peak areas of hydroxyl and aliphatic C-H stretch vibrations respectively vary directly with O/C and H/C except for P3. Meanwhile, P1 displays a carbonyl shoulder peak of aryl ketone at 1640 cm-1, and P2 has an obvious ester carbonyl shoulder peak at 1715 cm-1. For SL PA sub-fractions, the changes of the peak areas of hydroxyl and aliphatic C-H stretch vibrations are agreement with their O/C and H/C, but there is no obvious proportion to be observed. The aryl ketone is the principle formation of carbonyl in all SL PA sub-fractions, and a certain amount of ester carbonyl can also been found in P3 and P4. It suggests that the structure differences between PA sub-fractions from SF coal are more than those from SL lignite. In the major sub-fraction P1(SF) and P2(SL), the carbonyl groups consist of aryl-ketone with a little of ester carbonyl. Further, Fig.5 shows that the GPC curves of sub-fractions from different PA, respectively. Although the retention times of different sub-fractions from two PA decrease following P1, P4, P2 and P3, but all sub-fractions from SF coal have shorter retention time than corresponding sub-fractions from SL lignite. Compared with AS sub-fractions, the PA sub-fractions from SF coal display shorter retention time but the PA sub-fractions from SL lignite display similar retention time. It indicates that the PA sub-fractions from SF coal are larger effective scale than their AS sub-fractions, but the PA sub-fractions from SL lignite have similar effective scales to their AS sub-fractions. SF P1
P1
P2
P2 T/%
T/%
SL
P3
P3
P4
4000
3500
3000
2500
P4
2000
1500
Wavenumber / cm
1000
500
4000
3500
3000
2500
-1
2000
Wavenumber / cm
-1
1500
1000
500
Fig.4 FTIR spectra of PA sub-fractions from SF coal and SL lignite respectively 0.20
0.20
SF
0.15
Intensity a.u.
0.15
Intensity a.u.
SL
0.10
0.10
P1
P1
P2
P3
0.05
P2
0.05
P4
P3
P4
0.00
0.00 4
6
8
Retention time / min
10
12
4
6
8
10
Retention time / min
Fig.5 GPC curves of PA sub-fractions from SF coal and SL lignite respectively
12
0.14
0.14
SL
285.2
0.12
0.12
0.10
0.10
Relative Intensity
Relative Intensity
SF
0.08
P1 P2 P3 P4
0.06
0.04
285.0
P1 P2 P4 P3
0.08
0.06
369.0 0.04
335.8
370.4 334.8 368.6
0.02
0.02
0.00 200
0.00
300
400
Wavelength / nm
500
600
200
300
400
500
600
Wavelength / nm
Fig.6 SFL spectra of PA sub-fractions from SF coal and SL lignite respectively
Fig.6 shows the SFL spectra of PA sub-fractions. Compared with the SFL spectra of AS sub-fractions shown in Fig.3, the strength of peak at 285 nm relative to other fluorescence band distinctly increases though all PA sub-fractions show also a multiple peaks SFL spectrum in the range between 250 nm and 550 nm. Meanwhile, the strength of fluorescence band above 400 nm relative to the spectrum band in the range of 300~400 nm also increase, especially for the P1 from SF coal. In addition, two weak peaks near to 335 nm and 370 nm can also be observed in other minor PA sub-fractions similar to the AS sub-fractions. According to above assignment of fluorescence peak, all PA sub-fractions contain 1~4 rings FAN, in which the content of single ring aromatic nucleus is the highest. For PA sub-fractions from SF coal, beside single ring aromatic nucleus, P1 consists from a certain amount of 2~5 rings FAN, in which the contents of 3~4 rings FAN are more than those of 2 and 5 rings FAN. P2, P3 and P4 show similar distribution of aromatic nucleus, in which 2 and 3 rings FAN can only be observed besides single ring aromatic nucleus. For PA sub-fractions from SL lignite, P1 and P2 have a similar distribution of aromatic nucleus to P1 from SF coal, and P3 and P4 are like as the P2~P4 from SF coal too. However, the differences between P1(SF) and P1(SL) or P2(SL) are that former contains more single ring and 4~5 rings FAN, and less 2 and 3 rings FAN than the latter. Based above SFL characterization results, it can be obtained that highly fused aromatic ring structures mainly consist in the weak polar sub-fraction, such as A1(SL), A2(SF) and P1, and the PA and AS from SF coal show more multi-ring FAN than those from SL lignite. Furthermore, compared with the corresponding AS sub-fractions, the PA sub-fractions show more single ring structure and multi-rings (more than 3) FAN, suggesting that the single ring aromatic structure in PA was linked with FAN by weak bridge bonds, such as ether bond, methylene etc. In the liquefaction process, the single ring fragment is hydro-cracked into oil and FAN fragment is partially hydrogenated into AS. Conclusion Different sub-fractions were respectively separated from PA and AS liquefied from two Chinese coals by column chromatography. The results of element analysis, FTIR, GPC and SFL of their sub-fractions suggest that the differences between PA and AS are the distribution of aromatic nucleus scale and the linkage formation of
aromatic nucleus. Due to the salvation of THF, it is difficult to determine the real scale of PA and AS sub-fraction by GPC with THF as eluent. Acknowledgement This study was supported by the National Natural Science Foundation of China (20776001, 20876001, 20936007). The authors gratefully acknowledge the International Cooperative Project of Anhui Province (07080703001). The authors also appreciated the financial support of innovative group of Anhui Province “Coal Resource Processing & Cleaning Utilization”. Reference [1] Kanda N, Itoh H, Yokoyama S, Ouchi K. Mechanism of hydrogenation of ocal-derived asphaltene. Fuel 1978; 57:676-80. [2] Kershaw J R, Sathe C, Hayashi J-i, Li C-Z and Chiba T. Fluorescence spectroscopic analysis of tars from the pyrolysis of a Victorian brown coal in a wire-mesh reactor. Energy Fuels 2000; 14:476-82. [3] Goncalves S, Castillo J, Fernandez A, Hung J. Absorbance and fluorescence spectroscopy on the aggregation behavior of asphaltene-toluene solutions. Fuel 2004; 83:1823-8. [4] Ghosh A K, Srivastava S K, Bagchi S. Study of self-aggregation of coal derived asphaltene in organic solvents: A fluorescence approach. Fuel 2007; 86:2528-34. [5] Seshadri K S, Young D C and Cronauer D C. Characterization of coal liquids by 13C-NMR and FTIR spectroscopy-fractions of oils of SRC-I and asphaltenes and preasphaltenes of SRC-I and SRC-II. Fuel 1985; 64:22-8. [6] Mckay J F, Latham D R. Fluorescence spectrometry in the characterization of high-boiling petroleum distillates. Analytical Chemistry 1972; 44:2132-6. [7] Katoh T, Yokoyama S and Sanada Y. Analysis of a coal-derived liquid using high-pressure liquid chromatography and synchronous fluorescence spectrometry. Fuel 1980; 59:845-50. [8] Kister J, Pieri N, Alvarez R, Diez M A and Pis J J. Effects of preheating and oxidation on two bituminous coals assessed by synchronous UV fluorescence and FTIR spectroscopy. Energy Fuels 1996; 10:948-57. [9] Li W, Ye C, Feng J, Xie K. Influence of column chromatography and soxhlet extraction on the composition of coal pyridine-soluble. Chinese Journal of Analytical Chemistry 2006; 34:905-9. [10] Wang Z, Cui X, Shui H, Wang Z, Lei Z, Kang S. Fluorescence spectroscopy characterization of asphaltene liquefied from coal and study of its association structure. Spectroscopy and Spectral Analysis 2010; 30:1530-4 (in Chinese). [11] Wang Z, Li L, Shui H, Wang Z, Cui X, Ren S, Lei Z, Kang S. Study on the aggregation of coal liquefied preasphaltene in organic solvents by UV-vis and fluorescence spectrophotometry. Fuel 2011; 90:305-11. [12] Wang Z; Shui H; Zhang D; Gao J. A comparison of FeS, FeS+S and solid Superacid catalytic properties for coal hydro-liquefaction. Fuel 2007; 86, 835-42. [13] Benkhedda Z, Landais P, Kister J, Dereppe J-M and Monthioux M. Spectroscopic analyses of aromatic hydrocarbons extracted from naturally and artificially matured coals. Energy Fuels 1992; 6:166-72.
Oviedo ICCS&T 2011.
Jurassic Perhydrous Coals from the Lusitanian Basin, Portugal. Petrography, geochemical and textural characteristics. A. Costa1, C. Tomás2, I. Suárez-Ruiz3, P.P. Cunha2,4, D. Flores1,5, B. Ruiz3 1
Centro de Geologia, Faculdade de Ciências, Universidade do Porto, Rua do Campo Alegre, 687, 4169-007 Porto, Portugal.
[email protected] 2 Departamento de Ciências da Terra, Faculdade de Ciências e Tecnologia da Universidade de Coimbra, 3000-272 Coimbra, Portugal 3 Instituto Nacional del Carbón, (INCAR-CSIC), Ap. Co. 73, 33080- Oviedo, Spain. 4 IMAR – Marine and Environmental Research Centre; Grupo de Investigação em Sistemas Sedimentares, Hidrodinâmicos e Transformações Globais, Coimbra, Portugal 5 Departamento de Geociências, Ambiente e Ordenamento de Território e Centro de Geologia, Faculdade de Ciências, Universidade do Porto, Rua do Campo Alegre, 687, 4169-007 Porto, Portugal. Abstract The petrographic, geochemical and textural characteristics of Upper Jurassic perhydrous coals from the Lusitanian Basin, Portugal, were studied in order to obtain information about this coal which is little known in Portugal. The coal samples were taken from different Formations and Members of this basin ranging from the Oxfordian to the Titonian age. During the Upper Jurassic, the Lusitanian Basin underwent an extensional episode, which led to significant changes in its paleogeography, including variations in sedimentary environments. The six selected coals are dark with a homogeneous appearance. They do not have a clear vegetable structure. They are hard, compact, bright black with a glassy structure and show a conchoidal fracture with straight and sharp edges. Huminite is the main organic component and reflectance lower than expected ("suppressed reflectance"). The mineral matter content of these coals is very low; and their carbon and hydrogen contents are similar and higher to the non-perhydrous coals, respectively. They have relatively low porosity and pore volume.
Keywords: Jurassic Perhydrous coal, Reflectance, FTIR, Textural properties
1
Oviedo ICCS&T 2011.
Introduction Perhydrous coals are anomalously enriched in hydrogen by natural causes (Stach et al., 1982; Taylor et al., 1998, among others) and are characterized by discrepancies in their chemical and petrographic composition. Several authors (Newman and Newman, 1982; Gentzis and Goodarzi, 1994; Gurba and Ward, 1998; Petersen and Rosenberg, 1998) consider that hydrogen enrichment is caused by the organic matter deposition in a sedimentary palaeoenvironment where anaerobic conditions prevail. However, other authors (Teichmüller and Teichmüller, 1982; Diessel, 1992; Diessel and Gammidge, 1998) pointed out that the perhydrous coals can be originated by the presence of liptic material resulting from the bacterial degradation of the vitrinite precursor. The Jurassic Formations from Portugal have small interbedded strata of coals, which are known to show perhydrous characteristics. Taking into account that, the objective of this work is to determine the physical, chemical and textural properties and their influence on the petrographic composition. For this propose, six coals, representing different Formations and Members of the Jurassic of the Lusitanian Basin, ranging from the Oxfordian to the Titonian age were selected for this study. During the Upper Jurassic, the Lusitanian Basin underwent an extensional episode, which led to significant changes in its paleogeography, including variations in sedimentary environments.
Analytical procedures Petrographic analyses of the seven perhydrous coals were carried out on polished blocks of whole rock prepared according to standard procedures (ISO 7404-2, 2009). The petrographic characterization was performed following the ISO 7404-5 (2009) standard and the ICCP nomenclature (ICCP, 1971, 2001; Sýkorová et al., 2005) was used. Proximate analyses were carried out in accordance with ISO 1171 (1997); ISO 589 (2008); ISO 562 (1998) standards and ultimate analyses were determined on a LECO S2000 and LECO S-632 for CHN and total sulphur, respectively. SEM observations were made in a “FEI Quanta 400FEG / EDAX Genesis X4M” microscope. FTIR analyses were performed in a DRIFTS – Diffuse Reflectance Infrared Fourier Transform Spectroscopy technique, using a Nicolet 8700 FTIR Thermo Scientific equipment and a DTGS-TES detector. For this technique samples were previously ground to 63μm and spectra were acquired doing 200 scans between 4000 and 450 cm−1 frequency at a resolution of 4 cm−1. Textural characterization was carried out by measuring the N2 2
Oviedo ICCS&T 2011.
adsorption isotherms at -196ºC in an automatic apparatus (Micomeritics ASAP 2420). The isotherms were used to calculate the specific surface area SBET, and total pore volume, VTOT, at a relative pressure of 0.95.
Results and Discussion The six selected coals (LJS10, LJS12, LJS13, BALJS2, BJV and CMJS2) are dark with a homogeneous appearance. They do not have a clear vegetable structure. They are hard, compact, bright black with a glassy structure and show a conchoidal fracture with straight and sharp edges. Microscopically, the maceral group huminite is the main organic component of these coals and their mineral matter content is very low. Inside the huminite, ulminite is the predominant component (93.2 to 99.4 % vol., mmf), followed by corpohuminite (0.2 to 6.6 %, vol., mmf) and textinite with well-defined cellular walls and in some cases this cavities filled with resinite resinite (0.2 to 4.0 %, vol., mmf). In some cases, the presence of fluorescent globules in the ulminite is noted, due to the interactions of hydrocarbons. The occurrence of hydrocarbons was visible by fluorescence in the form of oil drops. (Fig. 1 A,1B). The random reflectance of ulminite and corpohuminite ranges between 0.16 - 0.34% and 0.31 – 0.38%, respectively. A
B
50
50
Figure 1. A- Ulminite and corpohuminite; B- same field in fluorescent mode, showing a fracture infilled by hydrocarbons. SEM examination shows mainly an homogeneous material, made up of an amorphous matrix occasionally porous with lumens corresponding to a well-preserved cellular structure. Epigenetic minerals such as pyrite, dolomite and calcite were also identified. The content in ash of these coals is low (2.57 to 9.84%, dry basis) The FTIR spectra, show the predominance of aliphatic structures over aromatic ones (Fig.2) which agrees with
the high hydrogen content (5.41 to 6.40%,) and the
corresponding high H/C atomic ratio.
3
Oviedo ICCS&T 2011. CH aliphatic CH aromatic
Con formato: Fuente: 8 pt, Color de fuente: Negro Con formato: Fuente: 8 pt, Color de fuente: Negro, Portugués (Portugal) Con formato: Fuente: 8 pt, Color de fuente: Negro Con formato: Fuente: 8 pt, Color de fuente: Negro, Portugués (Portugal)
Figure 2. FTIR spectra where is visible the predominance of aliphatic structures over aromatic ones. The real density varies between 1.26 to 1.37g/cm3, which makes these coals less dense than non-perhydrous coals with a similar carbon content. All the samples have a low porosity (3.9 to 12.7 %) and the total pore volume ranges between 1 to 11mm3/g, which suggest that the generated hydrocarbons might be trapped in the porosity of the ulminite. The textural characterization shows that Lusitanian coals do not exhibit the same properties as those of the commercial coals.
Conclusions The seven Jurassic perhydrous coals from Portugal exhibit anomalous physical, chemical and textural properties when compared to non-perhydrous coals. These properties include: (i) a lower huminite/vitrinite reflectance than expected ("suppressed reflectance") and lower density values than those of non-perhydrous coals with a similar carbon content; (ii) a higher hydrogen content; (iii) a relatively low porosity and pore volume.
Acknowledgements The current work was prepared under a PhD scholarship Ref: SRH/BD/45564/2008 granted by FCT, from which the first author benefits.
References
4
Oviedo ICCS&T 2011. Cuesta, M.J., 2004. Caracterización de vitrinitas perhidrogenadas. Influencia del enriquecimiento en hidrógeno en su comportamiento durante la evolución térmica y la oxidación. PhD Thesis. Universidad de Oviedo. Diessel, C.F.K., 1992. Coal-Bearing Depositional Systems. Springer-Verlag. Diessel, C.F.K., Gammidge, L., 1998. Isometamorphic variations in the reflectance and fluorescence of vitrinite — a key to depositional environment. International Journal of Coal Geology 36, 167–222. Gentzis, T., Goodarzi, F., 1994. Reflectance suppression in some Cretaceous coals from Alberta, Canada. In: Mukhopadhyay, P.K., Dow, W.G. (Eds.), Vitrinite Reflectance as a Maturity Parameter. Applications and Limitations. : American Chemical Society Symposium Series, 6. ACS Books, pp. 93–110. Gurba, L.W., Ward, C.R., 1998. Vitrinite reflectance anomalies in high-volatile bituminous coals of the Gunnedah basin, New South Wales Australia. International Journal of Coal Geology 36, 111–140 ICCP, 2001. The new inertinite classification (ICCP system 1994). Fuel 80, 459–471. International Handbook of Coal Petrography, Supplement to the 2nd Ed., Centre National de la Recherche Scientifique, Academy of Sciences of the USSR. Paris, Moscow. 1971. ISO 1171, 1997. Solid mineral fuels — determination of ash. International Organization for Standardization, Geneva, Switzerland. 4 pp. ISO 562, 1998. Hard Coal and Coke — Determination of Volatile Matter. International Organization for Standardization, Geneva, Switzerland. 7 pp. ISO 589, 2008. Hard Coal — Determination of Total Moisture. International Organization for Standardization, Geneva, Switzerland. 9 pp. ISO 7404-2, 2009. Methods for the petrographic analysis of coals — Part 2: Methods of preparing coal samples. International Organization for Standardization, Geneva, Switzerland. 12 pp. ISO 7404-5, 2009. Methods for the petrographic analysis of coal — Part 5: Methods of determining microscopically the reflectance of vitrinite. International Organization for Standardization, Geneva, Switzerland. 14pp. Jiménez, A. 1995. Estudio del grupo vitrinita en carbones de distinto rango. Determinacion de sus propriedades y relationes com su génesis. Tesis Doctoral. Universidad de Salamanca. 255pp. Jiménez, A., Iglesias, M.J., Laggoun-Défarge, F., Suárez-Ruiz, I., 1998. Study of physical and chemical properties of vitrinites. Inferences on depositional and coalification controls. Chemical Geology 150, 197–221. Külaots, I., Gao, Y.M., Hurt, R.H., Suuberg, E.M., 1998. The Role of Polar Surface Area and Mesoporosity in Adsorption of Organics by Fly Ash Carbon. ACS Division of Fuel Chemistry Preprints, 43, 980. Külaots, I, Hurt, R.H., Suuberg, E.M., 2004. Size distribution of unburned carbon in coal fly ash and its implications. Fuel 83, 223–230. Newman, J., Newman, N.A., 1982. Reflectance anomalies in Pike River coals: evidence of variability in vitrinite type, with implications for maturation studies and “Suggate rank”. New Zealand Journal of Geology and Geophysics 25, 233–243. Petersen, H.I., Rosenberg, P., 1998. Reflectance retardation (suppression) and source rock properties related to hydrogen-enriched vitrinite in Middle Jurassic coals, Danish North Sea. Journal of Petroleum Geology 21, 247–263.
5
Oviedo ICCS&T 2011. Sykorova, I., Pickel, W., Christanis, K., Wolf, M., Taylor, G.H., Flores, D. 2005. Classification of huminite-ICCP System 1994. International Journal of Coal Geology 62, 85-106. Stach, E., Mackowsky, M.-T.h., Teichmüller, M., Taylor, G.H., Chandra, D., Teichmüller, R., 1982. Stach's Textbook of Coal Petrology, 3 rd Ed. Gebrüder Borntraeger, Berlin, Stuttgart.
Suárez-Ruiz, I., Jiménez, A., Iglesias, M.J., Cuesta, M.J., Laggoun-Defarge, F. 2006. El azabache de Asturias. Características físico-químicas, propiedades y génesis, Trabalhos de Geologia. Universidade de Oviedo, 26, 9-18. Taylor, G.H., Teichmüller, M., Davis ,A., Dissel, C. F. K., Littke, R., Robert,P.,1998. Organic petrology. A new handbook incorporating somerevised parts of Stach's Textbook of Coal Petrology, 1rd Ed. Gebrüder Borntraeger, Berlin, Stuttgart, 704pp. Teichmüller, M., Teichmüller, R., 1982. Fundamentals of coal petrology, In: Stach, E., Mackowsky, M.-T.h., Teichmüller, M., Taylor, G.H., Chandra, D., Teichmüller, R. (Eds.), Stach's Textbook of Coal Petrology, 3rd Ed. Gebrüder Borntraeger, Berlin, Stuttgart, pp. 5–83.
6
Rank distribution of the Cretaceous coals in the region of San Juan de Sabinas, Coahuila de Zaragoza, Mexico. N. Piedad-Sánchez1, I. Suárez-Ruiz2, F. R. Carrillo-Pedroza1, J. A. MorenoHirashi3, G. De la Rosa-Rodríguez3, K. Flores-Castro4, R. Corona-Esquivel5, 6, J. L. Cadena-Zamudio4, J. O. Navarro-Lozano4, B. Santiago-Carrasco3, F. González-Carrillo1. 1.Tecnología e Ingeniería de Materiales, DES Ciencias Extractivas, Unidad Norte, Universidad Autónoma de Coahuila, Monclova, Coahuila de Zaragoza.
[email protected],
[email protected]. 2. INCARCSIC, Ap. Co. 73, 33080 Oviedo, España. 3. Dirección Minerales Energéticos. Servicio Geológico Mexicano. Blvd. Felipe Ángeles km. 93.50-4, Col. Venta Prieta, C.P. 42080, Pachuca, Hidalgo, México. 4. Cuerpo Académico de Ciencias de la Tierra, Universidad Autónoma del Estado de Hidalgo, Pachuca, Hgo., México. 5. Unidad Ticomán, Escuela Superior de Ingeniería y Arquitectura, Instituto Politécnico Nacional. Ticomán, México, D.F., México. 6. Instituto de Geología, Universidad Nacional Autónoma de México. Ciudad Universitaria, México, D.F., México. ABSTRACT The Nueva Rosita coal-bearing region (Cretaceous age) is located in the Sabinas Sub-basin syncline (North of Mexico). The studied area presents a complex sequence of terrigenous layers of the Upper Cretaceous age, including interlayered coal beds. This work discusses the causes and events that led to the different evolution and the varied quality of Cretaceous coal in this region. Coal samples were collected from underground and open-pit mines to obtain data about the rank and composition of the coal by means of petrographic techniques. The maceral composition of the coals studied is dominated by vitrinite whose reflectance values range between 0.7 to 1.2% (medium rank, ISO 11760:2005). The vitrinite reflectance values indicate a variable thermal gradient, possibly attributable to the extrusive volcanism that outcrops near the studied area, and to the singular burial history of the Sabinas Sub-basin.
INTRODUCTION The San Juan de Sabinas coal-bearing region (Cretaceous age) is located in the Coahuila de Zaragoza state (North of Mexico), which is a part of the Sabinas Sub-basin syncline. The studied area included in this work presents a complex sequence of terrigenous layers of the Upper Cretaceous age, including
interlayered coal beds (between 2 and 5 coal beds) whose lateral facies are still not fully known. In order to gain knowledge about existing coal reserves for the electrical, iron and steel industry in Mexico, and to make up for the scarcity of detailed geologic data, a study is undertaken into the causes and events that led to the different evolution and, consequently the varied quality of Cretaceous coal in this region (Figure 1).
Figure 1. San Juan de Sabinas área in Google Earth™ image.
For this study coal samples were collected from underground and open-cast mines to obtain data about the rank and composition of the coal by means of petrographic techniques.
GEOLOGIC FRAMEWORK The rocks in the Sabinas Sub-basin to represent a complex sequence of terrigenous strata from the Upper Cretaceous, mainly in the Olmos Formation of Maestrichtian age (Robeck et al., 1956), which includes interbedded coal seams with unknown lateral relations, causing no consensus on the lithostratigraphy, as some authors like Antunano Eguiluz and Amezcua-Torres (2003) indicate that coal beds are located in the Escondido Formation of the same age. Lithologically, two main coal seams of this region are embedded between carbonaceous siltstones and shales of the so-called Olmos Formation. The thickness of the coal seams is variable (20 to 180 cm), and generally at the bottom of the stratigraphic sequence, the coal seams are divided by a refractory clay horizon (bentonite) 25 to 30 cm thick (sometimes, referred to as tonstein),
that is used as a level for lithological correlation and is the result of the alteration of a lithic tuff of dacitic composition. The study area is covered by volcanic effusions for the volcanic field Las Esperanzas (Mulleried, 1941, Valdes-Moreno, 2001). In the forties, Mulleried (1941) published studies in which he discusses the distribution and origin of the lava flows in the region of Sabinas, Muzquiz, Barroterán Mines and Las Esperanzas. From this last town was renamed Las Esperanzas Basalts by which they are known in geological literature. Mulleried (1941) indicates that the Las Esperanzas basalts overlying Reynosa Formation of Late Tertiary, and describes the volcanoes extravasated as "linear extrusions" (volcanic fissure) with some outcrops characterized by columnar joints and central extrusions (lava shields) evidenced by the presence of volcanic bombs near small topographic prominence. In addition, Mulleried (1941) makes a calculation of the area covered by basaltic lava flows (181 km²) and the volume issued (2.6 km³). Robeck et al. (1956), in a geological study of coal deposits, only cite the Las Esperanzas lavas in general. Las Esperanzas lavas are basaltic flows (sensu lato), brown on weathered surface, and dark gray to black on fresh surface. Many lavas tend to be phyric or microporphyritic with aphanitic to glassy matrix (Valdés-Moreno, 2001). The centers of emission of Las Esperanzas basalts are two different types: In the eastern, lavas were extravasated by the central-type shield volcanoes, called Agua Dulce (520 msnm), Kakanapo Grande (510 msnm), Kakanapo Chico (470 msnm), La Peña (420 msnm), andy El Barril (400 msnm); in the western, Las Esperanzas lavas were expelled along the eastern and southern edges of the anticline of the Sierra Santa Rosa as volcanic fissure, and there was no evidence of central volcanic apparatus (Valdés-Moreno, 2001). The geochemical results obtained by Valdes-Moreno (2001) suggest that the magmas, which gave rise to Las Esperanzas Volcanic Field, have a marked influence of the mantle, and with little or no contamination of continental crustal, with the combination of partial melting and fractional crystallization as dominant petrological processes.
METHODOLOGY
The geological data in the study area are scarce and/or very regional, so in this work, the samples were collected from outcrops of coal in underground mines and opencast mines, for information on organic petrography, including the identification of macerals groups, and measuring the vitrinite reflectance. Samples were prepared according to standard. ASTM D2797: Preparing coal samples for microscopical analysis by reflected light (ASTM, 2007), and ISO 7404-2:2009: Methods for the petrographic analysis of coals - Part 2: Methods of Preparing coal samples (ISO-ANSI, 2009). The determination of the vitrinite reflectance was performed according to ASTM D 2798: Microscopical determination of the vitrinite reflectance of coal (ASTM, 2007),and ISO 7404-5:2009: Methods for the petrographic analysis of coals -Part 5: Method of determining microscopically the reflectance of vitrinite (ISOANSI, 2009), using a microscope Leica DM RX™ coupled to a spectrometer 400 COAL MSP™. Calibration was carried out considering a xenon lamp with optic fiber connected to the spectrometer, and standards of 0.898% (Yttrium-Aluminum-Garnet), and 1.704% (Gadolinium-Gallium-Garnet) reflectance. The measurements and statistical analysis was carried out with the MSP200 ™ software supplied by Leica.
RESULTS The maceral composition of the coals is dominated by vitrinite whose reflectance values range between 0.7 to 1.2%. The coals are therefore of medium rank (ISO 11760:2005). These data fall within range average of results reported by regional authors as Ariciaga-Martínez (1987), Ariciaga-Martínez and Maycotte (1987), Eguiluz de Antuñano (2001), Piedad-Sánchez (2004), Corona-Esquivel et al. (2006, 2007), Carrillo-Pedroza et al. (2010). The maceral composition is dominated by vitrinite (mainly collotelinite and collodetrinite) indicates a terrigenous origin of organic matter.
Figure 2. Photomicrograph of polished section of coal in the area of San Juan de Sabinas, Coahuila de Zaragoza (N Mexico) where the center is observed liptinite, vitrinite in a mass with signs of degassing, and particles of inertinite in the top (reflected light, with oil immersion, objective 50X).
The reflectance values indicate that the coal seams have been affected by a variable temperature gradient. The presence of lithic tuff with altered dacitic composition, between the coal seams, indicates the influence of a volcanic event of Cretaceous age, or at least one event associated with magmatism (hydrothermalism?), which provided an additional heat flow to Sabinas Subbasin in the area of San Juan de Sabinas. This additional heat input could be very similar to the calculated heat flow of 3830 mW/m2 to dacitic magma in parallel cracks to the rift of Kilauea volcano (Teplow et al., 2008) considering the thermal conductivity K = 2.9 W/m2 °C for the basalt (Brigaud and Vasseur, 1989; Teplow et al., 2008). This heat flow is greater than the average of 270290 mW/m2, typical of mid-oceanic ridges. In this case, the problem of heat transfer is significant but difficult to quantify and dating, but do not forget that once a fluid is present, the heat transfer is more efficient and the magmatic mixture is spread through of sediment material. This volcanic pulse seems to be older than the one gave rise to the Las Esperanzas basalts, and could be related to magmatic episode documented in the Rio Grande Rift, Texas, aged between 53 and 23 ma (Seager and Morgan, 1979). These local temperature variations could explain the variation of coal quality both horizontally and vertically in the San Juan de Sabinas region, in addition to
differential burial of the other sub-basins of the region for its placement in different synclines.
CONCLUSIONS The vitrinite reflectance values indicate a variable thermal gradient, possibly attributable to the extrusive volcanism that outcrops near the studied area (e.g., Las Esperanzas Volcanic Field which erupted through fissures located at the border of the Santa Rosa anticline and covered the studied area), and to the singular burial history of the Sabinas Sub-basin. The results obtained using an organic petrology approach will contribute to the numerical modelling of the basin as a means of delimiting and evaluating more accurately, new areas of potential economic interest.
ACKNOWLEDGEMENTS The authors acknowledge the partial financing of CONACYT through 025355 and 67039 projects, to the company Minerales Monclova S.A. de C.V. for the samples provided. Piedad-Sánchez is grateful to José Luis Hernández-Michaca and Víctor Sánchez-Granados of Geología y Medio Ambiente S.A. de C.V. for logistical support to carry out this work.
BIBLIOGRAPHIC REFERENCES Adams, W.E., 1881. Coals in México, Santa Rosa District. Transactions of the American Institute of Mining Engineers. 10: 270-273. Ariciaga-Martínez, C., 1987. Censo de localidades con indicios de carbón en México: Exploración Regional. Informe, Superintendencia de Estudios Zona Norte, Comisión Federal de Electricidad. 84 p. Ariciaga-Martínez, C., and Maycotte, J.I., 1987. Resultados geológicos del censo de localidades con indicios de carbón en México. XVII Convención Internacional de Minería, Acapulco, Guerrero, México. 149-167. ASTM, 2007. Annual book of ASTM standards: Petroleum products, lubricants, and fossil fuels; Gaseous fuels; coal and coke, sec. 5, v. 5.06. ASTM International, West Conshohocken, PA. 711 pp. Brigaud, F., and Vasseur, G., 1989. Mineralogy, porosity and fluid control on thermal conductivity of sedimentary rocks. Geophysical Journal. 98: 525-542.
Carrillo-Pedroza, F.; Piedad-Sánchez, N.; Soria-Aguilar, M.; Pecina-Treviño, T.; Dávalos-Sánchez, A.; Villasana-Martinez, O.; Gauna-Arista, J.A.; EncisoCárdenas, J.J.; Garza-García, M.; 2010. Estudio comparativo de diversos tratamientos para la remoción de azufre en carbón mineral. Pruebas de laboratorio. GEOMIMET. XXXVII (287): 6-17. Corona-Esquivel, R., Martínez-Hernández, E., Tritlla, J., Benavides-Muñoz, M.E., Piedad-Sánchez, N., 2007. Principales yacimientos de carbón mineral en México. Geomimet. XXXIV (269): 8-40. Corona-Esquivel, R., Tritlla, J., Benavides-Muñoz, M.E., Piedad-Sánchez, N., Ferrusquía-Villafranca, I., 2006. Geología, estructura y composición de los principales yacimientos de carbón mineral en México. Boletín de la Sociedad Geológica Mexicana. Volumen Conmemorativo del Centenario: Revisión de algunas tipologías de depósitos minerales de México. LVIII (1): 141-160. Eguiluz de Antuñano, S., 2001. Geologic evolution and gas resources of the Sabinas Basin in northeastern Mexico. In: Bartolini, C., Bufler, R.T., CantúChapa, A. (eds.). The western Gulf of Mexico Basin: tectonics, sedimentary basins and petroleum systems. AAPG Memoir. 75: 241-270. Eguiluz de Antuñano, S., and Amezcua-Torres, N., 2003. Coalbed methane resources of the Sabinas Basin, Coahuila, México. in C. Bartolini, R. T. Buffler, and J. Blickwede, eds., The Circum-Gulf of Mexico and the Caribbean: Hydrocarbon habitats, basin formation, and plate tectonics. AAPG Memoir. 79:. 395–402. ISO 11760:2005. Classification of coals. International Organization for Standardization. American National Standards Institute. 9 p. ISO 7404-2:2009. Methods for the petrographic analysis of coals -- Part 2: Methods
of
preparing
coal
samples.
International
Organization
for
Standardization. American National Standards Institute. 12 p. ISO 7404-3:2009. Methods for the petrographic analysis of coals -- Part 3: Method of determining maceral group composition. International Organization for Standardization. American National Standards Institute. 7 p. ISO 7404-5:2009. Methods for the petrographic analysis of coals -- Part 5: Method of determining microscopically the reflectance of vitrinite. International Organization for Standardization. American National Standards Institute. 14 p.
Mulleried, F. K., 1941. Actividad Volcánica bastante reciente del Oriente del Estado de Coahuila, México. Revista Geográfica del Instituto Panamericano de Geografía e Historia.1 (2,3): 182-201. Piedad-Sanchez, N., 2004. Prospection des hydrocarbures par une approche intégrée de pétrographie, géochimie et modélisation de la transformation de la matière organique: analyse et reconstitution de l’histoire thermique des Bassins Carbonifère Central des Asturies (Espagne) et Sabinas-Piedras Negras (Coahuila, Mexique). Ph. Thesis. UHP Nancy I, Francia. 356 pp. Robeck, R. C., Pesquera, V. R., Arredondo, S. U., 1956. Geología y depósitos de carbón de la región de Sabinas, Estado de Coahuila. 20th International Geological Congress, Mexico City, 190 p. Seager, W.R., and Morgan, P., 1979. Rio Grande rift in Southern New Mexico, West Texas, and Northern Chihuahua. In: Riecker, R.C. (editor), Rio Grande rift: Tectonics and magmatism. American Geophysical Union. 87-106. Teplow, W. J.; Marsh, B. D.; Hulen, J.; Spielman, P.; Kaleikini, M.; Fitch, D. C.; Rickard, W., 2008. Dacite Melt at the Puna Geothermal Venture Wellfield, Big Island of Hawaii. American Geophysical Union, Fall Meeting 2008, abstract #V23A-2129. Valdez-Moreno, G., 2001. Geoquímica y petrología de las rocas ígneas de los campos volcánicos Las Esperanzas y Ocampo, Coahuila, México. Posgrado en Ciencias de la Tierra, Universidad Nacional Autónoma de México. 115 pp.
Oviedo ICCS&T 2011. Extended Abstract-Poster
Correlation between Fluidity and Structural Transformation of Three Coal Ashes Xiongchao Lin1, Jin Miyawaki2, Seong-Ho Yoon1,2* and Isao Mochida3 1
Interdisciplinary Graduate School of Engineering Sciences, Kyushu University, Fukuoka 816-8580, Japan 2 Institutes for Materials Chemistry and Engineering, Kyushu University, Fukuoka 816-8580, Japan 3 Research and education center of carbon resource, Kyushu University, Fukuoka 816-8580, Japan *Corresponding author: Tel: +81-92-583-7959; Fax: +81-92-583-7897; E-mail:
[email protected] Abstract Understanding of temperature-viscosity tendency of various mineral species with different composition and structure is still inevitable and important for stable operation of practical entrained-flow coal gasification systems. In this study, the compositions and structures of ashes and slags from 3 kinds of coals and their relations to the fluidity properties were investigated. Results demonstrated that viscosity was strongly influenced by the compositions. Further, mineral structure analysis indicated that the different viscosities were caused by the different framework structure of melting ashes. As the fluxing agent, Ca ion was considered to modify the Si-O-Si and Si-O-Al network structures. However, Fe ion probably might influence only on the Si-O-Al structure. Keywords: gasification, coal ash, viscosity, mineral 1. Introduction
In coal gasification, coal ashes are accumulated as a molten state (slag) on the internal wall of the gasifier and are discharged by flowing down to outside due to the gravity. For the stable operation of coal gasification system, the effective discharge of slag is a most important task. To achieve the stable discharge of molten slag on the operation using various coals, full understanding of the physical and chemical properties especially fusibility and fluidity characteristics of the molten slag are necessary. Regardless of the intensive researches over several decades, however the
Oviedo ICCS&T 2011. Extended Abstract-Poster
relationship between the compositional structures of molten slag and the viscosity is not fully understood. This study aims at obtaining a detailed understanding of compositional and structural influences on the molten viscosity of ash and slag. 2. Experimental section
Various coals (Datong (D), Malina (M), Adaro (A), Pinan (P) and Tsugunuyi (T)) were used in this study. Ashes were prepared according to the JIS standard. Slags were supplied from companies which were obtained from gasification plant. The compositional and structural analyses of ashes and slags were carried out by XRF, XRD, SEM-EDX and multi-nuclear solid state NMR. Moreover, the relative viscosities of ashes were measured under N2 atmosphere up to 1700 oC using a self-designed high temperature rotational viscometer with high purity alumina rotor and crucible. 3. Results and Discussion
3.1 Basic composition of coal and ashes Table 1 shows the compositions of ashes. The majority of specimens are SiO2, Al2O3, Fe2O3 and CaO. Compositions are usually divided into two parts as basic and acidic such as Ratios of basic components = %Fe2O3 + %CaO + %MgO + %Na2O + %K2O ;
Ratios of acidic components = %SiO2 + %Al2O3 + %TiO2 of primary
minerals, of which properties were strongly influenced by them. [1] High alkaline and alkaline-earth contents lead to the lower ST, HT and FT. The effects of iron oxide are also considered to decrease the ST, HT and FT, however it is more complex owing to the different temperatures and atmospheres than that of CaO.
Oviedo ICCS&T 2011. Extended Abstract-Poster
Table 1 Analyses of coal and ashes D M A Samples SiO2 55.56 29.57 33.3 Al2O3 28.12 19.69 17.6 Fe2O3 11.52 32.16 16.56 CaO 1.38 7.23 16.8 K2O 1.18 1.23 1.1 0.97 1.47 1.26 TiO2 SO3 0.57 5.35 6.27 c P2O5 0.36 ND NDc MgO NDc 2.68 6.64 c ZrO2 ND 0.03 0.01 Total 99.60 99.41 99.54 d 0.19 0.85 0.90 B/A Ash properties (Oxidizing Atmosphere) d ST (oC) 1310 1380 1210 e o HT( C) 1490 1400 1220 f o FT( C) >1500 1420 1260 a
ND not detected
b c
Trace elements (e.g. Sr, Mn, V, Zr, Y, Zn, Cu,Rb) were ignored.
b=%Fe2O3+%CaO+%MgO+%Na2O+%K2O; a=%SiO2+%Al2O3+%TiO2
d
ST as softening temperature; eHT as hemispherical temperature fFT as fusing
temperature. 3.2 XRD analyses Compositional and structural transitions of primary minerals in coal ashes and slags were systematically investigated at a temperature range of 300 oC to1600 oC. Kaolinite and/ or illite, quartz, pyrite, siderites, and ankerites etc., are the mainly composed minerals in a majority of coals as low temperature crystalline minerals. The transitions of such primary minerals with increasing temperature (300-1000oC) are principal decomposition and crystal transition at low temperature region. D ash with high SiO2 content underwent the transformation of quartz to cristobalite and mullite through the solid reaction of agglomerated silica and alumina. In contrast, M ash with high Fe2O3, these transformations were carried out through the reactions among silica, alumina
and
iron
oxide
to
Fe-rich
spinel.
Ca-rich
mineral
(primary
anorthite-CaAl2Si2O9) in A ash formed the low temperature eutectic. Rapid quenching
Oviedo ICCS&T 2011. Extended Abstract-Poster
slags show the states of amorphous glasses and mixed structures with some solid minerals, which make it difficult to identify their compositions by using XRD technique as shown in Fig.1. Moreover, three broad peaks were formed in the different regions. According to the components of ashes, we assumed that the broad peak at 13-33o in slag D is ascribed to the amorphous alumina-silicate and silica; broad peak at 15-35o in slag M is attributed to the amorphous iron-rich minerals and alumina-silicate; moreover, the broad peak at 17-37o in A ash is considered to be the amorphous (Ca, Mg)-bearing minerals that from the destruction of crystal structure by the rapid quenching.
Intensity (a.u.)
(a)
(b)
(c)
10
20
30
40
50
60
2 theta (deg.)
Fig.1. XRD patterns of (a) D, (b) M and (c) A Slag as received. 3.2 Molten structural of ashes with different compositions In the present study, the structural analyses were carried out by using multi-nuclear solid state NMR. Some of the amorphous matters derived from decompositions of minerals (such as alumina and meta-kaolinite) were detected by NMR analyses. Due to the complicated mixing states of mineral components, molten ashes and slags showed the relatively broad resonance signals. Resonances in A ash was obviously shifted to low magnetic field, reflecting the effect of the Ca2+ cutting of the large polymeric structure to segments (Fig. 2a). Different effects of Ca and Fe are considered as follows: Fe ion may only influence on the Si-O-Al chain of framework that result in the less shift of 29Si resonance in M molten ash (Fig. 2b). In contrast, Ca
Oviedo ICCS&T 2011. Extended Abstract-Poster
ion cut the network of Si-O-Si and Si-O-Al to result in the apparent shift of 29Si peaks to low magnetic field. 60
Fraction (%)
50
D1600 M1600 A1600
Si
(a)
2+
Ca O O
O
30 O
Al
O
O
10 0 -80 -85 -90 -95 -100 -105 -110 -115 -120 Chemical shift (ppm)
O
Ca2+O
Si
Q3
O
Major in A ash
40
20
Q2
O
Si O
O
Si
O
Si
O
O
Al O
Si
O Ca2+ O
O Si
O
Al
(b) OCa 2+
O Fe 2+
O O
Si
Q4
Major in M ash
Major in D ash
O
Fig.2. Structural analyses of ashes (a) Chemical shift as a function of structural fraction of ashes prepared at 1600 oC and (b) prediction of structural distribution in melting ashes. 3.3 Flow properties of ashes with correlate with structures Due to the co-existing of various micro-structures in melting ashes, they show different rheological behaviors at different temperatures. Fig. 3a shows the relationships between the viscosity and the composition of molten ashes. Molten D ash with high amounts of solids and/or crystal showed high viscosity, in which the melting matters were thought as silica and mullite eutectic. The isolated amorphous alumina was considered significantly increased the viscosity. A and M ashes have similar b/a ratio that contributed to high Ca and Fe contents, respectively. However, they show different viscosity tendencies. The distinctions should be caused by the different effects of Ca and Fe ions in the network of melting ashes. The combining force of alkaline and alkaline-earth must be stronger than that of iron, because the iron ions are considered to combine principally with Al tetrahedrons. Hence, alkaline and alkaline-earth cations decrease the viscosity and melting temperature of ash more effectively. Fig. 3b shows the effect of shear rate on the sensitivity of the melting viscosity to temperature variation. At low temperature, viscosity decline rapidly with the
Oviedo ICCS&T 2011. Extended Abstract-Poster
increasing of shear speed. This change performed less at high temperature after majority liquid phase forming. [2] Viscosity constant regardless with shear rate at high temperature region was considered totally liquid with relate to short molecule or chain structure (such as anorthite in A ash); moreover, viscosity became more distinct with shear rate changes may be induced by the layer structure in M ash (such as Fe-rich spinel) and framework structure in D ash (such as mullite).
140
60 40
A ash
B/A=0.85 High Fe
B/A=0.19 High Si+Al
B/A=0.90 High Ca
40 30 20 10
o
1200 C o 1250 C o 1300 C o 1350 C o 1400 C o 1450 C o 1500 C o 1550 C o 1600 C
10001100120013001400150016001700 o Temperature ( C )
00 1 2 3 4 5 6 Sh e a
r rat
( Co
0
)
20
7
e (S -1 )
8
9
o
1650 C
10
pe ra tu re
80
(b)
D ash Mash
100
Te m
(a)
Viscosity (Pa.s)
Viscosity (Pa.s)
120
0 6.250 12.50 18.75 25.00 31.25 37.50 43.75 50.00
50
Fig.3.Viscosity of (a) tendencies correlation with compositions and (b) as a function of shear rate at different temperatures. 4. Conclusions
In present works, various structure transitions were found with strong relationship to the ash properties, especially the fluidity behaviors at melting stage. Alkaline-earth ions certainly cut the framework to segments corresponding to the shift of 29Si spectra. Large frame work structure shows non-Newtonian flow and strong shear-thinning behavior; In contrast, segmental structures hold the Newtonian flow till low temperature and it contributes less shear-thinning behavior. Acknowledgement
The authors are grateful to the financial support from new energy development organization (NEDO) and globe center of excellence (GCOE). References
[1] Harold H. Schobert, Robert C. Streeter, Erle K. Dieh. Flow properties of low-rank coal ash slags: Implications for slagging gasification, Fuel 1985; 64:1611-1617.
Oviedo ICCS&T 2011. Extended Abstract-Poster
[2] Wenjia Song, Lihua Tang, Xuedong Zhu, et al., Flow properties and rheology of slag from coal gasification Fuel 89 (2010), pp. 1709-1715.
Oviedo ICCS&T 2011. Extended Abstract
Nanominerals and ultra-fine particles within coal ashes
Luis F. O. Silvaa; Frans Waandersd; Marcos L. S. Oliveiraa; Kátia da Boita a
Catarinense Institut of Environmental Research and Human Development – IPADHC,
Brazil. b
School of Chemical and Minerals Engineering North West University (Potchefstroom
campus) Potchefstroom 2531, South Africa * Corresponding author. E-mail address:
[email protected] (L.F.O. Silva)
Abstract Environmental and human health risk assessments due to nanoparticle effects from coal and coal ashes require thorough characterization of the nanoparticles and their aggregates. In this paper the nanosized particles from coal combustion sources are investigated and the complex micro mineralogy of these airborne combustion-derived nanomaterials are characterised. The investigation forms part of a larger experiment on the technical feasibility and environmental impacts of combustion in a Brazilian coalfired power station which uses coal with many potential damaging elements and pyrite, producing a high ash. The combination of Optical Microscopy with instrumental microscopic techniques like Electron Microscope coupled to Energy Dispersive X-Ray Spectroscopy (EDS), Confocal Microscopy and Micro-Raman Spectroscopy have demonstrated to be useful tools for the research of the mineralogical composition of coal ashes. Nanometre-sized crystalline phases in fly ash were characterised using FE-SEM, HR-TEM, and EDS. HR-TEM data reveal nanoscale C-deposits juxtaposed with and overgrown by slightly larger aluminosilicate glassy spheres, oxides, silicates, carbonated, phosphates, and sulphates. The nanoparticles include iron-rich oxide, Fesulphate, and Fe-aluminosilicate glass. Individual metalliferous nanoparticles have a heterogeneous microstructure in which elements such as iron, aluminium and silicon are not uniformly distributed. Iron oxides (mainly hematite and magnetite) are the main coal ashes products of the oxidation of pyrite, sometimes via intermediate pyrrotite formation. The presence of iron oxide nanocrystals mixed with silicate glass particles emphasises the complexity of coal and bottom ash micro mineralogy. Given the
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Oviedo ICCS&T 2011. Extended Abstract
potentially bio reactive nature of such transition metal-bearing materials, there is likely to be an increased health risk associated with their inhalation. The techniques provide a fast and powerful analytical technique in the study of fly ash nanoparticles, providing a better understanding of the detailed chemistry of this potentially strongly bio reactive component of atmospheric particulate matter, non-destructive and highly-selective analysis of both the surface and the coal inner bulk, and through chemical modelling simulations give the required information to confirm the stability of secondary minerals detected in the samples and helps to diagnose the potential environmental risks associated to their weathering.
1.
Introduction
The combustion of coal by amongst others power plants generate large quantities of fly ash that can have a significant impact on the environment. Although fly ash particles with sizes ranges between 1-10μm diameter, are readily identified using analytical electron microscopy and well documented in the literature [1, 2], there is relatively little information available on ultrafine (< 100 nm) particles, even though these are abundantly present in coal fly ash. The greater surface areas of ultrafine particles compared with larger particles with the same chemical compositions make them more environmentally active with respect to bio-uptake and associated health risks [1, 3-5]. In this paper it is demonstrated how FE-SEM, HR-TEM, and EDS can be used to investigate the elemental distribution, morphology, crystalline phase and electronic structure of individual coal fly ash particles, with emphasis on the ultrafine particles that may have the greatest impact on human life.
2.
Experimental
The coal ash and fly ash samples for this study were collected at the principal Brazilian power plant in the Santa Catarina, State, which uses coal for the generation of electricity. The incineration temperature in the combustion chamber varies between 1 000 ºC and 1 500ºC, and almost 98% of the fly ash is captured in the electrostatic precipitators, which then is then used in the cement industry. The composition of the crystalline minerals in the coal and coal ash were determined by means of a Siemens model D5005 X-ray diffractometer. Qualitative chemical analysis were performed with a LEO-435VP scanning electron microscope (SEM) fitted with an Oxford EDS with a resolution > 133eV. HR-TEM analyses were conducted on particles Submit before 31 May 2011 to
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Oviedo ICCS&T 2011. Extended Abstract
suspended in hexane. The suspension was stirred for ~1 min and subsequently pipetted onto lacy carbon films supported by Cu grids. The suspension was left to evaporate before inserting the sample into the TEM. The samples were studied by means of a 200 keV JEOL-2010F field-emission HR-TEM equipped with an Oxford energy dispersive X-ray detector (EDS) unit. EDS spectra were recorded in TEM image mode and then quantified using ES Vision software which uses the thin foil method to convert X-ray counts of each element into atomic or weight percentages.
3.
Results and Discussion
The common minerals and phases identified in the fly ash mainly contain inert silicates, oxides, and amorphous phases, and to a lesser extent, carbonates, sulphates, and hydroxides, which formed during heating. They are a result of some additional volatilization of elements and induced alteration, decomposition, crystallization, recrystallization, or amorphization of the actual minerals and phases present [6]. The energy-dispersive X-ray spectrometer data and high-resolution transmission electron microscopy (HR-TEM) images revealed the presence of fine crystalline phases, such as iron-rich oxide spinels and a Fe-aluminosilicate glass. The fate of such nanominerals during coal combustion is determined by competitive processes by which different iron compounds are produced and they typically exist either in iron-oxide form or in combination with other elements, forming multi-element oxides
Figure 1: HR-TEM image and EDS spectrum showing very-fine ordered Hematite structures in fly ash particles [7].
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Oviedo ICCS&T 2011. Extended Abstract
Figure 1: (A) Jarosite pseudomorph (pyrite-sulphur-jarosite assemblage) from fly ash (TEM image and EDS); (B) Hematite present in fly ash (TEM image) [8].
Fig. 4 (a) HR-TEM and Fourier transformation (FFT) confirm the size of a nanoquartz sphere; (b) FE-SEM of submicroscopic spheres, containing Zr, Ni, Mg and Al [1] 4.
Conclusions
The combination of FE-SEM and HR-TEM/ EDS as used in this study provides a powerful technique to characterise nanoparticles in coal ash and further demonstrates the complexity of mineralogical relationships between nanominerals present in coals and bottom ashes produced during coal combustion, yielding observations which are impossible to make using more traditional characterisation methods such as optical
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Oviedo ICCS&T 2011. Extended Abstract
petrography. Iron was found to be the most abundant transition metal in ambient particulate especially around the Santa Catarina power plant.
Acknowledgements One of the authors (FW) is greatly indebted to the main author (LFOS) who provided the information needed for this presentation.
The NWU and THRIP are thanked for
financial support to attend the conference.
References 1.
Silva, L.F.O. and da Boit, K.A. Nanominerals and nanoparticles in feed coal and bottom ash: implications for human health effects. Environ Monit Assess (2011) 174:187–197.
2.
Ribeiro, J., Flores, D., Ward, C.R., Silva, L.F.O., Identification of nanominerals and nanoparticles in burning coal waste piles from Portugal. Science of the Total Environment 408 (2010) 6032–6041.
3.
O’Connor, G.M., Dick, S., Miller, C., Linak, W. (2004). Differential pulmonary inflammation and in vitro cytotoxicity of size-fractionated fly ash particles from pulverized coal combustion. Journal of the Air & Waste Management Association, 54, 286–295.
4.
Oberdoerster, G., Oberdoerster, E., Oberdoerster, J. (2005).Nanotoxicology: An emerging discipline evolving studies of ultrafine particles. EnvironmentalHealth Perspectives, 113, 823–839.
5.
Xia, T., Lovochick, M., & Brant, J. (2006). Comparisons of the ability of ambient and manufactured nanoparticles to induce cellular toxicity according to an oxidative stress paradigm. Nano Letters, 6, 1794–1897.
6.
Vassilev, S.V., Vassileva, C.G., 2005. Methods for Characterization of Composition of Fly Ashes from Coal-Fired Power Stations: A Critical Overview. Energy Fuels, 19 (3), 10841098.
7.
Silva, L.F.O., Querol, X., da Boit, K.M., de Vallejueloc; S.F-O., Madariaga, J.M. Coal Mining Residues and Sulphides Oxidation by Fenton´s Reaction: an accelerated weathering procedure to evaluate the impact for environment and human health to be published in Science of the Total Environment (2011) Manuscript Number: STOTEN-D-10-00124
8.
Silva, L.F.O., Macias, F., Oliveira, M.L.S., da Boit, M.K., Waanders, F. Coal cleaning residues and Fe-minerals implications Environ Monit Assess (2011) 172:367–378.
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Oviedo ICCS&T 2011.
Properties Of Fly Ash From Biomass Combustion R. P. Girón, I. Suárez-Ruiz, B. Ruiz, E. Fuente, R. R. Gil Instituto Nacional del Carbón, CSIC. Francisco Pintado Fe, 26, 33011 Oviedo, Spain,
[email protected], phone: +34985119090, FAX: +34985297662
Abstract Nowadays the use of biomass as renewable energy source is increasing, however this application has the disadvantage that it produces large quantities of fly ash. The type of ash is different depending on the biomass source and combustion conditions. In this work fly ashes derived from fixed combustion and from fluidized bed combustion systems of forest biomass have been investigated. The raw fly ashes have been dry sieved, and then both the raw sample and the fractions were exhaustively characterized.
1. Introduction The increased demand for energy and the polluting nature of current sources of fossil fuel energy demonstrates the need for new energy technologies, which can offer greater efficiency and minimal environmental impact. Biomass combustion worldwide has a significant potential to meet this demand and great importance as regards global warming since the CO2 generated is considered neutral, for example, in regard to forestry wastes, wood chips, sawdust and bark as fuel will contribute to the conservation of fossil fuel resources and reducing greenhouse gas emissions [1, 2]. However, the disadvantage of incineration is that it produces large amounts of fly ashes. Landfill disposal has traditionally been the most widely used method of disposal in waste management. With the recent increase in the cost, acquisition and development of new waste disposal sites, the management of this fly ash represents a purpose today in energy generation. Biomass fly ashes have a predominantly inorganic fraction and an organic fraction (unburned carbon) minority. The proportion of both fractions is dependent on different parameters and conditions of the combustion process, such as the type of biomass, the load, the combustion, gasification and operating conditions. There are several types of combustion, industrial biomass combustion systems employ fixed bed and fluidized bed appliances. Fixed bed furnaces usually entail lower investment and operating costs for smaller power plants [3]. On the other hand, fluidized
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Oviedo ICCS&T 2011.
bed furnaces combine high fuel flexibility, good mixing and temperature control, with high conversion efficiencies and low pollutant emissions. This work aims to characterize of two types of fly ash obtained from burning forest biomass: i) fly ash obtained from a combustion furnace (750 ºC) and ii) in fluidized bed gasifier (550 °C). The objective of this characterization is to determine the possible uses of each of the fractions; for example, fractions with a high unburned carbon content can be used as precursors for obtaining activated carbons whereas fractions with a low unburned carbon content can be used as nutrient in ground.
Experimental section For this research, fly ashes were obtained from combustion of forest biomass in a paper mill industry located in northern Spain. The biomass used consisted of bark and chips of Eucaliptus globulus. The fly ashes were sampled in electrostatic precipitators, one from the fixed bed system that operates at a temperature of 750 ºC, referred to in this study as A and the second from fluidized bed at a working temperature of 550 ºC, called B. From the total sample (20 kg) representative subsamples were obtained for a quarter. The raw fly ashes samples were dry sieved yielding the following fractions: >500 μm, 500-212 μm and > 212 μm. Each fraction was labelled as follows; the fly ash from which it was derived followed by a subscript for the size fraction, thus the fraction >500 μm from A is referred to as A5. To characterize of the raw samples and the corresponding fractions a set of different techniques were used following in most cases the standard procedures. These techniques were: petrographic analysis for volumetric composition; XRF, atomic absorption and XRD for mineral matter characterization, and thermogravimetric analysis under N2 atmosphere for thermal behaviour.
2. Results and Discussion Petrographic composition. Table 1 shows the dry sieved yield and the total unburned carbon and mineral material contents in the raw fly ashes, A and B, and their fractions. The highest concentrations of unburned carbons were found in the two fractions >500 μm, but if the fly ash fractions are compared, the B5 content lower volume of unburned carbons. Fly ash B and its fractions have a higher mineral content, because the conditions provided by fluidized bed combustion appear to be more efficient than those
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Oviedo ICCS&T 2011.
of fixed bed combustion. Table1. Petrographic composition and dry sieved yield Dry Sieved Yield (%) Unburned (% vol.) Mineral Matter (% vol.)
A
A5
A2-5
A2
B
B5
B2-5
B2
16,0 84,0
2,17 95,2 4,8
4,17 9,5 90,5
93,66 7,6 92,4
4,0 96,0
4,01 28,4 71,6
8,63 4,0 96,0
86,76 1,7 98,7
With respect to the unburned carbons, Figure 1 and Figure 2 show the images obtained by both optical microscopy and SEM. Two types were identified: i) unburned carbons with a low level of transformation retaining the cellular structure of biomass, and ii) unburned carbons that are transformed, passing through a molten phase during which a large amount of volatiles are evolved resulting in a abundance of macroporosity. This pattern is repeated in both fly ash, A and B.
Fig 1. Unburned carbon weakly transformed
Fig 2. Unburned carbon transformed.
Chemistry and inorganic composition of fly ashes. The crystalline mineral species in fly ashes identified by XRD analysis are illustrated in Table 2. In both cases the inorganic fractions are dominated by Ca and Si, forming carbonates, sulphates and silicate components. The inorganic elements contents, expressed as oxides, are shown in Fig. 3. Both fly ashes are rich in Ca and Si, in B and their fractions the Si content is due to SiO2 which is part of the bed and is dragged together with fly ash. It is noted that the composition of the fraction of <212 microns is similar to that of its raw fly ash, this is due to the high inorganic yield. Table 2.Mineralogical analysis of the fly ash samples ( XRD analysis) A A5 A2-5 A2 B Mineral phases Quartz (SiO2) ++ ++ +++ ++ +++ Calcite (CaCO3) +++ ++ + ++ + Anhydrite(CaSO4) ++ + ++ Lime (CaO) + + Gehlenite (CaAl2SiO7) + ++ Sylvite (KCl) + + ++ + +++High intensity ++Medium intensity +low intensity
B5
B2-5
B2
+++
+++ +
++ ++ + + +
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Oviedo ICCS&T 2011.
90 0,14
120
80
A
A5
A2-5
A2
B
B5
B2-5
B2
0,12
70
100 0,1 80
A
B
0,08
%
W e ig h t( % )
50
40
0,06
60
0,04 40
30
D e r iv .W e ig h t (% /º C )
60
0,02
20 20 0
10 0
0
0
SiO2 (%)
Al2O3 (%)
Fe2O3 (%)
CaO (%)
MgO (%)
Na2O (%)
K2O (%)
TiO2 (%)
100
200
300
400
P2O5 (%)
Fig 3. Analysis of major elements as oxides of fly ashes and their fractions.
500
600
700
800
900
-0,02 1000
Temperature (ºC)
Fig 4 Thermogravimetric analysis in a nitrogen atmosphere
Thermogravimetric analysis. The thermograms of raw fly ashes A and B, Fig. 4, show three main temperature ranges of weight loss, apart from the first loss which is due to the moisture content. The first range of temperature is between 300 and 400 ºC with a maximum peak associated to the decomposition of magnesium carbonate. In the second range of temperatures (500-700 ºC) the weight loss is due to the decomposition of some of the unburned carbons and finally, at the last temperature (>700 ºC) weight loss is due to the decomposition of calcium carbonate and other carbonates. 3. Conclusions The results of the comparison of the two fly ash samples show that the unburned carbon content in fly ash from the fixed bed combustor is greater than that of the fluidized bed leading us to conclude that combustion is more effective in the second type of combustor. In both cases there are two types of unburned carbon, one type some more transformed than other. These unburned carbons will be investigated for potential precursor materials of activated carbons. The majority species in the fly ash are Ca and Si, in the case of A the predominant species is Ca in the form of CaCO3, and in B the predominant species is Si in form of SiO2. Acknowledgement. The financial support for this work was provided by research project from the PRIAsturias, PC 07-015.The Phstudent thanks the CSIC for a predoctoral research grant JAE (ref. PR2005-0168).
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Oviedo ICCS&T 2011.
References [1] Green, C., Byrne, K.A., Cutler, J.C. Biomass: Impact on Carbon Cycle and Greenhouse Gas Emissions. Encyclopedia of Energy. New York: Elsevier 2004:223-36. [2] Petersen Raymer, A.K. A comparison of avoided greenhouse gas emissions when using different kinds of wood energy. Biomass and Bioenergy 30, 605-17 (2006). [3] Warnecke, R. Gasification of biomass: comparison of fixed bed and fluidized bed gasifier. Biomass and Bioenergy 18, 489-97 (2000).
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Oviedo ICCS&T 2011. Extended Abstract
Characteristics of Crush Strength in Coal Briquette Molded with Polymer as a Binder Seung-Hyun Moon1, Seung-Jae Lee1, In-Soo Ryu1, Yong-Woo Kim1, Tae-In Ohm2 1
Korea Institute of Energy Research, 71-2 Jang-dong Yuseong-gu Daejeon, S.Korea. E-mail :
[email protected] 2 Hanbat National University, San 16-1 Duckmyung-dong Yuseong-gu Daejeon, S.Korea. Abstract A rapid increase in energy price makes low grade fuel to be a more promising one. The lower grade fuels containing high contents of moisture, ash and sulfur cause many problems in combustion such as low calorific value, particulate and sulfur dioxide. The problems by ash and sulfur can be solved by flue gas treatment technologies. High content of moisture in low rank coal is related to combustion efficiency as well as transportation. Therefore, low rank coal should be dried before transportation. As a mere drying of low rank coal tends to cause spontaneous ignition during mass storage, the coal has been dried by oil frying method and made in a form of briquette. In this study, we compared crushing strength of coal briquette in which polymer such as polyethylene and polypropylene is used as a binder. Indonesian low rank coal was fry-dried in the kerosene containing 0.5% polymer. A certain amount of the fry-dried coal was put into the mold, heated up to 60-100oC. Molding pressure and molding duration were varied to find optimum conditions. The crushing strength of a briquette formed from the fry-dried coal increased along with the content of polyethylene. The higher fry-drying temperature influenced to lower the briquette strength. Molding pressure had an optimum condition at 100 kgf/cm2 showing stronger briquette than the others molded at 80 kgf/cm2 and 120 kgf/cm2. Briquette strength increased with a mold temperature, indicating that heat is necessary to evenly disperse a binder between coal particles. Molding duration was varied from 20 to 60 seconds, besides which conditions briquette strength was much weakened. In conclusion, polymer could be used as a binder for coal briquette and conditions such as drying temperature and temperature, pressure and duration in molding should be optimized to obtain proper briquette strength.
Key Words: Low-rank coal, Coal briquette, Polymer binder, Crush strength
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Oviedo ICCS&T 2011. Extended Abstract
1. Introduction A sharp increase of oil price made coal, especially low-rank coal, to be more important. Low-rank coal cannot be used as itself because of lower heating value without upgrading process. The coal with high content of moisture brings about several problems in transportation and combustion. Thus, the coal has been effectively upgraded by drying processes such as UBC (upgrading brown coal), BCB (bindless coal briquette), MTE (mechanical thermal expression), HWD (hot water drying), K-fuel, Syn-coal and so on. However, even after drying the coal, the dried coal tends to re-adsorb moisture from atmosphere during transportation and storage, which can cause various problems in processes of mixing, pulverizing and screening as well as in transportation [1]. Many efforts have been made to prevent the dried coal from re-adsorbing moisture. One of promising technologies is the fry-drying process in oil, well known as the UBC process [2], where the pore surface of the dried coal is coated with oil component e.g. a mixture of kerosene with asphalt. The heavy oil component of asphalt can mainly play roles of the avoidance of moisture re-adsorption on the coal and as a binder for coal-molding process [3]. However, the developed process is disadvantageous in high energy consumption required to transfer moisture and heavy oil component in the coal pores. Recently, our research group has developed a novel coal-drying process (non-fried carbon briquetting process: NFCB process) lowering drying temperature, where vinyl and plastic wastes, asphalt and palm oil are utilized as a heavy oil component to avoid moisture readsorption of coal. Thus, this study deals with mechanical characteristics of briquettes molded with coal dried by the NFCB process with a polymer.
2. Experimental Coal sample used in this study is Indonesian lignite containing higher than 30 wt.% of moisture, which is summarized in Table 1 with the results in composition analysis and calorific value of the coal sample using Tru Spec Elemental Analyzer (LECO Co., USA), SC-432DR Sulfur Analyzer (LECO Co., USA) and TEA-701 Thermogravimeter (LECO Co., USA). Lump coal of Indonesian lignite was crushed by a hammer mill and then by a conventional grinder. 1~3 mm of coal granules obtained using sieve tray were sealed in a plastic bag and then stored in a globe box in order to maintain the moisture content of the
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Oviedo ICCS&T 2011. Extended Abstract
coal. The coal granules were dried by the non-fried carbon briquette (NFCB) process developed by our research group, where kerosene and polyethylene polymer were utilized as light and heavy oil components, respectively.
Table 1. Summary of composition and calorific value of Indonesian Lignite. Moisture 34.27
Proximate Analysis (wt%) Volatile matter Ash Fixed carbon 33.64 2.01 29.99
Elemental Analysis (wt%) C H N O S 70.50 5.14 0.99 21.33 0.03
15 g of dried coal granule was mounted in a heated molder as shown in Figure 1, and then was pressurized by using hydraulic cylinder for a fixed period. Crushing strength of molded coal sample was measured by using an instrument with a load cell as shown in Figure 2.
(a)
(b)
Figure 1. Hydraulic molding machine for dried coal granule: (a) coal molder and (b) molded coal.
Figure 2. An instrument measuring crushing strength measuring of molded coal sample.
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Oviedo ICCS&T 2011. Extended Abstract
In order to examine an effect of polymer content of molded coal on crushing strength, polymer content was varied from 0.2 wt.% ~ 0.8 wt.% of total coal weight, where dried coal sample for molding was prepared with 100 g of coal and 100 g of kerosene by the NFCB process heating up to 135 oC for 30 min followed by evaporation of kerosene in oven at 130 o
C for 15 min. The dried coal was pressed to 100 kgf/cm2 of pressure, and was heated up to
80 oC for 40 sec. After 20 min of cooling time, the molded coal sample was taken out from the molder for measurement of crushing strength. Temperature for kerosene evaporation in the NFCB process was also varied form 110 oC to 150 oC. Molding pressure was adjusted in a range of 80 kgf/cm2 and 120 kgf/cm2, and molding time was increased from 20 sec to 60 sec. On the other hand, molding temperature was raised from 60 oC to 100 oC. Finally, cooling time after molding the dried coal was increased by 10 min. The conditions except each molding variable was fixed to the same values mentioned above.
3. Results and Discussion When 100 g of coal and 100 g of kerosene were used in the NFCB process, the increase of vinyl content can generally enhance crushing strength of molded coal samples (see Figure 3).
Figure 3. Crushing strength of the coals molded with various vinyl contents
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Oviedo ICCS&T 2011. Extended Abstract It is expected that heating the dried coal to 100 oC of molding temperature would melt vinyl in the dried coal to pass through the gaps between coal particles and block their contact to air, which could lead to improve the binding strength of coal particles. It is also worth noting that 0.5 wt% and 0.8 wt% of vinyl content do not significantly increase crushing strength of molded coals, which reveals that there should be the maximum of vinyl content over which the crushing strength of molded coal is not affected as much as the addition of vinyl. An effect of kerosene evaporation temperature is displayed in Figure 4. It seems that the coal samples evaporated at 110 oC and 130 oC are not significantly influent on the crushing strength. On the other hand, the coal prepared at 150 oC of the evaporation temperature exhibits relatively lower crushing strength than the other samples. Thus, the enhancement of the crushing strength of the molded coal samples can be assisted by the presence of some oil component of kerosene rather than complete removal of kerosene from the dried coal.
Figure 4. Effect of the oil evaporation temperature in the NFCB process. As shown in Figure 5, 100 kgf/cm2 of molding pressure can produce briquettes having the
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Oviedo ICCS&T 2011. Extended Abstract highest crushing strength. It is considered that molding pressure below 100 kgf/cm2 is difficult to get properly together coal particles. However, molding pressure over 100 kgf/cm2 may stress coal particles in molder out to lower crushing strength of molded coal samples.
Figure 5. Crushing strength of the coals molded under various molding pressure.
In the variation of molding time, 40 sec of molding time can achieve the highest crushing strength of coal briquette. Thus, it is not necessary to delay molding time too much in a view point of crushing strength of the briquette. The crushing strength can increase with increasing molding temperature in a range between 60 oC and 100 oC, implying that the mobility of oil component and vinyl present in the dried coal can rise with molding temperature to bind strongly coal particles. On the other hand, it is observed that the crushing strength is maximized at 20 min of cooling time after molding, and is not significantly changed at longer than 20 min of the cooling time. Thus, it is suggested that coal briquette heated during the molding process should be cooled down in atmosphere for longer than 20 min in order to form strong coal briquette in crushing. Moisture re-adsorption of the coal briquettes prepared with various oil components is
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Oviedo ICCS&T 2011. Extended Abstract
exhibited in Figure 6. As-molded coal briquettes have less than 3 wt.% of water, but water content of the samples left in atmosphere for 10 days goes up to 8 wt.%. It is obviously observed that the addition of oil components such as kerosene, asphalt and vinyl can retard moisture re-adsorption of the briquette, even though the samples with the oil component are saturated with water at the same water content of the sample without the oil component.
Figure 6. Moisture re-adsorption of briquettes of coals dried with various oil components.
4. Conclusions It is certain that the addition of vinyl as a heavy oil component in the NFCB process is favorable to increase the crushing strength of the product of coal briquette. Moreover, it is suggested that evaporation temperature of kerosene as light oil component in the NFCB process should not exceed 130 oC, and thus kerosene left in the dried coal particles can help to increase binding strength during the molding process of the dried coal. Crushing strength of coal briquette can be affected by several molding conditions such as time, pressure and temperature, which should be controlled to be 40 sec, 100 kgf/cm2 and 100 oC, respectively,
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Oviedo ICCS&T 2011. Extended Abstract
in order to obtain the proper strength. Furthermore, the heated coal briquette should be cooled down for longer than 20 min in atmosphere to maintain the crushing strength. Finally, it is confirmed that the coal briquette having oil components can retard moisture re-adsorption in the exposure of the sample to atmosphere.
Acknowledgement We acknowledge funding from Korea Institute of Energy Technology Evaluation and Planning (KETEP) for a project of power generation & electricity delivery.
References [1] Moon, S.-H. and Lee, I.-C., Moisture control in application of coal, Korea Institute of Energy and Resources (1990) in Korean [2] Chun, B.-S., Kumar, S., Park, Y.-S., Kim, J.-Y., Shin, S.-U., Shin, B.-W., Do, J.-N. and Kim, Y.-I., Application of Alternative Materials for a Fossil Fuel of Fuel by Vacuum Frying, Korean Geo-Environmental Conference Sep. 19 (2008) 107 in Korean [3] Sugita, S., Deguchi, T. and Shigehisa, T., UBC (Upgraded Brown Coal) Process Development, Kobe Steel Engineering Reports 53(2) (2003) 41-45 in Japanese [4] Hickey H., MacMillan B., Newling B., Ramesh M., Eijck, P.V. and Balcom B., Magnetic resonance relaxation measurements to determine oil and water content in fried foods, Food Research International 39 (2006) 612 [5] Li X.; Song H., Wang Q., Meesri C., Wall T. and Yu J., Experimental study on drying and moisture re-adsorption kinetics of an Indonesian low rank coal, Journal of Environmental Sciences Supplement (2009) 127 [6] Peregrina C., Arlabosse P., Lecomte D. and Rudolph V., Heat and mass transfer during fry-drying of sewage sludge, Drying Technology 24 (2006) 797 [7] Farkas B.E., Singh R.P. and Rumsey T.R., Modeling heat and mass transfer in immersion frying I. Model development, Journal of Food Engineering 29 (1996) 211 [8] Peregrina C., Rudolph V., Lecomte D. and Arlabosse P., Immersion frying for the thermal drying of sewage sludge: An economic assessment, Journal of Environmental Management 86 (2008) 246
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Oviedo ICCS&T 2011. Extended Abstract
Characteristics of dried low-rank coal by hot oil immersion drying Method for the upgrading Tae-In Ohm1, Jong-Seung Chae1, Seung-Hyun Moon2 1 Hanbat National University, San 16-1 Duckmyung-dong Yuseung-gu Daejeon, 2
S. Korea. E-mail :
[email protected] Korea Institute of Energy Research, 71-2 Jang-dong Yuseung-gu Daejeon, S. Korea.
Abstract Coal is the most abundant fuel on earth, and Low-rank coal (LRC) such as subbituminous coal and lignite makes up about half of all coal deposits. LRC is inconvenient to use due to its low caloric value and high moisture content, and because these oxygen-rich coals tend strongly toward spontaneous combustion. Solving these problems would substantially improve the efficiency of LRC use. In this study, we describe a drying technique utilizing hot oil immersion. This upgrading process may be executed under relatively low temperature and pressure, greatly reducing its energy cost. Drying tests of Indonesian lignite were performed with refined oil and B-C heavy oil, which were heated to 120 °C, 130 °C or 140 °C. Following 10 min of treatment, the moisture content of the upgraded coal was improved from 32 wt.% to 2.0-3.2 wt.%, and its high heating value from 12,500 kJ/kg to 25,100 kJ/kg. The coal drying curve shows that most of the moisture evaporated within about 5 minutes at an oil temperature of 140 °C, much more rapidly than occurs using heated gas drying methods. Further, there was a stepwise reduction in weight indicating rapid moisture evaporation to its final level. This occurred because moisture was boiled off at the same time that high temperature and vapour pressure in the coal led to hardening and carbonization. BET analysis indicated that moisture occupying fine pores was replaced by oil after the drying process, and that pore diameter, surface area and volume were all reduced. In the drying process using high viscosity B-C heavy oil, oil absorption became more significant as coal size decreased. This was due to absorption of the viscous oil in the pores of the smaller particles, which on a relative basis have a wider surface area. Coal size had less influence on the absorption of refined oil, which was more readily recovered from the smaller coal by centrifugal separation. Key Words: Low-rank Coal, Coal Drying, Oil Immersion Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
1. Introduction A stable security plan requires low-cost energy sources, but the present substantial demand for energy has elevated the prices of gasoline as well as coal. The upgrading of Low-rank coal (LRC) offers a means to reduce total energy costs. LRC accounts for about half of the world's coal deposits, and is relatively inexpensive at just 20-30% of the price of high-rank coal. However, the use of LRC has been extremely limited for several reasons. Its high moisture content of 30-70% makes it more difficult to transport over long distances, and its heating value is only 4,500 kcal/kg at best. Also, due to its high porosity and high carbonyl content, LRC has a strong tendency toward spontaneous combustion during long-term storage. Overcoming these disadvantages through an upgrading process would confer a significant economic benefit. In LRC upgrading processes, water is evaporated from surface and interstitial spaces, the moisture being replaced with oil. Hence, the heating value of LRC is increased, and the conversion from a hydrophobic to hydrophilic condition serves to stabilize it[2,3,4,5]. Coal drying technologies can be categorized as evaporative or nonevaporative, the latter category including processes known as the K-Fuel process (U.S.A.), the Binderless Coal Briquettes process (Australia) and the Upgrading Brown Coal process (Japan). In this study, the technique of immersion in hot oil was used to dry LRC through heat and mass transfer mechanisms. Coals were added to oil heated above the evaporation temperature of moisture, generating a strong turbulent flow on the coal surface by boiling. Through heat transfer, both surface water and interstitial water evaporated rapidly. By this drying technique, hereinafter called 'fry drying', LRC could be upgraded to a moisture content of 5% or less and a heating value of 6,000 kcal/kg or more, owing to replacement of moisture in the coal with oil.
2. Experiment Deposits of Indonesian lignite are abundant, and its advantageous geographic distribution helps to keep transportation costs low. For these reasons, Indonesian lignite was selected as the test sample to be upgraded. For use in experiments, block coals were mechanically crushed and then passed through 2-3 mm, 6-7 mm and 10-11 mm screens, such that specific coal sizes could be tested. Basic characteristics of coal samples were measured using a Truspec CHN Elemental
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2
Oviedo ICCS&T 2011. Extended Abstract
Analyzer (LECO), SC -432DR Sulfur Analyzer (DIONEX), Analyze (LECO), IC-2000 Analyze (LECO) and TG A-701 Proximate Analyzer. Results of proximate, ultimate, and heating value analyses are shown in Table 1. Table 1. Characteristics of Indonesian Low Rank Coal. Proximate Analysis (as received base) Moisture (wt%)
Ash (wt.%)
Fixed carbon (wt.%)
Volatile matter (wt.%)
32.3
2.0
34.6
31.1
Ultimate Analysis (as received base) C (wt.%)
H (wt.%)
O (wt.%)
N (wt.%)
S (wt.%)
Cl (ppm)
43.5
4.1
17.5
0.6
not done
not done
Calorific Values High calorific value (kcal/kg)
Low calorific value (kcal/kg)
3,230.8
2,864.2
Refined waste oil and B-C heavy oil were used as drying media. Refined oil is recycled waste oil from which moisture, ashes, heavy metals and other contaminants had been removed. B-C heavy oil, the most viscous of the heavy oils with a minimum viscosity of 50 cst (50 °C), is used as fuel for burners equipped with a preheating and warming apparatus, such as large-scale boilers and low-speed diesel engines. The refined oil used in this study had a boiling point of approximately 340 °C and a specific gravity of 0.856-0.86; the B-C heavy oil had a boiling point of about 340 °C and a specific gravity of 0.92-0.95. Fig. 1 shows the design for a batch-type coal drying apparatus. The cylindrical reactor (height 23 cm, diameter 20 cm) and a square mesh net receptacle for coals to be fed into the reactor (width 10 cm, length 10 cm, height 2.5 cm) were made of stainless steel. The weight of the oil in the reactor was measured with an electronic scale placed under the drying apparatus. An automatic temperature controller was provided to precisely control the reactor temperature. A notebook computer recorded changes in temperature and weight as functions of heating time of the oil and drying time of the coal. The oil was preheated to a stable temperature of 120 °C, 130 °C, or 140 °C before the Submit before May 15th to
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3
Oviedo ICCS&T 2011. Extended Abstract
addition of 50 g of coal per litre of oil. Each fry-drying experiment was performed for 10 minutes. After drying, coals were transferred to a centrifugal separator for 10 minutes, and the amount of separated oil was then measured. To assess drying efficiency under each reaction condition, moisture content was measured by electric oven in samples collected before and after the drying process, and crushed to a uniform size of 0.25 mm. Moisture measurements were completed the same day to preclude changes in moisture content. For all parameters measured, the arithmetic mean of three experiments was calculated.
Fig. 1. Batch-Type Coal Drying Apparatus.
3. Results and Discussion 3.1 Drying curves The drying process of a general solid can be divided into three periods, namely, a preheating period, a constant-rate drying period and a falling-rate drying period. In the brief preheating period, the solid initially warms and the moisture content slowly decreases. In the subsequent constant-rate drying period, the evaporation rate at the surface of the solid is equal to the internal diffusion rate, and there is an abrupt, linear decrease in moisture content. Lastly in the falling-rate drying period, the evaporation rate at the surface increases while the internal diffusion rate diminishes, and thus moisture content decreases slowly until a minimum is reached. This scenario will vary according to the shape of the solid and conditions of the heated gas such as temperature and humidity.
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Oviedo ICCS&T 2011. Extended Abstract
A drawback in the use of heated gas for solid drying is the length of time required for both the constant-rate drying period and the falling-rate drying period. To solve this problem, in this study we have investigated the drying characteristics of an alternative process that utilizes boiling heat transfer. The coal drying curve in Fig. 2 shows that most of the moisture evaporated within about 5 minutes at an oil temperature of 140 °C, much more rapidly than occurs using heated gas drying methods. Further, there was a stepwise reduction in weight indicating rapid moisture evaporation to its final level. This occurred because moisture was boiled off at the same time that high temperature and vapour pressure in the coal led to hardening and carbonization. 860 850
Weight(coal+oil)
840 830 820 coal input
810
coal diameter : 2-3mm coal weight : 50g oil temp : 140 C
800 790 780
0
2
4
6
8
10
12
14
16
18
20
Fry-dry time (min)
Fig. 2. Drying Curve of Raw Indonesian Lignite. 3.2 Results of TGA analyses Coals 2-3 mm in diameter, before and after drying in B-C oil at 140 °C, were TGAanalyzed in nitrogen or in air at a heating rate of 20 °C /min. The results are shown in Figs. 3. Raw coal lost mass in either nitrogen or air both before reaching a temperature of 100 °C and after, as moisture phase-changed to vapour. In contrast, coal first subjected to the drying process lost mass gradually over a temperature range of 100150 °C. This was apparently caused by decomposition of organic material present in the coal. The reduction of mass in coal dried in B-C oil took place at slightly higher temperatures than with coal dried by refined oil, due to the characteristically high viscosity and high density of B-C heavy oil.
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5
Oviedo ICCS&T 2011. Extended Abstract
90
90
80
80
70
70
Weight (%)
100
Weight (%)
100
60 50 40 30
0
0
100
200
300
60 50 40
20
Temp : 20 oC / min Gas : N2
10
Temp : 20 oC / min Gas : air
30
raw-coal B-C oil drying Refined-oil drying
20
raw-coal B-C oil drying Refined-oil drying
10
400
500
600
700
800
0
900
0
100
200
300
400
500
600
700
800
900
Temperature ( oC)
Temperature ( oC)
Fig. 3. TGA Curves of Raw and Upgraded Indonesian Lignite in N2(left) and Air(right) (Heating rate 20 °C /min). 3.3 Effectiveness of fry-drying The coal contained 27-30% moisture before the drying process; moisture contents determined after the fry-drying process are shown in Fig. 4. Moisture content was reduced to 2.6-9% after drying with B-C oil, and 2-7% when dried with refined oil. The effectiveness of fry-drying was greater at higher oil temperature and with smaller coal size. 10
10 2-3mm
9
9
2-3mm
4-5mm 10-11mm
8
Moisture(%)
Moisture(%)
10-11mm
7
7 6 5 4
6 5 4
3
3
2
2
1
1
0
6-7mm
8
120
130
B-C Oil Temperature ( o C )
140
0
120
130
Refined Oil Temperature ( oC ) C
140
Fig. 4. Moisture Contents of Upgraded Coal Dried in B-C Oil (upper panel) or in Refined Oil (lower panel). 3.5 Oil absorption by coal In the drying process using high viscosity B-C heavy oil, oil absorption became more significant as coal size decreased (Fig. 5). This was due to absorption of the viscous oil in the pores of the smaller particles, which on a relative basis have a wider surface area. Coal size had less influence on the absorption of refined oil, which was more readily recovered from the smaller coal by centrifugal separation. Submit before May 15th to
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6
12
12
10
10
Oil Absorption Amount (g)
Oil Absorption Amount (g)
Oviedo ICCS&T 2011. Extended Abstract
8
6
4
2-3mm
2
8
6
4
2-3mm
2
6-7mm
6-7mm
10-11mm
0
120
130
o C ( C ) B-C Oil Temperature
10-11mm
0
140
120
130
Refined Oil Temperature ( o C)
140
Fig. 5. Amount of Oil Absorbed by Coal Upgraded in B-C Oil (upper panel) and in Refined Oil (lower panel). 3.6 Oil recovery rate and net oil consumption Oil was recovered from dried coal by centrifugal separation for 10 minutes. B-C heavy oil could not be recovered due to its high viscosity. Rates of refined oil recovery and net consumption are shown in Fig. 6. The rate of recovery was independent of both drying temperature and coal size. Net oil consumption increased with drying temperature, regardless of the coal size. 2
2 2-3mm 6-7mm
Net Oil Loss Rate (%)
Oil Seperation Rate (%)
10-11mm
1.5
1
0.5
1.5
1
0.5 2-3mm 6-7mm 10-11mm
0
120
130
Refined Oil Temperature ( o C)
140
0
120
130
Refined Oil Temperature ( o C )
140
Fig. 6. Oil Recovery Rates (upper panel) and Net Rates of Consumption (lower panel) for Centrifugal Separation of Refined Oil. 4. Conclusions We have developed a new, rapid method for the upgrading of LRC. Drying tests were performed on Indonesian lignite (moisture content 32%, high heating value 3,000 kcal/kg) using immersion in either refined oil or B-C heavy oil that had been heated to a
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7
Oviedo ICCS&T 2011. Extended Abstract
temperature of 120 °C, 130 °C or 140 °C. Using this fry-drying process, it was possible to obtain upgraded coal with a moisture content of 3% or less and a high heat value of approximately 6,000 kcal/kg.
Acknowledgement. This work was supported by the New & Renewable Energy of the Korea Institute of Energy Technology Evaluation and Planning(KETEP) grant funded by the Korea government Ministry of Knowledge Economy(No. KETEP : 20103020100010).
References [1] Sakaguchi, M. et al., Hydrothermal upgrading of Loy Yang Brown Coal - Effect of upgrading conditions on the characteristics of the products, Fuel Processing Technology 89 (2008) 391-396 [2] Renfu, X. et al., Effects of chemicals and blending petroleum coke on the properties of low-rank Indonesian coal water mixtures, Fuel Processing Technology 89 (2008) 249-253 [3] Nugroho, Y. and McIntosh, A., Low-temperature oxidation of single and blended coals, Fuel 79 (2000) 1951-19 61 [4] Küçük, A. and Kadioglu, Y., A study of spontaneous combustion characteristics of a Turkish lignite; particle size, moisture of coal, humidity of air, Combustion and Flame 133 (2003) 255-261 [5] Sonata, W. and Zhang, D., Low- temperature oxidation of coal studied using wiremesh reactors with both steady-state and transient methods, Combustion and Flame 117 (1999) 646-6516. Sugita, T. et al., UBC (Upgraded Brown Coal) Process Development, in, Kobe Steel Engineering Reports, p. 53, 2003 [6] Katalambula, H. et al., Low-Grade Coals, A Review of Some Prospective Upgrading Technologies, Energy & Fuels 23 (2009) 3392-34058. Morimoto, M. et al., Hydrothermal extraction and hydrothermal gasification process for brown coal conversion, Fuel 87 (2008) 546-551 [7] Lee, S. J., Shin, H. Y, Bae, I. K., Chae, S. C., Report on upgrading technology for low-rank coal, Journal of the Korean Society for Geosystem Engineering 45 (2008) 276 -282 [8] Ohm, T., Chae, J., Kim, J., Kim, H. and Moon, S., A study on the dewatering of the industrial waste sludges by fry-drying technology, Journal of hazardous materials 168 Submit before May 15th to
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8
Oviedo ICCS&T 2011. Extended Abstract
(2009) 445-450 [9] Ohm, T. I., Chae, J. S., Kim, H. K., Lim, K. S., Kim, M. J., Moon, S. H., An experimental study on the drying characteristics of the sludge in the EDIHO using drying curve, Journal of Korea Society of Waste Management 26 (2009) 312- 318 [10] Ohm, T. I., Chae, J. S., Kim, J. E., Kim, H. K., Moon, S. H., A study on dewatering characteristics of the industrial waste water sludge using fry-drying technology, Journal of Korea Society of Waste Management 25 (2008) 225-231 [11] Ohm, T. I., Chae, J. S., Kim, J. E., Kim, H. K., Lim, K. S., Moon, S. H., A study on the characteristics of evaporative drying in immersed hot oil for summer and winter sewage sludge, Journal of Korea Society of Waste Management 26, (2009) 78-85
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9
ENERGY AND EXERGY ANALYSIS OF CONTINUES MICROWAVE DRYING OF COAL M. B. Alvarado1∗, E. J. Muñoz1, S.C. Navarro2∗∗, F. Chejne1, H. Velazquez1 1 Grupo de Termodinámica Aplicada y Energías Alternativas, Escuela de Procesos y Energía, Universidad Nacional de Colombia, Sede Medellín, 2 Investigación y Desarrollo de Tecnologías y Procesos, Cementos ARGOS.
Abstract The aim of this study is evaluate the energetic and exergetic efficiency of coal drying process in a microwave tunnel-type oven, using a Colombian coal for the experiments. The sample has moisture content around 20% and the process remove an 8% by weight, it is assumed that the carbonaceous material does not vary in elemental composition, dielectric and physicochemical properties with increasing temperature and moisture change due to microwave radiation. The design parameters of the equipment taken for analysis are based on data from a research group TAYEA of the Universidad Nacional de Colombia associate with Investigación y Desarrollo de Tecnologías y Procesos,ARGOS Company and financed by them. It was found that the energy drying process efficiency is equal to the second law around 69% which is explained by the quality of electromagnetic waves exergy. Keywords: Microwave, coal, exergy, energy.
1. INTRODUCTION The moisture content in coal is an important parameter that affects the thermal combustion efficiency, handling and transportation cost, and allocation of the price. Some parameters affect the amount of moisture reduction with microwave treatment such as: density, particle diameter, mass processing, temperature, location in the dry chamber, material composition where are considering the water, ash and minerals amount, the presence of metallic elements, and volatile matter [11].
∗
Corresponding author:
[email protected] [email protected]
∗∗
Pyrite as water are the main source of heating of the carbonaceous structure during microwave exposition [2]; Microwave heating can lead fractures due to thermal stresses between the present phases in the material. Thermal stresses are result of the different capacities of absorber microwave in coal’s materials that generate differentials of temperatures inside the material. Electric permittivity and magnetic permeability of each phase are the properties that produce these differences. Coal is not a good microwave absorber because their relative electrical permittivity as a volumetric material is lower than 0.1 (at 2.45GHz and 25°C), otherwise for water is around 12 at the same conditions resulting more susceptible to microwave heating than coal [12]. The density of energy carried by microwaves is proportional to the square of electric field magnitude incident to the material and is represented by Lambert-Beer equation [9] Equation 1), therefore the physical strength ( between coal particles can be increased when high density of microwave energy is applied, this effect generates large intermolecular tensions that eventually cause fractures within coal matrix [13][2]. Equation 1 The instantaneous power in an electromagnetic wave is described by the Poynting vector, whose value at the material surface takes the form of Equation 2, which depends on the maximum field value. Equation 2 The attenuation factor β of heat generation presented in Equation 1 is determined by Equation 3: Equation 3 Where (
the frequency of electric field ( ), ), is the electric permittivity and
is the light speed in vacuum space indicated such a material can be
penetrated by an electric field, causing it to warm up, and as this heat is dissipated (
).
For the coal and indicates the electrical energy distribution within a microwave irradiated material and affects the efficiency of energy transfer from the microwave oven to the product. Thus the attenuation factor will determine the energy distribution of energy within a material [10]. The thickness of the depth of penetration ( ) of the applied power depend on the dielectric properties of the material [1], to coal it takes a value of 0.05m.
Previous studies have found coal drying efficiencies around 83%, in tunnel-type microwave oven where coal sample passes through a conveyor belt, equipped with insulation systems and Remote control [12]. The selection of this application depends on the market behavior of coal and the use the use will be given to it: pyrolysis, gasification, desulfurization, substance or catalyst absorber, etc. The Colombian coal provided by ARGOS has been studied in particle sizes less than 38mm, finding moisture removal up to 6% in times of 8 minutes without degrading other properties such as calorific value, Hard Grove Index, ash and sulfur content. We found that larger particles are those with higher moisture removal [14]. The interest aroused by the use of microwaves for drying coal justified the study of energy and/or exergy benefits in the process. The coal can be treated as a flow stream in the exergetic analysis of the process and also is discuss how the coal price affects the use of this technology; the above in order to stimulate the comparison of this technology with others conventional existing for drying coal. Some advantages of microwave heating application have been discussed through the literature: - Substantial reduction of heating times and energy costs. Heating times may be even reduced at less than one percent required by conventional diffusive or convective heating techniques [3]. - Possibility of selective heating depending on the responses of different phases present in a material [5] [6]. - Ability to improve product quality and improved chemical synthesis [7], [8]. The evaluating system of this study is presented in the Figure 1 where could see mass and energy flows involved in the microwave dry process.
Figure 1. Scheme of microwave dryer tunnel-type. Due to the difference in the effects occurred between water and carbonaceous structure present in the coal flow on the conveyor belt at the Figure 2 it was divided into two streams, one for water
(mWo) and the other for the carbonaceous mass (mCo) even when there are at the same temperature (To). To the operation, the conveyor belt requires a voltage (V2) and a current (I2) and due to energetic inefficient heat (QC2) is released. The magnetron for generating microwave power (Po) requires a electrical consumption of voltage (V1) and current (I1), overheating is product of the energy inefficiency of the magnetron to prevent it and damage the equipment the Fan 2 is used, which circulates air mass (mAf) through a T0 temperature to Tf2 releasing heat (QC4) and has a energy consumption of voltage (V4) and current (I4). The power output (Po) is supplied to the drying chamber, in which one part is absorbed by the coal moisture on the conveyor belt causing the warming of the material to the outlet temperature (Tf) and evaporation of water which gives result an outflow (mcf + mWf) and the other part is considered like lost power (Pp). Fan 1 is used to remove water vapor generated (mVf) to a temperature (Tf1), consuming a power voltage (V3) and current (I3) and heat loss (QC3). All equipment of the drying system operating at the same voltage, so V1 = V2 = V3 = V4 = V = 220V.
2. MODELO TERMODINAMICO Y TERMOECONOMICO
2.1.
PARÁMETROS USADOS EN EL MODELO
The coal mined present a moisture of W0=20% and the process is able to removed until the 8%, the environment temperature is set at 25°C. The thermodynamic properties of the materials involved in the process are summarized in Table No 1 which were taken to 25°C of temperature and 1atm of pressure using Enginnering Equation Solver program (EES 8.609) licensed by Universidad Nacional de Colombia [15]. 50% was added to latent heat of vaporization which refers to the energy required to break the interaction of coal-steam and calculation thus find a latent heat of sorption. The energy magnetron efficiency was considered 70%, and this is fed with a voltage of 220V. Table No 1. Materials properties. Water Air Coal (kJ/kg °C) 3.979 1.005 1.260 PCI (kJ/kg) (kJ/kg)
----2256.8
-------
24000 ----
Efficiencies were evaluated for microwave drying system. Energetic efficiencies for microwave
drying and to the system are shown in Equation 4 and
Equation 5. Equation 4 Equation 5
Exergetic efficiencies for microwave drying and to the system are shown in Equation 6 and Equation 7. Equation 6 Equation 7
2.2.
THERMODYNAMIC MODEL
2.2.1. Mass Balance To the solution is necessary to perform the mass balance in the drying chamber for water, because the carbonaceous mass is unchanged as shown in Equation 8: Equation 8 Is necessary to specify that this mass of water is equal to the percentage of moisture present on coal . (
2.2.2. Energy Balance The energy balance considering the system shown in Figure 1 is presented by
Equation 9:
Equation 9 Where are the current consumed by the microwave generator, fan 1, fan 2 and the conveyor belt, respectively. The second and third term on the left side of the equation refers to the energy contained in the stream of wet coal and air input to the system. The first, second and fourth term of
Equation 9 represents the energy contained in the air, water and dry coal respectively. The third term on the right side of the equation is the energy consumed in water evaporation. The following terms represent the heat lost in the magnetron, the fan 1, fan 2, the conveyor belt and the drying chamber, respectively, due to the inefficiencies of the equipment. The two main parts (magnetron and dryer chamber) is necessary to present the energy balance as shown in Equation 1011
Equation 12 and 12, respectively: Equation 10
Equation 11 Where, is the maximum power capable to produce by the magnetron. 2.2.3. Exergy Balance Exergetic balance taking to account the system presented in Figure 1 is showing by Equation 13:
Equation 12 The Equation 12 present the physic and chemical exergy associated with the initial and final state of coal and air, and the steam stream at the outlet of the dryer chamber. The last term is associated with the exergy destroyed accounted for the losses associated with the system irreversibilities. The physical exergy of air is calculated as shown in equation 14: Equation 14
Equal to energy balance is necessary to perform an exergetic balance for the two main parts (Magnetron and Dryer Chamber) presented in Equation 135 and 16, respectively: Equation 135 Equation 146
2.2.3.1.
Chemical Coal Exergy
The total chemical exergy [4] is calculated by equation 17: Equation 157 The
chemical
fuels
exergy
(coal)
was
calculated
by
the
Equation 168 for elemental analysis free of ash and moisture
Equation 168 Where the temperature should be in Kelvin degrees, PCS is the caloric value of coal and mole fractions of each component which are calculated using the following relationships:
The
entropy
and
caloric
value
of
coal
[4]
are
is
described
by
Equation
17
and 20: Equation 179 Equation 20 The Table No 2 shows the elemental analysis performed at the coal sample before and after drying using a microwave oven with reference LG MS-1146SQP, tests were made in the laboratory of the Universidad Nacional de Colombia using an elemental analyzer (CHN) marks Exeter. Water vapor exergy was calculated with Equation .
Equation 21 Table No 2. Ultimate Analysis Output Element Input H 0.0484 0.0531 O 0.1314 0.1441 S 0.0222 0.0243 C 0.5387 0.5910 Ash 0.0477 0.0526 H2O 0.1974 0.1196 N 0.0112 0.0123
3.
RESULTS Y DISCUSSION
The results of mass, energy and exergy balance with maximum power input conditions are
¡Error! No se encuentra el origen de la referencia.
shown
in
and 4.
Table No 3. Energy Results Energy fluxes Value (kW) 10,00 VIMW
Table No 4. Exergy Results Exergy Fluxes Value (kW) 459,40
3,00
QS QD QP
4,81 0,57 0,35 1,32
Po
10,00 7,00 0,27 464,20
Figure 2 and 3 represent the energy and exergy Sankey diagram for the coal microwave dryer working at full power (10.00 kW).
Figure 2. Sankey diagram of energy fluxes Figure 22 present the energy flows of streams involved in coal drying process, showing that most of the input energy is needed to evaporate the water contained in coal matrix, in addition the magnetron dissipates a significant amount of energy that is fed into the system, the system efficiency of the drying process according to Equation 5 is 48%, while the process efficiency of the drying chamber according to Equation 4 is 69% . This indicates that much of the microwave energy is transferred to the vapor and the rest is stored as heat in the coal.
Figure 3. Sankey Diagram of Exergy Flux Figure 3 present the Sankey diagram associated with exergy flux in the microwave drying system, where it is clear that the useful stream is associated with the exergy contained in coal, the system is able to process 70kg/h; according by Equation 7, the exergetic efficiency of microwave drying process is 48%, and the exergetic efficiency in the dryer chamber is 69% (Equation 6). It is remarkable that energetic and exergetic efficiencies are equal, which is understandable due to the exergetic forms that interact within the system have quality 1, and this makes the exergy equal to the energy transport by streams getting the maximum benefit, that is a maximum energy efficiency.
4.
CONCLUSIONS
Equality between energetic and exergetic efficiencies of microwave drying process occurs because the exergy of the microwaves is equivalent to its energy content, which does not happen with conventional convective or conductive drying processes, where required heating dry air or other conductive material. These do not transfer all the energy contained in the process, which reduces significantly the efficiencies of these processes, also microwave drying has another advantage because it does not require long times to reach operating conditions necessary to initiate the process. Acknowledgement Authors acknowledge the financial support of INCARBO, Investigación y Desarrollo de Tecnologías y Procesos, ARGOS and Universidad Nacional de Colombia. References [1] [2] [3] [4] [5] [6] [7] [8] [9] [10] [11] [12] [13] [14]
A. Datta, Handbook of microwave technology for food applications. New York: M. Dekker, 2001. Marland, S., A. Merchant, and N. Rowson, Dielectric Properties of coal. Fuel, 2001. Meredith, R., Engineers' Handbook of Industrial Microwave Heating. The Institution of Electrical Engineers, London, 1998. A. Bejan, Thermal design and optimization. New York: Wiley, 1996. Harutyunyan, A.R., et al., Purification of single-wall carbon nanotubes by selective heating of catalyst particles. The journal of physical chemistry B, 2002. 106(34): p. 8671-8675. Wang, Y., E. Forssberg, and M. Svensson, Microwave assisted comminution and liberation of minerals. Mineral processing on the verge of the 21st century, 2000. Jones, D.A., et al., Microwave heating applications in environmental engineering - a review. Resources, Conservation and Recycling, 2002. 34: p. 75-90. Whittaker, G. and D.M.P. Mingos, The application of microwave heating to chemical syntheses. Journal of microwave power and electromagnetic energy, 1994. 29(4): p. 195219. Swami, S., Microwave heating characteristics of simulated high moisture foods. 1982, MS thesis, University of Massachusetts: Amherst, MA, USA. Ayapa, K.G., et al., Microwave heating: an evaluation of power formulations. 1991, Editorial Pergamon Press plc. Vol. 46, No. 4: Gran Bretaña. p. p. 1005 -1016. Ponte, D.G., et al., Determination of moisture content in power station coal using microwave. 1995, Oviedo University: Spain. Kalra, A., Dewatering of fine coal slurries by selective heating with microwaves. 2006, College of Engineering and Mineral Resources West Virginia University: Morgantown, United States. Yag˘mur, E., S¸ims¸ek, E. H., & Taner Tog rul, Z. A., Effect of demineralization process on the liquefaction of Turkish coals in tetralin with microwave energy: Determination of particle size distribution and surface area. Fuel, 2005. 84: p. 2316–2323. M. Alvarado, J. Mejia, M. Vanegas, L. Hernandez. Estudio de secado con Energía Microondas del Carbón de Bijao–Córdoba, in Programa de Ingeniería Química. 2009, Universidad del Atlántico: Barranquilla, Colombia.
[15]
Robert Perry, Perry's chemical engineer's platinum edition Perry's chemical engineers' handbook, 7th ed. (New York: McGraww-Hill ;, 1999). [16] Anual Energy Review, 2009
NOMENCLATURA B BD W Cp Emax PCI
c
Exergía Exergía Destruida Humedad Capacidad calorífica (kJ/kg K) Campo eléctrico máximo (V/m) Calor latente de vaporización del agua (kJ/kg) Poder calorífico inferior Potencia absorbida Potencia aplicada inicial Distancia de penetración de la onda microonda Constante dieléctrica relativa (habilidad para almacenar energía eléctrica) Perdida dieléctrica relativa (habilidad para disipar energía eléctrica) Frecuencia de aplicación de la microonda Velocidad de la luz en el vacio Permitividad de la luz en el vacio Masa del carbón húmedo Masa del carbón seco Masa de agua a la entrada Masa de vapor evaporado Masa de agua que sale con el carbón Masa de aire a la salida del ventilador 2
Eficiencia energética Eficiencia exergética Calor perdido por la ineficiencia del magnetrón Calor perdido por la ineficiencia del ventilador 1 Calor perdido por la ineficiencia del ventilador 2 Calor perdido por la ineficiencia de la banda transportadora Calor perdido por las ondas no absorbidas en la cámara de secado Fracción molar T Temperatura (C) Subíndices w agua MW Generador de microondas A aire V1 Ventilador 1 C coal V2 Ventilador 2 S vapor (steam) CB Banda Transportadora DC Cámara de secado 0 inicial f final
Moisture Re-Adsorption Characteristics of Coal Samples Dried by a Pneumatic Dryer
Sangdo Kim, Youngjun Rhim, Sihyun Lee:
Clean Coal Center, Korea Institute of Energy Research, 152 Gajeong-ro, Yuseong-gu, Daejeon, South Korea (305-343):
[email protected]
Deposits of low rank coal (LRC) represent nearly half of the estimated coal resources in the world. However, LRC is difficult to use in thermal power plants for two reasons: high moisture content (30% or more) and capacity for spontaneous combustion. To avoid such problems, upgraded technology to reduce the moisture content of low rank coal has been developed; however, dried coal can re-adsorb moisture when exposed to outside air. This paper discusses the results of moisture re-adsorption of coal samples dried by a pneumatic dryer. In addition, the results on moisture re-adsorption and BET surface area of coal samples after drying and equilibrium moisture content are discussed. The moisture content of coal samples past 25 days did not change greatly compared to the moisture content of coal samples measured after drying. BET surface area of the dry coal was decreased compared to raw coal, and the equilibrium moisture content was lower than that of raw coal.
1. Introduction Lignite and sub-bituminous coal are classified as low rank coal, and their reserve is so abundant as to represent more than half of the entire coal deposits in the world. However, only about a quarter of the low rank coal reserve is currently in use or production. This is because of the low power-generating efficiency of low rank coal due to high moisture content; furthermore, a great deal of energy is required to dry low rank coal in a power plant setting. In addition, low rank coal has a high risk of spontaneous combustion, and hence has restrictions regarding its long-term storage and transport. To resolve this problem, many studies have been undertaken to dewatering low rank coal[17]. These include Australia's BCB (Binderless Coal Briquetting) [1-2], which uses pneumatic drying; Japan's UBC (Upgrading Brown Coal) [3-5], which uses drying in the oil phase; Germany's WTA (fluidized-bed drying with internal water heat utilization)[6], which uses a fluidized-bed of steam; and the DryFineTMprocess[7] of the U.S., which uses a fluidized-bed.
Dried coal tends to re-adsorb moisture, which can be critical for the long-term storage and transport of dried coal. Kartikeyan [8, 9] conducted research on the moisture re-adsorption characteristics of dried coal and explored how to minimize re-adsorption. Xianchun [10] carried out an experimental study on drying and moisture re-adsorption kinetics of low rank coal from Indonesia. This research examines moisture re-adsorption characteristics of coal dried by a pneumatic dryer.
2. Experimental 2-1 Experimental Equipment For the experiment, Meng Tai coal from Inner Mongolia was used. Table 1 describes the results of the proximate, ultimate, and heating value analysis. The moisture content and heating value were 29.74 wt% and 4,270 kcal/kg in ARB(as received base) condition, respectively. Through a pulverizing and distributing processes, the coal particles had a size of 100~2,000 ㎛. Figure 1 illustrates a pneumatic dryer used in the experiment, consisting of five parts: (1) an LPG burner to provide high-temperature gas, (2) a feeder to provide raw coal, (3) a riser tube to dry coal with high moisture content, (4) a cyclone to separate gas and solid matters, and (5) a bag filter. The diameter and height of the riser tube were 40 mm and 5,000 mm, respectively. The coal was provided at the rate of 5 kg/hr. Drying temperature was measured at the mouth of the feeder providing the coal. The temperature of inlet gas was 400~600 ℃, and the flow rate of the gas in the riser tube was set at 20 m/sec. During the experiment, the LPG burner generated high temperature gas, and when it reached a certain temperature, raw coal was supplied. The coal went through the riser tube and was dried. The dried coal was separated from the high-temperature gas in the cyclone and stored in the storage tank to be discharged. The dried coal was then put in containers: one sample was stored with a cover, and the other samples were stored indoors with their covers open. Although the storage conditions varied slightly, the humidity was kept at 60-80 % and the temperature at 20-25 ℃.
To measure moisture content of the raw coal and the dried coal, 841 KF Titrando (Metrohm, Switzerland) with Karl Fischer method was used.
Table 1. Analyses of coal sample. Coal Name
Meng Tai
Proximate analysis (wt%, ARB) Moisture
29.74
Volatile matter
27.83
Ash
10.51
Fixed carbon
31.92
Ultimate analysis (wt%, ADB) Carbon
62.73
Hydrogen
4.11
Nitrogen
0.95
Oxygen
21.42
Sulfur
0.28
Heating value analysis (kcal/kg) ARB
4,270
ADB
5,730
Note: ARB - as received base, ADB - air dried base
2-2 Measurement of Equilibrium Moisture To examine the characteristics of moisture re-adsorption, equilibrium moisture was measured. Equilibrium moisture is the moisture coal can hold in a humid atmosphere; it is effectively the sum of adsorbed moisture filling micro-pores and moisture filling large pores by capillary action[11]. For this research, equilibrium moisture was measured according to the ASTM D 1412-07 standard method; the measuring method is shown in Fig. 2. Five grams of coal with diameters less than 1.18mm were used, and the experiment was conducted at an absolute humidity of 96-97 % and temperature of 30 ℃.
Fig. 1 Experiment equipment.
Fig. 2 Measuring method of Equilibrium moisture
3. Results and Discussion Dried coal tends to re-adsorb moisture. For this reason, it is very important to study the readsorption characteristics of coal in order to ensure effective storage and transport. Table 2 shows the results of the ultimate, proximate, and heating value analysis for the coal dried by pneumatic dryer. After the experiment, the moisture content of the raw coal dropped below 7 wt%. Kartikeyan [8] conducted moisture re-adsorption experiments after completely drying the coal at temperatures 75 ℃, 100 ℃ and 150 ℃. Re-adsorption experiment is under an ambient environment of about 80% humidity at room temperature of 27℃. According to the experiment result, the moisture content of dried coal increased by 10-13% in two to four days, due to readsorption, and the amount depended on drying temperature. The moisture content did not change any more after 15 days. Xianchun [10] noted that the moisture desorption and re-adsorption isotherms of coal showed irreversibility in the desorption-adsorption cycle, indicating that irreversible coal structure changes occurred due to the shrinkage of the coal that took place during thermal drying, which is dependent on drying temperature. In addition, the higher the drying temperature, the more intensely the pore structure of coal collapsed, resulting in less surface area on coal particles to re-adsorb the moisture from the surrounding atmosphere [10]. Fig. 3 shows the results of re-adsorption moisture content when the dried coal was stored in closed condition. Dried with inlet gas at a temperature of 400℃, the coal re-adsorbed 30% of the moisture in two days; afterwards, the moisture content did not alter significantly. At drying temperatures of 500℃ and 600℃, the dried coal re-adsorbed some 5% of moisture, even after 24 days had elapsed. Fig. 4 shows the results of re-adsorption moisture content when the dried coal is in atmospheric condition. Dried with inlet gas at a temperature of 400℃, the coal re-adsorbed 55% of moisture in two days. After 24 days, about 75% of moisture was re-adsorbed. At drying temperatures of 500℃ and 600℃, there was no initial re-adsorption of moisture, but the moisture content gradually increased up to 55%.
The results of re-adsorption moisture content for dried coal showed similar results as the research outcome of Kartikeyan [8]. It was also shown that re-adsorption moisture content depended on drying temperature. Table 2. Analyses of dried coal. Inlet gas temperature(℃)
400
500
600
Proximate analysis (wt%, ARB) Moisture
6.61
6.54
6.1
Volatile matter
38.1
38.19
38.74
Ash
14.16
15.17
14.64
Fixed carbon
41.13
40.1
40.52
Ultimate analysis (wt%, ADB) Carbon
61.05
60.05
61.45
Hydrogen
4.05
4.03
4.22
Nitrogen
1.08
1.08
1.10
Oxygen
19.56
19.56
18.5
Sulfur
0.1
0.11
0.09
Heating value analysis (kcal/kg) ARB
5,537
5,586
Note: ARB : as received base, ADB : air dried base Fig. 2 Measuring method of Equilibrium moisture
5,636
Fig. 3. Re-adsorption moisture content of dried coal in closed condition.
Fig. 4. Re-adsorption moisture content of dried coal in atmospheric condition.
Table 3 shows the results of equilibrium moisture measurement. Five types of samples were used: raw coal, coal dried for 1 hour in the oven at 107℃, and three more samples dried with a pneumatic dryer at different temperatures. For the first two samples, the equilibrium moisture content was similar, at 25wt%. For the rest of the samples, the equilibrium moisture content decreased with rising gas temperature.
Table 3. Equilibrium moisture. Sample
Equilibrium moisture content (%)
Raw coal
25.52
Oven drying coal
25.06
400℃ dried coal
18.66
500℃ dried coal
18.59
600℃ dried coal
17.43
As Xianchun [10] presented, the equilibrium moisture content decreased compared to raw coal, because of internal structural changes that take place in coal dried with a pneumatic dryer. To verify this, BET surface areas of raw coal and dried coal were measured (Table 4). Coal dried with a pneumatic dryer showed greater decrease of BET surface area than raw coal. The amount of decrease also depended on drying temperature, which indicates that surface area shrinks as moisture rapidly evaporates upon contact with high-temperature gas.
Table 4. BET surface area. Coal sample
BET surface area(m2/g)
Raw coal
11.7035
400℃
6.9129
500℃
6.3596
600℃
5.5317
Conclusion In this research study, moisture re-adsorption characteristics of coal dried with a pneumatic dryer were examined. Compared with the initial moisture content, the dried coal re-adsorbed more than 55% of moisture, but after 24 days the overall moisture content did not change substantially. The equilibrium moisture content of dried coal was below 18.6wt%, and that amount decreased with higher gas inlet temperature. As the drying temperature increased, BET surface area decreased, because the moisture in the coal evaporated rapidly upon contact with the high-temperature gas of the pneumatic dryer. The experiment results suggest that, when dried high moisture content coal with a pneumatic dryer, the internal structural change of coal can help to effectively prevent moisture re-adsorption.
Acknowledgement This work was supported by the Power Generation & Electricity Delivery of the Korea Institute of Energy Technology Evaluation and Planning (KETEP) grant funded by the Korean government Ministry of Knowledge Economy.
References 1. Mangena, S.J., Korte, G.J., McCrindle,R.I. and Morgan, D.L., "The amenability of some Witbank bituminous ultra fine coals to binderless briquetting", Fuel Processing Technology, 85, 1647–1662(2004) 2. Keith C., "Commercial scale low rank coal upgrading using the BCB process", 2nd Coaltans Upgrading Coal Forum, Presentation(2010) 3. Sugita, S., Deguchi, T. and Shigehisa, T., "Demonstration of a UBC process in Indonesia", 神戸製鋼技報, 56(2), 23-26(2006) 4. Drtin, F.U., Hiromoto U. and Bukin, D., "Change of combustion characteristics of Indonesian low rank coal due to upgraded brown coal process", Fuel Processing Technology, 87, 10071011(2006) 5. Yukio A., "UBC Process - Upgrading the Future", 2nd Coaltans Upgrading Coal Forum, Presentation(2010)
6. Klutz, H.J., Moser, C. and Block, D. “WTA Fine Grain Drying – Module for Lignite-Fired Power Plants of the Future”, VGB Power Tech Report 11(2006) 7. Sarunac, N. Ness, M. and Bullinger, C., "One year of operating experience with a prototype fluidized bed coal dryer at coal creek generating station", National energy technology laboratory. 8. Karthikeyan, M and Mujumdar, A.S., "Factors affecting quality of deiced low rank coal", Drying Technology, 25(10), 1601-1611(2007) 9. Karthikeyan, M and Mujumdar, A.S., "Minimization of Moisture Re-adsorption in Dried Coal Samples Drying Technology", Drying Technology, 26(7), 948-955(2008) 10. LI Xianchun, Song Hui, Wang Qi, Meesri Chatphol, Wall Terry and YU Jianglong, "Experimental study on drying and moisture re-adsorption kinetics of an Indonesian low rank coal", Journal of Environmental Science Supplement, S127-S130(2009). 11. Barry Ryan, “A discussion on moisture in coal implications for coal bed gas and coal utilization”, ISSA(International Social Security Association), Summary of Activities, 139-149, 2006
Oviedo ICCS&T 2011. Extended Abstract
Moisture Readsorption Characteristics of Upgraded Low Rank Coal Hokyung Choi, Sangdo Kim, Jiho Yoo, Donghyuk Chun, Jeonghwan Lim, Youngjoon Lim, Sihyun Lee Korea Institute of Energy Research, 152 Gajeong-ro, Yuseong-gu, Daejeon 305-343, Republic of Korea:
[email protected] Abstract This study reports the moisture readsorption characteristics of dried coal produced from low rank coal using the upgraded brown coal (UBC) process. The moisture readsorption ratio rapidly increased only during the first 10 hours and formed equilibrium at a certain level thereafter. Raw coal showed the highest moisture readsorption ratio among the samples to form equilibrium at around 14 %. In the case of dried coal, as the amount of asphalt increased, the moisture readsorption ratio decreased. The appropriate portion of asphalt added during drying process is around 1.0 %. Further, increasing drying pressure negatively affects moisture readsorption characteristics of dried coal.
1. Introduction Low rank coal sees little use in spite of its large reserves and low price because it has high moisture content and low heating value. However, the recent high price of oil and high rank coal has encouraged many industries to use low rank coal. Upgraded brown coal (UBC), a well-known coal drying process, enables the utilization of low rank coals. The UBC process removes moisture in the coal using the slurry dewatering process, which involves the evaporation of moisture in the slurry of coal and light oil containing a small portion of heavy oil (such as asphalt) at temperatures of 130–160 °C under a pressure of 0.4 to 0.45 MPa [1,2]. Moisture in the air generates heat when adsorbed or condensed by coal, and this leads to temperature increases of the coal, promoting its spontaneous combustion. Drying coal usually improves the susceptibility to the spontaneous combustion of coal. However, moisture readsorption on the coal inevitably occurs in the absence of a stabilization treatment and causes the loss of the drying effects. To explore the efficient use of low rank coal, this study investigated the moisture readsorption characteristics of upgraded low rank coal produced by the coal-oil slurry dewatering process. To this end, proximate properties, BET surface area, and moisture readsorption characteristics of the coal were analyzed. Submit before 31 May 2011 to
[email protected]
1
Oviedo ICCS&T 2011. Extended Abstract
2. Experimental This study used an Indonesian lignite (KBB coal) as raw coal. The raw coal was upgraded as follows, referring to the procedure of the UBC process [3]. The raw coal was ground and sieved to obtain 0.5–1.3 mm particle size. It became slurry by mixing with kerosene solvent. Up to 4.0 wt% of asphalt, as a heavy oil additive, was added to the slurry. The weight ratio of the solution (kerosene + asphalt) and the raw coal was adjusted to 1:1. The slurry was retained at 140 °C for 30 minutes to evaporate the moisture in the coal. The pressure of the reaction vessel varied in the 0.1–0.3 MPa range. After the evaporation, the solid coal was filtered and dried in an oven at 130 °C to remove the remaining solvent. Before the analyses, other than proximate analysis and BET surface area analysis, raw coal and upgraded coal samples were dried again at 105 °C for 5 hours in nitrogen to remove the moisture that might have been adsorbed while the sample was being preserved. The moisture readsorption of a coal sample was characterized by monitoring a weight change of the coal sample caused by moisture readsorption in a humidity chamber and expressed as a ratio of the changed weight to the initial weight of the coal sample. The coal sample was placed in the chamber and maintained at a temperature of 30 °C and a relative humidity of 100 %. The weight change was monitored every hour for the first 10 hours and then every 24 hours for the following five days.
3. Results and discussion Table 1 shows the results of proximate analysis and BET surface area analysis of coal samples. In the table, the upgrading pressure refers to the pressure of the moisture evaporator in the upgrading process and the asphalt concentration denotes the weight percentage of the asphalt from the total weight of the slurry. Significantly, while going through the UBC process, the moisture content in the coal decreased from 15.96 % in the raw coal to 3.22 % at the minimum in the upgraded coal. As the moisture content in the coal decreased, the heating value of the coal increased from 5448 kcal/kg in the raw coal to 6623 kcal/kg at the maximum after upgrading. As upgrading pressure increased, the moisture content of the upgraded coal slightly increased. This occurs because of the increase in boiling temperatures along with the increase in upgrading pressure.
Submit before 31 May 2011 to
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2
Oviedo ICCS&T 2011. Extended Abstract
Table 1. Proximate and BET surface area analysis results of coals (air dried basis). Coal type
Upgrading pressure (MPa)
Asphalt concentration (%)
Moisture
Ash
(%)
raw
-
-
0.1
upgraded
0.2
0.3
(%)
Volatile matter (%)
Fixed carbon (%)
Heating value (MJ/kg)
BET Average surface pore width (m²/g) (nm)
15.96
4.25
46.78
33.01
27.14
26.20
10.31
0.0
9.45
4.30
49.94
36.31
28.57
20.11
13.09
0.5
4.41
4.63
52.65
38.31
28.31
17.70
13.80
1.0
4.11
4.48
52.91
38.50
28.35
16.13
14.23
2.0
3.26
4.56
53.72
38.46
28.22
12.18
15.59
4.0
3.22
5.08
53.56
38.14
28.65
10.71
16.09
0.0
9.53
4.43
50.06
35.99
27.61
21.21
11.55
0.5
4.85
4.41
52.60
38.14
27.67
17.68
12.57
1.0
3.92
4.99
53.61
37.48
27.68
16.85
12.61
2.0
3.87
4.17
53.95
38.01
27.85
10.92
14.56
4.0
3.79
4.80
53.56
37.85
27.95
8.99
15.20
0.0
9.96
4.38
50.11
35.55
27.67
17.00
12.79
0.5
5.58
4.93
52.21
37.29
27.54
15.67
12.01
1.0
5.41
4.25
53.07
37.27
27.20
13.83
12.38
2.0
5.35
4.05
52.71
37.89
27.46
9.20
14.10
4.0
5.33
4.53
52.37
37.77
27.44
8.68
14.45
When comparing changes in moisture content after adding the different amounts of asphalt, whereas the moisture contents ranged around 9–10 % in the case of the upgraded coal with no asphalt addition, they stayed lower than 6 % in all cases of upgraded coal with a 0.5 % or higher percentage of asphalt added. Upgraded coal with no asphalt addition shows high moisture content probably because of the readsorption of moisture in the process of handling after upgrading. The BET surface area of raw coal reached 26.2 m²/g. In upgraded coal, with the upgrading pressure at 0.1 MPa, as the amount of asphalt increased from 0.0 % to 4.0 %, the BET surface area decreased from 20.1 m²/g to 10.7 m²/g. In addition, the average pore size increased from 13.09 nm to 16.09 nm because the asphalt blocks the fine pores more easily than it blocks the larger pores. No large changes in BET surface area and pores resulted from increases in upgrading the pressure. Figure 1 shows the moisture readsorption characteristics of the raw and upgraded coal. As shown in the figure, the moisture readsorption ratio rapidly increased during the first 10 hours only and reached equilibrium at a certain level thereafter. Raw coal showed the highest moisture readsorption ratio among the samples to form equilibrium at around
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3
Oviedo ICCS&T 2011. Extended Abstract
1.140. In the case of upgraded coal, as the amount of asphalt increased, the moisture readsorption ratio decreased. This would take place because of the suppression of moisture readsorption by the asphalt with its hydrophobic nature. When the upgrading pressure increased, moisture readsorption ratios increased slightly even in the upgraded coal with asphalt added. When the content of the asphalt reached 4.0 %, the equilibrium moisture readsorption ratio of the coal upgraded at 0.1 MPa reached 1.104, that of the coal upgraded at 0.2 MPa reached 1.124, and that of the coal upgraded at 0.3 MPa reached 1.130. The effect of upgrading pressure on the moisture readsorption remains unclear. The drying pretreatment seemed to expose the adsorption sites of the upgraded coal to the air by removing moisture that had not been completely
Figure 1. Moisture readsorption ratios of coals upgraded at (a) 0.1 MPa, (b) 0.2 MPa, and (c) 0.3 MPa.
removed under enhanced pressure.
References [1] Sugita S, Deguchi T, Shigehisa T, Katsushima S, Makino E, Otaka Y. Demonstration of UBC process in Indonesia–Upgrading of low rank coal. Proc of the International Conference on Coal Science and Technology, Okinawa 2005. [2] Kinoshita S, Yamamoto S, Deguchi T, Shigehisa T. Demonstration of upgraded brown coal (UBC) process by 600 tonnes/day plant, Kobelco Technology Review 2010;29:93-98. [3] Umar DF, Usui H, Daulay B. Change of combustion characteristics of Indonesian low rank coal due to upgraded brown coal process. Fuel Processing Technology 2006;87:1007-1011.
Submit before 31 May 2011 to
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4
SOLAR DRYING TECHNOLOGY OF COAL IN THE OPEN IN THE IRON AND STEEL RESEARCH CENTER Authors: Andrey Leyva Mormul (
[email protected]) Ariel Díaz Castillo (
[email protected]) Oscar Sinecio Leyva González (
[email protected]) José Anival Trotman Gavilán (
[email protected]) Oscar Figueredo Stable (
[email protected])
Address: Centro de Investigaciones Siderúrgicas, Calle 72, No.8-F, S/C, La Pasa, Nicaro-Levisa, Mayarí, Holguín, Cuba
ABSTRACT The job presents the results achieved by the Iron and Steel Research Center after the implementation of solar drying of coal in the open air, prior to conventional drying oven or rotary drum cylinder. This has brought an increase drying efficiency in the production process, and involves a parallel decrease in the volume of products of combustion gases, propellants maximum greenhouse effect, acid rain and desertification of soils. From the quantitative evaluation, we determined that for each 1 % moisture lost by the coal exposed to sunlight, saves 4 liters of diesel fuel per ton of coal produced with the prior application of solar drying, systems, lowering the moisture content (W) of coal to be processed in about 8 % to reach equilibrium moisture content is 2 %, while minimizing the fuel consumption per ton of coal drying, 40 to between 19 and 22 l, and increase furnace productivity by 0.5 to 1.3 t/h, which brings economic benefit of 58215.46 USD over a year.
INTRODUCTION Energy education not only contributes to a better and more efficient use of fossil fuels available to mankind, but also a guarantee in the transition to a sustainable energy economy that rests on the available solar energy [Retired, et al ., 2007]. The coal used as loading, swelling and adjustment in the steel company José Martí Antilles Steel and Stainless Steel Enterprise ACINOX Tunas, is prepared mechanically in DSIT coal plant. Swelling coals and setting are subjected to the process of milling, sorting and drying in a rotary drum furnace for the removal of moisture contained in these materials, then incorporated into a crushing and screening equipment CMD 27, with which guaranteeing the fineness required by users. Among the most important aspect we have raised, in terms of reducing fuel consumption, is the efficient implementation of the drying process in the rotary cylinder. To contribute to this, we implemented the solar drying of coal in the open air, prior to drying in the oven rotating cylinder what has been successful with regard to the efficiency of the production process and obtain a higher quality product. Before proceeding to explain the procedure used and presents the results, some interesting elements are required:
•
In each square meter of Cuban territory, we have already received a daily amount of solar energy of 5 kWh, roughly equivalent to half a kilogram of fuel oil, average virtually unchanged throughout the year [Bérriz and Alvarez, 2008].
•
Coal, according to their physical characteristics, is a hygroscopic mineral, it has the ability to impregnate or exhale moisture depending on the environmental conditions in which they are placed, and is clearly black, which absorbs between 90 and 98 % sunlight without reflecting it later.
TECHNOLOGY FLOWS WITH AND WITHOUT SOLAR DRYING
Final Product Hopper
Hopper 2
Big Bag
Final Product Hopper
Big Bag
Balance
Hopper 1 Combustion Chamber
Sieve Shaker Dryer CM B E
Coal
CM: Cone Mill BE: Bucket Elevator
Fig 1. Preparation of coal without the application of solar drying.
Big Bag
Final Product Hopper
Hopper 2
Big Bag
Final Product Hopper
Big Bag
Big Bag
Balance
Hopper 1 Combustion Chamber
Sieve Shaker Dryer CM B E
Solar Radiation
Coal Coal
CM: Cone Mill BE: Bucket Elevator
Fig 2. Preparation of coal to the implementation of solar drying.
Comparing Figures 1 and 2, we can see that both the coal stacked in the reception area of mineral, receives solar radiation and air flow that promotes natural convection. But when the ore is spread in hours either early morning (Fig. 2), across the yard surface, forming a platform about 0.15 or 0.20 m tall, and is removed every 3 h until very Late in the afternoon, which is collected and stored in a conical pile in the shed, to the day after being subjected to drying in the oven rotating cylinder. There is a significant reduction of moisture in coal, which brings a minimization of fuel consumption per ton of dried material. Note: The recreation operations removal and collection of coal are performed by a loader TO-18 mark.
MATERIALS AND METHODS To take a representative sample, sufficient and minimal systematic sampling was used where an imaginary overlaid network and selected 30 points, which was taken the sample, it is an analysis of moisture in triplicate in glass test tubes 0.10 kg of coal inside. This procedure is performed three times a day (7:00 am, 12 am and 5:00 pm). In the development of work using various tools and materials allowed the execution of the experiments (oven 300 ºC C HOL 3524-5, dried and Analytical Balance BЛKT-500 gM). All tests were part of the moisture. To conduct the study in DSIT climate, were used the following equipment: ¾ SIC 100 Powermeter & Integrator (measure solar radiation and temperature) ¾ mercury thermometer -10 to 200 ºC (Measuring temperature) ¾ Fluke 62 Mini IR Thermometer (Measure temperature of solids)
RESULTS Figures 1 and 2 appear, it shows the results of changes Coal Moisture, Solar Radiation and Ambient Temperature. Observing the percentage decrease of the mass of coal as it dries in the course of time. 100
120 100
96
60 40
Moisture (%)
100·W/m 2
98 80
94 20 0 7:55 8:31
92 9:07 9:43 10:19 10:55 11:31 12:07 12:43 13:19 13:55 14:31 15:07 15:43 16:19 16:55 Hora (h) Solar Radiation
Coal Moisture
Graph 1. Variation of Coal Moisture and Solar Radiation vs. Time. 100
35 33
ºC
31 96 29 94
27 25 7:55
Moisture (%)
98
8:31
92 9:07 9:43 10:19 10:55 11:31 12:07 12:43 13:19 13:55 14:31 15:07 15:43 16:19 16:55 Hora (h) Ambient Temperature
Coal Moisture
Graph 2. Variation of Coal Moisture and Ambient Temperature vs. Time. The coal in the open air exposed to the sun, reduces its mass by drying, approximately 8%, reaching to achieve a 2% moisture content, and take a temperature of 60 ºC, lower by 5 °C the temperature reached coal to exit the rotary dryer. Note that rainy seasons and rainfall in the year, where the cloudiness does not allow a good stock rays of the sun on the mineral. During this period, coal remains piled in a cone-shaped battery and the main drying agent, an even
greater amount of solar radiation that is air. Normally in this geographical area of the Cuban archipelago, these seasons are not as significant, droughts are more representative. TABLE I. Comparative results of the efficiency of the dryer before and after implementing the previous solar drying of the mineral. Productivity (t/h)
Diesel consumption per ton of Quality of the product (acording to coal drying (l/t)
moisture %)
Before
0,5
≈ 40
0,8-1,3
After
1,3
≈ 22
0,8-1,3
Difference
+0,8
≈ -18
-
TABLE II. Environmental Analysis. Chemical concentration of the gases of combustion products [%] CO2 CO
O2
1 liter
11
0,2
6
40 liters
440
8
240 3312
22 liters
242
4,4
132 1821,6
Difference 198
3,6
108 1490,4
N2 82,8
The more dry the mineral is subjected to the oven, shorter it stays in it, resulting in lower emission of fines or dust into the atmosphere, among which the particles of aerodynamic diameter is less than 10 μm (PM10), these are very dangerous for the humans health, because they are able to reach the bottom of the lungs. TABLE III. Economic Impact. BEFORE Dry mass (t/day) Diesel consumption (l)
AFTER Cost (USD)
Dry mass (t/day) Diesel consumption (l)
Cost (USD)
1
40
22,46
1
22
12,36
18
720
404,35
18
396
222,39
OTHER COSTS
OTHER COSTS
Loader (l/day)
-
-
Loader (l/day)
40
22,46
Total
720
404,35
Total
436
244,86
Drying 18 t, our average production per day, currently has an economic cost equal to 244.86 USD per day. So every day we are saving 159.49 USD, that in a period of one year it would be 58215.46 USD.
CONCLUSIONS 1. The coal in the open air exposed to the sun, reduces its mass by drying, approximately 8%, reaching to achieve a 2% moisture content, and take a temperature of 60 ºC, lower by 5 °C the coal temperature reached the exit rotary dryer. 2. According to the analysis developed, it was determined that to each 1% moisture lost by the coal exposed to sunlight, saves 4 liters of diesel fuel per ton of coal produced. 3. With the prior application of solar drying in the rotary drum dryer has been an increase in productivity of 0.8 t/h, a significant reduction in diesel fuel, approximately half and the product presents the quality required by customers. 4. From an environmental perspective: •
The reduction of diesel consumption per ton of mineral drying brings parallel minimization of products of combustion gases (see Table II).
•
When using a coal with less humidity, less time is retained in the furnace, which contributes to the emission of fine powders, among which (PM10), there are particles of aerodynamic diameter less than 10 μm. This value has not yet been quantified.
5. The previous solar drying has resulted in the production process a significant economic impact, and now we are saving 159.49 USD daily in the period of one year, it would be 58215.46 USD.
BIBLIOGRAPHY 1. Arun S. Mujumdar. Handbook of Industrial Drying. Jerzy Pikon. Drying of Coal. Silesian Technical University. Gliwice, Poland. 2007. Págs. 977-981. 2. Bérriz Luis. Cuando el sol seca (Plantas medicinales). Energía y Tú. Nº 7. 1999. ISSN: 1028-9925. 3. Clariana Josep A.; Rougé Philippe; Arata Paola; Acevedo Sebastián; Segovia Javier. Postratamiento de Biosólidos en Era de Secado de la Estación Depuradora de Aguas Residuales El Trebal (Santiago de Chile). Tomado de: http://aca-web.gencat.cat/aca/documents/ca/jornadatecnica003/15_clariana_segovia.pdf (20/03/09 a las 14:30) 4. Corp Sergio. El secador solar de polen. Energía y Tú. Nº 15. 2001. ISSN: 1028-9925. 5. Cortéz Cádiz Elvira del Carmen. Fundamentos de ingeniería para el tratamiento de los biosólidos generados por la depuración de aguas servidas de la región metropolitana. Memoria para optar al titulo de Ingeniero Civil Químico. Año 2003. Tomado de: http://cabierta.uchile.cl/revista/21/articulos/pdf/rev3.pdf (23/03/09 a las 13:28) 6. Corvalan R., Horn M., Roman R., Saravia L. Ingeniería del Secado Solar. Ciencia y Tecnología para el Desarrollo CYTED. CDROM. Salta, República Argentina, abril de 2006. 7. Cuevas Fernando. Estrategia Energética Sustentable Centroamericana 2020. III Taller Contaminación Atmosférica vs. Desarrollo Sostenible. Ciudad de La Habana, CUBAENERGIA; 2008.
8. Helmer, W. A. and A. Abbas. nd. Solar drying of coal wastes from slurry ponds. Tomado de: http://pdf.aiaa.org/jaPreview/JE/1981/PVJAPRE62537.pdf. (19/05/08 a las 14:50) 9. Estenoz Severo; Espinosa Marianny; Pérez Niurka. Uso de energías renovables en la industria cubana del níquel. ECOSOLAR (Rev. Científica de las Energías Renovables), Nº 8, 2004. ISSN 1028-6004. 10. Fonseca Fonseca Susana, Abdala Rodríguez Jorge Luis, Ferro Fernández Victor R., Pantoja Enríquez Joel, Torres Yen Alonso. Estudio comparativo del secado solar de café en plazoletas tradicionales y ennegrecidas.
Tecnología
Química
Vol.
XXIII,
No.
3,
2003.
Tomado
de:
http://www.uo.edu.cu/ojs/index.php/tq/article/view/520/392. (23/05/08 a las 08:00) 11. Lara Coira Manuel. Escenario Energético Mundial. DYNA 82(9): 471-478, 2007. 12. Leyva Andrey; Díaz Ariel; Leyva Oscar. Secado solar del carbón mineral a la intemperie. V Taller Internacional de Educación, Energía y Desarrollo Sostenible, Ciudad de La Habana; 2008. 13. Retirado Yoalbys; Góngora Ever; Torres Enrique; Rojas Arturo L. Comportamiento de la humedad durante el secado solar del mineral laterítico. Minería y Geología. Vol. 23, Nº 3, 2007. ISSN 1993 8012. 14. Suárez Rodríguez José A., Beatón Soler Pedro A. Estado y perspectivas de las energías renovables en Cuba.
Tecnología
Química
Vol.
XXVII,
No.
3,
2007.
Tomado
de:
http://www.uo.edu.cu/ojs/index.php/tq/article/view/1255/916. (23/05/08 a las 08:00) 15. Teske Sven, Zervos Arthouros, Schäfer Oliver. Revolución energética. Perspectiva mundial de la energía renovable. Greenpeace Internacional, Consejo Europeo de Energías Renovables (EREC). Enero 2007. 16. Ficha Técnica (Secado Solar). Tomado de: http://www.itdg.org.pe/fichastecnicas/pdf/FichaTecnica13Secado%20solar.pdf. (17/05/08 a las 12:08) 17. La energía solar. Tomado de: http://www.dinero15.com/site_images/PDF/energiasolar.pdf. (19/05/08 a las 08:00) 18. 05 Desecación y deshidratación. Tomado de: http://www.caempa.com.ar/Presentaciones/05%20Desecaci%C3%B3n%20y%20deshidrataci%C3%B3n.pdf. (19/05/08 a las 08:06)
A New Supercritical Solid Acid for Breaking Car-Calk Bond in Di(1-naphthyl)methane Xiao-Ming Yue,† Xian-Yong Wei,†* Bing Sun,† Ying-Hua Wang,† Zhi-Min Zong, † and Zi-Wu Liu, ‡ †
Key Laboratory of Coal Processing and Efficient Utilization (Ministry of Education), China University of Mining & Technology, Xuzhou 221008, Jiangsu, China, and ‡School of Chemistry and Chemical Engineering, South China University of Technology, Wushan 381, Guangzhou 510640, Guangdong, China Introduction Catalysis in coal hydroliquefaction (CHL) has been exten- sively studied. Metal oxides, metal sulfides, metal halides, and acidic species, are well used as catalysts for CHL[1]. Rapid development in the area of carbocation chemistry began after the pioneering work of Olah, who utilized antimony pentafluoride as a strong Lewis acid[2]. Solid acids have advantages of corrosion resistance, higher safety than liquid acid and mineral acids, store, and handle easily and produce no wastes. Due to the increasing awareness of environmental protection and safety factors in industry, solid acid alternatives such as zeolites, heteropoly acids, acid clays, acid ion exchange resins, and sulfonated polystyrenes have been developed[3-8]. The complexity of coal structures leads to difficulty in understanding CHL mechanisms using coals themselves. Alternatively, the reactions of coal-related model compounds proved to be a powerful tool to reveal CHL mechanisms on the molecular level[9-12]. The cleavage of bridged bonds in coals is one of the most important reactions in CHL. Wei et al.[13] found that metal sulfides significantly catalyze the hydrocracking of di(1-naphthyl)methane (DNM) via monatomic hydrogen transfer. Activated carbon (AC) has similar catalysis in DNM hydrocracking but shows lower activity than metal sulfides[14]. In the present study, we investigated a new solid super- acid (SSA)-catalyzed DNM hydrocracking. Experimental Materials. DNM was synthesized by heating naphthalene with 1-chloromethylnaphthalene in the presence of zinc powder. Cyclohexane was commercially purchased and distillated before use. SbCl5, trimethylsilyl trifluoromethanesulphonate (TMSTFMS) and AC were also commercially purchased. SSA preparation and characterization. After being grounded to < 75 µm and dried for 24 h in a vacuum at 80 oC, AC was impregnated with SbCl5 and TMSTFMS under micro- wave irradiation. Then the mixture was magnetically stirred at room temperature for 12 h followed by filtration through a membrane filter with 0.45 μm of pore size and dried at 120 oC for 24 h. The SSA prepared was characterized with Nicolet Magna IR-560 FTIR, Hitachi S-3700N EDS/SEM, and TP-5000 II versatile adsorption detector with QIC-20 gas analysis system. General Procedure. DNM (1 mmol), catalyst (0.4 g), and cyclohexane (30 mL) were put into a 60 mL stainless, magnetically stirred autoclave. After being pressurized with hydrogen to 5 MPa at room temperature, the autoclave was heated to a reaction temperature (170-300 °C) in 15 min and maintained for a prescribed period of time (1 to 10 h). Then the autoclave was immediately cooled to room temperature in an ice-water bath. The reaction mixture were taken out from the autoclave and identified with HP 6890/5973 GC/MS and quantified with Agilent 7890 GC.
1500
1100
700
579.4
809.3 900
649.4
1040.7
1176.6 1300
1099.8
1358.0
Absorbance
1259.8
Results and Discussion The existing of sulfonic group, Si-O, and Si-C can be confirmed from the absorbances at 1040.7, 1099.8, 1176, 1259.8, and 1358.0 cm-1, while the absorbances at 579.4 and 649.4 could be related to the stretch vibration of C-Cl bond, as shown in Fig. 1, i.e., the reaction of SbCl5 with AC could occurred to some extend.
500
Wavenumbers (cm-1)
Fig. 1 FTIR spectrum of the SSA prepared
The SEM micrographs reveal the dispersion of active components on AC. As Fig. 2a displays, the surface of AC is smooth and slightly concave. However, the SEM micrograph of SSA (Fig. 2b) exhibits a rough surface. There were irregular grains with diameter less than 5 µm adhered to the surface of AC. The elemental composition from SbCl5 and TMSTFMS can be confirmed by EDS measurement (Fig. 3).
(a)
(b)
Fig. 2 SEM micrographs of AC (a) and the SSA (b) The TPD profile of SSA in Fig. 4 displays that one well-resolved desorption peak of NH3 appears at around 220 oC, which is attributed to NH3 adsorbed on weak acid sites. A narrow desorption peak at the range of 590–670 oC corresponds to the NH3 adsorbed on strong acid sites.
TPD signal
Cl
Sb C O
0
S
Si
F
1
2
Fig. 3
Sb Cl 3
Sb
Energy (keV)
4
5
EDS spectrum of the SSA
6
0
100
Fig. 4
200
300
400 500
600
Temperature (oC)
700 800
NH3-TPD profile over the SSA
As shown Table 1 and Fig. 5, the reaction of DNM over the SSA only afforded naphthalene and 1-methylnaphthalene (1-MN), indicating that the SSA specifically catalyzed DNM hydrocracking under the reaction conditions. Both raising reaction temperature and prolonging reaction time increase DNM conversion. Noteworthily, the yield of naphthalene is significantly higher than that of 1-MN in most of cases, suggesting that demethylation of the resulting 1-MN significantly occurred. As an acidic catalyst, the SSA may catalyze the heterolytic splitting of H2 to mobile H+ and immobile H-. The addition of mobile H+ to ipso-position of naphthalene ring induced DNM hydrocracking and subsequent 1-MN hydrocracking followed by stabilization of the resulting naphthyl-1-methyl and methyl cations with immobile H- (Scheme 1).
o
temp. [ C] 150 170 200 250 250 250 250 300 300 300 300
Table 1. The SSA-Catalyzed Hydrocracking of DNM yield [mol%] time [h] conv. [%] naphthalene 1-MN 3 0 0 0 3 0.1 0.1 0.1 3 6.4 11.1 1.7 1 8.7 12.9 4.5 3 9.1 13.9 4.3 6 10.2 15.1 5.3 10 10.5 15.2 5.8 1 20.9 21.6 20.2 3 50 58.4 41.6 6 69.3 75.1 63.5 10 74.9 82.4 67.4
DNM conversion (%)
50
50
40
40
30
30
20
20
10
10
0 120
Fig. 5
60
naphthalene 1-MN
160 200 240 280 Reaction temperature (oC)
320
Yield (mol%)
60
0
DNM conversions and product yields over the SSA at different temperatures for 3 h Scheme 1. Possible pathway for hydrogen transfer to DNM over the SSA H-
H
+
CH3
CH 4
+ + H+ - H+
H
+ +
+ H+ - H+ H-
Conclusions The SSA we prepared shows high activity and extremely high selectivity for DNM hydrocracking. The catalyst may heterolytically split H2 to mobile H+ and immobile H-. The addition of mobile H+ to ipso-position in DNM could be crucial step for DNM hydrocracking and subsequent steps include cleavage of bridged bond in hydroDNM cation, stabilization of the resulting naphthyl-1-methyl cation, mobile H+ to ipso-position in 1-MN, and stabilization of the resulting methyl cation. Acknowledgments. This work was subsidized by the Special Fund for Major State Basic Research Project (Grant 2011CB201302), National Natural Science Foundation of China (Grants 20936007 and 51074153), the Fund from the Natural Science Foundation of China for Innovative Research Group (Grant 50921002), the Program of the Universities in Jiangsu Province for Development of High-Tech Industries (Grant JHB05-33), and the Fundamental Research Funds for the Central Universities (China University of Mining & Technology, Grant No. 2010ZDP02B03 and 2010LKHX09). References [1] Matsuhashi H., Nakamura H., Arata K., Yoshida R., Maekawa Y. Fuel 1997; 76(10): 913-918. [2] Olah G. A., Praqkash G. K. S. Encyclopedia of Physical Science and Technology 2004: 175-188. [3] Smith G. V., Notheisz F. Heterogeneous Catalysis in Organic Chemistry. Academic Press, San Diego, 1999. [4] Sheldon R. A., van Bekkum H. Fine chemicals through hetero- geneous catalysis. Wiley-VCH, New York, Weinheim 2001. [5] Yadav J. S., Reddy B. V., Eeshwaraiah B., Srinivas A. Tetrahedron 2004; 60: 1767-1771. [6] Yadav J. S., Reddy B. V., Raju A. K., Guaneshwar D. Adv. Synth. Catal. 2002; 344: 938-940. [7] Thomas J. M., Raja R., Lewis D. W. Angewandte Chemie International Edition 2005; 44: 6456-6482. [8] Olah G. A., Iyer P. S., Prakash G. K. S. Synthesis 1986: 513-531. [9] Murata, S., Nakamura, M., Miura, M., Nomura M. Energy Fuels 1995; 9 (5): 849-854. [10] Zong Z. M., Wei X. Y. Fuel Process. Technol. 1994; 41 (1): 79- 85. [11] Grigorieva E. N., Panchenko S. S., Kovrobkov V. Yu. Kalechitz, I. V. Fuel Process. Technol. 1994; 41 (1): 39-53. [12] Kamiya Y., Ogata E., Goto K., Nomi T. Fuel 1986; 65 (4): 586- 590. [13] Wei X. Y., Hai N. Z., Zong Z. M., Lu Z. S., Chun X. Y., Wang X. H. Energy Fuels 2003; 17 (3): 652-657. [14] Sun L. B., Zong Z. M., Kou J. H., Zhang L. F., Ni Z. H., Yu G. Y., Chen H., Wei X. Y., Lee C. W. Energy Fuels 2004; 18 (5): 1500-1504.
Oviedo ICCS&T 2011. Extended Abstract
Briquetting of carbon-containing wastes from steelmaking for metallurgical coke production M.A. Diez, R. Alvarez, J.L.G. Cimadevilla Instituto Nacional del Carbón, INCAR-CSIC, Apartado 73, 33080-Oviedo. Spain.
[email protected]
Abstract This work focuses on the manufacture of briquettes by using carbon-containing wastes from steelmaking as fillers and binders for use in coke ovens to produce metallurgical coke. Coal-tar sludges from the tar decanter of a by-products coking plant were used individually as a binder or combined with other wastes, such as oils from the steel rolling mills and deposits from the coke oven gas pipelines. Another objective of this study was to find alternative low-cost fillers such as the coal generated after routine cleaning operations in the coal stockyards, so as to reduce the overall cost of briquette manufacture. The feasibility of using carbon briquettes with different formulations produced in a roll-press machine was tested in a semipilot movable wall oven by adding them to a coking blend at a ratio of 10 wt%. The quality of the cokes produced was assessed by measuring of their reactivity towards CO2 and mechanical resistance before and after gasification with CO2. In general, the coke quality parameters did not show any significant deterioration as a result of the addition of carbon briquettes when the amount and the nature of the binder and the particle size of the filler are optimized. Partial briquetting of the charge enabled cokes to be produced according to the specific requirements of blast furnace.
Keywords: coal, wastes, briquetting, carbonization, metallurgical coke
1. Introduction Briquetting processes that use coal as filler and binders such as pitches and tars have been industrially implemented in the past for metallurgical coke production [1-4]. Due to increasingly stringent regulations governing waste disposal practices and the need to reduce fossil fuel consumption, cold coal briquetting is regarded as a useful technology for recycling low-value and carbon-containing wastes generated in steelmaking.
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Oviedo ICCS&T 2011. Extended Abstract
Wastes generated in different parts of an integrated steel factory such as tar sludge and residual pitch from the benzol distillation column have been successfully used as minor components in coal blends or as binders in partial charge briquetting for metallurgical coke production [5]. The use of a combination of several heterogeneous carbon-containing wastes from steelmaking of different composition and origin to manufacture briquettes is proposed as an environmentally friendly way which has several advantages associated with a utilization of the wastes in situ and their removal in single operation making it unnecessary to dispose them. The objective of this work was to determine the effects of partial briquetting on the overall carbonization behaviour of the charge and on the quality of the resultant cokes in order to establish the viability of using several carbon-containing wastes to manufacture briquettes.
2. Experimental section Coal blend and wastes The coal blend P was prepared and supplied by ArcelorMittal-Spain. Coal blends are prepared every month by the company by mixing bituminous coals of different rank, thermoplastic properties and geographic origin. The main characteristics of the coal blend P are as follows: ash, 7.6 wt.% db, volatile matter, 25.5 wt.% db, S, 0.72 wt.% db, Gieseler maximum fluidity, 544 ddpm. Two wastes from different parts of the integrated steel installation were selected to be used in combination with coal-tar sludges from the tar decanter of the by-products coking plant (Mx) as a binder: oily wastes from the steel rolling mills (Lo) and deposits from the coke oven gas (COG) pipelines (Tu). Due to fluctuations in the composition of Mx, two different samples of Mxa and Mxb were used, the difference between them being the amount of waste which is soluble in powerful solvents such as quinoline. While Mxa contains about 47 wt% of quinoline soluble, Mxb has a very low content nearly 15 wt%. The coal generated during routine cleaning operations in the coal stockyards –B(a particle size of < 5 mm and a volatile matter content of 21 wt% db) was used as a low-cost filler.
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Oviedo ICCS&T 2011. Extended Abstract
Manufacture and characterization of briquettes Carbon briquettes were produced in a Komarek B050 roll-press machine using different combinations of carbon-containing wastes from the steel industry (Table 1). All of the briquettes produced were pillow-shaped 39 mm long, 19 mm wide and 10 mm thick. The physical properties of the briquettes (density and mechanical strength) were measured in order to assess the effectiveness of the agglomeration process. The density of the briquettes was determined by water immersion at 20 ºC using 100 g briquettes. The mechanical strength of the briquettes was tested with an I-type rotating drum (the same to determine CSR and described below) applying a mechanical treatment of varying severity from 20 to 150 revolutions at a rotation rate of 20 rpm. After each set of rotations, the sieving process and weighing of the size fractions after which all of the material was returned to the drum. The amount of unbroken briquettes and broken briquettes with a size >10 expressed as a percentage of the initial weight of the briquettes was used as an indicator of the strength of the briquettes.
Semipilot carbonization tests The amount of briquettes added to the industrial coal blend P was 10 wt%. All the carbonization tests were carried out in an electrically-heated movable wall oven of about 17 kg capacity with the following dimensions: 790 mm height, 250 mm length and 150 mm width. During the carbonisation tests, the temperature of the wall was kept constant at 1010 ºC. The coking time was nearly 3 h 30 min, the temperature in the centre of the charge rising to 950 ºC. After the hot coke was pushed from the oven, it was quenched with a water spray. The oven used produces enough coke to enable it to be characterized by the standard procedures used by the steel industry. As the bulk density of the charge varies as a function of grain size and moisture content, both of these parameters were kept as close as possible in each carbonization test. The bulk density was kept at 782 ± 6 kg/m3 db.
Coke characterization The quality of the resultant cokes was evaluated in terms of their reactivity to CO2 (CRI) and the mechanical strength of the partially-gasified coke (CSR) by the NSC method, following the ASTM D5341 standard procedure. To determine CRI, 200 g of coke with a particle size between 22.4 and 19 mm was exposed to the action of CO2 at flow rate of
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Oviedo ICCS&T 2011. Extended Abstract
5 L/min h at 1100 ºC for 2 h. The weight per cent of the initial coke mass lost during the reaction is defined as CRI. The mechanical degradation of the partially gasified coke (CSR) was measured as the weight of coke remaining on a 9.5 mm sieve after 600 revolutions at a rotation rate of 20 rpm in an I-type drum. In addition, the cold mechanical strength of the coke was evaluated from a sample of 10 kg with an initial size of > 50 mm, employing a JIS drum and rotating it for 150 revolutions at a rotating rate of 15 rpm (JIS K2151 standard procedure). Two indices were derived from this test: the DI150/15 and DI150/5 indices which are defined as the amount of coke with sizes >15 mm and <5 mm, respectively, after the mechanical treatment.
3. Results and Discussion Characterization of the briquettes Table 1 shows the composition of the briquettes prepared using coal left behind after routine cleaning operations in the coal stockyards (B) as filler and bituminous or oily wastes as binder. The oily waste (Lo) is a complex mixture of several lubricating oils of petrochemical and synthetic origin, whereas the deposits from the COG pipelines are a heterogeneous mixture of solid particles and light tar components.
Table 1. Series of carbon briquettes prepared with different wastes. Briquettes
Filler
Amount of filler (wt%)
Binder/s
Amount of binder (wt%)
B30Mxa
B
70
Mxa
30
B13Mxa13Tu
B
74
Mxa Tu
13 13
B13Mxa7Lo
B
80
Mxa LO
13 7
B30Mxb
B
70
Mxb
30
B15Mxb15Tu
B
70
Mxb Tu
15 15
Mechanical resistance is dependent on the amount and type of binder in the briquette [6-8]. The briquettes manufactured with the coal-tar sludge Mxa exhibit a high mechanical resistance and when it is combined with other wastes (Lo and Tu) they retain a high stabilization index (CS20) (Table 2). However, other wastes produce weaker briquettes with decreasing cohesion and increasing abrasion as the severity of the
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Oviedo ICCS&T 2011. Extended Abstract
mechanical treatment is increased. The density of the briquettes in water ranges from 1.22 to 1.28 g/cm3. On the other hand, the use of a poor binder such as tar sludge Mxb clearly undermines the mechanical strength and water resistance of the briquettes. The briquettes break after only a small number of revolutions and generate a high amount of fines. They also disintegrate when immersed in water. Due to its composition, the tar sludge will perform two roles in the briquette. On the one hand, the material that is soluble in quinoline and exhibits good binding properties will act as a binder, while, on the other hand, the insoluble fraction which is composed of very fine solid particles will contribute to reducing interparticulate friction. Basically, the two types of tar sludge differ in the amount of material with good binding properties. Whereas the quinoline soluble fraction of Mxa represents about 47 wt%, the other waste Mxb only contains about 15 wt%. This basic difference may explain the divergent mechanical behaviours of the two series of briquettes. An increase in the binder components is achieved by the combination of Mxb with Tu waste at the relative binder proportion of 1:1 w/w. Although the mechanical strength of the briquettes is improved, the high resistance levels of the Mxa series are not attained. It is important to point out that high strength is not necessary for coking. However, the briquettes made with Mxb would not resist prolonged storage in open stockpiles and the operations of loading and unloading before being charged into a coke oven without breaking and generating fine particles.
Table 2. Density and mechanical strength indices of the briquettes. Briquettes
B30Mxa B13Mxa13Tu B13Mxa7Lo B30Mxb B15Mxb15Tu
dw (g/ml)
CS20 (%)
CS40 (%)
CS80 (%)
CS150 (%)
1.28 1.25 1.22 nd nd
99.6 94.5 94.6 8.5 63.7
99.1 86.8 86.2 1.8 52.9
98.7 76.7 70.0 0.0 39.2
98.2 64.2 45.0 0.0 25.5
dw: density determined in water a 20 ºC; CSX: amount of briquettes (unbroken or broken at a size > 10 mm) after X revolutions in a I-type drum; nd: not determined, most of the briquettes disintegrate when immersed in water.
Partial briquetting carbonization and coke quality All the briquettes were carbonized with the industrial coal blend P at an addition rate of 10 wt%. To avoid the secondary effect of an increase in bulk density on coking pressure
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Oviedo ICCS&T 2011. Extended Abstract
generation and metallurgical coke properties, the series of carbonization tests were carried out with only minor variations in bulk density. No relevant effect on the generation of coking pressure was observed during partial briquetting (1.4 kPa for the blend P vs. 0.8-1.4 kPa for carbonizations with partial briquetting). Thus none of the formulations used for briquetting were detrimental to cokemaking. When the cold mechanical strength indices (DI150/15 and DI150/5) of the resultant cokes are compared with those of the coke produced from the blend P, similar cohesion and abrasion indices are observed for cokes from the charges which contain Mxa (Table 3). In addition, the high-temperature properties of the cokes (CRI and CSR) remain constant, whether Mxa is used, individually or combined with the oily waste from the steel rolling mills (Lo). In contrast, the incorporation of the waste left behind in the COG pipelines (Tu) clearly produces a more reactive coke (CRI) and a less resistant coke after the reaction with CO2 (CSR). The increase in reactivity by about 3 points is attributed to the catalytic effects of the iron oxides present in the solid fraction of this waste.
Table 3. Characteristics of the cokes produced from the coal blend P and partial briquetting of the charge (briquette addition rate: 10 wt%). Coke
P P+10B30Mxa P+10B13Mxa13Tu P+10B13Mxa7Lo P+10B30Mxb P+10B15Mxb15Tu
DI150/15
DI150/5
CRI (%)
CSR (%)
P (%)
77.8 76.7 77.2 78.0 62.9 69.3
16.5 15.7 15.3 15.5 22.7 20.8
30.9 31.6 33.7 31.6 31.4 32.3
56.2 55.6 52.6 56.8 55.4 53.0
52.9 51.4 51.7 51.4 52.2 51.7
The poor agglomeration achieved in the briquettes manufactured using the low quinoline-soluble sludge Mxb is also reflected in the cold mechanical strength of the cokes. A loss of cohesion and resistance to abrasion in the coke is a result of the partial briquetting of the charge. However, the CRI and CSR indices are not greatly affected. Thus, variation of the high-temperature properties seems to be a consequence of the type of binder used in the manufacture of briquettes and not the degree of agglomeration achieved.
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4. Conclusions Partial briquetting can be considered as a recycling route for wastes of different types, origin and composition in an integrated steel factory. An evaluation of the amount of tar sludge necessary to act as binder in the briquette is essential to ensure that the cold mechanical strength of the cokes resulting from partial briquetting is adequate and to retain the amount of breeze coke. No negative effect was observed in the coke hightemperature properties, reactivity to CO2 and strength after reaction. By controlling the amount of effective binder used to make the briquettes, the partial briquetting of the coal charge will produce cokes with the desired characteristics for use in blast furnaces.
Acknowledgements The financial support provided by PCTI-Asturias-Spain through the research project PEST08-07 is gratefully acknowledged. We are also grateful to ArcelorMittal-Spain for its collaboration.
References [1] Akamatsu K, Nire H, Miyazaki T, Nishioka K, Influence of non-coking coal on the quality of metallurgical coke, Coal, coke and the blast furnace, The Metals Society, 1977:55-65. [2] Nakamura N, Togino Y, Adachi T, Philosophy of blending coals and cokemaking technology in Japan, Coal, coke and the blast furnace, The Metals Society, 1977:93-106. [3] Schinzel W, Briquetting. In: Chemistry of Coal Utilization, Second supplementary volume, Elliot MA (ed.), John Wiley and Sons, New York, 1981, Chap. 11, pp.609-664. [4] Braun NV, Glushchenko IM, Panchenko NI, Ivchenko AY, On the possibility of using coking plant wastes as a binder for briquetting a coal charge, Coke and Chemistry USSR 1986;5:28-33. [5] Álvarez R, Barriocanal C, Diez MA, Cimadevilla JLG, Casal MD, Canga CS, Recycling of hazardous waste materials in the coking process Environ. Sci. Technol. 2004;28:1611-15. [6] Clarke DE, Marsh H, Influence of coal/binder interactions on mechanical strength of briquettes, Fuel 1989:68:1023-30. [7] Clarke DE, Marsh H, Factors influencing properties of coal briquettes, Fuel 1989:68:1031-38. [8] Taylor JW, Hennah, The effect of binder displacements during briquetting on the strength of formed coke. Fuel 1991;70:873-76.
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Oviedo ICCS&T 2011. Extended Abstract
Co-carbonization behaviour of coal and biomass-derived products and its effect on coke structure and properties M.A. Diez1, R. Alvarez1 and M. Fernández2 1
2
Instituto Nacional del Carbón, INCAR-CSIC, Apartado 73, 33080 Oviedo, Spain Centro Nacional de Investigaciones Metalúrgicas, CENIM-CSIC, Avda Gregorio del Amo 8, 28040 Madrid, Spain.
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Abstract Co-carbonizations were carried out to study the feasibility of using biomass-derived products as additives to coking coals for the production of metallurgical cokes. Blends of a coking coal with Eucalyptus wood and the products resulting from its carbonization at 415 ºC (charcoal, light and heavy tars) were prepared at an addition rate of 2 wt%. Moreover, the effect of the biomass products was compared to that of wood constituents such as hemicellulose, cellulose and lignin which were used as reference additives. Wood biomass, its carbonization products and its model compounds were observed to reduce the Gieseler fluidity of the coal. In general, the overall effect of Eucalyptus wood and its carbonization products (char and tars) greatly resembled that of the parent biomass component. As regards coke quality, not all of the biomass-derived products affected the quality parameters in the same way. In general, the cokes were found to be less mechanically resistant and more reactive, but the degree of deterioration was depended on the type of additive. A common feature of these cokes was their higher total pore volume which was accompanied by a shift to smaller pores of 10-50 µm. These preliminary laboratory results may contribute to a better prediction of the behaviour of biomass/coal blends in the production of coke.
Keywords: coal, biomass, wood by-products, charcoal, co-carbonization, coke
1. Introduction The environmental benefits of wood or other forms of vegetable biomass are associated with the reduction of CO2 in the atmosphere due to their CO2-neutrality, their contribution to the preservation of natural resources by partially replacing coal in conversion processes and their flexibility in the production of solid, liquid and gaseous fuels. Biomass can be used to produce chemicals and liquid fuels, charcoal for use in
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metallurgy, carbon adsorbents from wood and biomass wastes and for co-firing with coal in energy generation [1-5]. Charcoal from hardwood species in small blast furnaces is employed for iron production in the Brazilian steel industry [1,2,6]. In the last few years, the use of charcoal from sustainable biomass in modern blast furnaces has opened up an environmentally friendly way to reduce CO2 emissions and the use of fossil fuels [7-9]. In the steel industry, charcoal is also seen as a potential additive to coal blends in the production of coke that is used to feed blast furnaces [9]. When hardwood or other types of biomass decompose by the action of heating in an oxygen-deficient atmosphere, a carbon material (char) is obtained together with a complex mixture of volatile products. These are allowed to escape as gases and other volatiles or they are condensed and converted to useful by-products (permanent gases and tar). The relative yields of each product (char, gas and tar) depend on the type of biomass, the design of the carbonization reactor and the carbonization conditions. When tar condensers are installed, two types of tar are recovered, a heavier tar that settles at the bottom of the column and a watery tar, the so-called pyroligneous acid that remains at the top. These tars account around 7 and 35 wt% of the initial wood, respectively, the raw material serving as a base for the vegetal carbochemistry [1-3,10]. The aim of the present work is to study the effects of biomass-derived carbonization products on the thermoplastic behaviour of coal and the structure and properties of the resulting cokes as a preliminary evaluation of the performance of coalbiomass blends in cokemaking. To understand the effects of Eucalyptus sawdust and its carbonization products on coal and coke properties, xylan –as being representative of different types of hemicelluloses-, cellulose and lignin were used as model additives.
2. Experimental section Coking coal A which was used to prepare the blends has a volatile matter content of 21 wt% db and a Gieseler maximum fluidity of 389 ddpm, yielding a high-temperature coke with a high mechanical strength and a low reactivity to carbon dioxide. Eucalyptus wood and its carbonization products obtained at 415 ºC (charcoal –Ch- and the two tar fractions -water-soluble tar, ST and water-insoluble tar, IT-) were added to coal A at an addition rate of 2 wt%. Three biomass model compounds (commercial xylan, cellulose and lignin from Sigma-Aldrich) were also tested.
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The thermoplastic properties of the coal and blends were tested in a constanttorque Gieseler plastometer, R.B. Automazione PL2000, following the ASTM D2639 standard procedure. This instrument measures the rotation of a stirrer inside a compacted 5 g sample of particle size <0.425 mm. At the same time, the sample is heated from 300 °C up to 550 °C at a heating rate of 3 °C/min and the fluidity is recorded in dial divisions per minute (ddpm) as a function of the temperature. The parameters derived from this test are: (i) coal softening temperature (Ts); (ii) temperature of maximum fluidity (Tf); (iii) resolidification temperature (Tr) of the fluid mass into a semicoke; (iv) the plastic or fluid range (Tr-Ts); (v) maximum fluidity (Fmax), expressed as ddpm. The blends were carbonized in a horizontal oven under nitrogen up to 1000 ºC at a heating rate of 5 ºC/min employing a soaking time of 15 min. The volatiles were swept away by nitrogen at a flow rate of 300 ml/min. The total amount of sample was about 25 g and the particle size <1 mm, except for model compounds. In this case, the particle size was <20 µm. The quality of the resultant cokes for use in a blast furnace was assessed in terms of reactivity to CO2 and mechanical microstrength before and after reaction with CO2. Reactivity towards CO2 was measured as the weight loss of a coke sample of 7 g with a size of 1-3 mm at 1000 ºC for 1 h using a CO2 flow of 120 ml/min, following a procedure based on the ECE-INCAR method [11]. The initial and partially-gasified cokes were subjected to mechanical treatment in cylindrical drums, according to the microstrength procedure described by Ragan and Marsh [12]. While the microstrength of the cokes before reaction with CO2 (initial cokes) was measured using samples with a particle size of between 0.6-1.18 mm, the initial particle size of the partially-gasified cokes was > 1.18 mm. All of the cokes were subjected to 800 revolutions at a speed of 25 rpm. The weight per cent of coke fines (<0.212 mm in size) was taken as an indication of the degree of abrasion of the cokes (R3 index for the initial cokes and PR3 for partially-gasified cokes). The helium density of the cokes was measured with a Micromeritics Accupyc 1330T pycnometer. Their apparent density was determined using a Micromeritics Autopore IV 9500 porosimeter with mercury pressures of 0.005 and 0.1 MPa. The open porosity was calculated from the helium and apparent densities. Mercury porosimetry was also used to determine the pore volume fraction of coke over various diameter ranges: supramacropores or devolatilization pores (dp >10 µm), macropores (10 µm >
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Oviedo ICCS&T 2011. Extended Abstract
dp > 50 nm), mesopores (50 nm > dp > 5.5 nm). The microporore volume (dp <5.5 nm) was estimated by subtracting the macropore and mesopore volumes (obtained by mercury porosimetry) from the total pore volume, calculated from the mercury and helium densities.
3. Results and Discussion Raw materials The solid (lump charcoal) and liquid products (water- soluble and –insoluble tars) obtained by carbonization at 415 ºC from Eucalyptus wood were studied as potential additives for co-carbonization with coking coal in coke manufacture. Table 1 shows the data from the proximate and elemental analyses of coking coal A, the Eucalyptus sawdust and the woody biomass-derived products. Eucalyptus sawdust (SD) is characterized by a high volatile matter content, low carbon and high oxygen contents and low ash content. The insoluble tar, a viscous and dark red-brown liquid (IT), is very similar to the raw material in its chemical composition. In contrast, the water-soluble tar (ST) is a complex mixture of water and water-miscible oxygenated compounds. Due to the high amount of water and low carbon content (13 wt% db) in this soluble tar fraction, its calorific value is negligible. Charcoal (Ch) with its high carbon content still retains a high amount of oxygen in its chemical composition and an appreciable amount of tarry compounds.
Table 1. Proximate and elemental analysis (expressed on a dry basis). Ash (wt%) VM (wt%) C (wt%) H (wt%) O (wt%)
Coal A 9.0 21.6 82.5 4.5 1.7
Sawdust 0.28 83.7 46.7 5.9 46.8
Water-insoluble tar 48.3 7.4 43.2
Charcoal 0.37 30.4 83.3 3.1 13.2
Thermoplastic behaviour of blends High-temperature plasticity is one of the most important parameters for evaluating the coking ability of a coal and predicting coke quality. Biomass products reduce the coking capacity of coal when they are added at a rate of 2 wt%, as reflected by the decrease in the Gieseler maximum fluidity (Fmax) (Figure 1). The fluidity ranges from a value similar to that of coal A (388 ddpm for A2XYL) to 180 ddpm for A2Sd and the extent of
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the reduction is very much dependent on the biomass product and component. Despite the reduction in fluidity (increase in viscosity) caused by the additives, the coal blended with 2 wt% biomass-derived product is sufficiently fluid during carbonization as to lose virtually all its solid structural identity during thermal treatment and produce a graphitizable carbon. 450 400
389
388 338
Gieseler Fmax (ddpm)
350
321
300
271 247
250
200
180
200 150 100 50 0 A
A2Sd
A2ST
A2IT
A2Ch
A2XYL
A2CEL
A2LIG
Figure 1. Variation of Gieseler maximum fluidity of coal A with the addition of biomass products and components at a rate of 2 wt%.
In accordance with the above results the following order from a weak to a strong effect on coal thermoplasticity has been established: XYL < ST < Ch < LIG < IT < CEL < Sd In general, the overall effect of Eucalyptus wood and its carbonization products (char and tars) greatly resembled that of the parent biomass component. It is well known that lignocellulosic biomass mainly consists of various forms of organic polymeric networks such as hemicellulose (20-30 wt% db), cellulose (40-50 wt% db) and lignin (20-30 wt% db). Structurally, the three components of the woody biomass are different and, consequently, the distribution and composition of the carbonization products differ from one component to another. Lignin, a phenolic polymer, is the major precursor of charcoal, whereas the other components mainly decompose to become tar and gas fractions with a small amount of char, ranging from 6 to 30 wt%. On the other hand, hemicellulose, such as xylan, is made up of polysaccharides of varying composition with five and six carbon monosaccharide units and yields more char but fewer gas and tar fractions than cellulose, a linear crystalline polysaccharide formed by β-linked glucose moieties. Depolymerization of the woody-biomass components by thermal degradation produces a complex mixture of oxygenated compounds which condense as tar fractions.
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Among the components of the water-soluble tar (ST), furfural and its derivatives predominate in the chromatographiable fraction, suggesting that the major precursor is hemicellulose, while the most abundant components of the water- insoluble tar (IT) are methoxyphenol derivatives produced mainly by the cleavage of β-O-4 aryl ether bonds in lignin [13]. D-allose from cellulose is present in both tar fractions. In view of the key characteristics of the tar fractions and the major contribution of lignin to charcoal, the fact that their effects on the thermoplastic properties of coal is similar to that of the lignocellulosic components is not surprising. The type of biomass that is added to the blend seems to have little effect on the temperatures which define the maximum fluidity (Tf), the softening of the coal particles (Ts) and the solidification of the overall fluid mass into a semicoke (Tr). Any differences amount to less than 4 ºC.
Coke yield and properties Table 2 gives the yield, the reactivity towards CO2 and the abrasion index before (R3) and after reaction with CO2 (PR3) of the cokes produced at 1000 ºC. If we take into account the contribution of coal A in the blends to the coke yield (78.9 wt%), most of the additives reduce the amount of coke by less than 1 %, whereas charcoal and lignin, maintain their coke yield at a similar level.
Table 2. Yield, mechanical strength and reactivity to CO2 of the resultant cokes. CY (wt% db) R3 (%) RECE (%) PR3 (%)
A A2Sd A2ST A2IT A2Ch A2XYL A2CEL A2LIG 80.5 79.0 78.6 78.6 80.5 79.0 78.3 79.7 56.3 56.4 61.7 55.7 57.7 53.1 61.0 64.3 5.9 7.2 6.9 6.9 8.6 6.2 7.4 9.2 43.1 50.2 47.4 43.1 49.9 43.2 51.8 51.6
CY: coke yield; R3: abrasion of coke; RECE: reactivity to CO2 as determined by the ECE method; PR3: abrasion of coke after reaction with CO2.
When water-soluble tar (ST), charcoal (Ch), cellulose (CEL) and lignine (LIG) are blended with coal A, these additives cause the cokes to deteriorate drastically, as reflected by the generation of coke fines (R3 index). As regards reactivity to CO2, the biomass-additives produce more reactive cokes. From the results, it is possible to classify the cokes in order of decreasing reactivity. The more reactive cokes are produced from both charcoal and its main precursor, lignin, whereas the addition of sawdust and its major constituent, cellulose, produces cokes
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with intermediate reactivity indices between LIG-Ch and tars (ST and IT)-xylane. These data are in agreement with the molecular organization of the carbon in the coke. Charcoal and lignin contribute to the coke structure by creating isotropic carbon which is more reactive. The generation of fine particles during the mechanical treatment of the partiallygasified cokes is lower than that of the cokes before the reaction with CO2. This means that the most fragile part of the coke at the surface has been removed by the action of CO2. It is also well known that reactivity to CO2 and the strength of the partially-gasified coke are two closely related coke properties: the higher the reactivity to CO2, the lower the post-mechanical strength. This relation was also observed in the cokes studied here. It should be noted, however, that two families can clearly be distinguished (Figure 2). The first group of cokes are those produced from the co-carbonization of the coal A and sawdust -Sd-, cellulose –CEL- and xylane –XYL- (two of the constituents of wood) and water-soluble tar –ST-, a complex mixture of water and oxygenated compounds mainly derived from the decomposition of hemicellulose
and cellose. The second group
includes three of the additives: lignin –LIG- and the solid product from its carbonization (charcoal –Ch-) and the condensable fraction (water-insoluble tar –IT-). 54
abrasion PR3 index
52
LIG
CEL Ch
Sd
50 48
ST
46 44
A
XYL
IT
42 5.5
6.0
6.5
7.0 7.5 8.0 Reactivity ECE (wt% loss)
8.5
9.0
9.5
Figure 2. Relation between reactivity towards CO2 and the abrasion index of the partially-gasified cokes (PR3). Symbols denote the additive used in the blend prepared with coking coal A.
The porous structure of cokes is also modified by biomass additives. A common feature of cokes is a high total pore volume which is accompanied by a shift to smaller supramacropores or devolatilization pores in the range of 10-50 µm, except for the coke obtained with xylan as additive (Figure 3).
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600 Vtotal
V50‐10
Pore volume (mm3/g)
500
400
300
200
100
0 A
A2Sd
A2ST
A2IT
A2Ch
A2XYL
A2CEL
A2LIG
Figure 3. Variation in total pore volume and pores with size between 10 and 50 µm in the cokes produced from the blends of coal A with biomass additives.
Specific modifications of the porous structure are also observed. Cellulose and sawdust induce a higher proportion of devolatilization pores with a diameter larger than 100 µm, whereas charcoal produces the opposite effect and favours the formation of micropores. As in the case of Gieseler fluidity, the effect of the type of biomass on the coke porous structure resembles that observed in the case of the biomass constituents.
4. Conclusions The use of the wood constituents (xylan, cellulose and lignin) as model additives allows some simple guiding principles to be drawn on the modifications caused by biomassderived products in coal thermoplasticity and coke structure and properties when they are blended with coal. All biomass additives reduce the fluidity of the coal and yield more reactive cokes and becoming less resistant after reaction with CO2. In general, the modifications induced by Eucalyptus sawdust and its carbonization products are closely related to those observed for wood constituents: xylan, cellulose and lignin.
References [1]
FAO Forestry Department, Industrial charcoal making.1986, Paper 63.
[2]
Campos Ferreira O, Emission of greenhouse effect gases in vegetal coal production, Economy & Energy 2000:21.
[3]
Antal Jr MJ, Gronli M. The art, science and technology of charcoal production. Ind Eng Chem Res 2003;42:1619–40.
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[4]
Syred C, Griffits AJ, Syred N, Beedie D, James D. A clean, efficient system for producing charcoal, heat and power (CHaP), Fuel 2006;85:1566-78.
[5]
Nzihou A, Toward the valorization of waste and biomass, Waste Biomass Valor 2010;1:3-7.
[6]
Ferreira-Leitão V, Gottschalk LMF, Ferrara MA, Nepomuceno AL, Molinari HB, Biomass residues in Brazil: Availability and potential uses. Waste Biomass Valor 2010; 1:65-76.
[7]
ULCOS: Ultra-low CO2 steelmaking. http://www.ulcos.org.
[8]
Korthas B, Peters M, Schmöle P. Back to the future-Ideas for new blast furnace concepts. Proc. of the 5th European Coke and Ironmaking Congress, Stockholm (Sweden), 2005, paper Mo 1:4.
[9]
Hanrot F, Sert D, Delinchant J, Pietruck R, Bürgler T, Babich A, Fernández M, Alvarez R, Diez MA. CO2 mitigation for steelmaking using charcoal and plastics wastes as reducing agents and secondary raw materials, In: López A, Puertas F, Alguacil FJ, Guerrero A, editors. Proc. 1st Spanish National Conf. on Advances in Materials Recycling and Eco–Energy. Recimat 09, Madrid, Spain, 2009, paper S05-4. Available at: http://hdl.handle.net/10261/18433.
[10] Prauchner MJ, Pasa VMD, Molhallem NDS, Otani C, Otani S, Pardini LC, Structural evolution of Eucalyptus tar pitch-based carbons during carbonization, Biomass and Bioenergy 2005;28:53-61. [11] Menéndez JA, Pis JJ, Alvarez R, Barriocanal C, Canga CS, Diez MA. Characterization of petroleum coke as an additive in metallurgical cokemaking. Influence on metallurgical coke quality. Energy and Fuels 1997;11:379-84. [12] Ragan S, Marsh H. Carbonization and liquid-crystal (mesophase) development. 22. Micro-strength and optical textures of cokes from coal-pitch co-carbonizations. Fuel 1981; 60:522-528. [13] Diez MA, Alvarez R, Fernández M. Biomass derived products as modifiers of the rheological properties of coking coals. In: López A, Puertas F, Alguacil FJ, Guerrero A, editors. Proc. 1st Spanish National Conf. on Advances in Materials Recycling and Eco–Energy. Recimat 09, Madrid, Spain, 2009, paper S01-1, Available at: http://hdl.handle.net/10261/18359.
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Oviedo ICCS&T 2011. Extended Abstract
Evolution of volatile products of coal and plastic wastes during co-pyrolysis S. Melendi, M.A. Diez, R. Alvarez Instituto Nacional del Carbón, INCAR-CSIC, Apartado 73, 33080-Oviedo. Spain.
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Abstract The aim of this study was to evaluate the pyrolytic characteristics of blends made up of coking coals and several plastics by determining the evolution pattern of selected evolving products during their co-pyrolysis. For this purpose, simultaneous thermogravimetric–mass spectrometric analyses (TG-MS) of the blends were performed by thermal treatment up to 1000 ºC in a helium atmosphere under dynamic conditions at a heating rate of 25 ºC/min. The plastics used for blending with coal at an addition rate of 5 wt% were: five thermoplastics commonly found in municipal wastes (LDPE, HDPE, PP, PS and PET) and two mixtures with different compositions. Different characteristic ion fragments from selected families of evolving products during the copyrolysis process such as hydrogen, aliphatic hydrocarbons with one to four carbon atoms, aromatic hydrocarbons and carbon dioxide were monitored together with their thermogravimetric parameters (temperature, mass) at different times. Hydrogen evolution profiles have similar shapes and the maximum evolution temperature was not greatly affected by the addition of plastic wastes. In contrast, the evolution patterns of aliphatic hydrocarbons (alkanes and alkenes) were characterized by a low temperature of evolution and a high relative proportion of these components. The delay in the decomposition of the plastics together with the changes in the composition of volatiles promoted interactions between the components and had negative effects on coal fluidity.
Keywords: coal, plastic wastes, co-pyrolysis, evolved gas, TG-MS
1. Introduction Currently, the co-pyrolysis of single or mixed plastics with fossil fuels (coal and petroleum) are being investigated in order to recover chemicals, to partially replace fossil fuels in well-established industrial conversion processes of fossil fuels and to
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contribute to the protection of the environment. Among the different routes based on copyrolysis, the co-processing of coking coals with plastics from municipal wastes for metallurgical coke production has been implemented at industrial scale [1,2]. The composition of the plastic waste added has been shown to be a critical factor in controlling the effect on the coal thermoplastic properties, coking pressure generation during coking and the structure and properties of metallurgical coke [3-5]. In accordance with the different structure and thermal behaviour of the plastics contained in municipal wastes, opposite effects have been observed: polyolefins which cause a slight decrease in fluid coal properties [6-10], improve or maintain coke strength and reactivity and increase the wall pressure generated during coking up to extremely high values. Aromatic polymers such PET and PS, however, which are the strongest modifiers of coal fluidity, cause a deterioration in coke reactivity towards CO2, and help to balance the wall pressure [3-5]. The aim of this study is to gain additional information about the interactions between coal and plastics in order to explain the different effects on the fluidity of the coal and the generation of the wall pressure during coking. For this purpose, TG-MS was employed to obtain information on the different events which take place during copyrolysis and the chemical composition of the evolved gases.
2. Experimental section An industrial coal blend, PA, was used to prepare mixtures with several plastics at an addition rate of 5 wt%. The mixed plastic waste Wa was composed of 73 wt% highdensity polyethylene (HDPE), 20 wt% polypropylene (PP), 5 wt% polyethylene terephthalate (PET) and 2 wt% cellulose and the mixture Wb was a more heterogeneous waste containing the six thermoplastics: PP, 39.2 %; PET, 18.8 %; PS, 16.6 %; HDPE, 0.7 %; LDPE, 5.4 %; PVC, 1.2 % and 6.9 % of non-identified plastics. Both plastic wastes were provided by the Spanish recycling company Abornasa. The powdered mixtures (7 g, < 0.212 mm size) were subjected to thermogravimetric-mass spectrometric analysis (TG-MS) in a simultaneous TA Instrument SDT2960 analyzer coupled to a quadrupole mass spectrometer (Balzers, Thermostar GSD-300T) by a fused silica transfer line heated at 200 ºC. About 7 mg of the mixtures was heated from room temperature up to 1000 °C at a heating rate of 25 °C min-1 using a helium flow rate of 100 ml/min to sweep out the volatile products. The
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evolution of the temperature of the evolved gaseous products and the intensity of the selected ion fragments were monitored together with the thermogravimetric parameters (temperature, mass) at different times.
3. Results and Discussion The DTG curves of the plastic wastes, the coal blend PA and their mixtures at an addition rate of 5 wt% of the plastics are shown in Figures 1 and 2. The temperature of maximum decomposition of plastics varies in the following order: PS < PET < PP < LDPE < HDPE (Figure 1). The position of the DTGmax of the two plastic mixtures, Wa and Wb, are slightly lower than what might be expected from their composition. In the case of Wa (73 wt% HDPE, 20 wt% PP and 5 wt% PET), the Tmax value is located between PP (475 ºC) and PET (449 ºC) whereas Wb presents a similar value to that of PET and PS (442 ºC), which make up about 35 wt% of the waste. 80
HDPE LDPE PP PET PS Wa Wb
70 60 DTG (%/min)
50 40 30 20 10 0 250
300
350
400
450
500
550
600
Temperature (°C)
Figure 1. DTG curves of plastic wastes studied.
The DTG profile of the coal blend shows a main peak at about 505 ºC, which is also shown in blends containing the three polyolefins (HDPE, LDPE and PP). Under the pyrolysis conditions applied, polyolefins have the narrowest decomposition temperature ranges with a Tmax inside the thermal degradation of the macromolecular network of the coal, whereas the degradation of PS and PET takes place close to the early stages of coal decomposition (Figure 2). An examination of the DTG profiles of the blends shows that blends with HDPE, LDPE and PP show a single peak at a temperature slightly lower than that of the coal blend PA (495-499 vs. 505 ºC). However, when PS and PET are added to coal, these blends present a bimodal evolution of volatiles, with the first peak
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Oviedo ICCS&T 2011. Extended Abstract
being attributed to plastic decomposition and the high-temperature peak to coal devolatilization. When comparing the profiles of the blend and the corresponding plastic, a shift in the evolution of volatiles from the plastic in the blend towards a higher temperature can be clearly observed. This suggests that some degree of physical and chemical interaction may occur during the co-pyrolysis of plastics with coal. PET is the
6 5 4 3 2 1 0
450 Temperature (ºC)
500
550
80
7
70
6
60
PA PA+5HDPE HDPE
5 4
50 40
3
30
2
20
1
10
0
600
0 300
350
400
450 Temperature (ºC)
500
550
600
80
7
70
6
60 PA PA+5LDPE LDPE
40
3
30
2
20
1
10
0
0 350
400
450 Temperature (ºC)
500
550
80 70
6
60 PA PA+5PS PS
50 40
3
30
2
20
1
10
0
0 300
350
400
450 Temperature (ºC)
500
550
90 80
7
70 60
PA PA+5Wa Wa
5 4
50 40
3
30
2
20
1
10
0
0 300
350
400
450 Temperature (ºC)
500
550
600
40 30
2
20
1
10 0 350
400
450 Temperature (ºC)
500
550
600
6
60
5
50
4
40
PA PA+5PET PET
3
30
2
20
1
10
0
0 300
8
50
3
600
9
6
4
300
7
4
60 PA PA+5PP PP
5
600
8
5
6
0
DTG of coal and blend (%/min)
4
50
350
400
450 Temperature (ºC)
500
550
600
6 DTG of coal and blend (%/min)
5
DTG of coal and blend (%/min)
8
70 DTG of plastic (%/min)
80
7
DTG of plastic (%/min)
DTG of coal and blend (%/min)
400
90
8
8
300
DTG of coal and blend (%/min)
350
DTG of plastic (%/min)
DTG of coal and blend (%/min)
300
9
DTG of plastic (%/min)
DTG (%/min)
7
DTG of coal and blend (%/min)
PA PA+5HDPE PA+5LDPE PA+5PP PA+5PET PA+5PS PA+5Wa PA+5Wb
DTG of plastic (%/min)
8
60
5
50
4
40
PA PA+5Wb Wb
3
30
2
20
1
10
0
DTG of plastic (%/min)
9
DTG of plastic (%/min)
exception to the general tendency (Figure 2).
0 300
350
400
450 Temperature (ºC)
500
550
600
Figure 2. DTG profiles of the coal blend PA, the plastics and their mixtures at an addition rate of 5 wt%.
When plastic decomposition via radical chain reactions occurs close to the decomposition of the macromolecular network of the coal, there is greater opportunity for the small size species from coal decomposition which are responsible for the
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Oviedo ICCS&T 2011. Extended Abstract
development and maintenance of fluidity to volatilize and then to be stabilized by hydrogen transfer or cross-linking reactions. As a consequence of this, the fluidity decreases drastically. PS and PET are good examples of strong modifiers of coal thermal behaviour [7-11]. Then decrease the fluidity of the coal and give rise to more disordered carbon structures in the semicokes [7,8]. However, if the degradation products of the plastics are produced close to the range of maximum evolution of volatiles from the coal, when the maximum amount of gas and tar is produced and solidification sets in, the decomposition products from plastic will be trapped in the co-pyrolysis system and, then, incorporated into the semicoke [5,7,8]. This behaviour is exhibited by the polyolefins which overlap over a wide interval as the coal volatiles evolve. The interactions between the coal and plastics are also influenced by the chemical composition of the volatiles. By means of TG-MS analysis, it is possible to obtain information about the more volatile species evolved during co-pyrolysis. The ion fragment signals presented in Table 1 represent different families of compounds which were monitored during the co-pyrolysis.
Table 1. Ion fragments monitored by TG-MS analysis. m/z
Assignment
2
H2+
15
CH3+
29, 43, 57
Alkane series: C2H5+, C3H7+, C4H9+…CnH2n+1+
27, 41, 55
Alkene series: C2H3+, C3H5+, C4H7+ … CnH2n-1+
77, 78, 91
Aromatic series: C6H5+, C6H6+, C7H7+
44
CO2+
Above 450 ºC, the most abundant hydrocarbon during the pyrolysis of the coal blend and its mixtures with plastics was methane which was accompanied by the release of other aliphatic and aromatic hydrocarbons. Paraffinic and olefinic fragments always evolve in the temperature range of 495 to 550 ºC, the temperature increasing as the number of carbon atoms increases. In most of the blends containing plastics the maximum temperature of hydrocarbons occurs at a lower temperature than the coal blend PA. Hydrogen was detected in more intensity in the last stages of thermal decomposition of the blends and coming mainly from the condensation of the aromatic
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Oviedo ICCS&T 2011. Extended Abstract
structures at a high temperature (782-790 ºC). When the compositions of the light pyrolysis products from the coal blend and its mixtures with plastics are compared, some relevant features of the co-pyrolysis of the blends made up of coal and plastics, (which were derived from the normalized areas of the corresponding peaks to that of hydrogen), are shown to be related to: (i) a higher proportion of hydrogen; (ii) a higher amount of aliphatic compounds from C2 to C4 in the form of both alkanes and alkenes; (iii) a higher relative proportion of alkenes versus alkanes, with the exception of the PA5PS blend (Figure 3).
175
0.8
170
0.7 0.6
C3
Intensity of alkene ion fragments
10 8 6 4 2
C2
18
C3
a
+5 W PA
PA
+5 W
T +5 PE
PS
PP
PA +5
PA
PA
20
C4
PA +5
PA
b PA
+5 W
+5 W
T PA
PA
+5 PE
PP PA +5
PA +5
PE
E
+5 LD PA
+5 HD P PA
C2
12
PE
0.0
a
0.1
130
PS
0.2
135
b
0.3
140
14
C4
16 14 12 10 8 6 4 2
b PA +5 W
a
T
PA +5 W
5P E PA +
+5 PS PA
+5 PP PA
E LD P PA +5
+5 HD PE PA
W b PA +5
W a PA +5
T 5P E PA +
+5 PS PA
+5 PP PA
E LD P PA +5
PA
PA
0
+5 HD PE
0
PA
Intensity of alkane ion fragments
0.4
E
145
+5 LD
150
0.5
+5 HD P
155
PA
Paraffin/Olefin ratio
160
PA
+
+
CH3 /H2 ratio
165
Figure 3. Variation of hydrocarbons evolved during the pyrolysis of the coal blend PA and its mixtures with the different plastic wastes.
As a consequence of the polymer structure, blends made up of PS and PET behave in a different way to polyolefins and they increase the amount of aromatic hydrocarbons. As expected, the addition of oxygen-containing polymers such as PET increases the CO2 content in the gas, which is released at low temperatures of approximately 465 and 635 ºC. Although it is difficult to attribute the origin of the two peaks of CO2, the low-temperature peak could be linked to functional groups in the PET polymeric chain, whereas the second peak may be assigned to a cross-linking reaction during the formation of tar and coke.
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4. Conclusions A delay in the evolution of volatile species from the plastics was observed when they are blended with coal. Depending on the thermal stability of the plastics, the shift of the evolution of volatiles from the plastic towards higher temperatures and the increase in overlapping between the components may explain the degree of reduction in fluidity caused by plastics. In general, the relative proportion and the temperature of emission of light gases such as hydrogen, methane, aliphatic hydrocarbons with up to four carbon atoms (including paraffin and olefin pairs), aromatic hydrocarbons and carbon dioxide was consistent with the functional groups of the plastic added to the coal. The thermal events during co-pyrolysis and the chemical families of compounds in the gas are in agreement with the modification of the coal fluidity, the degree of ordering of the carbon structure of the semicokes and the evolution of gas pressure during the coking process.
Acknowledgements The financial support provided by the Ministerio de Ciencia e Innovación through project CTM2004-03254 is gratefully acknowledged.
References [1] [2] [3] [4]
[5] [6]
[7]
Kato K, Nomura S, Uematsu H. Development of waste plastics recycling process using coke ovens. ISIJ International 2002;42:S10. Nomura S, Kato K. Basic study on separate charge of coal and waste plastics in coke oven chamber. Fuel 2005;84:429-35. Diez MA, Alvarez R, Melendi S, Barriocanal C. Feedstock recycling of plastic wastes/oil mixtures in cokemaking, Fuel 2009;88:1937-44. Melendi S, Diez MA, Alvarez R, Barriocanal C. Plastic wastes, lube oils and carbochemical products as secondary feedstocks for blast-furnace coke production, Fuel Processing Technology2011;92: 471-8. Melendi S, Diez MA, Alvarez R, Barriocanal C. Relevance of the composition of municipal plastic wastes for metallurgical coke production, Fuel 2011;90:1431-8. Vivero L, Barriocanal C, Alvarez R, Diez MA. Effects of plastic wastes on coal pyrolysis behavior and the structure of semicokes. J. Anal. Appl. Pyrolysis 2005;74:327-36. Dominguez A, Blanco CG, Barriocanal C, Alvarez R, Diez MA. Gas chromatographic study of the volatile products from co-pyrolysis of coal and polyethylene wastes. J. Chromatography A 2001;918:135-44.
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Oviedo ICCS&T 2011. Extended Abstract
[8]
Nomura S, Kato K, Nakagawa T, Komaki I. The effect of plastic addition on coal caking properties during carbonization, Fuel 2003;82:1775-82. [9] Sakurovs R. Interactions between coking coals and plastics during co-pyrolysis. Fuel 2003;82:1911-16. [10] Diez MA, Barriocanal C, Alvarez R. Plastic wastes as modifiers of the thermoplasticity of coal, Energy and Fuels 2005;19:2304-16. [11] Castro Díaz M, Edecki L, Steel KM, Patrick JW, Snape CE. Determination of the effects caused by different polymers on coal fluidity during carbonization using high-temperature 1H NMR and rheometry. Energy & Fuels 2008;22:471-79.
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Oviedo ICCS&T 2011. Extended Abstract
Role of selected coal- and petroleum-based additives in low- and hightemperature co-pyrolysis with coal blends E. Rodríguez, S. Melendi, R. García, R. Alvarez, M.A. Diez Instituto Nacional del Carbón, INCAR-CSIC, Apartado 73, 33080-Oviedo. Spain.
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Abstract The co-pyrolysis of a coal blend with coal- and petroleum-based additives was investigated under slow-heating rate conditions and final temperatures of 600 and 900 ºC. A series of four additives was selected on the basis of their proportion of aliphatic hydrocarbons in the composition of the additive and/or in the primary tar obtained from the pyrolysis. They included a low-rank coal (HVC), a deposit from coke oven gas pipelines (TUB), high-density polyethylene (HDPE) from domestic containers and a lubricating oil of petrochemical origin (LUB). Each additive was added to a coal blend used in the production of blast-furnace coke at addition rates of 2 and 5 wt%. The main objective of this study is to determine how these additives affect the pyrolytic and rheological behaviour of the coal blend, the composition of tars and the microstructure of the semicokes and cokes obtained at lab-scale. All the additives were observed to decrease the semicoke and coke yields at the expense of tar and gas formation. Both HVC and TUB exhibited similar trends in the yields of the major pyrolysis products (semicoke/coke, tar and gas), enhancing the formation of gas species, whereas LUB and HDPE promoted the molecular species that make up the tar. Although other factors also need to be considered, the amount of heavier hydrocarbons in the primary tars obtained from every coal+additive mixture is related to the reduction of coal fluidity caused by the additive. All the additives produced cokes with a more disordered carbon microstructure, as was detected by Raman spectroscopy. Keywords: coal, wastes, co-pyrolysis, fluidity, tars, semicoke, coke
1. Introduction The use of carbon sources as inert and reactive additives to coal blends is a subject that is continually arising in cokemaking, in relation with the improvement of the
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Oviedo ICCS&T 2011. Extended Abstract
coking properties of blends, to safeguarding of good-quality coal, the recycling of wastes and the reduction of costs of raw materials. Chemical and physical properties of cokes are known to be closely related to their structure, which clearly depends on the fluidity of the parent coal [1-3]. For this reason, fluidity developed between 350 and 500 ºC is one of the coal properties that is used to predict the blending potential of additives used in cokemaking. The degree of interactions between components determines the influence of a specific additive not only on the thermoplastic properties of coal, but also on the composition of the carbonization by-products (tar and gas) and on the structure of high-temperature coke. Among the wide spectrum of coal- and petroleum-derived additives for cokemaking [4-9], three wastes were selected to evaluate the modifications in thermoplastic behaviour of a coal, the distribution of the primary products from cocarbonization, the composition of the primary tar and the coke structure. Two of the wastes were generated in an integrated steel plant: a deposit from coke oven gas pipelines and a lubricating oil of petrochemical origin used in the steel rolling mills. The other waste used was a high-density polyethylene (HDPE) from domestic containers. For comparison purposes, a non-coking coal of high volatile matter content was also used. These four additives were selected on the basis of the presence of aliphatic hydrocarbons in their composition (lub oil and deposit from COG pipeline) and/or in the primary tar obtained from their pyrolysis (non-coking coal and HDPE).
2. Experimental section Raw materials and wastes Coal blend A was prepared and supplied by ArcelorMittal in Spain. The main characteristics of coal blend A are as follows: ash, 8.9 wt% db, volatile matter, 22.5 wt% db, S, 0.57 wt% db, Gieseler maximum fluidity, 312 ddpm. The additives used included a low-rank coal (HVC) used for pulverized coal injection in blast furnaces, a deposit from the coke oven gas pipelines (TUB), highdensity polyethylene (HDPE) from domestic containers and a lubricating oil of petrochemical origin (LUB). Each additive was added to coal blend A at addition rates of 2 and 5 wt% and, then, subjected to Gieseler plastometry and co-pyrolyzed in a labscale horizontal oven.
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Oviedo ICCS&T 2011. Extended Abstract
Gieseler fluidity development The Gieseler fluidity was determined in an R.B. Automazione PL2000 plastometer, following the ASTM D2639 procedure. The specific parameters use to measure the fluidity development of the coal blend and its mixtures are: (i) the softening temperature at which the coal starts to be fluid; (ii) the temperature of maximum fluidity reached during the thermal heating; (iii) the resolidification temperature at which the fluid mass resolidifies into a semicoke; and (iv) maximum fluidity, expressed as dial divisions per minute (ddpm).
Co-pyrolysis and products characterization Co-pyrolysis experiments were performed on mixtures of coal blend A and the selected additives in a Gray-King type furnace. A sample of about 8 g with a particle size of < 0.212 mm was pyrolyzed from ambient temperature to a final temperature of 600 or 900 ºC at a heating rate of 5 ºC/min in an atmosphere of evolving gases, with a soaking time of 15 min. The condensable products (primary tar) obtained during the pyrolysis experiment were collected by means of an ice-cooled trap. Primary tars were separated from the decomposed water by decantation before subsequent analysis. The semicoke/coke and primary tar yields were calculated relative to the starting material on a dry basis and the non-condensable gas fraction was calculated by difference.
Characterization of the pyrolysis products The primary tars were characterized by Fourier Transform infrared spectroscopy (FTIR) in transmission mode and by gas chromatography using a flame ionization detector and a mass spectrometer (GC-FID-MS). FTIR spectra were recorded on a Nicolet IR8700 spectrometer equipped with a DTGS detector. The sample was deposited as a thin film between NaCl windows and subjected to 64 scans at a resolution of 4 cm-1. For the semiquantitative analysis, the ratio between the integrated areas of the characteristic absorption bands corresponding to aliphatic hydrogen (2990-2750 cm-1) and aromatic hydrogen (3100-2990 cm-1) was calculated. Gas chromatographic analyses of the primary tars were carried out on an Agilent Technologies Model 6890N Series II gas chromatograph coupled to a mass selective detector 5973 N. The experimental conditions used were described elsewhere [9]. The Raman spectra of the cokes obtained at 900 ºC were performed on a LabRam Submit before 31 May 2011 to
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Oviedo ICCS&T 2011. Extended Abstract
HR UV spectroscope from Jobin Yvon Horiba equipped with a CCD camera and an argon laser excitation source (λ = 514.5 nm). The power source used was 24.3 mW. An Olympus M Plan optical microscope with a 100x objective lens was used to focus the laser beam. Each first-order Raman spectrum for the coke was deconvoluted to obtain four main components at 1595 -G band-, 1520, 1345 -D band- and around 1200 cm-1 [10]. The D/G ratio was used as a measure of the degree of structural order in the carbon matrix of the cokes.
3. Results and Discussion Effect of the additives on the development of fluidity Table 1 shows the main thermoplastic parameters of coal blend A and its mixtures with the additives at an addition rate of 2 and 5 wt%. The additives differ in their capacity to modify the plastic properties of coal. Except for the lubricating oil (LUB) that enhances the fluidity and extends the plastic temperature range, the other additives produce a decrease in the caking capacity of the coal blend. TUB is the strongest inhibitor of coal fluidity, whereas the behaviour of HDPE is very similar to that of the non-coking coal HVC. An additive which modifies the fluidity of the coal, but keeps it within the range of optimum values established for cokemaking has a good chance to produce a coke with an acceptable strength [11,12]. This is not the case for the deposits from the COG pipelines (TUB), due to the fact that the fluidity is lower than 200 ddpm and the temperature fluid interval is too narrow. Table 1. Thermoplastic parameters of coal blend A and its mixtures with the selected additives. Blend A
Fmax (ddpm) 312
Ts (ºC)
Tf (ºC)
Tr (ºC)
Tr-Ts (ºC)
Fmax variation (%)
414
452
488
75
-
A2HVC
248
414
453
486
72
-20
A5HVC
222
411
452
484
73
-28
A2TUB
137
416
452
485
69
-56
A5TUB
30
426
456
483
57
-90
A5HDPE
197
409
454
487
78
-37
A5LUB
684
402
450
489
87
+119
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Oviedo ICCS&T 2011. Extended Abstract
The data also indicate that additives reduce Fmax more than can be accounted for a simple dilution of the components. In order to understand the interchemical reactions between the coal and additives, co-pyrolysis of the mixtures was also performed.
Co-pyrolysis of the coal blend and its mixtures with the additives Table 2 shows the distribution of the pyrolysis products obtained at 600 and 900 ºC. The major pyrolysis product at 600 ºC is semicoke, its yield varying from 83.8 wt% for coal blend A to nearly 79 wt% for the mixtures containing HDPE and LUB. The lower semicoke yield is a consequence of the distillation of the hydrocarbons which make up the lubricating oil (LUB) and the transformation of HDPE into hydrocarbons which enhances the tar fraction yield. The other two additives, the low-rank coal and TUB, have different effects on the distribution of the co-pyrolysis products, which increases the yield of the non-condensable products (gas fraction). The effect on the distribution of the co-pyrolysis products at 600 ºC is confirmed at the higher temperature of 900 ºC. In the temperature interval between 600 and 900 ºC, no more tar is formed and the transformation of semicoke to coke involves the release of small molecules which pass to the gas fraction. Table 2. Distribution of pyrolysis products at 600 and 900 ºC. 600
Temperature (ºC)
A A2HVC A5HVC A2TUB A5TUB A5HDPE A5LUB
900
Semicoke (wt%)
Tar (wt%)
Gas (wt%)
Coke (wt%)
Tar (wt%)
Gas (wt%)
83.8 81.4 81.0 82.6 81.7 79.2 79.1
8.9 8.2 8.4 8.5 8.5 11.9 13.6
7.4 10.4 10.6 8.9 9.8 8.9 7.3
77.3 77.8 76.9 78.1 77.7 75.0 73.8
8.9 8.0 8.6 7.2 7.4 11.2 13.9
13.8 14.2 14.5 14.6 14.9 13.7 12.3
The chemical composition of the primary tars obtained at 600 ºC was determined by GC-FID-MS and FTIR. In complex mixtures, an evaluation of the different chromatographic regions or hydrocarbon families is a common practice and is used to explain the differences in fluidity induced by the addition of several organic additives to coal [13-14]. Submit before 31 May 2011 to
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Oviedo ICCS&T 2011. Extended Abstract
In terms of chemical families, tars are mainly composed of phenol derivatives (14-18%, except for A5HDPE with about 8%), aliphatic compounds -paraffinic and olefinic hydrocarbons- (20-36%) and aromatic hydrocarbons and their highly-alkylated derivatives. The latter fraction is the most abundant in the tar, 46-67%, and it is mainly constituted by the hydrocarbons range from benzene to those containing three aromatic rings such as phenanthrene and anthracene. The aliphatic hydrocarbons, both the long-chain saturated and unsaturated types, dominate the tar obtained from the blend with HDPE (Figure 1). This tar contains a greater amount of this family of hydrocarbons. The amount and the distribution of aliphatic compounds in the A5HDPE tar resembles that obtained of single coal HVC which has a higher amount of aliphatic compounds, representing about 35% of the chromatographiable fraction of the tar. In general, the amount of aliphatic compounds evaluated by GC is consistent with the aliphatic hydrogen content estimated by FTIR. 40 paraffinic
35
olefinic
Aliphatic fraction (%)
30 25 20 15 10 5 0 A
A2HVC
A5HVC
A2TUB
A5TUB A5HDPE A5LUB
HVC
Figure 1. Distribution of aliphatic hydrocarbons, paraffinic and olefinic compounds, in the tars obtained at 600 ºC. Three chromatographic regions were also defined in the chromatograms on the basis of the compositional ranges used to characterize coal and petroleum fuels: a lighter fraction of tar which includes the low-boiling point compounds with a retention time lower than that of n-dodecane (C12); an intermediate fraction which is composed of medium-boiling point compounds with a retention time between that of C12 and n-nonadecane (C19); and the heaviest molecular-weight hydrocarbons which elute from C19 to the end of the chromatogram [11-12]. The tars are dominated by the intermediate fraction C12-C19 (Figure 2), except for the A5HDPE tar that contains a similar percentage of this fraction to the heaviest hydrocarbons (>C19), and a very low Submit before 31 May 2011 to
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Oviedo ICCS&T 2011. Extended Abstract
proportion of the lighter fraction (
60
C12-C19
>C19
Proportion of tar fraction (%)
50 40 30 20 10 0 A
A2HVC
A5HVC
A2TUB
A5TUB
A5HDPE
A5LUB
HVC
Figure 2. Distribution of the fractions of tars produced in Gray-King pyrolysis at 600 ºC. Although no clear relationship was found between any parameter relating the chemical composition of all the tars studied and the reduction in fluidity of the parent blend, there is a tendency towards an increase in the heavier hydrocarbon fraction (>C19) as the fluidity decreases (Figure 3). However, other chemical families present in the two tars A5HDPE and A5LUB should be considered to explain the different behaviour. 45
Amount of >C19 fraction (%)
40 35 30 25 20 15 A Fmax (ddpm) 312
A2HVC
248
A5HVC
A2TUB
A5TUB
224
137
30
A5LUB
A5HDPE
HVC
684
197
8
Figure 3. Variation of the amount of heavy fraction (>C19 ) in the tars obtained at 600 ºC with the parent blend fluidity.
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Oviedo ICCS&T 2011. Extended Abstract
The development of the fluidity of the coal in the presence of the additives clearly influences the structure of high-temperature cokes. The decrease in the Gieseler maximum fluidity in the parent blend results in a less organized carbon structure with various forms of structural defects and imperfections in the graphitic microcrystallites as is reflected by the increase in the D/G band ratio of the cokes obtained at 900 ºC (Figure 4). The exception to the general trend is the coke produced from the blend with the lubricating oil (A5LUB, not included in the graph). The increase in the fluidity of the coal blend from 312 ddpm to 684 ddpm when the lube oil gives rise to a coke with a similar carbon structure to that of the coal blend. 2.0 A5TUB
D/G ratio of coke
1.9
A5HDPE
1.8
A2TUB A2HVC
1.7
A5HVC
1.6 A 1.5 0
50
100
150 200 Gieseler Fmax (ddpm)
250
300
350
Figure 4. Variation of the band area ratio (D/G) for cokes obtained at 900 ºC (D and G bands at 1320 and 1595 cm-1, respectively) with Gieseler maximum fluidity of the parent blends.
4.
Conclusions
All the additives tested increase the volatility of the products by limiting the cokeforming processes and consequently enhancing the formation of tar and gas fractions. Gray-King pyrolysis provides useful information on the effect of the additives on the distribution of coke, tar and gas. HDPE and lubricating oil mainly contributed to the formation of a tar with heavier hydrocarbons, whereas the other two additives are mainly recovered as the non-condensable fraction. The degree of inhibition in Gieseler fluidity induced by the additives is related to the more disordered structure of the cokes produced at high-temperature as was detected by Raman spectroscopy.
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Oviedo ICCS&T 2011. Extended Abstract
Acknowledgements The financial support provided by PCTI-Asturias, Spain, through project PEST08-07 is gratefully acknowledged. We are also grateful to ArcelorMittal in Spain for its collaboration.
References [1]
Clarke DE, Marsh H, Mechanisms of formation of structure within metallurgical coke and its effect on coke properties, Erdöl und Kohle 1986;39:113-122.
[2]
Patrick JW. The coking of coal, Sci. Prog. 1974:61:375-399.
[3]
Loison R, Foch P, Boyer A (eds). Coke quality and production; Butterworth London; 1989.
[4]
Diez MA, Alvarez R, Barriocanal C, Coal for metallurgical coke production: predictions of coke quality and future requirements for cokemaking. Int. J. Coal Geology 2002;50:389-412.
[5]
Mochida I, Marsh H, Grint A. Carbonization and liquid-crystal (mesophase) development. 12. Mechanisms of the co-carbonization of coals with organic additives. Fuel 1979; 58:803-808.
[6]
Valia HS, Hooper W. Use of reverts and non-coking coals in metallurgical cokemaking, ISS Ironmaking Conf. Proc. 1994;53:89-105.
[7]
Menéndez JA, Pis JJ, Alvarez R, Barriocanal C, Fuente E, Diez MA. Characterization of petroleum coke as an additive in metallurgical cokemaking. Modification of thermoplastic properties of coal. Energy and Fuels 1996:10, 12621268.
[8]
Barriocanal C, Álvarez R, Canga CS, Diez MA. On the possibility of using coking plant waste materials as additives for coke production, Energy and Fuels 1998;12:981-989.
[9]
Diez MA, Alvarez R, Melendi S, Barriocanal C. Feedstock recycling of plastic wastes/oil mixtures in cokemaking, Fuel 2009;88:1937-1944.
[10] Cuesta A. Dhamelincourt P, Laureyns J, Martinez-Alonso A, Tascón JMD, Raman microprobe studies on carbon materials, Carbon 1994;32:1523-1532. [11] Miyazu T, The evaluation and design of blends using many kinds of coals for cokemaking. In: Int. Iron Steel Cong., Dusseldorf, 1974, Pap. 1.2.2.1. [12] Nakamura N, Togino Y, Adachi T. Philosophy of blending coals and cokemaking technology in Japan. In: Coal, coke and the blast furnace, The Metals Society, Submit before 31 May 2011 to
[email protected]
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1977, UK, p. 93-106. [13] Diez MA, Domínguez A, Barriocanal C, Alvarez R, Blanco CG, Casal MD, Canga CS, Gas chromatographic study for the evaluation of the suitability of bituminous waste material as an additive for coke production. Journal of Chromatography A, 1998;823:527–536. [14] Domínguez A, Blanco CG, Barriocanal C, Alvarez R, Diez MA. Gas chromatographic study of the volatile products from co-pyrolysis of coal and polyethylene wastes, Journal of Chromatography A 2001;918:135-144.
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Oviedo ICCS&T 2011. Extended Abstract
Effect of volatile matter evolution on optical properties of macerals from different rank coals A. Guerrero, M.A. Diez, A.G. Borrego Instituto Nacional del Carbón (INCAR-CSIC). Francisco Pintado Fe, 26, 33011 Oviedo, Spain Abstract The behavior of coal macerals in carbonization has been a topic of research for many years since the early recognition that the macerals behaved differently during the coking process. In addition, the coke optical texture and the amount of unreactives in the coke matrix have shown to be related to parameters relevant for coke properties such as reactivity and strength. The coals involved in this study are Polish coals of the Upper Silesian coal basin ranging in rank from high volatile to medium volatile bituminous . The experimental approach used in this study consisted on analyzing both the raw coal and the heated coal using optical microscopy and reflectance analysis. For each measured spot, both the reflectance and the component identified were recorded. The coals were heated in a thermogavimetric analyser at slow heating rate under inert gas flow up to temperatures in the range 450 to 1000 °C. Once the sample reached the desired temperature the oven was cooled down to stop the reaction and the cokefied residue recovered for petrographic analysis. Only the coals with reflectance over 1.0 % passed through a fluid stage in which integrity of the particles was lost. The size of the optical texture of the re-solidified material increased with the rank of the parent coal. The reflectance of all the macerals increased with increasing heating temperature, being the treatment final temperature the most relevant factor in determining the reflectance of inertinite macerals. The higher the heating temperature, the higher the reflectance with little influence of the parent inertinite reflectance. For those coals in which vitrinite did not pass through a fluid stage the distinction between vitrinite- and inertinite-derived material in the heated samples was very difficult. Inertodetrinite was virtually indistinguishable, whereas only relic cellular structure allowed distinguishing between vitrinite and semifusinite.
1. Introduction The behavior of coal macerals in carbonization has been a topic of research for
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Oviedo ICCS&T 2011. Extended Abstract
many years since the early recognition that the macerals behaved differently during the coking process (Shapiro and Gray 1964). In addition the coke optical texture and the amount of unreactives in the coke matrix has shown to be related to parameters relevant for coke properties such as reactivity and strength (Mackowsky, 1976). Despite the numerous studies dealing with the transformation of coal macerals upon heating, the issue is far from being solved. Different authors have identified as relevant factors determining the plastic behaviour of coal macerals the reflectance of the coal to which they belong (Taylor et al., 1967; Falcon and Snyman, 1986; Diessel and Wolff-Fischer, 1987), their individual reflectance (Diessel, 1983; Shapiro and Gray, 1960) or the characteristics of the whole coal reflectogram (Kruszewska, 1990). Other aspects not exclusively related with maceral reflectance are the type of maceral (Shapiro et al., 1965; Nandi et al., 1977) the fluorescence intensity (Diessel, 1985) or the degree of maceral association (Rentel, 1987), the latter recognized as related with the plastic properties of macerals. This topic is being revisited in a project of the European research fund for coal and steel focused on optimizing coking coal blends (RATIO-COAL).
2. Experimental section Six Polish coals in the rank interval of those used for metallurgical coke production and their cokefied products at different temperatures have been studied by optical microscopy. The experimental approach used is study consisted on performing a combined maceral-reflectance analysis at maceral level. For each coal 500 reflectance values were recorded and saved together with the corresponding maceral assignment. In this way detailed information of the maceral composition and also of maceral reflectance distribution is obtained. The coals were heated in a thermogavimetric analyser under a N2 flow of 5 0 mLmin-1 at a heating rate of 5 °Cmin-1. The maximum heating temperature ranged from 450 to 1000 °C. Once the sample reached the desired temperature the oven was cooled down to stop the reaction and the cokefied residue recovered for petrographic analysis. In the cokefied material around 250 instead of 500 reflectance points were recorded and also the corresponding maceral was identified. The degree of detail in the analysis of cokefied material was less than for the raw coal, first because the temperature obscured some of the maceral distinguishing features and also because the amount of sample recovered from the thermobalance was rather small.
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Oviedo ICCS&T 2011. Extended Abstract
3. Results and Discussion The coals involved in this study are from the Upper Silesian coal basin (Fig. 1) ranging in rank from high volatile to medium volatile bituminous (Table 1).
Figure 1 – Upper Silesian Coal Basin map showing the location of the samples studied
The coals have ash content below 10% and low sulfur contents. The Giseler test indicated a very low fluidity for all the coals, lower than expected for their rank. It should therefore be considered that the coals have lost the plastic properties due to alteration or weathering. Fluidity is very sensitive to exposure of the coals to air and prolonged storage times.
Table1. Proximate, ultimate and petrographic analyses of the studied coals. Ash V.M. C H N Odif ST Fmax VRr V I L Ck db % daf % ddpm (%) vol % mmf MR 3.3 33.82 85.34 5.09 1.40 7.59 0.59 8 0.81 68.6 27.4 3.6 0.4 SZ 7.13 33.06 86.35 5.16 1.55 6.18 0.76 248 0.89 68.4 27.6 3.0 0.2 PW 6.98 25.85 89.03 4.81 1.44 4.13 0.59 120 1.12 65.4 33.0 1.2 0.4 ZF 7.31 25.21 89.11 4.73 1.41 4.19 0.56 93 1.14 62.6 32.0 2.4 3.0 BY 6.42 25.05 89.21 4.84 1.59 3.80 0.56 183 1.17 66.4 31.6 2.0 0.0 JS 7.16 21.28 90.28 4.60 1.29 3.33 0.50 63 1.29 50.4 46.4 1.0 0.2 VM=Volatile matter content, Fmax=Maximum Giseler Fluidity, VRr=Vitrinite reflectance, V=vitrinite, I=inertinite, L=liptinite; Ck=natural coke, db=dry basis, daf=dry-ash-free basis ddpm=dial divisions per minut,. vol= volume, mmf=mineral matter free basis
The maceral analysis of the coals indicated a moderately high vitrinite content for most of the coals (around 65%) and inertinite content around 30%, except for JS coal, which Submit before January 15th to
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Oviedo ICCS&T 2011. Extended Abstract
exhibited balanced inertinite and vitrinite contents. All the coals had low liptinite contents in all of them the presence of coke was detected in very low amounts.
Figure 2 shows the volatile matter lost at different temperatures of the studied coals (a) and the vitrinite reflectance reached as a function of the normalized volatile release (b). Three groups can be established with the thermal behavior of the studied coals: Those with vitrinite reflectance below 0.9% lose volatiles at lower temperature and exhibited a higher devolatilization rate. The coals with vitrinite reflectance between 1.12 and 1.17% exhibited nearly identical thermograms and have similar volatile matter content despite their difference in rank due to the differences in maceral composition. The lowest rank coal of the group having the highest inertinite content and viceversa. The highest rank coal released the volatiles at higher temperatures and with the lowest devolatilization rate. 35
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Figure 2. a).Volatile matter lost at different temperatures for the studied coals and b). Vitrinite reflectance reached as a function of the normalized volatile release (b). The effect of devolatilization on the reflectance of vitrinite is seen in Figure 2b. It is observed that there is a low increase in reflectance until the volatile release has not reached around 67-75%. Below this threshold vitrinite reflectance is maintained below 2% and above this threshold the reflectance raised quickly reaching values around 7% at 1000 ºC. Figure 3 shows an example of the aspect of coal JS and the same sample heated at 550 and 1000 °C. The main features found for the different samples can be summarized as follows: -
Vitrinite from ZF, PW, BY and JS fused upon heating passing through a plastic stage in which a mosaic structure was formed. The size of the mosaic-forming grains was
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Oviedo ICCS&T 2011. Extended Abstract
larger in JS than in the other coals. Coals MR and SZ did not melt and the individual grains maintained their integrity although significant changes occurred in the vitrinite whose reflectance approached to the inertinite reflectance. Incipient anisotropy was only observed in the highest temperature sample of the MR series, whereas in SZ series incipient anisotropy was observed in lower temperature samples. -
The reflectance of all the macerals increased with increasing heating temperature and the treatment final temperature was the most relevant factor in determining the reflectance of inertinite macerals. The higher the heating temperature, the higher the reflectance with little influence of the parent inertinite reflectance.
-
For those coals in which vitrinite did not pass through a fluid stage the distinction between vitrinite- and inertinite-derived material was very difficult. Inertodetrinite was virtually indistinguishable, whereas only relic cellular structure allowed distinguishing between vitrinite and semifusinite.
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In the lowest ranked coals it was observed the formation of a high relief and high reflectance, supposedly carbon-rich structure in the inner part of the vitrinite grains.
550 ºC Rr=1.29%
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i
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Figure 3. Appearance of Coal JS and its cokefied material at different temperatures. Rr=random reflectance, i=inertinite, v-d=vitrinite-derived. Optical micrographs under incident white light. The evolution of the two maceral groups reflectance which are identifiable during the whole heating process is shown in Figure 4. The figure clearly shows a convergence of the reflectance values of both maceral groups upon heating. Vitrinite and inertinite reflectances become very close at 550 ºC and typically at 1000 ºC the reflectances of vitrinite surpass that of inertinite. In addition, as observed in the case of vitrinite, the main factor determining the inertinite reflectance is the peak temperature at which the sample was heated. The increase in inertinite reflectance up to 550 ºC was rather low, but further increase in temperature resulted in a drastic increase in the reflectance values.
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Oviedo ICCS&T 2011. Extended Abstract
8
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Figure 4. Evolution of vitrinite and inertinite reflectances with temperature for the studied coals.
4. Conclusions ·The lowest rank coals (MR and SZ) did not pass through a fluid stage leading to the formation of a matrix, whereas coals with vitrinite reflectance 1.12-1.29% passed through a plastic stage forming an anisotropic matrix in which the size of the mosaics increased with the rank of the coals.
·The reflectance of all the macerals increased with increasing the treatment temperature obscuring some of the distinguishing features of the macerals. For vitrinite which is maintained isotropic, the distinction between vitrinite and inertinite is difficult in the heated samples and the formation of a carbon-rich secondary structure is observed in the highest temperature samples. Semifusinite loses part of the cellular structure from 550 ºC becoming difficult to distinguish from macrinite. In addition inertodetrinite is also difficult to distinguish from the matrix.
·The reflectance reached by both vitrinite and inertinite is more closely related to the treatment temperature than to the reflectance of the parent coal, nevertheless the vitrinite optical texture is very different for different coal ranks. The high volatile vitrinites lose faster their volatile matter remaining isotropic whereas the lower volatile vitrinite have more time to re-organize the carbonaceous residue to yield anisotropic ordered structures.
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Oviedo ICCS&T 2011. Extended Abstract
·The drastic increase in reflectance occurs from 550ºC and corresponds to a volatile loss of 67-75% of the coal volatile matter.
Acknowledgements: The research leading to these results has received funding from the Research Programme of the Research Fund for Coal and Steel (Grant Agreement number RFC-PR- 09024). I. Jelonek (U. Silesia) is gratefully acknowledged for providing the coal samples for this study.
References: Diessel C.F.K.., 1983. Carbonization reactions of inertinite macerals in Australian coals.. Fuel 62, 883-892 Diessel, C.F.K., 1985. Fluorometric analysis of intertinite, Fuel 64, 883–892. Diessel, C.F.K.; Wolff-Fischer, E., 1987. Coal and coke petrographic investigations into the fusibility of Carboniferous and Permian coking coals Int. J. Coal Geol. 9,87-. Falcon, R.M.S. and Snyman, C.P., 1986. An introduction to coal petrography: atlas of petrographic constituents in the bituminous coals of Southern Africa. Geol. Soc. South Afr. Rev. Paper 2, p. 27 Kruszewska K., 1989The use of reflectance to determine maceral composition and the reactive-inert ratio of coal components Fuel, 68, 753-757 Mackowsky, M-Th., 1976. Prediction methods in coal and coke microscopy. J. Microscopy. 109, 119-137 Nandi B.N., Brown T.D., Lee G.K. 1977. Inert coal macerals in Combustion Fuel, 56, 125-130. Rentel, K., 1987. The combined maceral-microlithotype analysis for the characterization of reactive inertinites. International Journal of Coal Geology, 9, 77-86 Schapiro N.; Gray, R.J. 1960 Petrographic classification applicable to coals of all ranks, Proc. Ill. Mining Inst., 68th Year 83–97. Schapiro, N., Gray, R.J., Eusner, G.R. (1965). Recent developments in coal petrography. Blast Furnace, Coke Oven and Raw Materials Committee, Proc, 20, 89-112. Schapiro, N.; Gray, R.J., 1964. The use of coal petrography in coke making, J. Inst. Fuel 37, 234–242. Taylor G.H., Mackowsky M.Th.; Alpern, B. 1967. The behaviour of inertinite during carbonisation. Fuel 46, 431–440.
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Oviedo ICCS&T 2011. Extended Abstract
Semi-pilot scale carbonization to assess blast furnace coke quality E. Díaz-Faes, R. Alvarez, C. Barriocanal, M.A. Díez Instituto Nacional del Carbón (INCAR), CSIC, Apartado 73, 33080 Oviedo. Spain
[email protected] Abstract A series of coals of different rank and geographical origin were carbonized in two movable wall ovens (MWO) of 320 kg (pilot oven) and 15 kg (semi-pilot oven) at INCAR. All the coals used were chosen from among those commonly employed by the cokemaking industry in blend preparation. The quality of the cokes obtained from the two MWO`s was assessed in terms of reactivity towards CO2 by the Nippon Steel Corporation test (NSC test) and the ECE-INCAR test. A good correlation was found between the reactivity to carbon dioxide as determined by the NSC and ECE-INCAR methods. Furthermore, using a semi-pilot oven in preference to large-capacity ovens yields valuable results due to the small amount of coal employed (15 kg vs. 300-400 kg) since it is quicker, more flexible and leads to lower costs.
1. Introduction Metallurgical coal is a macroporous material obtained from coking coals or blends of coals by heating them in the absence of oxygen. In the blast furnace, coke plays three roles: as a fuel, providing the energy necessary for the process; as chemical agent, acting as reducing agent of the iron ore and as a permeable support of the burden, allowing the gases to pass through the burden to the upper part of the blast furnace and enabling the liquids to pass down to the crucible. For the two first functions, coke can be replaced by other fuels and carbon sources such as pulverized coal, plastics, oils, etc. [1-3], but there is no other material that can fulfill the role of permeable support of a blast furnace charge. This substitution has led to a reduction of the coke rates in the blast furnace over recent years. In other words, the quality of the coke has acquired greater importance in our efforts to keep the blast furnace operating in optimal conditions [4]. One of the most important properties of coke is its ability to perform well in oxidizing conditions and at the high temperatures characteristics of a blast furnace [5]. Japanese industry has developed a method of measuring the reactivity of a coke towards carbon dioxide and the resistance to the abrasion of the partially gasified coke (NSC test, ASTM-D 5341). INCAR has been using a modification of the ECE reactivity test which 1
Oviedo ICCS&T 2011. Extended Abstract
requires a smaller amount of coke and is less expensive and time consuming. The aim of this work was to establish relationships between the quality of the cokes obtained at two different scales, semi-pilot and pilot, using two movable wall ovens (MWO) of different capacity, 15 and 320 kg, respectively and between reactivity indices determined by the ECE and NSC methods.
Experimental section 2.1. Carbonization tests at pilot and semi-pilot scale Carbonization tests at pilot scale were carried out in an electrically-heated moveable wall oven of 320 kg capacity (MWO320) with the following dimensions: 915 mm length, 840 mm height and 455 mm width. The initial coke-oven wall temperature during charging was 880 ºC, rising at a rate of 14 °C/h up to 1200 ºC by the end of the process and to 1050 °C in the centre of the charge. The coking time required was 18 h.
Coals were also carbonised in a semi-pilot moveable-wall oven of 15 kg of capacity (MWO15) fitted with electrical heating and having the following dimensions: 150 mm length, 750 mm height and 250 mm width. During the carbonisation tests, the temperature of the wall was kept constant at 1010 ºC. The coking time required to reach a temperature in the centre of the charge of 950 ºC was 3 hours. In the two-scale carbonizations tests, after the coke had been pushed from the oven, it was quenched with water. The moisture content of the coals to be carbonized in the two ovens was kept as close as possible to 5 wt%.
2.2. Coke quality 2.2.1. NSC test Coke reactivity towards CO2 (CRI) and coke strength after reaction (CSR index) were assessed by the test developed by the Nippon Steel Corporation (NSC) and standardized afterwards by ASTM D5341-99 (ISO 18894:2006). Briefly, a sample of coke (200 g) with a particle size between 19 and 22.4 mm was reacted at 1100 ± 5 ºC for 2 h with CO2 at a flow rate of 5 L/min. The partially gasified coke was weighed and subjected to the tumbler test. The CRI was calculated as the percentage of weight loss. The mechanical degradation of the coke after CO2 reaction (CSR) was measured as the weight of coke remaining on a 9.5 mm sieve after 600
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Oviedo ICCS&T 2011. Extended Abstract
revolutions at 20 rpm and it was calculated as the weight percentage of coke larger than 9.5 mm relative to the weight of the coke after the reaction with CO2. 2.2.2. ECE-INCAR reactivity test The procedure used is based on a recommendation by the European Commission in 1965. Whereby, reactivity is measured as the weight loss after 2 hours of reaction instead of by analysing the gas composition during the reaction [6]. To carry out the test a sample of 7 g coke with a particle size between 1 and 3 mm was introduced in a quartz reactor and kept at 1000 ºC in an electrical oven. The heating and cooling of the sample was carried out under N2. Once the temperature of the sample had stabilized, a flow of CO2 of 7.2 L h-1 was made to pass through the sample for 1 hour. After the reaction had finished, the sample was left to cool before being weighed. Coke reactivity was calculated as the percentage of weight loss.
2. Results and Discussion The coals were carbonized in two MWO`s available at INCAR with the aim of comparing the quality parameters of the cokes obtained using two scales: reactivity and hot mechanical strength. It is important to point out that the quality indices of the cokes produced in the semi-pilot oven did not numerically match those from the pilot oven, mainly due to wall effects and the different conditions applied during the coking process (i.e. coking rate). The relationship between the reactivity of the cokes produced in the two ovens yields good regression coefficients: 0.943 for reactivity assessed by means of ECE-INCAR reactivity test (Figure 1) and 0.965 for the CRI indices (Figure 2a), the values of MWO15 being slightly higher than those of MWO320. With some exceptions, the cokes produced in the semipilot oven are more reactive than those produced at pilot scale. The same trend was observed in the case of the mechanical strength of the partially gasified cokes (CSR indices), where the regression coefficient was 0.963 (Figure 2b).
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Oviedo ICCS&T 2011. Extended Abstract
18
RECE-INCAR (%), MWO320
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Figure 1. Relationship between RECE-INCAR indices for cokes produced in the two MWO`s.
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Figure 2. Relationship between CRI (a) and CSR (b) indices for cokes produced in the two MWO`s. Relationships between the ECE-INCAR reactivities and CRI and CSR indices are shown in Figures 3 and 4 for the two moveable wall ovens. The best correlation coefficients are obtained when the ECE-INCAR reactivities and CRI indices are compared in the two MWO`s, a value of r = 0.952 being obtained for the cokes produced in the pilot oven and r = 0.911 for the cokes produced in the semipilot oven. The correlations are worse when CSR index is compared with ECE-INCAR reactivity index, a correlation coefficient of r = 0.882 being obtained for the cokes of the 320 kg capacity oven and r = 0.830 for the cokes of the 15 kg capacity oven. This is due to the influence that bulk density has on the strength of the partially gasified cokes during coking.
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Oviedo ICCS&T 2011. Extended Abstract
70
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Figure3. Relationship between the RECE-INCAR and CRI (a) and CSR (b) indices for the cokes produced in the pilot oven.
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Figure4. Relationship between the RECE-INCAR and CRI (a) and CSR (b) indices for the cokes produced in the semi-pilot oven.
Conclusions The use of a semi-pilot carbonization oven enables coke quality to be determined by means of reactivity towards CO2 and coke strength after reaction and valuable results to be obtained with only a small amount of coal (15 kg vs. 320 kg). Additional advantages include a considerable saving of time, a reduction in cost and a more flexible process. ECE-INCAR test can be used as a guide to estimate coke quality in terms of its reactivity to CO2, although it cannot replace the NSC procedure. Acknowledgement. The authors thank the European Coal and Steel Community –ECSC- (project 7220PR/119) for financial support.
References [1] Cross, WB. The future of the European steel industry and its demand for coal. The Coke Oven Managers’ Association (COMA). Year-Book, Mexborough, United
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Kingdom, 1994, 97-108. [2] Asanuma M, Ariyama T, Sato M, Murai R, Nonaka T, Okochi I, Tsukiji H, Remoto K. Development of waste plastics injection process in blast furnace. ISIJ Int. 2000, 40: 244-251. [3] Ohji, M Production and technology of iron and steel in Japan during 1999. ISIJ Int. 2000, 40: 529-543. [4] Díez, MA, Álvarez R, Barriocanal C. Coal for metallurgical coke production: predictions of coke quality and future requirements for cokemaking. International Journal of Coal Geology 2002, 50: 389-412. [5] Sakawa M, Sakurai Y, Hara Y. Influence of coal characteristics on CO2 gasification. Fuel 1982, 61 (8): 717-720. [6] Menéndez JA, Álvarez R, Pis JJ. Determination of metallurgical coke reactivity at INCAR: NSC and ECE-INCAR reactivity test. Ironmaking and Steelmaking 1999: 117121.
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Oviedo ICCS&T 2011. Extended Abstract
Fundamental Investigation Pyrolysis Behaviour of Low Rank Coals Tatsuro Harada1), Seiichiro Matsuda1), Nozomi Wada2), Yohsuke Matsushita2)*, Isao Mochida2) 1) Research Laboratory, Kyushu Electric Power Co., Inc, 12-1-47 Shiobaru, Minami-ku, Fukuoka, Fukuoka, 815-8520 Japan 2) Research and Education Center of Carbon Resources, Kyushu University, 6-1 Kasuga-koen, Kasuga, Fukuoka, 816-8580 Japan Abstract Pyrolysis of cellulose, brown and bituminous coals were examined by TGA at variable heating rates of 50 to 999 k / m up to 1473K. The volatile contents of these samples were around 92.9 % of cellulose, 53.5 % of brown coal and 28.2 % of bituminous coal. The heating rates influence the contents. The temperature range of the pyrolysis was most large with Loy Yang coal, indicating that the contents of remaining volatile of the char produced from that coal which is believed very influential on the combustion reactivity of the char are easily controllable.
TGA data give us the
Arrhenius plot and kinetic of pyrolysis at the heating rates. Remaining volatile contents of the chars produced from Loy Yang were calculated from the kinetics at the variable pyrolysis temperature as the solid fuel at the fixed heating rate. According to the calculation, targeted char is suggested to be produced by controlling the temperature and time at the selected heating rate, although chemistry of volatiles and char are also needed to evaluate the combustion properties of char as the fuel for the pulverized coal combustion.
1. Introduction Low rank coals such as brown coal are largely deposited but their use is limited domestically or even locally. Their issues are low calorific values and instability of the coals, when dried, for spontaneously combustion. Supply of black coals will be restricted as their price will increase, reflecting on demand / supply balance. Hence it is very necessary for coal demanding countries to convert low ranking coals into handy fuel as solid fuel through low cost procedures. In the present study, pylolysis properties of Victorian brown coal were studied to find ways to convert it into power generation fuel in coal demanding countries. The three major issues are low calorific value due to huge water and oxygen contents, too much high reactivity to form long flame when burn in pulverized
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Oviedo ICCS&T 2011. Extended Abstract
coal firing, and huge possibility of spontaneous combustion during the transportation of dried coal even of it is of low ash and contamination. The pyrolysis treatment of brown coal can solve to such issues. Hence we attempt to clarify the pyrolysis reaction of the brown coal among the natural resources, nature of the volatile as well as, pyrolytic kinetics to control structural characteristics and combustion properties of the char in comparison with the dried coal.
2. Experimental section
Table 1 summarizes the samples examined in this study. Brown coal, Loy Yang and bituminous coal Newlands came from Australia, Victoria and Queens land, respectively. Cellulose was a chemical regent and supplied by SigmaAldrich Corporation. Pyrolysis of the samples was measured by a thermogravidity analyzer (TGA, TG / DTA 200SA Bruker) at heating rates of 50 to 999 K/min and a final temperature up to 1473K. The samples were heated in an infrared furnace to vary the heating rate of a wide range. The sample tube was first vacuumed for 20 min and flushed with nitrogen flow for 30 min to lower the oxygen concentration less than 10 ppm. Table 2 summarizes the TGA conditions. Activation energy and pre-exponential factor of the pyrolysis were calculated by analyzing weight reducing profile at every heating rate. The volatile contents of products char were calculated at several temperatures based on the Arrhenius equation at the heating rate of 50 K / min.
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Oviedo ICCS&T 2011. Extended Abstract
3. Results and Discussion 3.1 Pyrolysis profiles of cellulose, brown and bituminous coals The
final
weights
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the
samples were basically around the fixed carbon amounts summarized in Table 1. The weight decreased markedly in order of Newlands < Loy Yang < Cellulose, reflecting oxygen and hydrogen contents since they are lost in principally forms of CO2, H2 and CH4. The larger heating rate tends to reduce the final weight, regardless the samples. Figure 1 shows weight loss profile of the samples relative to the final loss at 1473 K by heating rate at 50 K / min. Figure 1 indicates the regions of weight loss. Cellulose started weight loss at about 700 K and completed at 800 K, whole 95 % weight loss occurring in the range of 100 K. The very rapid weight loss was noted in a narrow range.
In contrast, Loy Yang coal lost its weight in a very broad
temperature range of 300 to 1473 K. The rapid weight loss was observed at 600 – 800 K where ca 40% was lost. Higher ranking coal of Newlands coal lost in a little narrower range. The principle loss was observed in the range of 600 to 800 K where 60% was lost. The brown coal appears to have the larger variety of components in comparison with both cellulose and Newlands, being located between them in the
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Oviedo ICCS&T 2011. Extended Abstract
coalification progress. Figure 1 suggested feasibility for the control of remaining volatile contents in the product char from brown coal. The control of the remaining volatile content less than 30% is rather easy because its decrease is mild against the temperature in the range of 800 to 1423 K.
3.2 Arrhenius analysis of weight loss profile Figure activation
2
summarizes
energy
of
the
pyrolysis
weight loss at the heating rates. The activation energy increased with larger heating rate.
The
energy of Loy Yang coal ranged in 2.5×104 to 3.5×104 J / mole.
The
energy of Loy Yang coal is basically smaller than that of Newlands coal by 0.2 - 0.5×104 J / mole. The largest difference was observed at 400-600 K / min.
3.3 Product Char from Loy Yang Coal The amounts of produced gas & tar and char were ca 50.2 wt% and 49.8 wt%, respectively at 1173 K. Gaseous products consisted of CO2, CO, H2, CH4 and C2H6. Pyrolysis
of
Loy
Yang
Coal
changed its color from brown to black. The sizes of coal and char were mach the same although the large amount of volatile was lost according
to
the
pyrolytic
temperature and time. The char appears combusted in different character from those of row coal due to different amount of remaining volatile. The char looks still much more reactive than the Newland coal or its char. Submit before 31 May 2011 to
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Oviedo ICCS&T 2011. Extended Abstract
3.4 Discussion The rate of pyrolysis is estimated from the Arrhenius plot by assuming the reaction order. Reaction rate was investigated by differential of weight loss profile at various heating rates.
Activation energy and frequency factor of pyrolysis
reaction were determined by the Arrhenius plots slope and intercept, respectively. Based on the rate equation, activation energy and frequency factors, the yield of char and its volatile content are calculated at the pyrolysis temperature and time by the fixed heating fate. Figure 2 summarizes the yield of char from the Loy Yang coal and its volatile content at every temperature and time on the heating by the rate of 50 K / min. The volatile content was 44.5 at 573K to 6.29 at 1173K where the necessary times of heating were 250 to 1,500 sec, respectively by heating at 50K / min. Table 3 summarized the remaining volatile contents of chars produced at several temperatures by variable heating rates. The contents of char produced at 573K increased slightly by increasing heating rate from 50 K/min to 999 K/min while those at 773 K were almost constant around 20%. The contents at 1173 K decreased significantly from 6.6 to 3.1 % according to the heating rate.
The combustion of char is influenced by the volatile contents and composition as well as char reactivity. The exact result and influence on the combustion may be useful to design the pyrolytic process to produce char to substitute steam for the current use of brown coal. Coal volatile products at a pyrolysis are also important to count the feasibility of the pyrolytic process because the volatiles are supplied to fuel and chemical industries as the wanted feedstock.
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Oviedo ICCS&T 2011. Extended Abstract
4. Conclusions Pyrolytic characteristics of the Newlans coal of bituminous coal, the Victorian brown coal of low rank coal and the cellulose of a regent under the various heating rate from 50 to 999 K/min were investigated by using the thremogravimetry analyzer. Each sample’s characteristics of volatile matter released and pyrolytic reaction were investigated. The activation energy in the pyrolysis of Loy Yang coal was lower than that of anstandard bituminous Newlands coal. Our approach will contribute to the easy and economical use and transportation to accelerate the utilization of brown coals.
Accurate chemical analysis of pyrolytic reaction is
necessary to find ways of advanced utilization and safe transportation of brown coal. Our work suggests the possibility of expanding utilization of cheap young coal such as brown coal or lignite in the future. The results will be very useful to design advanced utilization of low rank coal.
Acknowledgement This research is partially supported by international advanced utilization of Victorian brown coal based on clean coal technology development project for fundamental international collaboration research in the energy innovation program funded by New Energy and Industrial Technology Development Organization (NEDO).
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Oviedo ICCS&T 2011. Extended Abstract
High pressure pyrolysis of different coal types – Influence of pressure on devolatilisation characteristics using TGA/MS M. Klinger, B. Meyer TU Bergakademie Freiberg, Department of Energy Process Engineering and Chemical Engineering, Fuchsmühlenweg 9 Haus 1, D-09596 Freiberg
[email protected] Abstract The influence of pressure on devolatilisation characteristics like yields and compositions of pyrolysis products was investigated using a combined thermo gravimetric/mass spectrometric device (TGA/MS). Six coals of different ranks, including German brown coals, German anthracite and two world-market sub-bituminous coals, were pyrolyzed at temperatures of up to 1100 °C with a constant heating rate of 5 K/min and system pressures of 1, 5, and 10 bar, respectively. Gas analysis involved permanent gases like H2, CO, CO2, N2, CH4, COS, H2S, H2O, O2, and Ar. The latter two were not taken into consideration, since both are no pyrolysis products. As expected, higher coal ranks lead to increased yields of char, whereas the influence of pressure is low but steady – increased pressure results in increased yields for all coal types investigated. Gas yields do not show a clear trend, since most coals display a minimum at 5 bar and a maximum at 10 bar. Also gas compositions fluctuate in terms of maxima values as well as start and peak temperatures of different gas species and coals.
1. Introduction Pyrolysis is a key process in the utilisation of solid fuels (e.g. coal, biomass) in energetic and non-energetic applications. It is not only relevant as an individual process for the production of char/coke or tar/oil and for fuel conditioning, but also as a sub-step of gasification and combustion. Pyrolysis is controlled by various parameters like process conditions (temperature, pressure, heating rate) and feedstock properties. The relevant parameters, which have an impact on gasification, are defined by the behaviour of the feedstock during pyrolytic decomposition [1]. Most important are char yield and reactivity [2] as well as the yield of gases/vapours and their composition. The former two have a significant effect on the carbon conversion and therefore on the gasification temperature (at fixed residence time), which is reflected in the efficiency of the gasifier.
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Oviedo ICCS&T 2011. Extended Abstract
The latter two, yield of gases/vapours and their composition, strongly influence the quality of the product gas and the formation of tars (in case of fixed and fluidised bed gasifiers). Since modern gasification plants are operating at elevated pressures [3][4], basic knowledge and understanding of thermal devolatilisation at those conditions are needed [5][6][7]. In the literature only few studies of combined TGA/MS investigations at high pressure can be found [5], and no research in that field has been done on German brown coals.
2. Experimental Section 2.1. Samples Six different coals, namely one brown coal from Lusatia, two brown coals from Rhineland, anthracite from Westphalia (all four Germany), a world-market subbituminous coal from Colombia and a low-rank coal from Puertollano (Spain) were investigated. Samples were milled and sieved to a fraction of < 100 µm and oven dried until a constant mass was reached. Proximate and ultimate analyses as well as the calorific values are given in Table 1. Table 1 Results of ultimate, proximate analysis and heating value determination of coal samples. Brown coal sub-bituminous coal/anthracite Lusatia Rhineland Rhineland Puertollano Westpahlia Colombia Germany Germany Germany Spain Germany Sample identification LB3 DK HKN KOL PSK IA Proximate analysis in wt.-% Moisture (r) 11.91 *) 49.59 51.12 14.97 13.55 3.95 Ash (d) 6.05 9.69 5.47 8.31 48.16 11.20 Volatile Matter (d) 52.30 48.41 50.70 38.79 21.67 6.07 Fixed Carbon (d) 41.65 41.90 43.83 52.90 30.17 82.73 Sum (d) 100.00 100.00 100.00 100.00 100.00 100.00 Ultimate analysis in wt.-% (daf) Carbon 68.27 68.88 69.04 79.10 77.43 93.31 Hydrogen 5.01 4.98 5.01 5.35 5.41 3.14 Nitrogen 0.74 0.81 0.79 1.58 1.82 1.19 Combustible Sulphur 0.30 0.02 0.05 0.71 1.32 0.99 Oxygen 25.68 25.31 25.11 13.26 14.02 1.37 Sum 100.00 100.00 100.00 100.00 100.00 100.00 O/C 0.376 0.368 0.364 0.168 0.181 0.015 H/C 0.074 0.072 0.073 0.068 0.070 0.034 Heating value in kJ/kg (d) HHV 25,059 23,975 25,267 29,125 15,714 31082 LHV
24,031
22,995
24,233
28,054
15,103
30,474
*) Coal was delivered as pulverized and dried sample.
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2.2. Pyrolysis The pyrolysis experiments were carried out in a TGA/MS-system, manufactured by the Rubotherm Company, which offers several benefits like temperatures of up to 1100 °C, pressures from atmospheric to 40 bar, and the use of different gases including corrosive ones. Due to a special magnetic coupling a closed reaction chamber and thereby contactless weighing can be afforded. For this study, argon was used as purge gas with adapted flow rates for each pressure – 1 bar: 15 ml/min, 5 bar: 63 ml/min, and 10 bar: 125 ml/min – to assure the same gas velocities. Approximately 1 g (±0.002 g) of tried and pulverized coal was sampled into an alumina crucible of 15 mm in diameter and 20 mm in height. After closing the reactor the whole system was evacuated for 20 min, where at the end a reference weight, without buoyancy effects, was taken. With a flow of 500 ml/min of argon, the system was refilled and pressurized to the selected values. The sample was heated from room temperature to 1100 °C with a heating rate of 5 K/min. The process and pyrolysis gases pass a two-stage cold trap operating at 0 °C (ice water) and -10 °C (cryostat), respectively. An automated pressure controlling valve keeps the system pressure at a defined and constant level. The mass spectrometer (MS) samples gas directly from the exhaust tube via a quartz capillary, 75 µm in diameter. For each run a blank measurement at the same conditions like the pyrolysis experiment (pressure, flow and heating rate) was recorded, to be subtracted from the TG curve.
2.3. Gas analysis The composition of the exhaust gas (pyrolysis gas + purge gas) was analysed by means of mass spectrometry. The MS is a quadrupole type (IPI GAM 200) with a secondary electron multiplier (SEM) detector to determine concentrations down to few ppm. Volume concentrations for single gas species were recalculated into ml/min for each temperature, knowing the volume flow of argon and the measured concentrations of all gas species. MS analysis was started simultaneously with the heating of the sample, enabling a temperature allocation to each gas concentration. Determination of time lag was carried out in pre-experiments using a tracer gas and taken into consideration.
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3. Results and Discussion 3.1. Char yield and mass loss rate The six coal samples show typical mass loss curves for each coal rank, whereas the three brown coals have lower char yields at 1100 °C than the sub-bituminous coals and the anthracite. The higher char yield of the hard coals results from a lower volatile content (compare Table 1). For LB3 an additional mass loss at around 800 °C can be observed, which may be assigned to mineral decompositions, like carbonates or sulphates. With increased pressure a slight increase in char yield can be observed within all coals, most prominent at LB3 and KOL. A pressure increase from 1 to 5 bar results in more significant changes then a further increase to 10 bar (see Figure 1 and Table 2). 100
100
80
90
b)
a)
80
70
70
IA 60
60
PSK
50
KOL
40
HKN DK
30
LB3
1 bar
50
5 bar
40
10 bar
30
char yield in g/g coal (d)
char yield in wt.‐% (d)
90
20
20 200
400
600
800
1000
LB3
DK
HKN
KOL
PSK
IA
temperature in °C
Figure 1 a) Mass loss curves for all coals investigated at 1 bar, normalized to a start weight at 250 °C to eliminate first mass loss effects due to water release. b) Char yields at 1100 °C for 1, 5, and 10 bar. Table 2 Char yields for different pressures and coals, normalized to a start temperature of 250 °C. pressure in bar LB3 DK HKN KOL PSK IA 1 41.8 53.1 53.2 61.9 79.8 91.0 5 51.7 – 52.4 65.3 81.8 88.9 10 54.0 55.2 54.8 66.8 82.7 93.3
Curves of mass loss rate (DTG) display two maxima (see Figure 2). The first (100– 200 °C) is related to primary decomposition like water release, whereas the second (300– 500 °C; 400–700 °C) reflects the main pyrolysis. Figure 2 b) shows a shift of main peak maxima to higher temperatures with increasing coal rank – from 350 to 400 °C for brown coals, up to 500–600 °C for sub-bituminous coal PSK and anthracite IA. With decreasing peak maxima a widening of peaks takes place, meaning that the decomposition reactions proceed over a broad temperature range, best seen at PSK and IA. In contrast, a sharp peak form arises from higher reactivity and thus from faster devolatilisation – thermal decomposition takes places in a narrow temperature range of just 250 K for all brown
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Oviedo ICCS&T 2011. Extended Abstract
coals and the Colombian sub-bituminous coal KOL. This behaviour can be explained by a general higher amount of volatiles (see Table 1) and therefore of functional groups for low rank coals, compared to sub-bituminous or even higher rank coals. The influence of pressure is shown for Lusatian brown coal LB3, where an increase leads to a higher char yield, as already depicted. Additionally, the peak of main pyrolysis is shifted to lower temperatures (from ca. 350 to 300 °C) and rises with higher pressure. Again, the effect is more prominent for the pressure increase from 1 to 5 bar than further to 10 bar (see Figure 2 a). This trend is not seen significantly in all coals! Similar results were found by Yang et al. [6] and Yun & Lee [5] for one of their studied coals. In contrast, Tomeczek & Gil [8] found decreasing char yields with increasing pressure for rapid heating pyrolysis (20–100 K/s) of sub-bituminous coal. They listed other references showing similar trends to their results, but noted that a comparison of influence of pressure is difficult due to different types of coal and definitions of volatile matter (water and ash free or analytical state) – significantly different heating rates might also be mentioned as a reason for contrary trends. 0.50
a)
char yield in wt.‐%
80 70
1 bar
0.45
5 bar
0.40
10 bar
60
0.35 0.30
50
0.25
40
0.20
30
0.15
20
0.10
10
0.05
0.30 LB3
b)
0.25
DK HKN
0.20
KOL PSK
0.15
IA 0.10 0.05 0.00
0.00
0 0
200
400
600
800
1000
mass loss rate in wt.‐%/K
90
mass loss rate in wt.‐%/K
100
0
100
temperature in °C
200
300
400
500
600
700
temperatur in °C
Figure 2 a) Influence of pressure on mass loss curve and first derivate of brown coal LB3. b) First derivate of mass loss (DTG curves) of all coals investigated at 1 bar.
3.2. Gas yield and composition The start temperature of gas release indicates the beginning of the formation (≥50 ppm) of the assigned gas species. Total gas yield is calculated by summation of yields over temperature. In Table 3 values for start temperature and total yield for H2, CO, CO2, N2, and CH4 are given in respect of coal type and pyrolysis pressure.
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Oviedo ICCS&T 2011. Extended Abstract
Table 3 Start temperature for the release of different gas species and their total yield, depending on type of coal and system pressure. T start in °C p in bar LB3 DK HKN KOL PSK IA 1 561 475 499 573 624 615 H2 5 656 – 648 652 808 618 10 0 631 594 638 0 689 1 329 351 360 417 525 641 5 379 – 374 407 440 673 CO 10 290 356 281 163 387 191 1 190 434 435 397 328 432 5 515 – 561 640 703 769 CO2 10 442 164 426 191 333 634 1 374 510 497 424 461 477 5 399 – 383 445 451 0 N2 10 310 463 325 159 139 148 1 379 380 398 387 448 510 5 346 – 361 390 429 461 CH4 10 325 331 322 347 387 488 total yield in ml/gcoal LB3 DK HKN KOL PSK IA 1 349.1 1090.3 1267.9 182.1 112.8 349.8 H2 5 180.8 – 131.2 291.6 44.6 1335.7 10 0 39.9 171.9 134.3 0 73.8 1 971.6 831.8 836.2 399.1 202.3 226.1 CO 5 358.8 – 418.5 362.2 263.8 94.9 10 326.3 865 607.1 1038.5 647.4 284.6 1 2422.5 183 195.4 585.7 447.1 375.8 CO2 5 275.5 – 219.1 123.1 76.5 98 10 339.9 2223.8 397.5 1711.3 1190.7 113.7 1 126.2 37.8 36 69.8 45.7 35.3 N2 5 199.8 – 268.9 106.9 223.1 0 10 117 189.1 136.9 572.9 582.3 311.2 1 260.1 262.9 316 273.2 127.5 103.6 CH4
5
206.6
–
197.8
301
107.6
136.1
10 191.8 218 248.2 287.3 117.5 85.5 As can be seen from Figure 3 gas yields for 1 bar do not correlate with the amount of volatiles for each coal (compare Table 1), but show rather similar values of around 15–20 wt.-%. LB3 describes an exception, since it produces significant higher yields (43 wt.-%), which arise from large yields in CO2 (see
Table 3 and Figure 4 a), which in turn can be attributed to carbonate decomposition. The influence of pressure on gas yields does not follow a clear trend. For most coals 5 bar experiments display a yield minimum (except anthracite IA), whereas 10 bar leads to significant higher values for DK, KOL, and PSK (see Figure 3 b). So no distinction between brown coals and hard coals can be drawn in terms of gas yields. That LB3 shows its highest amounts at atmospheric pressure can also be explained by carbonate decomposition, since higher pressure depresses the CO2 formation for this reaction.
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Oviedo ICCS&T 2011. Extended Abstract
a)
yield in wt.‐%
45
50
b)
1 bar
45
40
LB3
5 bar
35
DK
10 bar
30
HKN
30
25
KOL
25
20
PSK
20
15
IA
15
40 35
10
10
5
5
gas yield in g/g coal
50
0
0 0
200
400
600
800
1000
LB3
DK
HKN
KOL
PSK
IA
temperature in °C
Figure 3 a) Gas yield curves for all coals investigated at 1 bar. b) Char yields at 1100 °C for 1, 5, and 10 bar. The gas yield minimum at 5 bar is also reflected in CO and CO2 yields (see
Table 3) for almost all coals investigated (except PSK, where 1 bar displays the lowest CO value). Nevertheless, CO2 contributes stronger to the total gas yield due to its higher density (ρCO2=1.98 g/l, ρCO=1.25 g/l). This can also be seen in case of sub-bituminous coal PSK, where the minimum for CO yield occurs at 1 bar and for CO2 at 5 bar – same as for the overall gas yield. Note that total gas yields are given in g/gcoal, enabling the direct comparison with char yields, whereas yields of gas species are given in ml/gcoal. Start temperatures of CO formation at 1 bar lie between 330 and 360 °C for brown coals, rising up to 410 (KOL), 440 (PSK), and 670 °C (IA) with increasing coal rank. CO formation temperatures for 5 bar show almost the same trend and values as for 1 bar, whereas 10 bar values are generally lower, but fluctuate strongly between 160 and 390 °C over all coals. Start temperatures of CO2 display a maximum at 5 bar (following the same trend of coal ranks as CO), but display no trend for 1 and 10 bar experiments. The influence of coal rank on CO and CO2 start temperature indicates different formation mechanism, depending on coal structure, e.g. volatile matter. In general, for 5 bar experiments CO2 formation temperatures are 100–250 K higher than those of CO. In contrast, Chen et al. [9] found for two coals (low rank and sub-bituminous) CO2 to be produced between 300 and 700 °C and CO at 500 to 1100 °C, showing that CO formation starts later than CO2, but over a wider range (at 1 bar). Compared to our data at 1 bar one can find lower start temperatures for CO2 than for CO, at least for the three hard coals (KOL, PSK, and IA).
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Sum PyGas H2
yield in (ml/min)/g Coal
2.5
1600 1 bar
a)
b)
5 bar
CO
1400 1200
10 bar
CO2
2.0
1000
N2
1.5 1.0
CH4
800
COS
600
H2S
400
H2O
0.5
H2 total yield in ml/g Coal
3.0
200 0
0.0 0
200
400 600 800 temperatur in °C
1000
LB3
DK
HKN
KOL
PSK
IA
Figure 4 a) Yield curves for singles gas species, exemplarily shown for LB3, 1 bar. b) Total hydrogen yields in respect of coal type and pressure. The hydrogen yield shows a strong influence of pressure for the brown coals, whereat 1 bar leads to higher amounts than 5 and 10 bar (e.g. HKN, see Table 3 and Figure 4 b), which is in good agreement with Tao et al. [7]. For sub-bituminous coals (KOL and PSK) H2 yields are not much affected by pressure. In contrast, anthracite shows a contrary behaviour than brown coals and produces most hydrogen at 10 bar (which depicts the highest value at all), followed by 1 and 10 bar. This shows that crucial different mechanisms for hydrogen formation take place. Maximum temperatures for H2 are almost in all cases at 1100 °C, only at 10 bar LB3, DK and IA have lower temperatures of 820, 720, and 780 °C, respectively. Porada [10] found the formation maximum of H2 to be at ca. 720 °C for a middle rank coal investigated at 1 bar, which is in contradiction to our 1 bar data. Hydrogen formation for 1 bar pyrolysis starts at 480–560 °C for brown coals with slightly higher values (570–620 °C) for hard coals (see
Table 3). With rising pressure start temperatures increase (650–810 °), but show its maxima at 5 bar for all coals (except the anthracite, where the highest H2 start temperature occur at 10 bar). Methane shows also an increase in start temperatures with higher coal ranks, but a contrary behaviour to hydrogen in respect of pressure influence. Lowest formation temperatures are found at 10 bar (320–490 °C) and highest at 1 bar (ca. 50 K above), whereas 5 bar values lie in between. Hence, a clear and steady influence of pressure on the formation temperature of CH4 can be detected. Methane formation displays a yield maxima, whose temperature decreases from approximately 800 to 700 to 600 °C with increasing pressure (1, 5, 10 bar, respectively) for all coals investigated (more or less independent on coal rank). Arenillas et al. [11] noticed the maximum at 550 °C, Porada 10[10] at 460 °C, and Jong et al. [12] at 605 °C, which are all at lower temperatures than our findings for 1 bar. CH4 yields at 1 bar are higher for brown coals (260–320 ml/gcoal) than for sub-bituminous coals and anthracite (100–130 ml/gcoal), whereas KOL shows with ca. 270 ml/gcoal also relative high amount (see
Table 3). This can be explained by the low state of coalification, reflected in the high volatile content of brown coals and KOL. Pressure increase leads to decreasing CH4 yields for most coals, except KOL and IA, whereas the effect of pressure is more prominent for brown coals. Sub-bituminous coal KOL and anthracite IA display a maximum in methane yield at 5 bar. Tao et al. [7] found slightly increasing CH4 amounts with higher pressure and explained that by methane formation reactions from CO2 + 4 H2. Since methane values in our study drop with increasing pressure, this homogeneous reactions cannot be taken into consideration. Furthermore, secondary reactions between
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Oviedo ICCS&T 2011. Extended Abstract
char and gas seem to be the reason, since increased pressure hinders gases to exhaust from the char particle. Other gases like N2, H2S, COS, or H2O show no clear trends in formation temperature and total yields or occur in too small concentrations to be evaluated in a good quality.
4. Conclusions The influence of pressure on devolatilisation characteristics, e.g. char and gas yields as well as gas composition was investigated at coals of different rank. Following conclusions can be made: The char yield increase with coal rank and with increasing pressure. Also the peak of mass loss rate becomes higher and is shifted to lower temperatures with increased pressure. Gas yields cannot be correlated to coal rank. Additionally, the impact of pressure fluctuates for the different coals in terms of gas yields. It can be stated that most coals display a yield minimum for 5 bar, but also highest values for 10 bar are found. So no clear trend in pressure effect on total gas yields can be drawn. Yields of CO and CO2 reflect the 5 bar minimum of the total gas yield with changing pressure dependencies for different coals. H2 yield is strongly affected by pressure in case of brown coals, where maximum values are found for 1 bar. Anthracite, in contrast, has its maximum and minimum at 10 and 5 bar, respectively. For methane an increase in pressure causes decreasing start and maximum temperatures. With higher coal rank the start temperature rises for all pressures investigated. It can be summarized that pressure affects char and gas yields as well as gas composition, whereas for the latter no consistent trends for coal ranks or gas species can be detected. Acknowledgement. The investigations were financially supported by the German Federal Ministry of Economics and Technology, EnBW, Eon, RWE, Siemens, and Vattenfall within the research project “HotVeGas”.
References [1]
Cai HY, Güell AJ, Chatzakis IN, Lim JY, Dugwell DR, Kandiyoti R. Combustion
reactivity and morphological change in coal chars: effect of pyrolysis temperature, heating rate and pressure. Fuel 1996; 75:15–24 [2]
Roberts DG, Harris DJ, Wall TF. On effects of high pressure and heating rate during coal
pyrolysis on char gasification reactivity. Energ Fuel 2003; 17:887–95
Submit before 31 May 2011 to
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Oviedo ICCS&T 2011. Extended Abstract [3]
Roberts DG, Harris, DJ. Char gasification with O2, CO2 and H2O: effects of pressure on
intrinsic reaction kinetics. Energ Fuel 2000; 14:483–9 [4]
Wall TF, Liu G, Wu H, Roberts DG, Benfell KE, Gupta S et al. The effect of pressure on
coal reactions during pulverized coal combustion and gasification. Prog Energ Combust 2002; 28:405–33 [5]
Yun L, Lee GB. Effects of pressure in coal pyrolysis by high pressure TGA. Korean J
Chem Eng 1999; 16:798–803 [6]
Yang H, Chen H, Ju F, Yan R, Zhang S. Influence of pressure on coal pyrolysis and char
gasification. Energ Fuel 2007; 21:3165–70 [7]
Tao W, Zou YR, Carr A, Liu J, Peng P.Study of influence on enhanced gaseous
hydrocarbon yield under high pressure–high temperature coal pyrolysis. Fuel 2010; 89:3590–7 [8]
Tomeczek J, Gil S. Volatiles release and porosity evolution during high pressure coal
pyrolysis. Fuel 2003; 82:285–92 [9]
Chen L, Zeng C, Guo X, Mao Y, Zhang Y, Zhan X et al. Gas evolution kinetics oft wo
coal samples during rapid pyrolysis. Fuel Process Technol 2010; 91:848–52 [10]
Porada S. The reaction of formation of selected gas products during coal pyrolysis. Fuel
2004; 83:1191–6 [11]
Arenillas A, Ruberiea F, Pevida C, Pis JJ. A comparison of different methods for
predicting coal devolatilisation kinetics. J Anal Appl Pyrol 2001; 58–59:685–701 [12]
Jong W, Nola G, Venneker BCH, Spliethof H, Wójtowicz MA. TG-FTIR pyrolysis of
coal and secondary biomass fuels: determination of pyrolysis kinetic parameters for main species and NOx precursors. Fuel 2007; 86:2367–76
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Oviedo ICCS&T 2011. Extended Abstract
Brown coal and rape cake co-pyrolysis products in the range 5 to 40 per cent Authors: Josef Vales, Jaroslav Kusy, Lukas Andel, Marcela Safarova Brown coal research institute j.s.c. Budovatelu 2830, 434 37 Most, Czech Republic Corresponding author:
[email protected], +420476208627
Abstract This article summarises experimental results of thermal reprocessing – co-pyrolysis of the mix of brown coal and defined biomass amount (rape cake) and its influence on the yield and quality of pyrolysis products – solid carbon semi-coke, liquid and gaseous products.
1. Introduction Decreasing resources of noble energy raw materials [1] make for searching for alternative energy source for the future. Domestic energy sources and their qualitative parameters are reassessed in relation to available and technologically effectively advanced procedures of their possible processing and using while respecting development trends of legislatively allowable limit loads of environment. Coal deposits become more important because of the size of geologic and mineable resources (240 years) [1], their homogenous distribution over all continents and because of their possible conversion to gas and liquid fuels. As a matter of fact, every carbonaceous raw material (coal, carbonaceous waste, and biomass) can be transformed to gaseous medium. Innovative ways of processing are oriented to thermal re-processing of coal mass together with hydrogen and carbon rich substances, so called co-processes [2, 3].
2. Experimental section Pyrolysis is a technological process of a thermal re-processing of suitable charge without air access. Solid pyrolysis residue (pyrolysate) is a product of the pyrolyse. Liquid phase created from a mixture of pyrogenous water, oil, or tar (organic liquid), and a mixture of reaction gases is the product of the pyrolysis. The technology was used for producing oil and paraffin of shale in France in 1832 [5]. The process was significantly developed technologically at liquid fuel production by tar hydrogenation of coal in Germany before 1940 [4]. Submit before May 31st to
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Oviedo ICCS&T 2011. Extended Abstract
Aim of experiments Experimental works were focused to the determination of yield changes during the thermal reprocessing (co-pyrolysis) of the mixtures of brown coal with the different additions of rape cake under the same conditions.
Experimental laboratory testing unit and raw materials Laboratory pyrolysis unit (figure 1) with a rectangle shaped retort, indirect electric heating, and electronically setting of temperature program run was used for testing [6]. Gaseous products are discharged through an outlet and cooler to collecting box. Uncondensed gas is combusted in a burner. Solid pyrolysis carbonaceous residue remains on the retort bottom. Samples of granularity prepared charge raw material, i.e. pressed rape cake, brown coal, and mass defined charge mixtures – coal: cake with 1000 g total mass were input individually to the retort and tested. Tests ran under the same process conditions: temperature increase gradient, final temperature and the delay time. Qualitative parameters were assigned to charge raw materials and products [7, 8, 9, 10, 11, 12, 13, 14].
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Oviedo ICCS&T 2011. Extended Abstract
For the co-pyrolysis experiments, brown coal mined in the coal mine named Důl Československé armády (CSA) and rape cake from the pressing of oil seed rape from the Preol, Lovosice. Basic qualitative parameters of the input raw materials are summarized in the table 1. Determination of the yield of the low temperature distillation according to the standard CSN ISO 647 [7] was also made. Results are in the table 2. Table 1: Qualitative parameters of the input raw material Parameter Wa Ad Std Cd Hd Nd Od Vd Qs d Qi d
[wt. %] [wt. %] [wt. %] [wt. %] [wt. %] [wt. %] [wt. %] [wt. %] [MJ/kg] [MJ/kg]
brown coal, ČSA 5,98 4,06 0,75 72,30 5,65 0,88 16,36 54,83 31,34 30,11
rape cake, Preol a.s. 6,31 6,36 0,68 49,12 7,20 5,26 31,38 78,11 21,99 20,42
Table 2: Yield of the tar, water, gas and semicoke from the low temperature distillation sample Brown coal ČSA Rape cake Preol a.s
TsKd [wt. %] 25,39 34,08
sK
d
[wt. %] 57,31 29,82
WsKd [wt. %] 8,43 20,80
GsKd [wt. %] 8,87 15,30
Pure brown coal, pure rape cake and their blends with the ratio 5, 10, 15, 20, 30 a 40 wt. % were tested. All samples were pyrolised under the same conditions – initial temperature 25 °C, heating rate of 2,47 °C.min-1, terminal temperature 650 °C with the 60 min of delay on this temperature.
Yields of pyrolysis tests Yields of laboratory pyrolysis tests obtained from the experiments with laboratory testing unit are shown in the table 3. Gaseous products were not measured, but they are calculated as a balance between 100 % and the sum of another products.
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Oviedo ICCS&T 2011. Extended Abstract
Table 3: Yields of pyrolysis tests Pyrolysis product
Batch, % mixture of coal : rape cake related to the yields of pyrolysis products
[wt. %] C-
mixture
mixture
mixture
mixture
mixture
mixture
R-
brown
C:R
C:R
C:R
C:R
C:R
C:R
rape
coal
95 : 5
90 : 10
85 : 15
80 : 20
70 : 30
60 : 40
cake
solid carbon
49,79
51,80
50,58
49,50
45,43
43,49
40,75
27,16
pyrogenetic water
12,64
13,70
14,55
15,06
16,49
17,06
18,60
22,33
tar
18,44
17,85
17,75
18,57
17,43
18,46
19,17
31,63
gas + loses
19,13
16,65
17,12
16,87
20,65
20,99
21,48
18,88
100
100
100
100
100
100
100
Sum of the wt. %
Solid carbonaceous products were analysed
according
to
Czech
100
standards
[8,9,10,11,12,13,14]. Elementar analysis was made on the analyzer Elementar Vario EL 3. Results of these analyses are showed in the table 4. Table 4: Qualitative parameters of the solid carbonaceous products of co-pyrolysis Parameter
Qualitative parameters of the solid carbonaceous products of co-pyrolysis
of the solid carbonaceous
C-
mixture
mixture
mixture
mixture
mixture
mixture
R-
product
brown
C:R
C:R
C:R
C:R
C:R
C:R
rape
[wt. %]
coal
95 : 5
90 : 10
85 : 15
80 : 20
70 : 30
60 : 40
cake
0,08
<0,01
<0,01
<0,01
0,62
0,52
0,61
0,94
Wa [wt. %] d
[wt. %]
8,09
8,02
7,94
8,25
9,01
10,66
11,43
16,06
Std
[wt. %]
0,71
0,31
0,41
0,25
1,31
1,71
0,86
0,49
C
d
[wt. %]
85,02
84,55
84,52
84,04
83,82
81,94
80,91
63,68
H
d
[wt. %]
2,09
2,24
2,12
2,31
2,20
2,22
2,25
3,90
N
d
[wt. %]
1,18
1,35
1,50
1,62
1,67
2,48
2,54
7,11
d
[wt. %]
2,92
3,53
3,51
3,53
1,97
1,60
2,13
8,76
[wt. %]
4,30
5,01
4,70
4,71
5,20
5,05
6,49
27,17
[MJ/kg]
32,22
32,39
32,27
31,98
30,89
30,25
29,70
27,01
Qi d [MJ/kg]
31,77
31,90
31,81
31,47
30,41
29,76
29,21
26,16
A
O
Vd Qs
d
Analysis of gaseous products Table 5 and figure 2 summarizes results of the gaseous products analysis from all laboratory tests. Gas were sampled between the temperatures of 510 and 520 °C to the antidiffusive bag [15]. Samples were analysed on the gas chromatograph GC 82 TT Submit before May 31st to
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Oviedo ICCS&T 2011. Extended Abstract
Labio Praha with the dual thermoconductivity detector. Results are summarized on the table 4 and on the figure 4. Hydrocarbons are given as the sum of lower hydrocarbons. Table 5: Qualitative parameters of the gaseous products Gaseous products of the pyrolysis
Gaseous products C-
mixture
mixture
mixture
mixture
mixture
mixture
R-
brown
C:R
C:R
C:R
C:R
C:R
C:R
rape
coal
95 : 5
90 : 10
85 : 15
80 : 20
70 : 30
60 : 40
cake
hydrogen [vol. %]
25,44
25,42
24,01
23,84
22,72
17,27
16,95
7,49
oxygen [vol. %]
0,19
0,06
0,07
0,05
<0,01
0,28
<0,01
0,09
nitrogen [vol. %]
1,37
0,69
0,99
0,94
0,86
2,84
0,66
0,95
methane [vol. %]
35,50
36,96
35,73
37,00
36,41
34,43
33,93
31,24
carbon oxide [vol. %]
8,83
10,25
9,92
10,30
9,73
10,60
9,83
8,29
carbondioxide [vol. %]
9,83
10,49
10,69
10,62
20,07
24,74
25,43
26,47
hydrocarbons [vol. %]
18,84
16,13
18,59
17,25
10,21
9,84
13,19
25,47
Gaseous products analysis 40 Gaseous products (wt. %)
35 30 25 20 15 10 5 0 0
10
20
30
40
50
60
70
80
90
100
Rape cake ratio in the mixture (wt. %)
hydrogen
oxygen
nitrogen
carbon oxide
carbondioxide
hydrocarbons
methane
Figure 2: Qualitative parameters of the gaseous products
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Oviedo ICCS&T 2011. Extended Abstract
3. Results and discussion Laboratory pyrolysis testing of the coal and rape cake ant their mixtures should be summarized: •
By the thermal re-processing of thy rape cake with the terminal temperature 620 °C, 31,6 wt. % of oil should be obtained. Brown coal pyrolysis under the same conditions provides 18,4 wt. % of the tar.
•
Solid carbonaceous product (semi-coke) reaches 49,8 wt. % from the coal and 27,2 wt. % from the rape cake. By the addition of 5 wt. % and 10 wt. % of the rape cake to the coal, the yield of the semi-coke will be higher by 1,5 wt. %. Higher levels of the rape cake addition should lower the yield by the 9 wt. % compared to coal.
•
Higher levels of the rape cake addition leads to the highest yield of the pyrogenetic water and modest yield of the oil – tar mixture..
•
Yield of the gaseous products are on the same level about 19 wt. % and rises very slowly with the rape cake addition.
•
Qualitative changes of the solid carbonaceous products from the pyrolysed batches with the higher ratio of the rape cake shows the lowering of sulphur, carbon and heating value in dry matter. On the other hand, concentration of ash, hydrogen and nitrogen is rising. No changes occures with the oxygen.
•
Change of the weight ratio between the rape cake and coal leads to the changes of gaseous products composition. With higher rates of the rape cake, methane and hydrogen concentration is lowering, as well as the carbon oxide. Carbon oxide is rising slowly until 15 to 30 wt. %, where the rising is very fast. Other components are approximately on the same levels.
4. Conclusions Except the pyrolysis conditions, the quality and yields of pyrolysis products can be affected by the composition of pyrolysed batch. Under the same conditions of described experiments, addition of the rape cake to the coal leads to the lowering of the solid carbonaceous product (semi-coke) yield. Qualitative changes of these semi-cokes made from the rape cake – coal mixtures are showing lower concentrations of sulphur, carbon and lower heating value, as well as rising of nitrogen, hydrogen and ash. Gaseous
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Oviedo ICCS&T 2011. Extended Abstract
pyrolysis products are affected by the addition of the rape cake to the mixture with the change of their composition. The most important change is between the ratios of 15 wt. % and 30 wt. %, where the concentration of carbon dioxide is rising very fast, which demonstrates its reaction with ammonia to the ammonium carbonate.
Acknowledgement: The work is a partial output of the solution of the project VaV SP/2f2/98/07 and it was made under financial support of Ministry of the Environment of the Czech Republic.
Abbreviations Wa
Content of analytical water
Wtr
Content of water in original matter
Ad
Ash content in dry matter
Std
Total content of sulphur in dry matter
Cd
Content of carbon in dry matter
d
H
Content of hydrogen in dry matter
Nd
Content of nitrogen in dry matter
Vd
Content of volatile matter in dry matter
Qsd
Combustion heat in dry matter
Qsr
Combustion heat in original matter
Qid
Heating value in dry matter
Qi
r
Heating value in original matter
TSKd
tar content (converted to dry condition)
WSKd
pyrogenetic water content (converted to dry condition)
GSKd
gas content (converted to dry condition)
d
SK
daf
semi-coke content (converted to dry condition) in superscript – results converted to the dry and ash free condition
References: [1]
ROUBÍČEK, V., BUCHTELE, J.: Coal – sources, processes, utilisation, Montanex a.s. Ostrava 2002
[2]
ŠAFÁŘOVÁ, M., KUSÝ, J., ANDĚL,L.: Distribuce stopových prvků v produktech pyrolýzy hnědého uhlí, All for Power 4/2009, ISSN 1802 – 8535, http://www.allforpower.cz/rubrika/uhlí
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7
Oviedo ICCS&T 2011. Extended Abstract [3]
STAF, M.: Výzkum termické konverze odpadní biomasy na plynná a kapalná paliva, http://biom.cz/odborné clanky
[4]
BLAŽEK, J., RÁBL, V.: Základy zpracování ropy, 2.vydání, VŠCHT Praha 2006, ISBN 807080-619-2,
[5]
JÍLEK. J.: Nízkotepelná karbonizace a tepelné zpracování hnědého uhlí, SNTL Praha, 1954, ISBN 301-05-118
[6]
VALEŠ J., KUSÝ J., ANDĚL L., ŠAFÁŘOVÁ M.: Možnosti využití pokrutiny z výroby rostlinného oleje pro energetické účely, Odpadové fórum 2010, Kouty nad Desnou, 04/2010,přednáška 066, 7 stran, ISBN 978-80-85990-12-6
[7]
ČSN ISO 647: Tuhá paliva. Metoda stanovení výtěžku dehtu, vody, plynu a koksového zbytku při nízkoteplotní destilaci,
[8]
ČSN ISO 351: Tuhá paliva. Stanovení obsahu veškeré síry – Vysokoteplotní spalovací metoda,
[9]
ČSN 44 1377: Tuhá paliva. Stanovení obsahu vody,
[10]
ČSN 44 1378: Tuhá paliva. Stanovení popela,
[11]
ČSN ISO 1928: Tuhá paliva. Stanovení spalného tepla kalorimetrickou metodou v tlakové nádobě a výpočet výhřevnosti,
[12]
ČSN ISO 609: Tuhá paliva. Stanovení uhlíku a vodíku – Vysokoteplotní spalovací metoda,
[13]
ČSN 44 1310: Tuhá paliva. Označování analytických ukazatelů a vzorce přepočtů výsledků na různé stavy paliva,
[14]
ČSN 44 1351: Tuhá paliva. Vážková metoda stanovení prchavé hořlaviny,
[15]
Kusý, J., Brejcha, J., Hautke, P., Antidifuzní vak pro odběr a uchování plynů a vzdušnin, autorské osvědčení č. 245959, Úřad průmyslového vlastnictví, Praha, 1988
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8
Oviedo ICCS&T 2011. Extended Abstract
Integrated coal pyrolysis with methane aromatization over Mo/HZSM-5 catalyst for improving tar yield Xun Zhou, Haoquan Hu*, Lijun Jin State Key Laboratory of Fine Chemicals, Institute of Coal Chemical Engineering, School of Chemical Engineering, Dalian University of Technology, Dalian 116024, China, *Email:
[email protected] Abstract In this paper, a new process to integrate coal pyrolysis with methane aromatization (MAP) over Mo/HZSM-5 catalyst was put forward for improving tar yield, and a Chinese Shenmu coal was used to confirm the validity of the process. The effect of pyrolysis temperature (600-800 ºC), CH4 flow rate (15-200 mL/min), Mo loading (2-8wt.%) on tar, water and char yields was investigated in an atmospheric fixed-bed reactor containing upper catalyst layer and lower coal layer. The tar, water and char yield is 21.5, 3.3 and 65.6 wt.%, respectively, in the MAP process at the optimum conditions of 700 °C pyrolysis temperature, 25 mL/min CH4 flow rate, 4 wt.% Mo loading and 30 min holding time. Compared with those in the pyrolysis of coal under N2 and H2, the tar yield increases by 7 and 6 wt.%, respectively. And the synergistic effect between coal pyrolysis and methane aromatization was also investigated. The tar and hydrocarbons yield of coal pyrolysis under CH4 (without catalyst) and methane aromatization (without coal) is 15.3 and 2.9wt.%, respectively, at the optimum conditions, and the sum of which is less than that in the MAP process. Based on the results above, it has been shown that the tar yield can be improved in the integrated process of coal pyrolysis and methane aromatization.
1. Introduction Coal is the relatively abundant and inexpensive energy source compared with petroleum. Coal pyrolysis to produce hydrocarbons as an alternative fuel source of crude oil has been received considerable attention recently. However, the tar yield in pyrolysis in inert gases is limited owing to the low ratio of hydrogen to carbon in coal [1]. Hydropyrolysis is an effective method to improve tar yield because of chemical 1
Oviedo ICCS&T 2011. Extended Abstract
reactions between H2 and the free radicals cracked from coal, but the high cost of pure hydrogen hinders its practical application. Thus, it is necessary to look for a cheaper hydrogen-rich gas to replace the expensive hydrogen. Methane is the main composition of natural gas and coal bed gas, and also accounts for about 25 vol.% of coke oven gas[2]. It could be the hydrogen substitute in coal conversion into liquid fuel because of its high hydrogen to carbon ratio and low cost. In coal pyrolysis, methane has been shown to act as inert gas at low temperature[3]. A study on the uncatalyzed conversion of subbituminous coal with methane in an entrained tubular reactor at short residence times gave a significant increase in coal conversion as well as in total liquids and hydrocarbon gases was observed [4]. For methane, the scission of a C-H bond is the first step that a methane transformation reaction encounters. Due to its perfect symmetry, methane has the most stable C-H bonds. Thus, the activation of C-H bond of methane needs relatively higher temperature or the presence of catalyst [5,6]. A methane aromatization reaction provides a practical method for methane activation and utilization. The impregnation method to prepare the Mo/HZSM-5 catalyst has been widely studied [7]. In this paper, a new process to integrate coal pyrolysis with methane aromatization (MAP) over Mo/HZSM-5 catalyst was put forward for improving tar yield. Although methane aromatization over Mo/HZSM-5 catalyst has been widely studied, the literatures on MAP have never been reported. Here we studied the effect of pyrolysis temperature, CH4 flow rate and Mo loading on tar, water and char yields in the MAP, and the synergistic effect between coal pyrolysis and methane aromatization.
2. Experimental section 2.1. Coal samples Chinese Shenmu coal was ground to -100 mesh and used in this study. The proximate and ultimate analysis of SM coal sample is shown in Table 1. 2.2. Apparatus and procedures The coal pyrolysis at different atmospheres was carried out in a fixed-bed reactor, as shown in Figure 1. The fixed-bed reactor, constructed of stainless steel with a length of 2
Oviedo ICCS&T 2011. Extended Abstract
Table 1 Proximate and ultimate analyses of SM coal sample Proximate analysis/wt% Ultimate analysis/wt%, daf Mad Ad Vdaf C H N S O* 4.83 5.61 37.22 80.20 5.62 1.11 1.83 11.24 *By difference.
12
11 3
2
P
5
GC
P
2 3
B
A
5
8 Ⅰ 9
CH4
N2
Ⅱ
Ⅳ 1
1
4
4
6
7
Ⅲ
10
Ⅴ C
1. Gas cylinder, 2. Valve, 3. Mass flow controller, 4. Mass flow display, 5. Pressure gauge, 6. Temperature controller, 7. Temperature indicator, 8. Furnace, 9. Reactor, 10. Cool trap, 11. Decompress valve, 12. Gas flowmeter;
. Gas Entrance,
. Catalyst Bed,
. Coal Bed,
. quartz wool,
. Gas Exit;
A. Methane, B. N2, C. Products
Figure 1. Schematic diagram of apparatus for pyrolysis. 150 mm and an internal diameter of 18 mm, contains an upper catalyst layer and a lower coal layer. About 1 g of catalyst and 5 g of coal were loaded in the two layers, respectively, which was separated by quartz wool. The height of the catalytic bed is about 4 mm. A thermowell, made of stainless steel of 2.5 mm outer diameter and equipped with a movable thermocouple, was placed in the center of the catalyst bed to monitor the pyrolysis temperature. The reactor was heated to a desired temperature within 8 min and held at the temperature for 30 min. The experiments of coal pyrolysis under N2 and H2 were performed at the same pyrolysis conditions as those under CH4 except without catalyst.
3
Oviedo ICCS&T 2011. Extended Abstract
The liquid products, tar plus water, were collected in a cold trap at -17 ℃, and then the water in the liquid products was separated according to ASTM D95-05e1 (2005) using toluene as solvent. In this way, tar and water yield could be calculated, respectively.
2.3. Catalyst preparation Mo/HZSM-5 catalysts were prepared by incipient wetness impregnation method. HZSM-5 was impregnated with aqueous solution of ammonium molybdate at room temperature for 16 h. After being dried at 120 ℃ for 4 h and calcined at 550 ℃ in air for 6 h, the catalyst powder was pressed, crushed and sieved to 20-40 mesh.
3. Results and Discussion 3.1. Effect of pyrolysis temperature Figure 2 shows the effect of pyrolysis temperature on tar, water and char under different atmospheres. It can be seen that the tar yield in the MAP process increases with increase of pyrolysis temperature and reaches the maximum at 700 °C, and then decreases with the further increase of pyrolysis temperature. And the tar yield under N2 and H2 decreases with the increase of pyrolysis temperature. The tar yield in the MAP process is 21.5 wt.% at 700 °C, which increases by 7 % and 6 % as that in pyrolysis under N2 and H2, respectively. The water yield decreases with the increase of pyrolysis temperature, and there is no big difference under different atmospheres, but more water was produced under H2. The char yield also decreases with the increase of pyrolysis temperature, and has the highest under N2 and the lowest under H2 at different pyrolysis temperature, but the char yield in the MAP process is lower than that under H2 at 700 to 800 °C. During coal pyrolysis, the bridge bonds ruptured in coal structure result in the formation of large amount of free radicals [8]. When combining with other small free radicals such as H radical, these free radicals are stabilized and volatiles including tar and gas products are formed, otherwise the free radicals themselves will recombine to form char or tar [9]. There are enough H free radicals in pyrolysis under H2, so the tar yield is 4
Oviedo ICCS&T 2011. Extended Abstract
higher than that in pyrolysis under N2. A rise in reaction temperature thermodynamically favors methane aromatization. However, catalyst deactivation is also more severe [10].The are enough CHx groups in the MAP process at 700 °C, so the tar yield in the MAP process reaches maximum at 700 °C.
10
N2
20
15
10
600
650
700
750
MAP H2
8
N2
6
4
2
0
800
75
Char yield(wt.%, daf)
MAP H2
Water yield(wt.%, daf)
Tar yield(wt.%, daf)
25
MAP H2 N2
70
65
60 600
650
Temperature(oC)
700
750
600
800
650
700
750
800
Temperature(oC)
o
Temperature( C)
Figure 2. Effect of pyrolysis temperature on tar, water and char yields under different atmospheres (conditions: flow rate 25 mL/min; 4 wt.% Mo/HZSM-5 catalyst; holding time 30 min) 3.2. Effect of gas flow rate Gas flow rate is also a significant factor in coal pyrolysis process. The tar, water and char yields under different atmospheres at different flow rates are shown in Figure 3. Compared with the MAP process, the tar yield under N2 and H2 change a little at different flow rates. The tar yield increases with an increasing CH4 flow rate from 15 to 25 mL/min and reaches the maximum 21.5 % at 25 mL/min, and then decreases obviously with the further increasing flow rate from 25 to 200 mL/min in the MAP process. A high gas flow rate shortens the contact time of CH4 with the Mo/HZSM-5 catalyst, which makes CH4 conversion decrease. So there are no much CHx groups in the MAP process, which results in the decrease of the tar yield. However, too low CH4
15
0
50
100
150
Rate of flow (mL/min)
200
MAP H2 N2 5
0
0
50
100
150
Rate of flow(mL/min)
200
Char yield (wt.%, daf)
H2
20
10
70
10
MAP N2
Water yield(wt.%, daf)
Tar yield(wt.%,daf)
25
MAP H2 N2 65
60
0
50
100
150
200
Rate of flow (mL/min)
5
Oviedo ICCS&T 2011. Extended Abstract
Figure 3. Effect of gas flow rate on tar, water and char yields under different atmospheres (conditions: pyrolysis temperature 700 oC; 4 wt.% Mo/HZSM-5 catalyst; holding time 30 min) flow rate makes much carbon deposition, which causes catalyst deactivation, and it can be seen that the tar yield at 15 mL/min is lower than that at 25 mL/min in the MAP process. The water yield increases with the increase of flow rate, and the char yield under different atmospheres decreases slowly with the increase of flow rate.
3.3. Effect of Mo loading The effect of Mo loading in Mo/HZSM-5 catalyst on tar and water yields in the MAP process is shown in Figure 4. The tar yield increases with increasing Mo loading to 4 wt.%, but higher Mo loading than 4 wt.% has slightly effect on tar yield. Excessive Mo loading might be attributed to the partial blockage of zeolite channels by agglomerative molybdenum oxide particles [11]. So the tar yield decreases with Mo loading from 4 to 8 wt.%. However, the water yield remains nearly unchanged with different Mo loading in the MAP process. 30 Tar yield Water yield
Yield(wt.%,daf)
25 20 15 10 5 0
2
3
4
5
6
7
8
Mo loading (wt.%)
Figure 4. Effect of Mo loading on tar and water yields in the MAP process (conditions: pyrolysis temperature 700 oC; flow rate 25 mL/min; holding time 30 min)
3.4. Synergistic effect between coal and methane aromatization The synergistic effect between coal and methane aromatization is shown in Figure 5. Coal pyrolysis under CH4 (without catalyst), coal pyrolysis in the MAP process and methane aromatization (without coal) at different flow rate from 15 to 65 mL/min are studied. It can be seen that the sum of the tar yield of coal pyrolysis under CH4 (without catalyst) and the hydrocarbons yield of methane aromatization (without coal) is less 6
Oviedo ICCS&T 2011. Extended Abstract
than that of coal pyrolysis in the MAP process at different CH4 flow rate. The high tar yield in the MAP process may be explained that H and CHx groups detached from CH4 activated by catalyst at high temperature takes part in the reaction, and provides more opportunity to form tar by combination with free radicals cracked in coal pyrolysis. There are lots of H and CHx groups detached from CH4 at low flow rate from 15 to 35 mL/min because of the longer contact time of CH4 with the Mo/HZSM-5 catalyst, so the tar yield can be improved evidently at low flow rate, while methane aromatization reaction occurs little at CH4 flow rate of 65 mL/min, so the tar yield in the MAP process is just improved a little. This is a further illustration of the synergistic effect of coal and methane aromatization. 25
Aromatization(without coal) CH4(without catalyst)
21.50
Tar yield (wt.%, daf)
20.77 19.65
20
MAP
18.98 17.20
15.19
15.28
14.96
15
10
5 2.90 1.79
1.79
0
15
1.12
25
35
65
Rate of flow (mL/min)
Figure 5. Synergistic effect between coal and methane aromatization (conditions: pyrolysis temperature 700 oC; 4 wt.% Mo/HZSM-5 catalyst; holding time 30 min)
4. Conclusions It has been shown that the MAP process can obviously improve the tar yield. Compared with those in the pyrolysis of coal under N2 and H2, the tar yield increases by 7 and 6 wt.%, respectively, at the optimum conditions of 700 °C pyrolysis temperature, 25 mL/min CH4 flow rate, 4 wt.% Mo loading and 30 min holding time. The part of the increase in tar yield of the MAP process can be explained as the interaction of free radicals formed from methane aromatization and the cracking of the coal chemical structure during pyrolysis.
7
Oviedo ICCS&T 2011. Extended Abstract
Acknowledgement This research was performed with the support of the National Natural Science Foundation of China (No. 20576019, 20776028), the National High-Tech R&D Program (863 Program), the Ministry of Science and Technology, China (No. 2008AA05Z307), and the National Basic Research Program of China (973 Program), the Ministry of Science and Technology, China (No. 2011CB201301).
References [1] Zhang L, Xu SP, Zhao W, Liu SQ. Co-pyrolysis of biomass and coal in a free fall reactor. Fuel 2007; 86:353-359. [2] Liu JH, Hu HQ, Jin LJ, Wang PF, Zhu SW. Integrated coal pyrolysis with CO2 reforming of methane over Ni/MgO catalyst for improving tar yield. Fuel Processing Technology 2010; 91:419-423. [3] Cypres R, Furfari S. Low-temperature hydropyrolysis of coal under pressure of H2-CH4 mixtures. Fuel 1982; 61:721-724. [4] Steinberg M, Fallon PT. Make ethylene and benzene by flash methanolysis of coal. Hydrocarbon Process 1982; 61:92-96. [5] Lunsford JH. Catalytic conversion of methane to more useful chemicals and fuels: a challenge for the 21st century. Catalysis Today 2000; 63:165-174 [6] Wolf D. High yields of methanol from methane by C-H bond activation at low temperatures. Angewandte Chemie International Edition 1998; 37:3351-3353. [7] Wang L, Tao L, Xie M, Xu G, Huang J, Xu Y. Dehydrogenation and aromatization of methane under non-oxidizing conditions. Catalysis Letters 1993; 21:35-41. [8] Solomon PR, Fletcher TH, Pugmire RJ. Process in coal pyrolysis. Fuel 1993; 72:587-597. [9] Tromp PJJ. Coal pyrolysis. Ph.D. Thesis, University Amsterdam, 1987. [10] Tan PL, Au CT, Lai SY. Effects of acidification and basification of impregnating solution on the performance of Mo/HZSM-5 in methane aromatization. Applied Catalysis A 2007; 324:36-41. [11] Shu J, Adnot A, Grandjean BPA. Bifunctional behavior of Mo/HZSM-5 catalysts in methane aromatization. Industrial & Engineering Chemistry Research 1999; 38: 3860-3867.
8
Oviedo ICCS&T 2011. Extended Abstract
Integrated process of coal pyrolysis and CO2 reforming of methane over Ni/Al2O3-MgO catalyst Jiahe Liu, Haoquan Hu*, Lijun Jin, Shengwei Zhu State Key Laboratory of Fine Chemicals, Institute of Coal Chemical Engineering, School of Chemical Engineering, Dalian University of Technology, Dalian 116024, P. R. China, *
[email protected] Abstract Integrated coal pyrolysis with CO2 reforming of methane was investigated over Ni/Al2O3 catalyst promoted by alkaline earth metal oxides. The effect of catalyst promoters (MgO, CaO and BaO), MgO contents (1-15 wt.%), calcination temperature (500-800 °C) and reduction temperature (550-850 °C) of Ni/Al2O3-MgO catalysts on tar, water yields and CH4 conversion in pyrolysis of Pingshuo coal, and the carbon deposition on the catalysts were investigated. It was found that the catalyst promoter has a remarkable effect on the integrated process and MgO is the best in the studied promoters. The tar, water yields, CH4 conversion and the carbon deposition on the catalyst is 32.2, 30.7 wt.%, 18.8% and 7.8 wt.%, respectively, over Ni/Al2O3-MgO catalyst with 8 wt.% MgO calcined at 800 °C and reduced at 850 °C.
1. Introduction Hydropyrolysis, thermal devolatilization of coal under H2, is a good method to improve tar yield and quality of coal pyrolysis, but is hindered by the high cost of H2. Methane with high H/C ratio is the main composition of natural gas and coal bed gas, and is regarded as H2 substitute in hydropyrolysis. By now, there has been some works on coal pyrolysis under CH4 or the gas mixture containing CH4 which improve the tar yield of coal pyrolysis [1-3]. Our previous work also found that the tar yield of integrated Pingshuo coal pryolysis with CO2 reforming of methane over Ni/MgO catalyst was 1.6 time as that of coal pyrolysis under H2 at 750 °C [4]. Ni/Al2O3 catalyst has been widely used in CO2 reforming of methane. However, this catalyst is sensitive to carbon deposition. Adding MgO into the catalyst is an effective way to improve the coke resistance of the catalyst [5,6]. The aim of this work was to examine the effect of catalyst promoter, MgO content, calcination temperature and reduction temperature of Ni/Al2O3-MgO catalyst on tar, water yields and CH4 conversion in integrated process of coal pryolysis and CO2
1
Oviedo ICCS&T 2011. Extended Abstract
reforming of methane. 2. Experimental section 2.1. Coal sample A Chinese coal, Pingshuo (PS), was ground to -100 mesh and used in the experiments. The proximate and ultimate analyses of the coal sample are listed in Table 1. Table 1 Proximate and ultimate analyses of PS coal sample Proximate analysis (wt.%) Mad Ad Vdaf 2.23 17.93 37.19
C 80.41
Ultimate analysis (wt.%, daf) H N S O* 5.20 1.38 1.06 11.95
* By difference.
2.2. Coal pyrolysis The experiments of coal pyrolysis under CH4/CO2 were performed in a fixed-bed reactor consisting of upper catalyst layer and lower coal layer [4]. About 1 g catalyst and 5 g coal was placed in the two layers, respectively. The gas mixture of CH4 and CO2 with a ratio of 1:1 was fed into the reactor at the flow rate of 800 ml/min from the upside. The reactor was heated to 750 °C within 10 min and kept at the temperature for 30 min. Details of products collection can be found elsewhere [4]. 2.3. Catalyst preparation and characterization The catalysts were prepared by impregnating industrial Ni/Al2O3 catalysts (20-40 mesh) with aqueous solution of Mg(NO3)2, Ca(NO3)2 and Ba(NO3)2 in the content of 1-15 wt.% MgO, CaO and BaO, respectively. After dried at 110 °C for 12 h and calcined at 500-800 °C for 4 h, Ni/Al2O3-MgO, Ni/Al2O3-CaO and Ni/Al2O3-BaO catalysts were reduced at 550-850 °C in 15% H2 nitrogen flow for 4h. X-ray diffraction (XRD) patterns of catalyst samples were obtained on a DMAX2400 diffractor with Cu Kα radiation. The amount of coke deposition on Ni/Al2O3-MgO catalyst with 5, 8 and 15 wt.% MgO was measured in a thermogravimetric analyzer (Mettler Toledo model TGA/SDTA851e) by heating the used catalysts to 650 °C at a heating rate of 10 °C/min in a flow of air. The amount of coke deposition on Ni/Al2O3MgO catalyst with 1 wt.% MgO was roughly determined without using thermogravimetric analyzer for the serious carbon deposition which made the samples nonuniform.
2
Oviedo ICCS&T 2011. Extended Abstract
3. Results and Discussion 3.1. Effect of catalyst promoters Table 2 summarizes the tar, water yields and CH4 conversion of PS coal pyrolysis under CH4/CO2, and carbon deposition on catalysts with different promoters. The carbon deposition on industrial Ni/Al2O3 catalyst is higher than 100 wt.%, which caused reactor plugging within 30 min. The carbon deposition on Ni/Al2O3 catalyst promoted by BaO and CaO is higher than 100 and 50 wt.%, respectively, and the carbon deposition on Ni/Al2O3-MgO catalyst is only 7.8 wt.%, which makes Ni/Al2O3-MgO catalyst show better performance, and higher tar yield over the catalyst is obtained. Therefore, Ni/Al2O3-MgO catalyst was further studied upon catalyst preparation. Table 2 Tar, water yields and CH4 conversion of PS coal pyrolysis under CH4/CO2 and carbon deposition on Ni/Al2O3 catalysts with different promoters Tar yield Water yield Conversion of Carbon deposition Promoter (wt.%, daf) (wt.%, daf) CH4 (%) (wt.%) 28.0 50.6 37.9 >100 BaO 28.5 43.6 32.1 >100 CaO 29.0 40.4 27.0 >50 MgO 32.2 30.7 18.8 7.8 Reaction conditions: 750 °C, 30 min, Ni/Al2O3 catalysts promoted by 8 wt.% BaO, CaO and MgO calcined at 800 °C and reduced at 850 °C.
3.2. Effect of MgO content The effect of MgO content in Ni/Al2O3-MgO catalysts on tar, water yields and CH4 conversion of PS coal pyrolysis under CH4/CO2 and carbon deposition on the catalysts is listed in Table 3. The carbon deposition on the catalyst with 1 wt.% MgO is higher than 50 wt.%, and the carbon deposition on the catalyst with 5, 8 and 15 wt.% is 0.1, 7.8 and 8.9 wt.%, respectively. With the increase of MgO content from 1 wt.% to 8 wt.%, the tar yield increases. And it decreases when the MgO content increases to 15 wt.%, because MgO covers the active center of the catalyst. The water yield and CH4 conversion over the catalyst with 1 wt.% MgO are higher than those over the other catalysts, which is concerned with the higher carbon deposition on the catalyst [4]. Figure 1 displays the X-ray diffraction patterns of Ni/Al2O3-MgO catalysts with 1, 5, 8 and 15 wt.% MgO. The peaks at 63º are identified as NiO-MgO solid solution in the XRD of the catalysts [6]. Compared with the XRD peak of NiO-MgO solid solution in the catalyst with 1 wt.% MgO, the intensities of the XRD peaks in the other catalysts increase, suggesting that more NiO-MgO solid solution was formed in the catalysts,
3
Oviedo ICCS&T 2011. Extended Abstract
which inhibits the carbon deposition on the catalysts [4]. Table 3 Effect of MgO content on tar, water yields and CH4 conversion of PS coal pyrolysis under CH4/CO2 and carbon deposition on catalysts MgO content Tar yield Water yield Conversion of Carbon deposition (wt.%) (wt.%, daf) (wt.%, daf) CH4 (%) (wt.%) 1 29.2 33.2 27.6 >50 5 30.0 26.8 15.4 0.1 8 32.2 30.7 18.8 7.8 15 30.7 28.2 16.1 8.9 Reaction conditions: 750 °C, 30 min, Ni/Al2O3-MgO catalysts calcined at 800 °C and reduced at 850 °C. 3000
Al2O3 NiO-MgO MgAl2O4 and/or NiAl2O4
15 wt.% MgO
Intensity (a.u.)
2500
2000
8 wt.% MgO
1500
5 wt.% MgO
1000
500
1 wt.% MgO
0
20
30
40
50 60 2θ (degree)
70
80
Figure 1 XRD patterns of Ni/Al2O3-MgO catalysts with different MgO contents
3.3. Effect of calcination temperature Table 4 shows the effect of calcination temperature of Ni/Al2O3-MgO catalysts on tar, water yields and CH4 conversion of PS coal pyrolysis under CH4/CO2 and carbon deposition on the catalysts. The higher tar yield and lower water yield, CH4 conversion and carbon deposition on Ni/Al2O3-MgO catalyst calcined at 800 °C are obtained than those over the catalysts calcined at 500-700 °C. Table 4 Effect of calcination temperature on tar, water yields and CH4 conversion of PS coal pyrolysis under CH4/CO2 and carbon deposition on catalysts Calcination Tar yield Water yield Conversion Carbon deposition (wt.%, daf) (wt.%, daf) of CH4 (%) (wt.%) temperature (°C) 500 27.3 45.1 36.7 >100 600
26.5
43.4
37.0
>100
700
30.7
35.2
28.1
>50
800
32.2
30.7
18.8
7.8
Reaction conditions: 750 °C, 30 min, Ni/Al2O3-MgO catalysts with 8 wt.% MgO reduced at 850 °C.
Figure 2 presents the XRD patterns of Ni/Al2O3-MgO catalysts calcined at 500-800 °C.
4
Oviedo ICCS&T 2011. Extended Abstract
With increasing calcination temperature, the intensities of XRD peaks of NiO-MgO solid solution in the catalysts increase, and Ni/Al2O3-MgO catalyst calcined at 800 °C has the most NiO-MgO solid solution. The Ni particles reduced from the catalyst highly disperse on the catalyst surface, leading to high tar yield and low carbon deposition on the catalyst [7]. 3500
o 800 C
NiO-MgO Al2O3 MgAl2O4 and/or NiAl2O4
3000
o 700 C
Intensity (a.u.)
2500 2000
o 600 C
1500 1000
o 500 C
500 0
40
50 60 2θ (degree)
70
80
Figure 2 XRD patterns of Ni/Al2O3-MgO catalysts calcined at different temperatures
3.4. Effect of reduction temperature Figure 3 shows the effect of reduction temperature of Ni/Al2O3-MgO catalysts on tar, water yields and CH4 conversion of PS coal pyrolysis under CH4/CO2. The tar, water yields and CH4 conversion over the catalyst reduced at 550 °C is 23.0, 5.2 wt.% and 0.8%, respectively. With the increase of reduction temperature, the tar, water yields and CH4 conversion increase. The formation of NiO-MgO solid solution in the catalyst reinforces the interaction of NiO and support, which increases the reduction temperature
Yield (wt.%, daf) and conversion (%)
of catalyst and results in the high tar yield [7].
30
20
10
0
Tar yield Water yield CH 4 conversion
550 600 650 700 750 800 850 o Reduction temperature ( C)
Figure 3 Effect of reduction temperature on tar, water yields and CH4 conversion of PS coal pyrolysis under CH4/CO2 (Reaction conditions: 750 °C, 30 min, Ni/Al2O3-MgO catalysts with 8 wt.% MgO calcined at 800 °C)
5
Oviedo ICCS&T 2011. Extended Abstract
4. Conclusions The Ni/Al2O3 catalyst promoted by MgO presents higher tar yield and lower water yield, CH4 conversion and carbon deposition than the catalysts promoted by CaO and BaO. The tar yield over Ni/Al2O3-MgO catalyst with 8 wt.% MgO is higher than that over the catalyst with 1 wt.% MgO for the formation of more NiO-MgO solid solution in the catalyst. With increasing the calcination temperature, more NiO-MgO solid solution is formed in the catalyst, so Ni/Al2O3-MgO catalyst calcined at 800 °C provides the highest tar yield. With the increase of reduction temperature, more Ni particles are reduced from the catalyst and highly dispersed on the catalyst, which results in the improvement of catalyst activity. Acknowledgement This research was performed with the support of the National Natural Science Foundation of China (No. 20576019, 20776028), the National High-Tech R&D Program (863 Program), the Ministry of Science and Technology, China (No. 2008AA05Z307), and the National Basic Research Program of China (973 Program), the Ministry of Science and Technology, China (No. 2011CB201301). References [1] Egiebor NO, Gray MR. Evidence for methane reactivity during coal pyrolysis and liquefaction. Fuel 1990; 69: 1276-82. [2] Liao HQ, Li BQ, Zhang BJ. Co-pyrolysis of coal with hydrogen-rich gases. 1. Coal pyrolysis under coke-oven gas and synthesis gas. Fuel 1998; 77: 847-51. [3] Smith GV, Wiltowski T, Phillips JB. Conversion of coals and chars to gases and liquids by treatment with mixtures of methane and oxygen or nitric oxide. Energy & Fuels 1989; 3: 536-7. [4] Liu JH, Hu HQ, Jin LJ, Wang PF. Effects of the catalyst and reaction conditions on the integrated process of coal pyrolysis with CO2 reforming of methane. Energy & Fuels 2009; 23: 4782-6. [5] Horiuchi T, Sakuma K, Fukui T, Kubo Y, Osaki T, Mori T. Suppression of carbon deposition in the CO2-reforming of CH4 by adding basic metal oxides to a Ni/Al2O3 catalyst. Applied Catalysis A: General 1996; 144: 111-20. [6] Mehr JY, Jozani KJ, Pour AN, Zamani Y. Influence of MgO in the CO2-steam reforming of methane to syngas by NiO/MgO/α-Al2O3 catalyst. Reaction Kinetics and Catalysis Letters 2002; 75: 267-73. [7] Yamazaki O, Tomishige K, Fujimoto K. Development of highly stable nickel catalyst for methane-steam reaction under low steam to carbon ratio. Applied Catalysis A: General 1996; 136: 49-56.
6
Programme topic: Coal pyrolysis and liquefaction
Kinetics of co-pyrolysis of high- and low-sulfur coal blends with additives L. Butuzova1, R. Makovskyi1, V. Bondaletova1, D. Dedovets2, G. Butuzov1 1
Donetsk National Technical University, 58 Artema str., Donetsk 83000, Ukraine, tel. fax: +38(0622) 55-85-24,
[email protected] 2
[email protected]
National Academy of Sciences of Ukraine, L.M. Litvinenko Institute of Physical Organic Chemistry and Coal
Chemistry, 70 R.Luxemburg str., Donetsk 83114, Ukraine
Abstract The thermogravimetric studies of a pyrolytic decomposition of blend based on low- and highsulfur coals with additives (components of coal-tar and radical polymerization initiator) were carried out. Thermokinetic analysis demonstrated that thermal decomposition of the chemically treated bland proceeds more intensely than for the original bland and permits variation in the gas evolution rate on the different stages of pyrolysis process. Keywords: sulfur coals, pyrolysis, kinetics 1. Introduction Previous investigations indicate that the dependence of the coal structure and reactivity on sulfur content is fairly strong [1-2]. But there no a comparative data in the scientific literature about the thermal behaviour of low- and high-sulphur coals of the same rank in coking blends and in presents of additives. Developments of pretreatment methods for sulfur coals are especially desirable for reduction of the sulfur content in pyrolysis products and for the control of caking ability. The components of coal-tar and radical polymerization initiator are deemed as the most effective additives for cokemaking and for a study of the pyrolysis mechanisms [3]. The effectiveness of such materials at the different stages of pyrolysis process was compared. The aim of this paper is a detailed study of the kinetic behaviour of low- and high-sulphur coals and their blend during pyrolysis with additives using thermogravimetry method and elucidation of usability of the thermokinetic analysis for coke properties determination. 2. Experimental Experiments were conducted on the pairs of petrographically homogeneous low- and highsulphur bituminous coals of Donets Basin. It was high-sulphur coal of J-Grade (JRC: Cdaf = 87,3; Vdaf =31,7; Sdt =2,81) and low-sulphur coal G-Grade (GLRC: Cdaf =85,1; Vdaf=36,0; Sdt =1,22) according to Ukrainian classification and their blend (50:50). The samples were treated by radical polymerization initiator (acrylic acid dinitrile - AAD) and by the components of coal tar (pitch, anthracene, phenanthrene). The radical polymerization initiator was introduced to affect the course of radical reactions. Other additives were used as possible analogues of the
components, of liquid semi-coking and coking products which are known to be responsible for synthesis reactions during coking. The thermal behaviour of coal blends were studied by thermogravimetric analyses and standard Sapozhnikov metods (GОSТ 1186-87). The thickness of plastic layer (y) and the contraction (x) by Sapozhnikov’s method was applied as characteristic of coal coking ability. Derivatogrammes were registered in a Q-1500D derivatograph of Paulic- Paulic-Erdei system at the rate 100C/min in a closed platinum crucible under the layer of quartz sand up to 10000C. The kinetic parameters, i.e. the activation energy E and the rate of decomposition in different periods of pyrolysis were calculated by the results of the continuous measurement of the weight loss. 3. Results and discussion Behaviour of the mass loss curves indicates that the blends decomposition process may be presented as a sum total of seven independent steps (linear parts on the curve TG). A Table 1, 2 shows the values of temperature intervals and corresponding mass loss for different steps of pyrolysis process. Table 1 – The temperature intervals for independent steps of coals pyrolysis process. The temperature intervals for different steps*, °C
Coals, blends
Additive
(50:50)
I
III
IV
V
VI
VII
J RC
–
100-140 350-415
415-480
480-595
595-867
867-930
G LRC
–
75-125
327-400
400-472
472-504
504-573
573-894
GLRC+JRC
–
70-140
328-400
400-477
477-542
542-700
700-900
GLRC+JRC
AAD
75-180
350-420
420-455
455-530
530-757
757-900
GLRC+JRC
pitch
60-150
340-400
400-485
485-542
542-720
720-880
GLRC+JRC
anthracene
30-110
190-395
395-480
480-588
588-700
700-885
GLRC+JRC
phenanthrene
50-110
170-400
400-473
473-573
573-700
700-900
* II step is occurred without mass loss
It can be seen from the Table 2, that the most intensive decomposition of J RC sample are occured at the IV-VI steps which is known to be related to formation of the main bulk of the semi-coking products. The periods of the most intensive decomposition of GLRC coal are the IV and VII steps. VII period (coking state) is characterized by a much higher rate of volatile products evolution from the solid phase as compared to the previous. Thermal decomposition of blend is characterized by a comparative deceleration of the mass loss at the V, VI steps in comparison with J RC.
Introduction of the radical polymerization initiator AAD results in acceleration of the gas evolution rate at the first - third steps (in a three or two-fold) and shift the temperature range of these stages to higher temperatures. Таблиця 2 – Kinetic of the mass loss during independent steps of coals pyrolysis process, %. Coals,
Additive
Mass loss at different steps Σ∆
I
III
IV
V
VI
VII
2
3
4
5
6
7
8
9
J RC
–
0,47
1,64
7,74
6,57
11,26
9,38
37,06
G LRC
–
2,11
3,29
8,92
1,41
5,87
15,96
37,56
GLRC+JRC GLRC+JRC GLRC+JRC GLRC+JRC GLRC+JRC
– AAD pitch anthracene
1,88 6,10 1,88 2,81 2,35
1,88 3,29 2,35 3,99 5,16
8,45 5,16 8,92 8,45 7,74
5,16 7,28 5,16 7,74 6,81
7,04 9,62 11,97 4,92 5,40
10,79 8,21 9,39 10,32 10,33
35,43 39,66 39,66 38,23 37,79
blend (50:50) 1
phenanthrene
There are reasons to believe that this acceleration appears due to scission of inter- and intramolecular bonds, including -C-S- bonds. This hypothesis is supported by the lowermost value for the activation energy for AAD-treated sample (Table 3). The removal of sulfur-and oxygen-containing groups causes a decrease in the rate of mass loss at the IV stage due to increases the thermostability of solid fuel. Decomposition rate increases at V and VI steps (semicoking) and decreases during coking (VII period) under the action of AAD. This is indicative of the formation of more condensed structures by polyrecombination reactions. Table 3 Thermokinetics parameters for the most intensive decomposition step in derivatograms of investigated coals and blends. Ee, kJ/mol
7,04
1,56
75,06
380-485
9,15
2,13
57,67
440
380-495
8,45
1,92
67,58
AAD pitch
430 450
380-512 390-505
11,03 8,92
2,57 1,98
40,20 54,66
GLRC+JRC
anthracene
440
400-505
10,79
2,45
42,29
GLRC+JRC
phenanthrene
440
395-512
11,26
2,56
40,81
Additive
Tm, Ka), °C
(Ti-Tf) b), °C
J RC
–
450
395-512
G LRC
–
430
GLRC+JRC
–
GLRC+JRC GLRC+JRC
blends (50:50)
a)
%
∆
r c) mg/(min ·g)
Coals,
Tm - temperature of maximum reaction/process rate; initial state; c) relative rate of thermal decomposition,
b)
at Tm
Tf - temperature of the final state, Ti - temperature of the
These data confirm that the addition of AAD to coal blend modifies the plastic layer: the thickness of plastic layer increases from 14.5 to 15.5 mm and the contraction increases
substantially from 27 up to 36 mm whereas the mechanical strength of coke is stable. As can be seen from Table 3, the influence of all additives results in a decrease the values of E and a change the ratio of the rates of destruction and synthesis reactions. The influence of pitch is more pronounced only at the VI step. The rate of organic sulfur decomposition was highest in the same temperature range [4]. Moreover the total sulfur content in the obtained cokes was less than 1.5 %. There are reasons to believe that the pitch intensifies desulphurization process and improves of the coking ability of blends. When anthracene and phenanthrene were added, an intensification of gas evolution processes at I, III and V stages and a significant deceleration of blend decomposition at VI stage are observed. The presence of these additives in the blend shifts the temperature region of the third stage to lower temperatures. Probably, highly condensed aromatic structures help to stabilize the free radicals present in reaction media with subsequent promotes of polyrecombination reactions at VI stage. Thermostability of linear structures is higher than angular. Therefore upon the effect of phenanthrene the conversion degree is increased from 1.88 to 5.16 at III stage. The reason for this is the lower value of the effective activation energy (Ee) in the process of vapour-gaseous products formation for phenanthrene in comparison with anthracene (depending on mutual arrangement of aromatic rings and variation in paramagnetic centers concentration). Thus, treatment with AAD, anthracene and phenanthrene has exerted considerable effects. Thermo-chemical destruction promotes impoverishment of solid products with sulfur- and oxygen-containing groups, i.e. it improves their quality. These reactants act as radical polymerization initiators, thus increasing the yield of semi-coke (Table 3) and coke (Table 2) as compared to untreated blend. Accordingly, there are many grounds for believing that these methods of coal pre-treatment are a promising for low quality coals processing. 4. Conclusions Chemical pretreatment has a considerable influence on the kinetics of co-pyrolysis of sulfur coalcontaining blends. The use of additives (components of coal-tar and AAD) shows the possibility to manage of the rate and mechanism of the separate stages of pyrolysis process. References [1] Butuzova, L., Safin, V., Marinov, S., Yaneva, N., Turchanina, O., Butuzov, G. The pathways for thermal decomposition of coals with high content of sulphur and oxygen Geolines, Academy of Science of the Czech Republic. 2009;22:15-9. [2] Mianowski A. Butuzova L., Radko T. Turchanina O. Thermokinetic analysis of decomposition of Ukrainian coals from Donets Basin. Bulletin of geosciences. 2005;80,№ 1:39-44. [3] Fernandez A.M., Barriocanal C., Diez M.A., Alvarez R. Influence of additives of various origins on thermoplastic properties of coal. Fuel. 2009;88:2365-72 [4] Gryglewicz G. Sulfur transformations during pyrolysis of a high sulfur Polish coking coal. Fuel. 2009;74,№ 3:356-61.
Oviedo ICCS&T 2011. Extended Abstract
Effect of elemental composition of various additives on the modification of coal thermoplastic properties
M.G. Montiano, C. Barriocanal, R. Alvarez
[email protected] Instituto Nacional del Carbón, CSIC, Apartado 73, 33080 Oviedo. Spain
Abstract A bituminous coal normally used in the cokemaking industry was selected as base coal for studying the modification of its thermoplastic properties due to the addition of two different sawdusts. 1. Introduction The iron and steel industry is a major greenhouse gas (GHG) emission source. The introduction of biomass into the steel industry either as a substitute for coal in the blast furnace or as a component of coal blends for coking has been considered as a way to reduce CO2 emissions [1-3]. Coal thermoplastic properties are considered to be of great importance for the formation of the structure of metallurgical coke and consequently their properties. The use of additives modifies this behaviour by either improving it or causing it to deteriorate [4,5]. The modification of thermoplastic properties of coal can be assigned to physical and chemical factors [4]. Our aim is to study the effect of the amount of heteroatms present in the biomass on the coal’s plastic properties. To that end, two waste sawdusts were heated up to 250°C in order to reduce their oxygen content and added to a bituminous coal to study the effect on coal thermoplasticity.
2. Experimental section 2.1.Materials One bituminous coal (G) and two waste sawdust were selected: chestnut sawdust (SC1) and oak sawdust (SR1). The Proximate analyses were performed following the ISO562 and ISO1171 standard procedures for volatile matter and ash content, respectively. The elemental analyses was carried out using a LECO CHN-2000 for C, H and N content (ASTM D-5773), a LECO S-144 DR(ASTM D-5016) for sulphur and a LECO VTF-900 for direct oxygen determination.
1
Oviedo ICCS&T 2011. Extended Abstract
2.2.Materials The TG/DTG analysis of the coal and the additives was carried out using a TA Instruments SDT 2960 thermoanalyser. Samples of 10-15 mg with a particle size of <0.212 mm were heated to 1000 ºC at a rate of 3 ºC/min under a nitrogen flow of 100ml/min.
2.3.Thermoplastic properties The thermoplastic properties of the coal and the blends containing additives in the proportions of 2 wt.% were tested by the Gieseler method in a R.B. Automazione Gieseler plastometer PL2000, following the ASTM D2639-74 standard procedure. The parameters derived from this test were: (i) softening temperature, Ts; (ii) the temperature of maximum fluidity, Tf; (iii) resolidification temperature, Tr; (iv) plastic range, Tr-Ts, defined as the difference between the resolidification and softening temperatures; and (v) maximum fluidity, MF, expressed as dial divisions per minute (ddpm).
3. Results and Discussion 3.1.Effect of heating on the elemental composition of the sawdust Table 1 shows the ash, volatile matter content (VM) content and elemental composition of the materials used. The biomass presents a high volatile matter and oxygen content and low ash and carbon content compared to coal. Heating the sawdust produces a decrease in volatile matter and oxygen content.
Table 1. Proximate and ultimate analysis of the materials studied. G
SC1
SC1-250
SR1
SR1-250
VM (wt.%)
32.0
79.8
58.1
88.2
65.1
Ash (wt.%db)
8.3
2.9
5.9
1.3
0.9
C (wt.%db)
78.3
49.6
58.8
52.3
59.8
H (wt.%db)
4.9
5.7
4.7
6.2
4.9
N (wt.%db)
2.5
0.6
0.8
0.5
0.5
S (wt.%db)
0.88
0.00
0.03
0.00
0.04
O (wt.%db)
4.5
41.4
27.5
39.7
33.8
2
Oviedo ICCS&T 2011. Extended Abstract
3.2.Thermogravimetric analysis of coals and blends In order to explain the different behaviours of the additives and possible interactions between the coal and the additives, the thermal behaviour was also studied. Figure 1 shows the derivative of the mass loss (DTG) of the coal and the additives. The curve shows three decomposition stages [6,7]. Hemicellulose started its decomposition over 220-315ºC (maximum mass loss rate at 268ºC). Cellulose decomposed at a higher temperature (the temperature of maximum mass loss rate being 315°C). Whereas lignine decomposed over a wide temperature range (100-900ºC). Figure 1 shows that the first decomposition step disappears for the sawdust that has been heated to 250 °C. At the same time the amount of volatile matter evolved by the heated sawdust in the plastic range of coal i.e. 400-500 °C increases, fomenting the possibilities of interaction between the volatiles evolved by the sawdust and the reaction system that is produced during coal plastic stage.
4.5
3.0 4.0 3.5
SC1 SC1 250 G
-1
2.0
DTG (% min )
DTG (% min -1)
2.5
1.5
1.0
SR1 SR1 250 G
3.0 2.5 2.0 1.5 1.0
0.5
0.5
0.0 100
200
300
400
500
600
700
0.0 100
Temperature (ºC)
200
300
400
500
600
700
Temperature (ºC)
Figure 1. DTG curves of the additives and the coal.
3.3.Influence of sawdust on coal plasticity Figure 2 shows the variation of fluidity with the temperature of the base coal G and its blends with 2 wt.% of each additive. Coal G is a high volatile matter content coal that presents a high maximum Gieseler plasticity. The addition of 2 wt% of the sawdust produced in all cases a reduction in coal plasticity. It can be seen from Figure 2 that for the blends maximum fluidity appears at higher temperatures than for the base coal. The softening temperature increases with the additives although no clear trend is observed for plastic range. SC1 produced a greater decrease than SR1 which could be attributed to its higher ash and oxygen content.
3
Oviedo ICCS&T 2011. Extended Abstract
4500
Gieseler fluidity(ddpm)
G G+2%SC1 G+2%SC1 250
3500 3000 2500 2000 1500 1000 500 0 400
Gieseler fluidity (ddpm)
4500
4000
4000
3000 2500 2000 1500 1000 500 0 400
410
420
430
440
450
460
G G+2%SR1 G+2%SR1 250
3500
410
420
430
440
450
460
470
470
Temperature (ºC)
Temperature (ºC)
Figure 2. Variation of coal fluidity with temperature.
Heating the sawdusts to 250 °C effectively reduced their oxygen content. The addition of the heat treated sawdust to the coal produced in both cases a greater reduction than that of the non-treated sawdust. Consequently the effect of the sawdusts which has been heat treated is more deleterious than the original one. Nevertheless the blends still maintain sufficient plasticity to produce coke of good quality. From Figure 1 it can be seen that the degree of chemical interaction between the sawdust and the coal is low. Consequently the physical effect of the additive that is, the presence of a material which has already lost most of its volatiles might be more important. This “inert” material may reduce coal fluidity by absorbing the small-molecular-mass compounds produced when the coal is heated to over 400ºC and which are responsible for the softening and melting of the coal.
4. Conclusions The addition of 2 wt% of sawdust was found to reduce the plasticity of the coal. Heat treatment of the sawdust to try to reduce its oxygen content produced an even greater reduction in the plastic properties of coal. The results obtained indicate that the physical effect of the presence of an “inert” material during the coal plastic stage has greater effect than the increase in the reactivity of the system due to the higher oxygen content.
Acknowledgements The research leading to these results has received funding from the European Union's Research Fund for Coal and Steel (RFCS) research programme under grant agreement n° RFCR-CT-2010-00007
4
Oviedo ICCS&T 2011. Extended Abstract
References [1] J.A. MacPhee, J.F. Gransden, L. Giroux l, J.T. Price. Possible CO2 mitigation via addition of charcoal to coking coal blends. Fuel Proc. Technol 2009; 90: 16-20. [2] F.G. Emmerich, C.A. Luengo. Babassu charcoal: A sulfurless renewable thermoreducing feedstock for steelmaking. Biomass and Bioenergy 1996; 10: 41-44. [3] S. Das, S. Sharma, R. Cloudhury. Non-coking coal to coke: use of biomass based blending material. Energy 2002; 27: 405-414. [4] Elliot M. A. Chemistry of Coal Utilization, 2nd Supplementary Volume. New York: Wiley Interscience; 1981. [5] A.M. Fernández, C. Barriocanal, M.A. Díez, R. Alvarez. Influence of additives of various origins on thermoplastic properties of coal. Fuel 2009; 88:2365-2372. [6] D.K. Shen, S. Gu, K.H. Luo, S.R. Wang, M.X. Fang. The pyrolytic degradation of wood-derived lignin from pulping process. Bioresource Technology 2010; 101:61366146. [7] Haiping Yang, Rong Yan, Hanping Chen, Dong Ho Lee, Chuguang Zheng. Characteristics of hemicellulose, cellulose and lignin pyrolysis. Fuel 2007; 86: 17811788.
5
Oviedo ICCS&T 2011. Extended Abstract
Influence of residual volatile matter in semicokes on coking pressure E. Díaz-Faes, C. Barriocanal, R. Alvarez Instituto Nacional del Carbón, CSIC, Apartado 73, 33080 Oviedo. Spain email:
[email protected]
Abstract Eleven coals of different rank and dangerousness were selected. The thermoplastic properties of the coals were tested by the Gieseler method and the composition of the residual volatile matter of the semicokes was assessed by means of a mass spectrometer attached to a thermobalance. Contraction of the semicoke was measured by means of the Koppers-INCAR expansion/contraction test. Coal devolatilization and fluidity were studied together with the amount and composition of the residual volatile matter from the semicokes in relation to contraction/expansion. Although not conclusive, there is evidence to suggest that the temperature of maximum evolution of the volatile products corresponding to the selected m/z ions of the semicokes is related to the contraction measured by Koppers-INCAR test.
1. Introduction The production of metallurgical coke for blast-furnace use is carried out in horizontal slot-type ovens, which are indirectly heated through the side walls. Consequently, a temperature gradient appears in the coal charge and two layers of coke and semicoke coexist with one layer of coal in the oven. The generation of excessive pressure during industrial coking may damage oven walls and operational difficulties can arise during coke pushing, resulting in substantial economic loss for the steel industry. Coal type is the most important parameter that influences the generation of coking pressure together with the type of additives and operational conditions employed [1]. There is no clear relation between the generation of coking pressure and specific coal properties. Several methods are currently used to study coking pressure generation, one of which involves the study of expansion/contraction during coking. The aim of this work is to study the composition of the residual volatile matter in semicokes obtained at resolidification temperature from a series of coals of different rank to see whether there is any relation between the composition of the volatile matter
1
Oviedo ICCS&T 2011. Extended Abstract
of the semicokes and the contraction/expansion of the parent coal.
2. Experimental 2.1 Materials Eleven bituminous coals of different rank and dangerousness on coking were selected from those used in the European cokemaking industry. The main characteristics of the coals are shown un Table 1. Proximate analyses were performed following the ISO562 and ISO1171 standard procedures for volatile matter and ash content, respectively. The elemental composition was determined using a LECO CHNS-932 and a LECO VTF900 set of equipment for direct oxygen determination. The thermoplastic properties were determined by the Gieseler method in a R.B. Automazione Gieseler plastometer PL2000, following the ASTM D2639-74 standard procedure. Table 1. Main characteristics of the coals studied Coal
C1
C2
C3
C4
C5
C6
C7
C8
C9
C10
C11
9.4
8.0
9.5
9.5
10.8
6.1
9.5
7.0
7.0
7.6
7.0
19.2
19.5 20.9 22.1 22.2
22.2
23.1
29.7
31.1
32.0
33.7
82.0
83.0 81.0 80.3 79.6
84.1
80.5
81.7
80.5
80.6
81.1
H (wt.% db)a
4.5
5.0
4.3
4.7
4.4
4.8
4.6
5.1
5.0
5.0
5.3
a
1.6
1.9
1.9
1.9
2.1
1.5
1.8
1.8
1.6
1.6
1.7
a
0.33
0.51 0.46 0.49 0.52
0.97
0.83
1.10
1.03
1.00
0.79
O (wt.% db)a
3.36
3.09 3.74 3.71 4.50
3.98
4.33
4.64
5.65
5.15
5.86
Gieseler MF (ddpm)b
240
436
235
928
230
1360
2712 29215 16306 21646 10157
Tr (ºC)c
495
500
488
491
492
495
494
Ash (wt.% db)a V.M. (wt.% db) C (wt.% db)
a
N (wt.% db) S (wt.% db)
a
483
480
477
474
a
Dry basis MF: maximum fluidity expressed in dial divisions per minute c Resolidification temperature obtained from the Gieseler plasticity test b
2.2 Semicoke contraction Contraction of the semicoke was measured by means of the Koppers-INCAR expansion/contraction test (KI) [2-5]. Briefly, 80 g of coal sample, ground to <1 mm size, with a bulk density of 820 kg/m3 was placed inside a stainless steel crucible and heated in a sole heated oven to 900 ºC. A constant pressure of 0.89 kg/ cm2 was exerted on the charge for 2 h. The change in charge volume compared to the initial state of the coal sample was recorded on a graph in millimetres. Contraction is expressed by negative values, while positive values indicate expansion. A coal is considered dangerous on coking when it presents a contraction value lower than 10 mm.
2
Oviedo ICCS&T 2011. Extended Abstract
2.3 Thermogravimetric analysis Semicokes obtained from the Gieseler test were subjected to thermogravimetric analysis (TG) in a simultaneous TA Instrument SDT2960 analyzer. In all the experiments, approximately 10 mg of sample was heated from room temperature up to 1000 ºC at a heating rate of 10ºCmin-1 using a helium flow rate of 100 mlmin-1 to sweep out the volatile products. The particle size used was the same as that employed for the proximate and ultimate analyses (< 0.212 mm). For the TG–MS runs of semicokes, a quadrupole mass spectrometer (Balzers, Thermostar GSD-300T) linked to the thermobalance was used to record the gas evolution profiles. Evolved gases pass through a quartz capillary of internal diameter 0.23 mm, kept at 200 ºC to avoid condensation. The ions selected (see Table 2) were monitored together with the thermogravimetric parameters. Table 2. MS signals followed during semicoke pyrolysis and their assignment to the fragmented ions. Alkyl fragments
Olefin fragments
CnH2n+1+
CnH2n-1+
Aromatic fragments
m/z
2
15
29
43
57
27
41
55
77
78
91
ion
H2+
CH3+
C2H5+
C3H7+
C4H9+
C2H3+
C3H5+
C4H7+
C6H5+
C6H6+
C7H7+
3. Results and Discussion The devolatilization of the semicokes obtained from the Gieseler test took place in two stages, presenting a first maximum in the range between 512 ºC and 568 ºC and a second peak between 720 ºC and 749 ºC (Table 3). In general, for the first maximum the rate of maximum volatile matter evolution (DTGmax1) increased and Tmax1 decreased as the rank of the parent coal decreased, whereas no apparent trend was found for Tmax2. DTGmax2 remained almost constant for all the semicokes studied. Although most of the volatile matter was emitted above 500 ºC, the amount emitted up to that temperature was higher for lower rank coals. Furthermore, volatile matter evolved in the ranges between 500 ºC and 750 ºC and 750 ºC and 950 ºC was lower for low rank coals. Additionally, it can be seen from Tables 1 and 2 that semicoke contraction is not directly related to the rank of coals. Nevertheless, it was found that for the safe coals, there was a linear relationship between the contraction values and the volatile matter evolved from the semicokes between 750 and 950ºC.
3
Oviedo ICCS&T 2011. Extended Abstract
Table 3. TG parameters and KI contraction values of semicokes Semicoke
C1
C2
VM500 (ºC)a
21.1
15.6 20.7 19.9 18.1 21.2 20.3 22.1 27.5 28.1 27.4
VM500-750 (ºC)a
60.7
62.6 61.6 62.4 63.0 59.6 62.3 61.1 57.1 58.5 57.6
VM750-950 (ºC)a
16.6
19.0 15.9 16.1 16.9 16.8 15.9 15.1 13.8 11.8 13.3
Residue (%)
87.9
89.9 87.0 87.7 87.4 87.4 88.2 87.1 85.3 85.6 84.3
DTGmax1 (%/min)b
0.42
0.36 0.47 0.44 0.41 0.39 0.40 0.44 0.48 0.52 0.52
Tmax1 (ºC)c
543
568
0.21
0.18 0.21 0.21 0.21 0.25 0.21 0.21 0.22 0.21 0.22
728
735
735
734
749
702
715
731
722
730
735
-15
-12
-18
-17
-17
-9
-7
-18
-21
-22
-21
DTGmax2 (%/min) Tmax2 (ºC)
c
KI (mm)d a b
b
C3
C4
538
532
C5
C6
541
C7
562
545
C8
543
C9
539
C10
521
C11
535
Volatile matter evolved up to a specific temperature or temperature interval and normalized to 100 % Rate of maximum volatile matter evolution. 1 and 2 refers to the first and second peak respectively
c
Temperature of maximum volatile matter evolution
d
Semicoke contraction measured by Koppers-INCAR test
For TG-MS coupling, the information obtained from the pyrolysis experiments is more complete. The evolution of different volatile products can provide information on the chemical reactions occurring during thermal decomposition. Arenillas et al. state that each compound detected in the MS has it own response factor and so the intensities of the same compound, for different samples, can be compared [6]. As an example, Figure 1 shows the evolution of the ions of m/z 29 (C2H5+) and 41 (C3H5+), corresponding to the ethyl and propenyl fragments, respectively. It can be seen that the amount of these two ions decreased as the rank of the parent coals increased, while the characteristic temperatures increased with rank.
C2 C8 C11
200
300
400
500
600
Temperature (ºC)
700
C2 C8 C11
m/z 41
Intensity (arbitraty Units)
Intensity (arbitraty units)
m/z 29
800
200
300
400
500
600
700
800
Temperature (ºC)
Figure 1. Ethyl and propenyl evolution curves followed by MS during the pyrolysis of three semicokes obtained from different rank coals. Although more investigation is necessary, there seems to be a relationship between rank
4
Oviedo ICCS&T 2011. Extended Abstract
of the parent coals and the intensity and characteristic temperatures of the ions studied. This can be attributed to the varying amounts of different functional groups present in the semicokes. It is well known that it is impossible to distinguish between dangerous and safe coals on the basis of their rank. Most low-volatile coals are dangerous on coking, while highvolatile coals generate very low coking pressure. Nevertheless, it is important to bear in mind that coals with a similar volatile matter content can behave differently as regards coking
pressure
generation,
e.g.
coal
C5
(VM = 22.2 wt %
and
KI
contraction = -17 mm) and coal C6 (VM = 22.2 wt % and KI contraction = -9 mm). The amount of volatiles and their rate of evolution during the process seem to be irrelevant for elucidating the behaviour of a coal during coking [7]. In this study there seems to be a relationship between the amount and characteristic temperatures of certain fragments measured by TG-MS and the KI contraction values, although it is necessary to broaden the study to include more coals that produce low contraction.
4. Conclusions The parameters obtained from semicoke devolatilization by thermogravimetric analysis are dependent upon the rank of the parent coals and some of them may be related to the contraction values obtained from the KI test, although more detailed studies are necessary. TG-MS appears to be a promising tool for helping to explain the factors that lead to the generation of dangerous pressures during the coking of some coals.
Acknowledgement The research leading to these results has received funding from the European Union's Research Fund for Coal and Steel (RFCS) research programme under grant agreement n° RFCR-CT-2010-00006
References [1] Tucker J, Everitt G. Coking pressures- its causes, measurement and control. 2nd International Cokemaking Congress, Londres. 1992; 2:40-61. [2] Escudero JB, Álvarez R. Influence of air oxidation on the pressure exerted by coking coals during carbonization. Fuel 1981; 60:251–3. [3] Alvarez R, Pis JJ, Barriocanal C, Lázaro M. Characterization of dangerous coals
5
Oviedo ICCS&T 2011. Extended Abstract
during carbonization. Effect of air oxidation and ash content of coals. Cokemaking Int 1991; 1:37–42. [4] Alvarez R, Pis JJ, Barriocanal C, Sirgado M. Practical application of a laboratory test to measure expansion and contraction during carbonization. Cokemaking Int 1992; 4:16-8. [5] Pis JJ, Alvarez R, Lázaro M, Barriocanal C. Application of a laboratory test to resolve the problem of coking a dangerous coal. Bull Soc Geolog France 1991; 2:443–5. [6] Arenillas A, Rubiera F, Pis JJ. Simultaneous thermogravimetric-mass spectrometric study on the pyrolysis behaviour of different rank coals. J. Anal. Appl. Pyrolysis 1999; 50:31-46. [7] Casal MD, Canga CS, Díez MA, Álvarez R, Barriocanal C. Low-temperature pyrolysis of coals with different coking pressure characteristics. J. Anal. Appl. Pyrolysis 2005;
74:96-103.
6
Oviedo ICCS&T 2011. Extended Abstract
Pyrolysis of wastes from tyre grinding B. Acevedo, C. Barriocanal, R. Alvarez Instituto Nacional del Carbón, CSIC, Apartado 73, 33080 Oviedo. Spain
[email protected] Abstract Two wastes from tyre grinding and their blends with coal and a residual pitch were pyrolyzed at various heating rates in a thermobalance. The curves derived from the analysis contributed to a better understanding of the decomposition steps of the wastes and the interactions between the components of the blends. 1. Introduction In recent years, the economic and environmental problems associated with the disposal of automotive tyres are increasing, and thereby the reuse and recycling of these materials has become extremely important. Developed countries generate around 6 million tons per year of this waste [1]. Moreover waste tyres are a potential fire hazard in landfills and they are not easily biodegradable material. Scrap tyres once collected are subjected to a grinding procedure that allows the separation of tyre granules, metallic parts and reinforcing fibres. Tyres are mainly made up of rubber, carbon black, steel and fibres acting as reinforcing materials. The most commonly used tyre rubber is styrenebutadiene copolymer (SBR) that contains styrene in various proportions, but also natural rubber and poly-butadiene. [2]. All of the materials used are 100% recyclable. Moreover, their chemical and physical properties make them a highly valuable resource. Pyrolysis is a process that allows the decomposition of waste tyres into gas, pyrolytic oil and carbon black, all of which are highly useful products [1,3]. The aim of this work was to study the pyrolysis of waste material derived from grinding scrap tyres and their blends with coal and a residual pitch. 2. Experimental section Reinforcing fibres obtained from shredding tyres are made of fibres of natural and synthetic origin [4]. The grinding process affects their final morphology and a heterogeneous mixture of two components is obtained: 1. fibre (original cord, F) and 2. microfibre (MF). Some rubber is also present in this waste due to difficulties in the separation process. Reinforcing fibres (RF) used for the present work were made up of 46.2wt. % microfibre and 53.8 wt.% of fibre. Both, the tyre crumbs (TC) and the fibres
1
Oviedo ICCS&T 2011. Extended Abstract
were cryogenically milled and sieved to a size of < 425 µm. A coal and a residual pitch with a volatile matter content of of 30.7 and 65 wt% db and ash content of 7.2 and 0.9 wt% db were used to prepare 1:1 blends. Thermogravimetric analysis of the wastes, the coal, the pitch and their 1:1 blends were carried out using a TA Instruments Q600 thermoanalyser. Samples of 10 mg were heated to 950 ºC at 5, 10, 20 and 50 ºCmin-1 under a nitrogen flow of 100 mlmin-1. 3. Results and Discussion 3.1. Thermogravimetric analysis of the two wastes The elemental composition, ash and sulphur content of the wastes are similar to those used in another research work [5]: Sulphur content 2 wt.%, ash content approx. 9 wt.% and volatile matter content approx. 64 wt% all expressed on a dry basis. The mass loss (TG) and the derivative of the mass loss (DTG) curves obtained at 5, 10, 20 and 50 ºCmin-1 have been plotted in Figures 1 and 2. The DTG curves shifted slightly to higher degradation temperatures with increasing heating rates. No variation in char yield was observed with the increase in heating rate. The evolution of volatile matter that occurs below 250 °C is related to the decomposition of rubber additives [1,6]. 30.0
5 ºC/min 10 ºC/min 20 ºC/min
20.0
50 ºC/min
15.0 10.0 5.0 0.0 100
mass loss (%)
DTG(%/min)
25.0
5 ºC/min 10 ºC/min 20 ºC/min 50 ºC/min
100.0
80.0
60.0
40.0
200
300
400
500
600
700
Temperature, ºC
Figure 1. DTG curves of rubber from tyres.
20.0 0
200
400 600 Temperature, ºC
800
Figure 2. TG curves of rubber from tyres.
The reinforcing fibres were also subjected to pyrolysis at different heating rates to determine whether any interaction occurs between the different components. Figures 3 and 4 show the DTG curves for 5 and 50 ºCmin-1. In the graphs the curves corresponding to the fibre, the microfibre and the fibre/microfibre mixture have been plotted together. At higher heating rates the decomposition steps appear less resolved. Nevertheless it appears that the microfibre presents two main decomposition steps while the fibre presents only one coinciding with the second one of the microfibre. Both
2
Oviedo ICCS&T 2011. Extended Abstract
present a shoulder at high temperature. RF F MF RF calculated
DTG(% /m in)
4.0 3.0 2.0 1.0 0.0 100
RF F MF RF calculated
40.0 35.0 30.0 DTG(% /m in)
5.0
25.0 20.0 15.0 10.0 5.0
300 500 Temperature, ºC
700
Figure 3. DTG of textile, fibre and microfibre. -1
Heating rate 5 ºCmin .
0.0 100
300
500
700
Temperature, ºC
Figure 4. DTG of textile, fibre and microfibre. Heating rate 50 ºCmin-1.
Again, the lower the heating rate, the easier it is to see the decomposition steps. This is because there is no overlapping in the decomposition of the different polymers that make up the fibres. Table 1 shows the temperature corresponding to the main decomposition steps (Tmax) of the wastes at the different heating rates studied. From the results presented in Table 1 it can be seen that the char yield is independent of the heating rate. There is a slight difference between the fibres heated at 5 ºCmin-1 and those heated at 50 ºCmin-1. This difference could be a consequence of variations caused by the presence of tyre particles in the fibres under analysis. There are also differences between the char yield obtained from the reinforcing fibres as a whole and the yield calculated from fibres and microfibres yields. The experimental is always higher than the calculated ones. 3.2. Thermogravimetric analysis of blends of tyre wastes with coal and pitch. In order to increase the char yield and also to try to obtain chars with lower ash contents, blends of the two wastes with coal and pitch were pyrolyzed. Coal gives a char yield of around 65wt.% and the pitch a yield of around 31 wt.% at all the heating rates employed. The results obtained for the four blends were similar, and so we will concentrate on the results obtained from the blends with coal. Figures 5 and 6 show the DTG curves corresponding to the 1:1 blends of the two wastes with coal at 5°Cmin-1, the DTG of the blend calculated from that of the individual components has also been included. The coal presents only one devolatilization step with a temperature of maximum volatile matter evolution of around 450 °C. The rubber/coal blend shows the influence of both components since the DTG curve has one smaller peak first, at around 368ºC, and a
3
Oviedo ICCS&T 2011. Extended Abstract
higher mass loss step at 440ºC which probably corresponds to the second decomposition step of the rubber and the coal. The char yield obtained (48wt%) is similar to the calculated value (50wt%). The DTG curve calculated is almost identical to the experimental one. Table 1. Char yield and temperatures of maximum DTG.
Heating rate Yield (wt.%) (ºCmin-1) Rubber 5 35.1 10 34.3 20 34.3 50 35.8 Fibre 5 22.8 10 24.3 20 24.3 50 29.8 Microfibre 5 17.9 10 19.8 20 20.8 50 18.9 Reinforcing fibre 5 24.9 (20.5) 10 27.7 (22.2) 20 29.7 (22.7) 50 27.2 (24.8)
Peak Tmax 1 (ºC)
Tmax 2 (ºC)
T Shoulder
368 377 391 405
420 427 441 -
440-450
387 400 412 418
-
410-427 431-448 428-445 -
338 347 356 370
381 398 406 420
409-428 ~432 -
335 347 356 368
379 378 400 418
406-416 414-427 -
As in the case of the previous blend the DTG curve corresponding to the RF/coal blend (see Figure 6) presents the same decomposition steps as those of its components and at the same temperatures. This blend has three mass loss steps. The first and second peaks correspond to the decomposition of RF (335 and 379 °C) while the third peak of the blend reflects the influence of coal. The calculated DTG curve is very similar to the experimental one. In addition, the experimental and calculated char yields are very similar (44 wt% and 45 wt%).
4
Oviedo ICCS&T 2011. Extended Abstract
RF
TC
2.5
3.0
Coal DTG(%/min)
TC/Coal 1:1
1.5
TC/Coal 1:1 calculated
1.0
RF/Coal 1:1
2.0
RF/Coal 1:1 calculated
1.5 1.0 0.5
0.5 0.0 100
DTG(%/min)
2.0
Coal
2.5
300 500 Temperature, ºC
700
0.0 100
300
500
700
Temperature, ºC
Figure 5. DTG, blend of rubber and coal 5ºCmin-1 Figure 6. DTG, blend of textile and coal 5ºCmin-1
4. Conclusions Tyre wastes and their blends with coal and pitch have been pyrolyzed in a thermoanalyser to determine the influence of heating rate on the decomposition process. Increasing the heating rate produces a greater overlapping of the decomposition processes. No apparent influence of heating rate was observed on char yield. The decomposition of both wastes occurs in similar temperature ranges. The blends of both wastes with coal produce a greater char yield. In the case of the reinforcing fibres the pitch also produces an increase in char yield. Acknowledgement. The authors are grateful to MICINN for financial support (Project CTM2009-10227). References [1] Murillo R, Aylón E, Navarro M V, Callén M S, Aranda A, Mastral A M. The application of thermal processes to valorise waste tyre. Fuel Processing Technology 2006; 87: 143-147. [2] Kyari M, Cunliffe A, Williams P T. Characterization of oils, gases and char in relation to the pyrolysis of different brands of scrap automotive tires. Energy and Fuels 2005; 19:1165-1173. [3] Seidelt S, Müller-Hagedorn M, Bockhorn H. Description of tire pyrolysis by thermal degradation behaviour of main components. J. Anal. Appl. Pyrolysis 2006; 75: 11-18. [4] Parres F, Crespo-Amoros J.E. Characterization of fibers obtained from shredded tires. J. Applied Polymer Science 2009; 113: 2136-2142. [5] Fernandez AM, Barriocanal C, Diez MA, Alvarez R.Influence of additives of various origins on thermoplastic properties of coal. Fuel 2009; 88: 2365-2372. [6] Williams P T, Besler S. Pyrolysis-thermogravimetric analysis of tyres and tyre components. Fuel 1995; 74: 1277-1283.
5
Oviedo ICCS&T 2011. Extended Abstract
6
Oviedo ICCS&T 2011. Extended Abstract
Secondary Reactions of HCl during Coal Pyrolysis: Studies on Reactions of HCl with Model Carbons N. Tsubouchi1, N. Ohtaka2, A. Kawashima2, Y. Ohtsuka2 Center for Advanced Research of Energy and Materials, Hokkaido University, N13 W8, Kita-ku, Sapporo 060-8628, Japan 2 Institute of Multidisciplinary Research for Advanced Materials, Tohoku University, Katahira, Aoba-ku, Sendai 980-8577, Japan E-mail:
[email protected] 1
Abstract In order to elucidate secondary reactions of hydrogen chloride (HCl) during coal pyrolysis, a model carbon prepared from phenol resin is first O2-activated at 500ºC and then impregnated with CH3COOK or (CH3COO)2Ca·H2O. When the sample is exposed to a stream of 100 ppmv HCl/N2 at 500ºC, HCl can readily react with all carbon samples in order to produce surface chlorine species. The extent of the reaction increases almost linearly as the number of carbon active sites determined by the temperature-programmed desorption (TPD) up to 500ºC increases, and it is the largest with the Ca-doped carbon. The X-ray photoelectron spectroscopy (XPS) measurements of the HCl-treated samples exhibit the distinct Cl 2p spectra, which can be identified to organic chlorine with pure carbon and to organic and inorganic chlorine with the K- and Ca-doped carbons. These observations suggest that HCl evolved during coal pyrolysis may react secondarily with carbon active sites and mineral components in the nascent char in order to form organic and inorganic chlorine. The formation mechnisms of organic C-Cl forms are discussed in term of reactions of HCl and carbon active sites.
1. Introduction The chlorine present in coal (coal-Cl) is usually in the range of 50-5000 μg/g-coal and emitted mainly as HCl during coal conversion processes, such as pyrolysis, combustion and gasification [1]. As well known, the HCl affects the speciation and vaporization of trace elements, such as mercury and alkali metals, in these processes [2], and it causes corrosion problems on gas turbine materials in IGCC and deterioration of fuel cell performance in IGFC [3]. It is thus important to understand chlorine chemistry for developing super clean coal technologies that should be targeted on zero emissions. However, only quite limited information on this topic has been provided so far [4, 5].
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The present authors’ group has been working on investigating the behavior of the release and retention of chlorine and fluorine during coal pyrolysis and subsequent char gasification [6-9], and on examining the functional forms of the halogens and carbon in pyrolysis chars [6, 7, 9], gasification residues [8] and dust particles released in a sintering process of iron ores [10, 11]. We have recently suggested that HCl evolved above 450ºC during temperature-programmed pyrolysis of many coals originates from inorganic and organic chlorides produced via secondary reactions of the devolatilizing chars with HCl once-released at a lower temperature [7, 9], and that O-functional groups (e.g. carboxyl groups and lactone/acid anhydride groups) on carbon surface play important roles in the formation of covalent C-Cl bonds [7-13]. These observations indicate that the reactions between HCl and chars are one of the key factors determining the release of HCl during pyrolysis and the fate of char-Cl during combustion and gasification. In the present paper, therefore, we first examine the dynamic behavior of HCl in the reaction process of HCl with model carbons produced from phenol resin, then make clear the chemical forms of HCl reacted with model carbons, and evaluate the role of surface O-species in the formation of organic C-Cl forms on the carbon surface.
2. Experimental Section 2.1. Carbon Samples.
Pure carbon without any minerals was prepared by carbonizing
powdery phenol resin in He at 950ºC for 60 min, followed by activation in 20%O2/He at 500ºC for 0-150 min. The activated carbon (AC) samples prepared from phenol resin are expressed according to the corresponding activation time as AC-0, AC-50, AC-100 and AC-150 throughout this paper. The C, H, N and O contents in AC samples with size fraction of 300-420 μm were 66-91, 1.4-2.2, 1.2-1.6 and 6.7-31 wt%-dry, respectively, the BET surface area analyzed by the N2 adsorption method being 5-790 m2/g. Alkali metal and alkaline earth metal ions, such as K+ and Ca2+, were loaded on AC-100 by the impregnation method using an ethanol solution of CH3COOK or (CH3COO)2Ca; these were denoted as K/AC-100 or Ca/AC100, respectively. The impregnation was made for 1 h at room temperature, followed by evaporation of ethanol under vacuum at 50ºC. Metal loading was 0.5 mass %. 2.2. HCl Treatment of Carbonaceous Materials.
Reactions of HCl with model
carbons were made with a cylindrical flow-type quartz reactor, which was heated with a glass-made transparent electric furnace. The details of the apparatus and procedure have
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been reported elsewhere [11]. About 500 mg of the carbon sample was first charged into the reactor, and special care was then taken to ensure that the whole reaction system was free from any leakage. After such prudent precautions, the reactor was heated at 5ºC/min up to 500ºC in pure He, and the resulting sample was immediately exposed to a flow of 100ppmHCl diluted with pure N2 at 500ºC. The change in HCl concentration in this HCltreatment process was analyzed at intervals of 1 min with an infrared analyzer. After 30 min exposure, the gas atmosphere was switched to pure N2, and the HCl-treated sample was quenched to ambient temperature. 2.3. TPD Runs.
To evaluate quantitatively carbon active sites on the surface of the
model carbons, the samples used were subjected to the temperature-programmed desorption (TPD) measurements. In the experiment, each carbon sample was heated at 2ºC/min up to 500ºC in a flow of pure He, and the concentration of CO or CO2 evolved was determined online at intervals of 2.5 min with a high-speed micro gas chromatograph. Carbon active sites were calculated based on total amount of CO and CO2 evolved up to 500ºC. No appreciable amounts of CH4 and C2 hydrocarbons were detectable in any cases. 2.4. XPS Measurements.
The X-ray photoelectron spectroscopy (XPS) analyses were
made with a non-monochromatic Mg-Kα X-ray source operating at 300W to investigate the functionalities of the chlorine present in carbon samples after the HCl treatment. The binding energies observed were referred to an In 3d5/2 peak of In2O3 at 444.9 eV. The analytical conditions have been described in detail earlier [12]. Least-squares curve-fitting analysis of Cl 2p spectra was carried out using Gaussian peak shapes. Upon deconvolution, the binding energy of each peak was fixed within ± 0.1 eV, and the full width at half maximum (fwhm) value was assumed to be 1.6 ± 0.1 eV.
3. Results and Discussion 3.1. Reactions of HCl with Carbonaceous Materials.
Figure 1 illustrates the time
changes in HCl concentration when AC samples without any metal ions are exposed to a stream of 100ppmHCl/N2 at 500ºC. The concentration decreased steeply with increasing time up to 4-6 min and lowered up to 95, 80, 65, 50 ppm for AC-0, 50, 100, 150, respecttively, and then approached to the original level. When K/AC-100 and Ca/AC-100 were treated with 100ppmHCl, the concentration lowered up to 55 ppm after 6 min for K/AC100 and 50 ppm after 7 min for Ca/AC-100, and then approached gradually to the initial
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level. It is evident that HCl can react with these carbon samples at 500ºC, and the extent of the reaction increases with increasing activation time and by metal doping. Total amount of HCl decreased (HClDecrease) during 30 min exposure and surface area (SBET) of each carbon sample are shown in Table 1, where the amount is calculated by integrating each curve provided in Figure 1. The HClDecrease increased with increasing activation time, and it was the largest with the Ca-doped carbon. As seen in Table 1, no distinct relationship between the HClDecrease and SBET could be observed. This result may indicate that SBET is not important as the factor for determining the extent of the reaction of HCl and carbonaceous materials.
Table 1.
Total Amount of HCl Decreased in the HCl Treatment Process
and Surface Area of Each Carbon Sample Used Carbon sample
HClDecreasea, b (μmol/g-sample)
SBETc (m2/g)
AC-0
2.5
5
AC-50
9.1
400
AC-100
12
650
AC-150
19
790
K/AC-100
16
600
Ca/AC-100
33
600
a
Total amount of HCl decreased during 30 min exposure.
b
Average value of the repeated experiments.
c
BET surface area.
3.2. Determination of Carbon Active Sites.
To evaluate quantitatively carbon active
sites formed from the present model carbon in the temperature region of HCl treatment, the samples used were heated up to 500ºC at a slow heating rate of 2 ºC/min. The results are given in Table 2. Total amounts of CO or CO2 from the four carbon samples without any metal ions were in the range of 30-200 or 60-140 µmol/g-sample, respectively, and these values increased with increasing activation time. As shown in Table 2, the K or Ca added to AC-100 promoted the formation of CO and CO2. This result may be reasonable, because it has been well-accepted that metal cations present as ion-exchanged forms on
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carbon surface promote the devolatilization reaction upon pyrolysis to change the distribution of volatile products [14]. Table 2.
Cumulative Amounts of CO and CO2 Evolved in the Temperature-
Programmed Desorption of Carbon Samples Used COa (μmol/g-sample)
CO2a (μmol/g-sample)
AC-0
30
60
AC-50
120
70
AC-100
160
80
AC-150
200
140
K/AC-100
180
100
Ca/AC-100
200
220
Carbon sample
a
Average value of the repeated experiments. Figure 2 illustrates the HClDecrease value (Table 1) as a function of the number of
carbon active sites (Table 2) determined by the present TPD method. Interestingly, there was an almost linear correlation between the two. According to earlier studies on surface functionalities of coal chars and carbons using the TPD technique [15, 16], CO and CO2 evolved at ≤ 500ºC arise mainly from decomposition reactions of acid anhydride groups and carboxyl groups, respectively. It is likely that the active sites from the O-functional groups play an important role in the reaction of HCl and carbonaceous materials. 3.3. XPS Spectra for HCl-Treated Carbon Samples.
To investigate the chemical
forms of the chlorine present in each carbon sample after HCl treatment, the Cl 2p XPS measurements were performed. Atomic Cl/C ratios measured by the XPS were 2.1×10-37.3×10-3, which were approximately 3-4 times those (0.61×10-3-2.0×10-3) determined by elemental analysis. This result means that most of the HCl reacted during HCl treatment is enriched on the surfaces of the carbon samples. Each Cl 2p spectrum observed was broad in the binding energy range of 196-204 eV, and it showed the main peak at 199-200 eV. It has been widely accepted that the Cl 2p3/2 spectra of inorganic chlorides (e.g. KCl and CaCl2) and chlorobenzenes are present at 198-200 and 200-201 eV, respectively [17]. On the basis of this information, all the Cl 2p spectra observed were deconvoluted into inorganic and organic chlorides by using the binding energy of each model compound mentioned above. The proportion of inorganic
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or organic chlorine was estimated to be 15-48 or 52-85 mol%. Since it has been reported that pyridinic nitrogen and aromatic amines are formed by the reaction of NH3 gas with activated carbon materials under the similar conditions as in the present work [18, 19], it is likely that HCl also reacts with carbon active sites to form covalent C-Cl bonds in the condensed aromatic structures. On the basis of the above-mentioned results and discussion, it may be reasonable to see that HCl evolved at a lower temperature in the temperature-programmed pyrolysis of coal can readily undergoes secondary reactions with mineral components and carbon active sites in the nascent char to form inorganic and organic chlorides. In the latter case, the HCl may react with the active sites derived from carboxyl and acid anhydride groups on coal surface to form Cl-containing intermediates, for example, HCl chemisorbed and/ or surface HCl species, which might subsequently be converted to organic chlorine. The role of K or Ca in the formation of covalent C-Cl bonds is not clear at present. To make clear this point should be the subject of future work.
4. Conclusions To investigate the possibility of secondary reactions of HCl with char during coal pyrolysis, a model carbon produced from phenol resin is O2-activated, followed by K- or Ca-doping, and exposed to a stream of 100 ppmv HCl/N2 at 500ºC. The HCl can readily react with all of these carbon samples. The extent of the reaction increases almost linearly as the number of carbon active sites determined by TPD up to 500ºC increases, and it increases by metal doping. The XPS analyses show the distinct Cl 2p spectra, which can be identified mainly as organic C-Cl forms.
Acknowledgement.
This work was supported in part by a Grant-in-Aid for Young
Scientists (A) from the Ministry of Education, Science, Sports and Culture, Japan.
References [1] Davidson RM. In: Chlorine and other halogens in coal. IEAPER/28, London: IEA Coal Research; 1996.
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Oviedo ICCS&T 2011. Extended Abstract
[2] Galbreath KC, Zygarlicke CJ. Mercury speciation in coal combustion and gasification flue gases. Environ Sci Technol 1996; 30: 2421-6. [3] Mitchell SC. In: Hot gas cleanup of sulfur, nitrogen, minor and trace elements. IEACCC/12, London: IEA Coal Research; 1998. [4] Herod AA, Hodges NJ, Pritchard E, Smith CA. Mass spectrometric study of the release of HCl and other volatiles from coals during mild heat treatment. Fuel 1983; 62: 1331-6. [5] Shao D, Hutchinson EJ, Cao H, Pan WP, Chou CL. Behavior of chlorine during coal pyrolysis. Energy Fuels 1994; 8: 399-401. [6] Tsubouchi N, Ohtsuka S, Hashimoto H, Ohtsuka Y. Several distinct types of HCl evolution during temperature programmed pyrolysis of high-rank coals with almost the same carbon contents. Energy Fuels 2004; 18: 1605-6. [7] Tsubouchi N, Ohtsuka S, Nakazato Y, Ohtsuka Y. Formation of hydrogen chloride during temperature programmed pyrolysis of coals with different ranks. Energy Fuels 2005;19: 554-60. [8] Takeda M, Ueda A, Hashimoto H, Yamada T, Suzuki N, Sato M, Tsubouchi N, Nakazato Y, Ohtsuka Y. Fate of the chlorine and fluorine in a sub-bituminous coal during pyrolysis and gasification. Fuel 2006; 85: 235-42. [9] Tsubouchi N, Saito T, Sato M, Suzuki N, Ohtsuka Y. Chlorine functional forms of Argonne premium coal samples and the changes with water washing and in slow heating rate pyrolysis. Am Chem Soc Div Fuel Chem Prepr 2007; 52: 78-9. [10] Tsubouchi N, Kasai E, Kawamoto K, Noda H, Nakazato Y, Ohtsuka Y. Functional forms of carbon and chlorine in dust samples formed in the sintering process of iron ores. Tetsu To Hagane - J Iron Steel Inst Jpn 2005; 91: 751-6. [11] Tsubouchi N, Kuzuhara S, Kasai E, Hashimoto H, Ohtsuka Y. Properties of dust particles sampled from windboxes of an iron ore sintering plant: surface structures of unburned carbon. ISIJ Int 2006; 46: 1020-6. [12] Tsubouchi N, Hashimoto H, Ohtaka N, Ohtsuka Y. Chemical characterization of dust particles recovered from bag filters of electric arc furnaces for steelmaking: some factors influencing the formation of hexachlorobenzene. J Hazard Mater 2010; 183: 116-24. [13] Tsubouchi N, Hayashi H, Kawashima A, Sato M, Suzuki N, Ohtsuka Y. Chemical forms of the fluorine and carbon in fly ashes recovered from electrostatic precipita-
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Oviedo ICCS&T 2011. Extended Abstract
tors of pulverized coal-fired plants. Fuel 2011; 90: 376-83. [14] Tyler RJ, Schafer HNS. Flash pyrolysis of coals: influence of cations on the devolatilization behavior of brown coals. Fuel 1980; 59: 487-94. [15] Zhuang Q-L, Kyotani T, Tomita A. DRIFT and TK/TPD analyses of surface oxygen complexes formed during carbon gasification. Energy Fuels 1994; 8: 714–8. [16] Zielke U, Huttinger KJ, Hoffman WP. Surface-oxidized carbon fibers: I. Surface structure and chemistry. Carbon 1996; 34: 983–98. [17] Wagner CD, Naumkin AV, Kraut-Vass A, Allison JW, Powell CJ, Rumble Jr JR. NIST X-ray photoelectron spectroscopy database; NIST standard reference database 20, version 3.5 (Web version). Gaithersburg: NIST; 2007. [18] Stöhr B, Boehm HP, Schlögl R. Enhancement of the catalytic activity of activated carbons in oxidation reactions by thermal treatment with ammonia or hydrogen cyanide and observation of a superoxide species as a possible intermediate. Carbon 1991; 29: 707-20. [19] Mangun CL, Benak KR, Economy J, Foster KL. Surface chemistry, pore sizes and adsorption properties of activated carbon fibers and precursors treated with ammonia. Carbon 2001; 39: 1809-20.
AC-0
HCl concentration, ppm
100
AC-100 AC-150
60
40 0
Figure 1.
AC-50
80
0
5
10 15 20 Time on stream, min
25
30
Time changes in HCl concentration during HCl treatment of AC samples
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without any metal cations.
40
HCl reacted, µmol/g-sample
Ca/AC-100 30
20
AC-150 K/AC-100 AC-100
10
AC-50 AC-0
0
Figure 2.
0
100 200 300 400 500 Carbon active sites, µmol/g-sample
Relationship between total amounts of HCl decreased in the HCl treat-
ment process and carbon active sites in carbon samples.
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Solvent modulation in getting high purity of anthracene and carbazole from crude anthracene Mingming Fan, Cuiping Ye, Huan Zheng, Tingting Wu, Jie Feng, Wenying Li* Key Laboratory of Coal Sci. Tech. (Taiyuan University of Technology) Ministry of Education and Shanxi Province, Taiyuan, 030024, China * Corresponding Author, E-mail:
[email protected] Abstract： ： As the raw material of dyes, pharmaceuticals, photoelectric materials etc, high purity (>98wt%) of anthracene and carbazole are required. Most of anthracene and carbazole are separated from coal tar now. And it is difficult to acquire high purity and high recovery rate. In this study, we developed a separation method to get the higher purity of anthracene and carbazole from crude anthracene: remove phenanthrene with xylene firstly, then refine anthracene and carbazole with solvent I (added 10% additive F) by using liquefying crystal. The result showed that with this new solvent separation method the recovery rate of anthracene and carbazole purity higher than 90wt% and 97wt% would be at more than 82% and 73% respectively in one pass. Besides, the developed solvent can be easily recycled, met with environmental requirements and is featured with products of good quality. Key words: coal tar; anthracene; carbazole; solvent; separation 1 Introduction Anthracene(An), phenanthrene(Phe) and carbazole(Car) are the important nonrenewable chemical resources for dyes, pharmaceuticals, photoelectric materials etc[1-3]. And their demands increased yearly. For fine chemicals manufacture, anthracene/carbazole with a purity >98% is required. Although anthracene and carbazole could be synthesized by chemical method now[4], they were separated and purified from crude anthracene of coal tar yet due to the cost and waste issues. So many researchers become concerned about the purity and recovery rate of products. And it is not an easy task to separate and collect them with high purity because the amounts of the main components of crude anthracene varied with the coals and coal processing, the purity of anthracene and carbazole could only reach about 97% at present. Most of the method used for getting high purity products must combine with solvent method. The key of solvent method is the selection for available and suitable solvent used in the processing. But so far, there are very few organic solvents used, such as benzole, dimethylformamide(DMF), pyridine and so forth, and a
few potential solvents were not reported. Furthermore, the essential difference between those solvents is lack of theoretical study. Here, we developed a separation method to get the higher purity of anthracene and carbazole from crude anthracene: remove phenanthrene firstly, then refine anthracene and carbazole by using liquefying crystal. The factors of having strong influences on the purity and recovery rate of products, such as solvent types, ratios of solvent, temperature of dissolution and crystallization were carefully investigated. 2 Experimental Refining of crude anthracene Crude anthracene used in this experiment is offered by One Steel factory, China (contained anthracene 40.94wt%, phenanthrene 21.01wt%, carbazole 18.94wt%). The solvent used here are acetophenone, DMF, xylene, methylene chloride (analytical grade) 200# solvent oil and oxygen containing solvent I. Add a certain amount of crude anthracene and solvent in flask equipped with blender, reflux condenser pipe and thermometer tubes with mixer, stir and heat until the solids are completely dissolved, heat up to a certain temperature and keep for 30min, then filter at about 40◦C under vacuum, and the cake was washed by solvent and dried under room temperature, and used for raw material of anthracene and carbazole. The filtrate was placed in a rotary evaporator (Buchi R-200, Switzerland) under vacuum until completely dry. And the residue used as raw material of phenanthrene. Gas chromatographic analysis. GC was performed on a Shellomadzu GC 2014 (Japan) with a Rtx®-5 (USA) capillary column (0.32mmх30m), wall-coated with 5% dipheny/95% dimethyl polysiloxane, film thickness 0.25µm. Using hydrogen flame ionization detector (FID). The temperature program was: 170◦C (0.4µL injection volume), keeping for 1min, then 3.5◦C/min to 200 ◦C, holding for 1min, then 10◦C/min to 250◦C. The internal standard curves method with fluorenone as internal standard substance is used for quantity. The chromatogram of the standards and the calibration curves is shown in Figure 1 and Figure 2, respectively.
200000
10.528 / Carbazole
50000
9.738 / Anthracene
100000
9.563 / Phenanthrene
8.780 / 9-Fluorenone
Intensity
150000
0 8
9
10
11
min
Figure 1 Gas chromatogram of the standard of anthracene, phenanthrene, carbazole and fluorenone
3.0
2.5
Anthracene Phenathrene Carbazole
Ai/As
2.0
1.5
1.0
yAn=1.6910x+0.0211 yPhe=1.9515x-0.0240 yCar=1.6206x-0.0321
0.5
0.0 0.0
0.2
0.4
0.6
0.8
1.0
R=0.99804 R=0.99918 R=0.99909 1.2
1.4
1.6
c ( mg/mL)
Figure 2 Calibration curves of anthracene, phenanthrene and carbazole 3 Results and Discussion 3.1 Choice of solvents By liquefying crystal, the separation and refining of anthracene, carbazole and phenanthrene is according to different solubility and volatility in different solvent. Phenanthrene has higher
solubility in most solvent, and it is easier to remove. And the solubility of anthracene is very low in almost all kinds of solvent; the solubility of carbazole is high in nirogen and oxygen cotaining solvents, such as DMF and N-methyl-2-pyrrolidone(NMP), but low in benzene solvent, so the separation of anthracene and carbazole depends on the solubility and selectivity. The purpose of using liquefying crystal is to remove the phenanthrene, fluorene and oil impurities to gather carbazole and anthracene components, so the problem is how to improve the solvent selectivity and dissolution of carbazole and anthracene. And the selectivity and solubility of the solvent could be reflected by the phase diagram and solubility curve. Carbazole o
A
0.0
45 C o 60 C o 70 C
1.0
0.2
0.8
0.4
0.6
0.6
0.4
ⅣⅣⅣⅣ ⅡⅡⅡⅡ
0.8
0.2 ⅠⅠⅠⅠ
1.0
DMF 0.0
0.0
0.2
0.4
0.6
0.8
1.0
Anthracene
Note: I- Crystallized area of anthracene, II –Crystallized area both of anthracene and carbazole, III- Crystallized area of carbazole, IV -Homogeneous area
Carbazole 0.94
B 0.95
o
30 C o 60 C
0.06 0.05
0.96
0.04
0.97
0.03
0.98
0.02
0.99
0.01
1.00
0.00
Xylene 0.00
0.01
0.02
0.03
0.04
0.05
Anthracene
0.06
Carbazole C
0.88
30oC 60oC
0.12
0.90
0.10
0.92
0.08
0.94
0.06
0.96
0.04
0.98
0.02
1.00 0.00
0.00 0.02
0.04
0.06
0.08
0.10
0.12
Acetophenone
Anthrancene
Fgure 3 Phase diagram of the ternary anthracene-carbazole-DMF(A)/ Xylene(B)/ Acetophenone(C) system In the ternary anthracene-carbazole-DMF system (Figure 3A), the crystallized area is large for both of anthracene and carbazole, especially for anthracene. And the selectivity of carbazole is higher than anthracene in acetophenone (Figure 3C). So DMF and acetophenone are relatively appropriate solvents to get high pure anthracene.
In the ternary anthracene-carbazole-xylene system (Figure 3B), the crystallized area is very small for both of anthracene and carbazole. Therefore, when using xylene, the necessary solvent ratio is very big. So xylene is not the ideal solvent for refining of anthracene. However, the crystallized area of carbazole is larger than that of anthracene, it could be used to refine carbazole. 3.2 Influence of the ratios of solvent for enrichment of anthracene and carbazole. The ratios of solvent affect the purity and recovery rate of products during processing[5], so we investigated the ratios of solvent from 0.5 to 3.0. The study used crude anthracene 50g, with xylene as solvent, dissolving at 100oC and crystallizing at 40oC, respectively. The results were presented in Table 1. With the ratios of solvent increasing, the purity of the products increased too. However, the recovery rate was maximum at 1.5. And it has a satisfactory result with xylene as solvent to remove phenanthrene in the first step, the contents of dried cake are as following: anthracene 57.72wt%, carbazole 27.92wt% and phenanthrene 14.33wt% in one pass. 3. 3 The selectivity of solvent I. The experiment had compared the refining effect of different solvent: DMF, acetophenone and solvent I (oxygen containing solvent). The results showed that (Table 2), when using solvent I, the purity of anthracene is slightly lower than using acetophenone and DMF, while the recovery rate is the highest. So we select solvent I for further study. Although the solvent I has good selectivity, the recovery rate of anthracene is still lower, about 69%. To make full use of anthracene resources and to improve the recovery rate of anthracene, amine additive (F) is introduced. And the addition proportion was investigated under two temperature of 70oC and 100oC. The results are as follows (Table 3). From Table 3, we can find when additive F is increased from 0%(V/V) to 5%(V/V), the recovery rate is increased about 10%, and the purity of anthracene is also increasing. The reason is probably that additive F can restrict the solubility of anthracene and increase the solubility of carbazole in solvent I [6]. Table 1 Influence of ratios of solvent on the purity and recovery rate of anthracene and carbazole Ratios of Solvent 0.5 1.0 1.5 2.0 3.0
An 45.32 53.56 57.72 59.65 63.15
Content/wt% Car 21.11 24.62 27.95 28.48 30.08
Phe 33.57 21.82 14.33 11.87 6.77
Recovery Rate /% An Car 79.15 75.78 88.85 83.95 92.22 91.82 80.67 79.17 70.06 68.60
Table 2 Experimental result of solvent I Heating temperature/◦C
Solvent Solvent I Acetophenone DMF
130 140 130
Recovery rate of An /%
Content/wt% An 84.56 86.42 85.64
Car 7.32 6.56 6.21
Phe 8.12 7.02 8.15
69.58 56.14 52.68
Table 3 The influence of addition of F in solvent I on refining anthracene F/% 0 5 10 0 5 10
Temperature/oC 70 70 70 100 100 100
Content/wt% 84.56 91.12 93.48 83.42 90.86 92.28
Recovery rate of An/% 69.58 80.42 83.86 68.32 79.84 82.72
Table 4 Refined effects of 200# solvent oil, methylene chloride and xylene for carbazole
No. of refining 1 2 3 Total
200# solvent oil Methylene chloride Recovery Recovery Content/wt% Content/wt% rate/% rate/% 53.10 89.00 71.40 52.90 88.90 86.50 80.60 78.60 90.80 88.40 87.50 86.70 62.49 33.10
Xylene Content/wt% 55.50 89.20 97.20
Recovery rate/% 96.70 91.90 89.20 72.78
3.4 Refining of carbazole. We used xylene to refine carbazole. The solvent ratio is 1.5 for the first pass, and the ratio increased to 2 for the second and third pass, with heating and filtering temperature at 40 oC, and the results are as follows (Table 4). The results suggested that the purity of carbazole is increased with the refining numbers. Meanwhile, the loss of carbazole is increased either. So the number of washing pass cannot be too many. When using xylene as the solvent, the purity of carbazole can reach 97.20% and total recovery rate is 72.78% after being washed three times. 4. Conclusions With the new solvent separation method: remove phenanthrene with xylene firstly, then refine anthracene and carbazole by solvent I with 10% additvie F using liquefying crystal, the recovery rate of anthracene and carbazole purity higher than 90wt% and 97wt% would be at
more than 82% and 73% respectively in one pass. And this new solvent system could improve the selectivity of anthracene and carbazole and changed their liquid-liquid partition. Besides, the developed solvent can be easily recycled, met with environmental requirements and is featured with products of good quality. Acknowledgment This work was supported by the Taiyuan Scientific and Technological project(11014907) and Jiangdu Refining Chemical Company. References [1] Zhang YD, Wada T, Sasabe H. J. Mater. Chem. 1998, 8(4), 809-828 [2] Grazulevicius JV, Strohriegl P, Pielichowski J, Pielichowski K. Prog. Polym. Sci. 2003, 28, 1297-1353 [3] Wang ZQ,Zheng CJ, Liu H,Ou XM, Zhang XH. Science China(Chemistry) 2011, 54(4), 666-670 [4] Vlčko M, Cvengrošová Z, Cibulková Z, Hronec M. Chinese Journal of Catalysis 2010, 31(12), 1439-1444 [5] Song YX, Hu ZQ, Gao JS. Coal Coversion 1999, 22(2), 94-96(In Chinese) [6] Gao JS, Zhou XP, Wang ZH, Hang YZ. Shanghai Chemical Industry 1994, 19(6), 1922(In Chinese)
STUDY OF COAL, CHAR AND COKE FINES STRUCTURS AND THEIR
PROPORTIONS
IN
THE
OFF-GAS
BLAST
FURNACE
SAMPLES BY X-RAY DIFFRACTION 1
A.S. Machado*, 2A.S. Mexias, 1A.C.F. Vilela, 1E. Osório.
1
Iron and Steelmaking Laboratory, UFRGS, Porto Alegre, Brazil.
2
X-Ray Diffractometry Laboratory, UFRGS, Porto Alegre, Brazil.
*
[email protected] Abstract Four dust samples were collected in the BF off-gas for this investigation, two at all coke and two at Pulverized Coal Injection (PCI) operations. The atomic structure of raw coke, chars and their parent coals used in PCI were investigated. This study has showed that the XRD technique could be used as a standard procedure to identify the char and coke structures. This technique was used to quantify the fines proportions of these carbonaceous materials in the BF flue dust samples. It was also showed the unexpected presence of coke fines in the flue dust fractions smaller than 63μm. The quantification of the carbonaceous material content in the BF off-gas samples could be used to improve the PCI performance in operating BF. Keywords: char, coke fines, X-ray diffraction, pulverized coal injection, Blast Furnace, graphitization.
1. Introduction The Blast Furnace (BF) is the most important technology for production of iron for steelmaking [1]. In the last decades, the PCI rates have increased in the most of BFs, achieving an injection rate between 150 — 220 kg/thm [2]. One of the problems during the PCI operation in BF, especially at high injection levels, is the increase in the formation of unburnt char, which could accumulate in the near raceway region. This material could harm the burden permeability. Higher char levels in the BF dust samples are related to insufficient coal combustibility and to an unstable furnace operation [3]. The quantification of the carbonaceous content in these samples could be used in the coal selection and to improve the PCI performance [3,4].
The BF off gas solid samples (flue dust and sludge) basically contain metallic oxides and carbonaceous materials. The carbon in BF dust samples is originated from coke fines, char and some unconsumed pulverized coal [4,5]. Chemical analyses in the BF dust samples do not reveal some carbonaceous material. Optical microscopy is used to study the off gas solid samples [5,6], but it can lead to some ambiguous results, since char mixed with ash are very fine and lack unique morphological features [3]. X-ray diffraction (XRD) technique has been utilized to determine the crystallite sizes (Lc, La, etc.) in carbonaceous materials [7]. Since the coke structure is more ordered than the char structure, it would be possible to quantify the proportion of these materials in the BF off-gas samples by chemical analysis in combination with XRD [3]. The aims of this work are to study coal, char and coke structures by XRD and to quantify carbonaceous components (char and coke fines) in the flue dust BF samples at all coke and PCI BF operations through the use of the XRD technique and ultimate analysis.
2. Experimental section The methodology conducted in this study is composed of: (1) chemical and granulometry analyses; (2) production of chars in a raceway simulator; (3) the chemical treatment used to demineralize the samples; (4) XRD and the mathematical characterization of their profiles to obtain Lc (average stacking height) values.
2.1 Raw and carbonaceous materials characterization Four flue dust samples collected in the BF off-gas (FD-AC1, FD-AC2, FD-AP1 and FD-AP2), three coal samples (CA, CB and CAB) and their laboratory produced chars (ChA, ChB e ChAB) and one metallurgical coke sample (CK), obtained from a Brazilian company, were selected for this work. The char ChAB and the coke fines of the 250-425μm fraction of the flue dusts (CKFD-AC1, CKFD-AC2, CKFD-AP1 and CKFD-AP2) were utilized as a standard for the quantification of the percentage of char and coke fines in the BF flue dust [4]. The char samples were made in a laboratory rig that simulates the behavior of fines injected into the raceway. The coal CAB is a mixture of coals CA and CB utilized in PCI. The flue dust samples FD-AC1, FD-AC2 were obtained from all coke BF operations and samples FD-AP1, FD-AP2 from PCI BF operations. Chemical analyses of the coal and coke samples were carried out. A granulometric analysis of coal CAB was conducted. The dust samples were mechanically separated in seven
different sizes ranging from <38μm to >425μm [4]. The separation was conducted to verify how the carbonaceous particles are distributed in the dust and to identify the possible fractions composed by char or coke. All coals, coke and flue dusts fractions were demineralized to avoid the effect of mineral matter on the quantitative analyses by XRD.
2.2 X-ray diffraction – carbonaceous structure characterization XRD analyses were carried out on an X-ray Powder Diffractometer using CuKα over an angular range of 5–115° (0.05° x 4s). The XRD traces were subjected to a series of mathematical treatments and the Lc value was determined according to the Scherrer equation [9]. The carbonaceous average stacking height value (Lc or L002) has been used to characterize the dimension of the crystalline carbon (crystallites) in all samples. Since the coke structure is more ordered than the char structure, the percentage of these materials in the BF flue dust samples were quantified on the basis of a suitable calibration using char and coke standards combining XRD and carbon elementary analyses [3,4].
3. Results and discussion
3.1 Samples characterization Chemical analyzes of the coal and coke samples are summarized in Table 1. The granulometric analyses of the coal CAB showed practically no particles above 150μm, and its char particles will probably not exhibit sizes above 250μm as described in literature [8]. Table 1 — Summary of chemical analyzes of the selected coals and coke
Item Proximate: (dry base), %
Ultimate: (dry base), %
CA
CB
CAB
Coke
Volatile matter
24,46
14,42
19,04
1,21
Fixed carbon
66,11
74,24
70,22
89,18
Ash
9,43
11,34
10,74
9,61
Carbon
75,92
76,84
76,82
90,17
Hydrogen
6,26
5,60
5,69
0,11
Nitrogen
1,72
1,64
1,70
1,39
Sulfur
0,53
0,71
0,60
0,61
Oxygen
6,14
3,87
4,45
—
Carbon content and inorganic matter content of the selected flue dusts are provided in Table 2. BF flue dusts basically contain metallic oxides and carbonaceous materials. Carbon in BF dust samples originated from coke, char and some unconsumed PCI coal [4,5]. Table 2 – Carbon and inorganic matter contents of the flue dusts FD: AC1, AC2, AP1 and AP2
BF — all coke
Item
BF — PCI
FD-AC1
FD-AC2
FD-AP1
FD-AP2
Carbon (dry base), %
38,91
40,15
43,57
36,66
Inorganic matter (dry base), %
59,97
55,23
57,06
61,60
Figure 2 compares the distribution of the carbon content in each size range of dust samples FD-AC1, FD-AC2, FD-AP1 and FD-AP2. The proportion of carbon fines in the dust samples less than 63μm is insignificant, smaller than 1%. According to Gupta et al. [4], the high carbon content found in size ranges greater than 90μm are associated to a higher generation of coke fines in these samples. The authors considered that the size range greater than 250μm is composed exclusively by coke fines.
Figure 1 – Weight of carbon in seven size ranges of dust samples FD-AC1, FD-AC2, FD-AP1 and FD-AP2
3.2 X-ray diffraction — coal-char and coke fines quantification Figure 2a shows the reduced intensity of XRD profile of coal CAB and its char ChAB. The L002 of char ChAB was higher than its parent coal CAB indicating slightly higher crystalline order of carbon. Figure 2b shows the reduced intensity without the XRD
background of the standards ChAB, CK, CKFD-AC1 and CKFD-AP1. It is clear in the figure that the structure of coke samples are more ordered than char structure (narrow peak Æ bigger L002). The coke fines in the flue dust size fraction 250–425μm is the most graphitized fraction in the dust (Tab. 3), and according to Sahajwalla et al. [8] this size fraction is a better representative of the residual coke fines (coke standard) in the BF flue dust.
(a)
(b)
Figure 2 – a) Comparison of reduced intensity of XRD pattern of the CAB and ChAB; b) Comparison of X-ray spectra and profile of 002 carbon peaks for the standards ChAB, CK, CKFD-AC1 and CKFD-AP1
The XRD spectra and profile of (002) carbon peaks in seven size ranges of dusts samples FD-AC1 and FD-AP1 are shown in Figure 3. The L002 values of each size range of FD-AP1 are indicated in Table 3. Figure 3 shows a small but perceptive (002) carbon peak in the flue dust fractions smaller than 63μm. Demineralization process allowed to observe and to calculate the L002 values of these small fractions. The L002 of the small fractions (<63μm) was greater than the L002 of char and CK, showing the unexpected [8] presence of coke fines in this fractions. The profiles of the (002) carbon peaks in the fractions of FD-AC1 showed little variation in L002, as expected of samples composed only by coke fines. The profiles of the (002) carbon peaks in the fractions of FD-AP1 showed considerable variation of L002, as expected of samples composed by char and coke fines.
(a)
(b)
Figure 3 – X-ray spectra and profile of (002) carbon peaks in seven size ranges of dusts samples (a) FD-AC1 and (b) FD-AP1
The proportion of char level in FD-AP1 size ranged fractions is provided in Table 3. The percentage of coal-char and coke fines were calculated on the basis of a suitable calibration using standards and combining XRD and carbon elementary analyses. The fractions >250µm were considered to be composed only by coke fines [8]. Table 3 – Char proportion in dust sample FD-AP1
FD-AP1 Dust particle size
L002
Total Carbon
PCI carbon —
Coke carbon
(µm)
(nm)
(wt, %)
Char (wt, %)
(wt, %)
< 38
2,55
0,08
0,03
0,05
38 – 63
2,91
1,05
0,24
0,81
63 – 90
2,93
1,98
0,42
1,56
90 – 180
3,18
11,7
1,08
10,62
180 – 250
3,32
12,62
0,31
12,31
250 – 425
3,37
7,75
—
7,75
> 425
3,12
5,22
—
5,22
FDtotal
3,28
40,4
2,08
38,32
Table 4 indicates the measured char and coke percentages in the dust samples FDAC1, FD-AC2, FD-AP1 and FD-AP2. The char content in dust samples (BF with PCI) was below 3%. The coke content in dusts samples ranged from 32% to 40% and was higher in flue
dust of the all coke BF. Considering the BF operational parameters and all BF off gas samples (flue dust and sludge), it would be possible to calculate the percentage of uncombusted char/thm (ton of hot metal). Table 4 – Summary of PCI and coke carbon in dust samples FD-AC1, FD-AC2, FD-AP1 and FD-AP2
Item
BF all coke
BF with PCI
FD-AC1
FD-AC2
FD-AP1
FD-AP2
Char (wt, %)
—
—
2,08
2,64
Coke (wt, %)
38,91
40,15
38,32
32,54
This study has showed that the XRD technique associated to chemical analyses could be used as a standard procedure to identify the coal-char and coke structures and to quantify the percentage of these carbonaceous materials in the BF flue dust samples. 4. Conclusions — The granulometric analyses of the coal CAB showed practically no particles above 150μm, and its char particles will probably not exhibit sizes above 250μm. — The L002 of char ChAB was higher than its parent coal CAB indicating slightly higher crystalline order of carbon. — The L002 of the small fractions (<63μm) was greater to L002 for char and CK showing the unexpected presence of coke fines in this fractions. — The char content in dust samples studied (BF with PCI) was below 3%. — This study showed that the XRD technique could be used to quantify the percentage of coal-char and coke fines in the BF flue dust samples.
Acknowledgement The authors wish to acknowledge the financial support of The National Council for Scientific and Technological Development (CNPq) and Brazilian Coal Net. The authors would also like to thank M.Sc. Eng. Henriquison M.B. Reis from USIMINAS for production of char samples.
References [1] Carpenter AM. Use of PCI in blast furnaces. CCC/116, London, UK, IEA Clean Coal Centre. Sep 2006, 66 pp. [2] Babich A, Senk D, Gudenau HW, Mavrommatis K. Ironmaking, Mainz GmbH, Aachen, 2008. [3] Sahajwalla V, Kong CH, Chaubal PC, Valia HS. Determination of proportions of coal char and coke fines in the off-gas blast furnace samples. In: IRONMAKING CONFERENCE, 59, 2000. Proceedings. Warrendale: ISS, 2000, p. 305-11. [4] Gupta SK, Sahajwalla V, Al-Omari Y, Saha-Chaudhury N, Rorick F, Hegedus G, et al. Atomic Structure of Coke Fines in Blast-Furnace Dust and Their Origin in Operating Blast Furnaces. In: IRONMAKING CONFERENCE, 62, 2003. Proceedings. Warrendale: ISS, 2003, p. 841-53. [5] Wu K, Ding R, Han Q, Yang S, Wei S, Ni B. Research on Unconsumed Fine Coke and Pulverized Coal of BF Dust under Different PCI Rates in BF at Capital Steel Co. ISIJ International, Vol. 50, 2010, No. 3, p. 390–5. [6] Yamaguchi K, Ueno H, Matsunaga S, Kakiuchi K, Amanoi S. Test on High-rate pulverized coal injection operation at Kimitsu n.3 blast furnace. ISIJ International, V.35, 1995, p.148-55. [7] Lu L, Sahajwalla V, Kong C, Harris D. Quantitative x-ray diffraction: Analysis and its application to various coals. In: Carbon, vol 39, 2001, p. 1821-33, [8] Sahajwalla V, Gupta SK, Al-Omari Y, Saha-Chaudhury, Rorick F, Hegedus G, et al. Combustion Characteristics of Pulverized Coals and Char Released in Blast Furnace OffGas. In: IRONMAKING CONFERENCE, 62, 2003. Proceedings. Warrendale: ISS, 2003, p. 775-85. [9] Cullity BD. In: Elements of X-ray diffraction, Reading, MA:56. Addison-Wesley, 1978, p.828.
Oviedo ICCS&T 2011. Extended Abstract
Yields of the Pyrolysis Tests of Brown Coals Mined in the Czech Republic
Authors: Jaroslav Kusý, Lukáš Anděl, Marcela Šafářová, Josef Valeš Brown coal research institute j.s.c. Budovatelu 2830, 434 37 Most, Czech Republic Corresponding author:
[email protected], +420476208627
Abstract In the framework of the project GA 105/09/1554 „Conversion of the Czech brown coals with the hydrogen rich materials as the process of liquefaction and gasification“, some pyrolysis tests were made with the samples of brown coal and lignite from all operational mines in the Czech Republic. The main objective was to find a suitable brown coal for the co-pyrolysis testing with other components, realized in the next stage of the project. This work is aimed to the pyrolysis tests parameters and comparison of the products yields. During this stage of the project, seven types of the brown coals and lignite were sampled. Those samples were analysed for the input parameters. All pyrolysis tests were made under the same conditions, which means the maximum temperature 650 °C. Mass balance from all tests were evaluated for all products – semi-coke, water, liquid organic phase and gas. By the comparison of all those parameters, the most suitable coal was chosen for the next stage of the project – co-pyrolysis with other materials.
1. Introduction Different pyrolysis products are prepared by the pyrolysis of the solid carbonaceous materials in the temperature range 600 °C to 900 °C. [1]. The most importatnt product is a semi-coke, which can be used as a fuel, or, after activation, as an adsorbent. Such adsorbents can be used in a wide spectrum of environmental purposes, for example for the waste-water treatment, permanent gases purification, carbon dioxide storage etc. [2, 3, 4, 5] The most important property of the semi-coke is a pore volume and specific surface area [6]. Another pyrolysis products are gas and liquid products, consisted from the water and organic phase – brown coal tar [7]. Brown coal tar is a material, which
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Oviedo ICCS&T 2011. Extended Abstract
can be used as a raw material for the most part of those products made from crude oil [17]. One of the possibilities, how to influence the yields of the pyrolysis process is a change of qualitative parameters of the pyrolysed material. The change should be realised by the addition of the carbonaceous material with the high level of hydrogen concentration, either from the waste, or from the renewable sources. For those co-pyrolysis tests, some pyrolysis tests should be made with the different types of brown coals, mined in the Czech Republic. Results of these tests will be used for the selecting of the brown coal suitable for co-pyrolysis tests. Raw brown coals were analysed and obtained parameters were compared to those, obtained by the analysis of output pyrolysis products. . 2. Experimental section Different types of brown coals, mined in the Czech Republic, were tested. Samples with the weight of 100 kg were quartered to 6 kg and grinded the particle size from 1 to 5 mm. Coals were sampled from the Sokolov coal company (1 - Sokolov), North Bohemian mines – DNT (2 - DNT), Vrsany coal complay (3 - Vrsany), Litvinov coal company (4 - ČSA), Kohinoor mine (5 - Kohinoor) and North Bohemian coal company – Bílina mines (6 - SD-DB). Localisation of those mines is in the Figure 1.
Figure 1: Localisation of the samled coals
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Oviedo ICCS&T 2011. Extended Abstract
Samples were tested for their basic parameters, results are summarized in the table 1 and table 2. Table 1 shows results of the low temperature distillation yields of the coke, gas, tar and water [8]. This analysis is basically pyrolysis and its results should estimate the yields from the laboratory testing unit experiments. Table 2 shows basic parameters of the coal samples according to the Chzech standards [9,10,11,12,13,14,15]
Table 1: Yield of the tar, water, gas and semicoke from the low temperature distillation Sample
TSKd % hm.
1 Sokolov
d
% hm.
WSKd % hm.
GSKd % hm.
22,02
60,55
9,8
7,63
2 DNT
8,53
69,27
10,1
12,10
3 Vršany
13,44
66,2
9,39
10,97
4 ČSA
25,39
57,31
8,43
8,87
5 Kohinoor
13,58
61,95
10,42
14,05
6 SD-DB
21,96
57,72
8,75
11,57
SK
Table 2: Basic parameters of the tested coals Parameter W
a
d
A
Input coal 1 Sokolov
2 DNT
3 Vršany
4 ČSA
5 Kohinoor
6 SD-DB
% wt. % wt.
14,85
6,67
18,00
5,98
6,67
5,31
5,06
13,3
19,66
4,06
2,28
7,28
Std daf
% wt.
0,48
2,63
1,23
0,75
1,02
1,10
% wt.
0,51
3,03
1,53
0,78
1,04
1,18
C
d
% wt.
70,62
58,74
55,60
72,30
66,50
65,70
C
daf
% wt.
74,38
67,75
69,20
75,36
68,05
70,86
H
d
% wt.
5,47
4,82
5,27
5,65
4,61
5,78
H
daf
% wt.
5,76
5,56
6,56
5,89
4,72
6,23
N
d
% wt.
2,45
1,45
0,92
0,88
0,9
0,83
N
daf
% wt.
2,58
1,67
1,15
0,92
0,92
0,9
O
d
% wt.
15,92
19,07
17,32
16,36
24,69
19,31
O
daf
% wt.
16,77
21,99
21,56
17,05
25,27
20,83
d
% wt.
52,92
41,13
43,45
54,83
50,66
52,45
daf
% wt.
55,74
47,44
54,08
57,15
51,84
56,58
d
MJ/kg
30,02
24,52
23,49
31,34
28,18
29,28
S
V V
Qs
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Oviedo ICCS&T 2011. Extended Abstract Qs daf
MJ/kg
31,62
28,29
29,23
32,67
28,84
31,59
Qi
d
MJ/kg
28,83
23,47
22,33
30,11
27,17
28,03
Qi
daf
MJ/kg
30,36
27,08
27,80
31,38
27,81
30,23
2.1 Pyrolysis tests Pyrolysis tests were performed on the laboratory pyrolysis unit [7] with the 2500 ml retort. This unit is depicted on the figure 2. All samples were pyrolysed under the same conditions – initial temperature 25 °C, heating rate of 6 °C.min-1, terminal temperature 650 °C with the 60 min of delay on this temperature. During the tests, gas products were sampled between the internal temperature 510 to 520 °C. those samples were analysed by the gas chromatography with the thermal conductivity detector. Yields of the pyrolysis products, which means gas, tar, semi-coke and water were observed. All results are summarized in the chapter 3 of this work.
Figure 2: Laboratory pyrolysis testing unit
3. Results and Discussion 3.1
Pyrolysis test balance
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Oviedo ICCS&T 2011. Extended Abstract
Table 3 sumarizes weight balance of the pyrolysis tests for the individual types of tested brown coal. Ballance are related to the input coal in original condition. The last line of the table is the sum of the gas and loses, calculated as the difference of the 100 % and the sum of other three products. This item consists with the main part from the gas, and the minor part is the tar, adhering to the apparatus surface. results are the average from 3 to 5 individual tests of the same coal Table 3: Mass balance of the pyrolysis products Coal type
1 Sokolov
2 DNT
3 Vršany
4 ČSA
5 Kohinoor
6 SD-DB
Component
% wt.
% wt.
% wt.
% wt.
% wt.
% wt.
Brown coal
100
100
100
100
100
100
Semi-coke
38,46
60,2
52,29
50,64
53,85
44,98
Water
35,01
16,65
25,48
12,65
17,94
26,29
Tar
9,42
5,2
6,64
17,98
8,18
10,93
Gas + loses
17,1
17,95
15,59
18,74
20,04
17,8
3.2
Gas analysis
Table 4 summarizes results of the gaseous products from all laboratory tests. Gas were sampled between the temperatures of 510 and 520 °C to the antidiffusive bag [16]. Samples were analysed on the gas chromatograph GC 82 TT Labio Praha with the dual thermoconductivity detector. Results are summarized on the table 4 and on the figure 4. Hydrocarbons are given as the sum of lower hydrocarbons. Table 4: Results of the gaseous products analysis Coal type Component hydrogene oxygen nitrogen methane carbon dioxide carbon oxide hydrocarbons
1 Sokolov
2 DNT
3 Vršany
4 ČSA
5 Kohinoor
6 SD-DB
% vol.
% vol..
% vol..
% vol..
% vol.
% vol.
16,35 0,09 1,74 28,80 9,90 23,35 19,77
28,57 <0,01 0,94 28,59 7,36 33,32 1,22
24,88 <0,01 1,03 29,30 8,25 31,65 4,88
25,44 0,19 1,37 35,50 8,83 9,83 18,84
13,49 0,04 0,79 33,39 16,40 25,46 10,43
15,39 <0,01 1,23 35,39 10,91 27,07 10,00
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Oviedo ICCS&T 2011. Extended Abstract
Brown coals 40 35
Volume per cent
30 25 20 15 10 5 0 Sokolov
DNT
Vršany
ČSA
Kohinoor
SD-DB
Localisation
hydrogen
oxygen
nitrogen
methane
carbon dioxide
carbon oxide
hydrocarbons
Figure 3: Results of the gaseous products analysis
4. Conclusions As a part of the research project GA 105/09/1554, 21 pyrolysis tests were performed under the defined conditions. During these tests, yields of the individual pyrolysis products were observed and the gaseous products were analysed. The main purposes of this work was to describe behaviour of the every type of brown coal mined in the Czech Republic during the pyrolysis and to choose the most appropriate type of coal for the follow-up co-pyrolysis tests with the biological or waste materials. The most important parameters are yields of gas and liquid organic phase (tar). With the regard to these parameters, the most suitable type of coal is from the ČSA mine with the highest level of the tar – 18 wt. %. Other types of coal have lower than 10 wt. %, so they are less suitable for this purpose. Yield of the gaseous products (including loses) are between 15 up to 20 wt %. For the further testing, composition of gaseous products is also important. For our purposes, the most interesting coal is from the ČSA mine again, with the 25 vol. % of
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Oviedo ICCS&T 2011. Extended Abstract
hydrogen, more than 35 vol. % of methane and almost 20 vol. % of another hydrocarbons. Table 4 shows the big difference of the concentration of carbon dioxide and hydrocarbons between tested coals. The highest levels of hydrocarbons are in the coal from Sokolov and ČSA (19 vol. %) the lowest from the DNT and Vršany mines. The lowest concentration of carbon dioxide is in the gaseous products from the ČSA mines (about 10 vol. %), another coals has this parameter between 23 to 35 vol. %. From this results can be concluded, that organic matter in some brown coal types is more susceptible to the oxidation process. Acknowledgement. This work has been solved in the terms of research project GA 105/09/1554 Conversion of Bohemian brown coal with hydrogen rich matters as a liquid and gaseous hydrocarbon obtaining process” with the support of Czech Scientific Foundation.
References [1]
Anděl, L., Kusý, J., Valeš, J., Šafářová, M., Pyrolysis process of waste polyethyleneterephtalate, Chemical Product and Process Modelling, Vol. 4, Iss. 1, Article 18, Berkeley Electronic Press, 2009
[2]
Ciahotný, K., Possile methods of CO2 removal from the flue gas and its safe sequestration, Aprochem 2009, Milovy – Sněžné n.M.
[3]
Vrbová, V., Ciahotný, K., Procházková, A., Adsorption of CO2 the special adsorbents, 2 nd CO2 NET EAST Workshop, CO2 Capture and Storage – Response to Climate Change, 3 –4 March 2009, Bratislava, 2009
[4]
Vrbová, V., Ciahotný, K., Procházková, A., CO2 removal from the biogas on the special sorbents, International conference Biogas 2009, České Budějovice,
[5]
Vrbová, V., Carbon dioxide removal from biogas, International Conference of students and young researchers, St. Petersburg, Russia, 2009
[6]
Valeš, J., Kusý, J., Anděl, L., Šafářová, M., Adsorption possibilities on the active carbon prepared from the brown coal, Reporter Brown coal 4/2009
[7]
Šafářová, M., Kusý, J., Anděl, L., Pyrolysis of brown coal under different process condition, Fuel 84 (2005), pp. 2280-2285, Elsevier Ltd., 2005
[8]
ČSN ISO 647 (441371) Brown coals and lignites, Yields of the tar, water, semi-coke and gas after the low temperature distilation, ČSNI, 1995
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Oviedo ICCS&T 2011. Extended Abstract
[9]
ČSN ISO 351: Tuhá paliva. Stanovení obsahu veškeré síry – Vysokoteplotní spalovací metoda,
[10]
ČSN 44 1377: Tuhá paliva. Stanovení obsahu vody,
[11]
ČSN 44 1378: Tuhá paliva. Stanovení popela,
[12]
ČSN ISO 1928: Tuhá paliva. Stanovení spalného tepla kalorimetrickou metodou v tlakové nádobě a výpočet výhřevnosti,
[13]
ČSN ISO 609: Tuhá paliva. Stanovení uhlíku a vodíku – Vysokoteplotní spalovací metoda,
[14]
ČSN 44 1310: Tuhá paliva. Označování analytických ukazatelů a vzorce přepočtů výsledků na různé stavy paliva,
[15]
ČSN 44 1351: Tuhá paliva. Vážková metoda stanovení prchavé hořlaviny,
[16]
Kusý, J., Brejcha, J., Hautke, P., Antidiffusive bag for the gas sampling, author certificate no. 245959, Office of Industrial Property, Praha, 1988
[17]
Schemes of the synthetic fuel production, Stalin’s plant, Záluží u Mostu, Praha 1959
Abbreviations Wa
Content of analytical water
Wtr
Content of water in original matter
Ad
Ash content in dry matter
Std
Total content of sulphur in dry matter
Cd
Content of carbon in dry matter
d
H
Content of hydrogen in dry matter
Nd
Content of nitrogen in dry matter
Vd
Content of volatile matter in dry matter
Qsd
Combustion heat in dry matter
Qsr
Combustion heat in original matter
Qi
d
Heating value in dry matter
Qi
r
Heating value in original matter
TSKd
tar content (converted to dry condition)
WSKd
pyrogenetic water content (converted to dry condition)
GSKd
gas content (converted to dry condition)
d
SK
daf
semi-coke content (converted to dry condition) in superscript – results converted to the dry and ash free condition
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8
Oviedo ICCS&T 2011. Extended Abstract
Gold-impregnated carbon materials as regenerable sorbents for mercury retention Jorge Rodríguez-Pérez, Federico Inguanzo-Fernández, Elena Rodríguez, M. Antonia López-Antón, Roberto García, Mercedes Díaz-Somoano and M. Rosa MartínezTarazona Instituto Nacional del Carbón (CSIC). C/ Francisco Pintado Fe Nº26, 33011, Oviedo, Spain. Phone: +34 985119090. Fax: +34 985297662. E-mail:
[email protected] Abstract Several technologies are under development for removing mercury from power plant flue gas streams, the main problem being the capture of elemental mercury which is hard to remove in the typical air-pollution control devices. Mercury control via adsorption processes in regenerable sorbents is an attractive way to reduce mercury emissions. A gold-containing sorbent can efficiently capture elemental mercury via the formation of a gold/mercury amalgam on the surface of the sorbent; then, as the sorbent becomes saturated, it is regenerated and the mercury recovered. Moreover, the process may be considerably improved if the cost of the sorbent and the amount of gold employed are low. In this study, a carbon material containing gold nanoparticles were employed as mercury sorbents. The sorbents were impregnated with 0.1 wt% gold using tetrakis(hydroxymethyl)phosphonium chloride (THPC). The elemental mercury retention capacity of the material was evaluated in a lab-scale device. The final objective of this study was to regenerate the sorbent and to evaluate the effects of multiple regeneration cycles on the sorbent performance. The regeneration process was assessed at experimental scale. For this purpose, the sorbents were heated for different periods of time and the amount of mercury in the regeneration process monitored by a continuous mercury emission analyzer (VM 3000).
1. Introduction In recent years many Governments, through international research projects and cooperation agreements, have been making efforts to reduce mercury emissions from coal combustion processes [1]. A variety of approaches for mercury control are under development, ranging from combustion modification to multipollutant technologies. Many of these unique technologies have passed the bench- and pilot-scale
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Oviedo ICCS&T 2011. Extended Abstract
developmental phases and are now being tested at full scale. Many of the developing technologies appear to be capable of achieving >90% control, but in most cases, this level of control has not been demonstrated over the longer term at a large scale [2]. Moreover, whereas existing wet or dry flue gas desulfurization (FGD) units are suited to capture oxidized mercury, elemental mercury continues being the main problem to solve.
Mercury control via adsorption process on regenerable sorbents positioned in the flue gas stream to adsorb elemental mercury is an attractive way to reduce its emissions. While a variety of regenerable sorbent materials could be used, most of the development work has focused on the use of gold-coated substrates [3-5]. Then, as the sorbent becomes saturated, the sorbent can be regenerated and the mercury recovered. Results from modelling studies and field testing of a single-plate, gold-coated MerCAP™, by the Electric Power Research Institute (EPRI), have indicated that high mercury removals can be achieved over relatively short plate lengths at very high flue gas velocities. The concept has moved beyond the bench-scale level and has been tested with a small slipstream device at a number of coal-fired power stations [5]. In studies carried out at lab-scale, the impregnation of activated carbons with gold nanoparticles has also resulted in high elemental mercury retentions, similar to those obtained with activated carbon impregnated with sulphur [4]. A key concern is not only the absorption ability of the gold for mercury but also the regeneration ability and lifespan of the material. Work to date has mainly focused on the capture of mercury from the flue gas and not the regeneration of the sorbents or the sequestration of mercury. A viable regeneration process, along with a method for mercury recovery, is a major development step that has not yet been addressed.
In this study a carbon material was doped with gold nanoparticles to be used as mercury sorbent with the final objective of regenerating the sorbent and evaluating the long-term effects that repeated regeneration and use cycles would have on the sorbents.
2. Experimental section A commercial activated carbon (peat-based, steam-activated) called Norit RB3 was used as
support.
Gold-loaded
RB3
sorbent
(RB3Au)
was
prepared
by
the
tetrakis(hydroxymethyl)phosphonium chloride (THPC) method, based on the formation Submit before 31 May 2011 to
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Oviedo ICCS&T 2011. Extended Abstract
of Au colloids [4]. The sorbent was impregnated with 0.1 wt% gold. The quantity of gold retained in the sorbent was determined by analyzing the gold in the solution after impregnation, by inductively coupled plasma mass spectrometry (ICP-MS).
The laboratory device used for mercury retention is shown in Figure 1. It consists of a glass reactor heated by a furnace. 80.0 mg of sorbent was placed inside the reactor. Elemental mercury in the gas phase was obtained from a permeation tube in a N2 atmosphere and passed through the sorbent bed in O2+N2 stream of 0.5 L min-1. The O2 concentration of the atmosphere was 12.6 vol%. The concentration of mercury in the O2+N2 stream was 100 µg m-3. The temperature of the sorbent was kept at 40 ºC. A continuous mercury emission analyzer (VM3000) was used to monitor mercury concentration at the reactor outlet. The mercury content after the retention experiments was determined by means of an automatic mercury analyzer (AMA254).
Mixture
N2 100 µg m-3 Hg 0.5 L min-1 Mercury continuous emission monitor Hg Permeation tube Activated carbon
O2+N2 Au-RB3 bed and furnace (40ºC) Dilution (Air)
Figure 1. Schematic diagram of experimental device for mercury retention
In order to test the regeneration cycles, the post-retention sorbent was heated in the same experimental device from room temperature to 450 °C with a heating rate of 10ºC min-1 in a stream of N2. Mercury emission was monitored as a function of temperature using the VM3000 analyzer. The gold-doped sorbent was then again used for mercury retention, in order to determine if any performance degradation occurred as a result of the regeneration process. Submit before 31 May 2011 to
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Oviedo ICCS&T 2011. Extended Abstract
3. Results and Discussion The mercury adsorption curves for RB3Au after regenerating the sorbent three times are shown in Figure 2. It can be observed that the mercury retention capacity decreases sharply after the first cycle of regeneration (Table 1).
[Hg] μg m
-3
20
1st cycle
2nd cycle
15
3rd cycle
10 5
0 cycle
0 0
1000
2000 3000 t (min)
4000
5000
Figure 2. Mercury adsorption curves of the activated carbon impregnated with gold (RB3Au) after three regeneration cycles.
Table 1. Mercury retention capacity.
1
Hg Retention After Regen. (mg g-1) 2
3
0.3±0.1
0.4±0.1
0.4±0.1
Hg Retention Before Regen. (mg g-1) Cycles RB3Au
2.5±0.3
The mercury retention capacity of the sorbent loaded with gold achieved values of 2.5 mg g-1 after approximately three days of the experiment with an efficiency of 100% for more than a day. As can be observed the maximum retention capacity was not reached during this time. However, after repeated regeneration cycles the mercury capture decreases to, approximately, 0.4 mg g-1, remaining constant in successive cycles of regeneration. Figure 3 shows the mercury desorption curves in the three cycles of regeneration. Similar expected peak temperatures were obtained with the repeated use of the sorbent. The desorption of mercury starts at approximately 100 ºC but its completely
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decomposition occurs at 240±10 ºC in the experimental conditions of this study. The thermal desorption of mercury during the first cycle of regeneration shows a broad signal over the 100-450 ºC range, pointing the higher mercury capture obtained before the thermal regeneration. The results suggest that the exposure of the activated carbon coated with gold to heating may lead to the formation of large aggregates of gold or to the coalescence of gold avoiding an efficient mercury capture.
3500 1st cicle
3000
[Hg] μ g m
-3
2nd cycle 2500
3rd cycle
2000 1500 1000 500 0 0
100
200
300
400
500
T (º C)
Figure 3. Temperature of mercury desorption for the sorbent RB3Au after repeated regeneration cycles.
4. Conclusions The effect of repeated regeneration and use cycles of an activated carbon containing gold nanoparticles has been tested to determine the mercury retention capacity. The sorbent has been thermally regenerated at approximately 240 ºC, showing a decrease of mercury capture efficiency after regeneration. Further research concerning the impact of the regeneration process on initial mercury removal performance is needed.
Acknowledgement The authors thank the CSIC (PIF-06-050) and the Spanish Ministerio de Ciencia e Innovación (CTQ2008-06860-C02-01) for financial support.
References [1] United Nations Environment Programme (UNEP). Process Optimization Guidance
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Oviedo ICCS&T 2011. Extended Abstract
Document for Reducing Mercury Emissions from Coal Combustion in Power Plants, Chemical Branch, DTIE, Geneva, Switzerland, Draft report, July 2010. [2] Mercury information clearinghouse; quarter 3 – advanced and developmental mercury
control
technologies,
2004;
Available
at:
http://mercury.electricity.ca/PDFs/MercuryResearchQuarterly3.pdf
[3]
National
Energy
Technology
Laboratory,
2006;
Available
at:
http://www.netl.doe.gov/technologies/coalpower/ewr/mercury/control-tech/mercap.html. [4] Rodriguez-Perez J, Lopez-Anton MA, Diaz-Somoano M, Garcia R, MartinezTarazona MR. Development of gold nanoparticle-doped activated carbon sorbent for elemental mercury. Energy Fuels 2011;25:2022-7. [5] Ebner T, Fisher K, Ley T, Slye R, Chang R, Richardson C et al. Large-scale demonstration of the MERCAP™ technology for mercury control, 2004; Available at: http://www.netl.doe.gov/technologies/coalpower/ewr/mercury/control-tech/pubs/7029% 20AQV%20Paper.pdf
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Oviedo ICCS&T 2011. Extended Abstract
A Cubic Mile of Oil: Realities and Options for Averting the Looming Global Energy Crisis Ripudaman Malhotra
[email protected] Chemical Science and Technology Laboratory, SRI International, Menlo Park, CA 94025 USA Introduction It is time to reframe the debate about our future energy supply, arguably the greatest challenge we face. This challenge is often portrayed as a tension between the moral imperative of protecting the environment and preserving the economy. However, this simplistic view misses the more difficult challenge that we face: balancing the need to protect the environment—which would require us to stop using fossil-fuels—against the equally important need to provide people around the world with sufficient and affordable energy so that they can all live healthy and productive lives. Energy plays a central role in alleviating poverty. It provides essential services such as sanitation, clean water, clean fuel for cooking, illumination, transportation, and so on. Between 1981 and 2005, the poverty level in China declined from 85% to 16%. This laudable achievement was made possible by almost quadrupling its energy consumption. Global statistics on poverty are stark: 1.4 billion people subsist below the poverty level, defined by the World Bank as $1.25/day; 2.4 billion people rely on wood charcoal, or dung as their primary source of energy, and the women and young girls among them spend over 6 hours each day collecting the fuel, water and other chores, which deprives them of opportunities for advancement through education and entrepreneurship. In addition, about 1.5 billion people have no access to electricity. Meeting the global demand for energy is going to be a daunting task, and the way we choose to do it—namely, the energy sources we decide to employ—will have a profound effect on the lives of billions of people. Compounding the challenge is the tenuous political situation facing many of the current energy suppliers. Choices must be made, and we must all engage in making them, lest choices are made for us.
With the goal of raising public awareness about energy, I joined my colleagues Hew Crane and Ed Kinderman to write, A Cubic Mile of Oil. As suggested by its subtitle, the book lays out the realities of a looming global energy crisis and options for averting it. It
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provides an unvarnished look at the different energy sources available and makes the technical discussion accessible to an interested layperson.
Current Status Developing a practical energy policy starts by understanding the scale of the problem. Currently, we have no clear way to talk about this challenge. Energy is a difficult Eliminado: s
subject, and its discussion is made even more difficult by the use of different units for each source of energy: tons or Btu for coal, barrels for oil, cubic feet for gas, and kWh for electricity. We decided to use a cubic mile of oil (CMO) as the metric for describing global production and consumption of energy from all sources. A CMO is a large unit of energy. It is equivalent to 26.2 billion barrels of oil, 153 quadrillion Btu, and
Eliminado:
15.3 trillion kWh (based on a heat rate of 10,000 Btu/kWh, which is more typical of thermal power plants).
It so happens that the current global oil consumption of roughly 80 million barrels a day adds up to almost exactly 1 CMO/year. Additionally, each year the world uses 0.8 CMO of energy from coal, 0.6 from natural gas, and roughly 0.2 each of hydro, nuclear, and wood for a grand total of 3 CMO. Solar, wind, and biofuels barely register on this scale; combined they produced a total of 0.03 CMO in 2008.
Asia/Pacific and North America are the two largest energy-producing regions, producing 0.6 and 0.7 CMO respectively. They are also the largest consumers of energy, and (along with Europe), the the largest net importers of energy. The Middle East/North Africa region produces a little over 0.5 CMO of energy, but it consumes only 0.15 CMO; it is thus the largest exporter of energy, followed by the Russia Group and Central/South America regions. These regional deficits and surpluses of total energy, as well as of specific resources, lie at the basis of many geopolitical tensions and alliances in evidence today.
There are large differences in the annual per capita consumption of energy across the regions around the world. A gallon of oil equivalent (GO) is a more appropriate volumetric measure than the CMO to describe per capita consumption. When humans were hunter-gatherers, our energy use was about 20 GO/yr but grew to around 100
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GO/yr with the establishment of an agrarian society. Depending on living conditions, per-capita energy use today varies from barely 100 GO/yr in rural sub-Saharan Africa to about 1800 GO/yr in North America. This increase in energy consumption was made possible by switching fuel sources from biomass to coal and then to oil and natural gas. The global per-capita average consumption currently stands at 450 GO/yr.
Projected Future Demand The global energy demand is expected to rise from 3 CMO/yr to somewhere between 6 and 9 CMO/yr by the middle of this century. The drivers for this increased demand are already in place. Large segments of China, India, and Brazil are poised to increase their standard of living—and the concomitant energy demand—substantially. If the average energy consumption in the Asia-Pacific region were to reach the current global average, the demand would increase by 0.7 CMO, an amount equal to the total energy consumption in the North America region.
Various scenarios for future energy requirements place energy demand at between 6 and 9 CMO by 2050. Our challenge is thus framed in terms of finding ways to provide or avoid the use of several CMO/yr of energy. Under these different scenarios, the cumulative energy consumption through 2050 is varies between 150 and 270 CMO. Where is that energy going to come from?
Available Resources Our dominant energy sources are fossil resources of coal, oil, and natural gas, and the proven reserves of oil and gas are about 80 CMO. These will surely be exhausted by 2050, and we will be tapping into our resource base. It is important to recognize that reserves and resources are not fixed quantities, but are affected by the state of technology and the prevailing price. Without this distinction, the public is often left wondering why for 60 years or more our reserves have always been predicted to last for an additional 40 years! Beyond the resource base of conventional fuels are unconventional sources, for which extraction technology has yet to be developed. While our petroleum reserves are only about 46 CMO, the resource base of conventional oil adds about 94 CMO to that, and the unconventional petroleum resources may add an additional 400 CMO. Coal and natural gas resource bases are even larger. Yet, an
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Eliminado:
Oviedo ICCS&T 2011. Extended Abstract
unmitigated exponential growth at the historic rates of 2.6%/yr would exhaust even Eliminado:
1000 CMO of energy resources in about 150 years.
Of course, the question of energy supply is not only about the total resource potential, but also the rate at which it can be produced. Global capacity to produce petroleum is only a few percent more than the current demand. No wonder that a potential for disruption in supplies on the order of half a million barrels a day out of the nearly Eliminado:
80 million barrels can send oil markets into a frenzy as was recently witnessed in the increase in crude oil prices when Egyptians overthrew Mubarak’s regime.
Even though our fossil resources provide over 85% of the total energy, increasing any of these sources by 1 CMO/yr is not an easy task, particularly when one considers the entire infrastructure. Producing and additional 1 CMO/yr from oil would require finding and developing about 4000 oil fields, building 250 refineries, and thousands of miles of product pipelines.
Nuclear energy will very likely play an important role in our future, but it is the most misunderstood of the energy technologies. Issues that have limited the growth of nuclear power include fuel supply, cost escalation, and fears of plants exploding, radiation exposure, as well as nuclear proliferation. Deploying nuclear power is not without its risks, but there are risks associated with all energy production technologies, and it boils down to comparing and managing risks. If we want to develop nuclear power to provide 1 CMO/yr, we will need about 2,500 1 GW plants. In other words, we need to commission one such plant each week for the next 50 years.
Our renewable resources, or income energy sources as we call them, include hydropower, geothermal, solar, thermal, photovoltaic (PV), wind, and biomass power. The total energy from the sun falling on Earth adds up to 23,000 CMO/yr, but very little of it is being used today. Our current technologies are woefully inadequate, and scaling them can exacerbate some of the other problems we are facing with respect to food and water supplies. In laying out the requirements for scaling up our income sources to the 1 CMO/yr level, we find that the numbers are truly staggering as it would require 150 Three Gorges dams; 2,500 1-GW nuclear reactors; 3 million 2-MW wind turbines
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covering 580,000 square miles, 70,000 Andasol solar parks and 4.5 billion rooftop 2-kW PV systems.
Path Ahead The task ahead is enormous and will only be exacerbated if we continue in our entrenched ways or make only small adjustments. While technological advancements and innovations are seriously needed, unless we are also willing to make some attitudinal changes, the goal of living on sustainable energy is likely to elude us. Increased efficiency and conservation can play an important role in mitigating the crisis. These offer the quickest means to avert the impending crisis for a period of a few decades, during which time we must gear up our income resources. However, improvements in efficiency generally result in increase of total energy use: Our refrigerators are more efficient today, but they are larger. TVs are more efficient today because they consume less power per diagonal inch, but they are larger and we have many more of them.
We have many opportunities to reduce both our energy consumption and greenhouse gas emissions, but many of these actions are only a greenwash. It is essential to pay attention to those activities that have the most impact so that we do not squander valuable resources in pursuit of approaches that, at best, make only a marginal difference. Assuming one drives 12,000 miles a year, switching from a car with a fuel economy of 30 mpg to one with 40 mpg would avoid emissions from 100 GO/yr. However, three times that amount could be avoided by switching from a standard red meat diet to a lacto-ovo vegetarian diet with the same total caloric content.
Planning and implementing approaches to resolve our energy dilemma can be aided by laying out a portfolio of short-, medium-, and long-term actions with the most important item being starting a campaign to improve energy literacy. A democratic society may make difficult decisions wisely if its citizens are educated about the issue; a democratic society definitely will not make wise decisions when the citizens in position to apply social pressure for change are ignorant of the issue. The public at large must get involved in an informed debate to choose the sources for energy. Information is key to such a dialogue, and in the energy arena, information is constantly evolving.
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Entrained-Flow Gasification of Coal under Slagging Conditions: Properties of Solid Wastes and Relevance of Char-Wall Interaction Phenomena Fabio Montagnaro1, Paola Brachi1 and Piero Salatino2 1
Dipartimento di Chimica, Università degli Studi di Napoli Federico II, Complesso Universitario del Monte di Sant’Angelo, 80126 Napoli (Italy), T: +39 081 674029; F: +39 081 674090; E:
[email protected] 2
Dipartimento di Ingegneria Chimica, Università degli Studi di Napoli Federico II; Istituto di Ricerche sulla Combustione, Consiglio Nazionale delle Ricerche, Piazzale Vincenzo Tecchio 80, 80125 Napoli (Italy), T: +39 081 7682258; F: +39 081 5936936; E:
[email protected]
Abstract The aim of this paper is to investigate the properties of solid wastes generated from an industrial-scale pressurized entrained-flow gasifier, by means of a combination of experimental techniques: elemental, granulometric and X-ray diffraction analyses, scanning electron microscopy and energy dispersive X-ray analysis. The results are critically discussed in the light of the different regimes of char-slag micromechanical interaction and of the different phases that are established in the gasification chamber. The discussion allows to give useful insights concerning the properties and partitioning of carbon among the three main sources (coarse slag, slag fines, fly ash) of solid residues coming from the gasifier: in particular, differences between coarse slag and slag fines are highlighted, though these wastes are generated from the same mainstream. Furthermore, it is observed that residual carbon in slag granules is present in a segregated embedded form, while slag fines are composed of both porous (high-carbon) and compact (low-carbon) material. Altogether, the properties of the three residues are consistent with a mechanistic framework of the bulk-to-wall transfer and partitioning of solids during entrained-flow gasification of coal developed in a recently published theoretical paper.
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Oviedo ICCS&T 2011. Extended Abstract
1. Overview Entrained-flow coal gasifiers of new generation are characterized by operating conditions (high operating temperatures and multi-stage feedings of coal and gaseous reactants) aimed at favoring ash migration/deposition onto the reactor walls, whence the molten ash (slag) flows and is eventually drained at the bottom of the gasifier [1,2]. Detailed studies concerning the fate of char particles as they impinge the wall slag layer have been only recently developed [3–6]. In a recently published paper, Montagnaro and Salatino [7] addressed the relative importance of the parallel pathways of coal conversion associated with entrained-flow of carbon particles in a lean-dispersed gas phase vs. segregated flow of char particles in a dense-dispersed phase in the near-wall region of the gasifier. Char segregation in the dense-dispersed phase is promoted by bulk-to-wall particle migration and by inelastic interaction of char particles with the molten slag wall layer. By taking into account characteristics such as char density, particle diameter and impact velocity, slag viscosity, interfacial particle-slag tension, theoretical criteria for either char particle entrapment inside or carbon-coverage of the wall ash layer have been developed. This is represented in Figure 1, which depicts the following plausible regimes of char-slag micromechanical interaction: regime E) (entrapment), in which char particles reaching the slag surface become permanently embedded into the molten layer and further course of combustion/gasification is hindered; regime S) (segregation), in which char particles reaching the wall adhere to the slag layer’s surface without being fully engulfed, so that the progress of combustion/gasification is permitted; regime SC) (segregation and coverage), in which the coverage of the slag layer with carbon particles is extensive. In this last regime a dense-dispersed annular phase is established in the close proximity of the wall ash layer, where the excess impinging char particles which cannot be accommodated on the slag surface accumulate. This phase is likely to be characterized by a velocity that is intermediate between that of the fast lean-dispersed phase and that of the slowly moving wall molten ash layer. This feature is beneficial to C conversion as it gives rise to a longer mean residence time of carbon particles belonging to this phase. Consistently, a schematic diagram of the entrained-flow gasifier is presented in Figure 2 in which, in particular, the presence of three different sources of solid wastes is underlined, that is: slag phase, yielding coarse slag granules upon interaction with the quench bath at the bottom of the gasifier; dense-dispersed phase, giving rise to slag fines upon interaction with the quench bath; lean-dispersed phase, giving rise to fly ash escaping the gasifier in the gas stream. 2
Oviedo ICCS&T 2011. Extended Abstract
coal (WF ) O2 (WOX) H2O (WS )
devolatilization/ combustion zone
lean‐ dispersed dense‐ dispersed slag
char, soot
syn‐gas fly ash (WFLY)
slag (WSLAG) syn‐gas slag fines (WSF )
Figure 1 (left). Regimes of C-slag micromechanical interaction (E=entrapment; S=segregation; SC=segregation and coverage). Figure 2 (right). Outline of flow patterns in the entrained-flow gasification chamber (the symbol ‘W’ stands for mass flow rates). 2. Rationale of the present investigation This investigation was stimulated by the operational experience from an industrial-scale pressurized (25 bar; cf. [8,9]) entrained-flow gasifier, operated in the slagging regime (at temperatures around 1700–1900 K; cf. [9,10]). The gasification chamber has an internal diameter of 3.8 m and a length of 13 m. Mass feeding ratios are: WOX/WF=0.8; WS/WF=0.1 (see Figure 2). Solid fuel feed rate is WF=30 kg s–1 and, typically, the solid fuel is a 50:50 coal:pet coke mixture. Practical operation of the gasifier revealed a value of the mass ratio WFLY/(WSLAG+WSF) smaller than expected at the design stage (circa 0.1 vs. 0.4; cf. [9,11]). This was primarily ascribed to an unexpectedly large value of the mass flow rate of slag fines (WSF) leaving the quench bath. Moreover, C content in fly ash resulted rather limited (circa 5%; cf. [11]), while that in slag fines was reported to be quite high. Finally, a non-negligible organic fraction was detected in the slag waste (cf. [12]), and this appears consistent with findings reported, for similar systems, by other authors [13–15]. In this paper, results concerning characterization of solid residues from the industrial-scale gasifier were interpreted in the light of the different regimes of char-slag micromechanical interaction (Figure 1) and of the different phases that are established in the gasification chamber (Figure 2), as presented in the Overview. In this context, it is useful to remind that the solid waste other than fly ash is quenched in a water bath generating, besides the slag (sometimes referred to as coarse slag), a black water whence slag fines are recovered by filtration [13]. The different thermal/conversion history of
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Oviedo ICCS&T 2011. Extended Abstract
these two solid wastes is likely to strongly influence their properties, in particular as far as carbon content, morphology and further reactivity is concerned. Only recently did the question concerning the differences between coarse and fine slag receive consideration: the reader is referred, for example, to the works [13–15].
3. Materials and experimental techniques Coarse slag and slag fines samples were generated in the ELCOGAS entrained-flow gasifier located in Puertollano, Ciudad Real (Spain). This material was provided in the summer of 2009 and, due to the considerable amount of wastes produced by the plant and to the variability of the operating conditions of the gasifier, it might not be fully representative of the ash generated by the industrial gasifier during normal operation. Coarse slag and slag fines were characterized by: carbon elemental analysis, performed by a LECO CHN-2000 instrument; granulometric analysis, performed by either a Malvern Instruments Master Sizer 2000 laser granulometer (operated down to a minimum particle size of 0.02 μm) or mechanical sieving (in 10 size ranges between 0 and 9.5 mm); X-ray diffraction (XRD) analysis, performed by a Bruker D2 Phaser diffractometer (operated at diffraction angles ranging between 10 and 60°2θ with a scan velocity equal to 0.02°2θ s–1); scanning electron microscopy (SEM), performed by a FEI Inspect microscope equipped with an energy dispersive X-ray (EDX) probe (operated up to magnifications of 3000×).
4. Experimental results A preliminary elemental analysis on coarse slag and slag fines revealed no appreciable carbon content (for the former) and a carbon content as high as 57.4% (for the latter). Figures 3 and 4 show absolute and cumulative particle size distributions for slag and slag fines, respectively. Due to the much coarser size of the slag material, the particle size analysis for this waste was carried out by mechanical sieving instead of laser granulometry. From Figure 3, it can be observed that slag particles size extended over a broad range, with maximum size equal to about 9 mm (cf. [12]). Nonetheless, a very distinct peak for the absolute distribution could be appreciated at 1.7 mm; the mean Sauter diameter for this distribution was equal to 1.18 mm. Finally, a median value (d50) of 1.27 mm was obtained from the cumulative distribution. Slag fines were characterized by much smaller values of particle size (Figure 4): particles coarser than 700 μm were not observed, the peak and the Sauter diameter for this distribution were equal to 100 and 20
4
Oviedo ICCS&T 2011. Extended Abstract
μm, respectively, and the d50-value for the cumulative distribution was 72 μm. 0.6
1.0
peak=1.7 mm Sauter=1.18 mm
d50=1.27 mm
0.9
Cumulative size distribution, -
Absolute size distribution, -
0.5
0.4
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5
6
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peak=100 μm Sauter=20 μm
d50=72 μm
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Cumulative size distribution, -
0.05
Absolute size distribution, -
4
Particle diameter, mm
0.04
0.03
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100
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Particle diameter, μm
Particle diameter, μm
Figure 3 (upper row) and Figure 4 (lower row). Absolute (left) and cumulative (right) particle size distributions for coarse slag (Figure 3) and slag fines (Figure 4).
Peak intensity, a.u.
A
P
slag fines
coarse slag
10
15
20
25
30
35
40
45
50
55
60
Diffraction angle, °2θ
Figure 5. XRD analysis for coarse slag and slag fines (A=anhydrite, CaSO4, CPDS#228; P=periclase, MgO, CPDS#2693). Figure 5 reports XRD patterns for the wastes under investigation. The results for coarse slag revealed no distinct peaks, due to the amorphous and vitreous nature of this material related to its very peculiar thermal history in the gasifier [10,12,16]. Though deriving from the same main-stream, slag fines separated from the slag after the quench bath showed a different crystalline microstructure, the peaks of anhydrite and periclase being 5
Oviedo ICCS&T 2011. Extended Abstract
clearly recognizable. While elemental analysis on coarse slag did not indicate the presence of carbon in this material, interestingly, different results were obtained when cross-sections of coarse slag granules were analyzed by SEM-EDX (Figure 6a and Table 1).
a)
c)
b)
Figure 6. SEM micrographs of cross-sections of: a) a whole coarse slag particle (magnification=50×); b–c) selected zones of slag particles (magnification=1600×), displaying carbon-rich patches. Table 1. EDX elemental analysis results (%wt) referring to SEM micrographs reported in Figures 6 and 7. C O Na Mg Al Si S K Ca Ti V Fe Ni
Figure 6a 9.27 26.95 0.14 0.79 14.91 32.62 n.d. 2.25 7.38 0.59 0.38 4.72 n.d.
Figure 6b 48.77 11.50 0.14 0.46 9.20 19.10 1.24 1.41 4.75 0.38 0.19 2.86 n.d.
Figure 6c 54.23 8.67 0.13 0.44 8.31 17.82 1.70 1.44 3.96 0.40 0.23 2.67 n.d.
Figure 7a 86.47 6.71 n.d. n.d. 0.31 0.51 3.03 0.20 n.d. n.d. 0.34 2.15 0.28
Figure 7b 82.27 8.18 n.d. 0.18 0.35 0.38 3.39 0.17 n.d. n.d. 0.28 4.32 0.47
Figure 7c 84.89 6.98 n.d. n.d. 0.26 0.29 2.91 n.d. n.d. n.d. 0.31 3.96 0.41
Figure 7d 18.36 39.21 0.33 0.64 11.00 22.08 0.69 2.86 2.52 0.70 0.46 1.14 n.d.
Figure 7e 13.90 40.94 0.25 0.64 10.92 24.13 n.d. 1.51 4.80 0.44 0.30 2.17 n.d.
The whole slag particle shown in Figure 6a appeared to be mostly vitreous and dense, in agreement with XRD results. While the inorganic fraction is primarily constituted by Si+Al (47.5%) together with Ca (7.4%, related to the feeding of limestone to the gasifier as fluxing agent), Fe (4.7%) and K (2.2%), a remarkable carbon content (9.3%) was also detected. Carbon could not be mainly ascribed to the possible presence of CaCO3: even if all the calcium were present as carbonate, the ‘free’ carbon would anyway be as large as 7.1%. Analyses performed on the cross-sections of other whole particles led to comparable observations. Figures 6b–c (and Table 1) report SEM-EDX results obtained carrying out the analysis at a greater magnification on the cross-section of two selected zones of slag particles. In both cases the occurrence of darker patches was observed.
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Oviedo ICCS&T 2011. Extended Abstract
Pointwise quantitative EDX results refer to these regions: C-contents as high as 48.8– 54.2% were obtained. This finding contributes to the assessment of the relevance of carbon entrapment (regime E), Figure 1) in slag particles. It is worth noting that elemental analysis did not show the presence of any organic fraction and this was essentially so because C was permanently entrapped into the slag matrix in a way that could not have been disclosed by thermal analysis. Only the cutting procedure associated with the SEMEDX analysis of granules cross-sections was able to disclose the unreacted carbon, which appeared to be somewhat segregated (in the patches) with respect to the inorganic slag matrix. By taking into account that quantitative results relevant to Figure 6a referred to the cross-section of a slag particle, an average carbon content around 3–4% was estimated for this waste. Figure 7 and Table 1 report the results of the SEM-EDX analysis performed on whole slag fines particles. In particular, particles having prevailing either porous (Figures 7a–c) or compact (Figures 7d–e) structures were observed. In any case, the carbon content was larger than the value obtained from the inspection of coarse slag particles, in line with results of elemental analysis. This is particularly evident for porous particles (Figures 7a–c): C-content ranged between 82.3 and 86.5%, and Fe (2.2–4.3%) and S (2.9–3.4%) could also be appreciated. Thus, these particles should be mainly associated with unreacted char present in the dense-dispersed phase (Figure 2) giving rise to slag fines upon impingement on the quench bath. On the other hand, dense particles (Figures 7d–e) display morphological and chemical features that are closer to those of coarse slag particles, at least as far as SEM-EDX results are concerned: C-content ranged between 13.9 and 18.4%, the Si+Al-fraction was as high as 33.1–35.0% and Ca (2.5– 4.8%), K (1.5–2.9%) and Fe (1.1–2.2%) could also be detected. It is also highlighted that elements such as Na, Mg, Al, Si, K, Ca and Ti, while revealed in smaller amounts (or not revealed at all) in high-C porous slag fines, were present in larger percentages in both coarse slag and low-C dense slag fines particles. The opposite is true for S and Ni. The results are consistent with the previously-reported C-content of slag fines (about 57%), obtained by standard elemental analysis carried out on waste containing both high-C porous and low-C dense materials. Moreover, the more compact slag fines should be regarded as having intermediate properties between porous slag fines and coarse slag: this observation, jointly with XRD results, highlights again the establishment of a densedispersed phase that, together with the slag phase, generates both streams: slag and slag fines (Figure 2).
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Oviedo ICCS&T 2011. Extended Abstract
a)
c)
b)
d)
e)
Figure 7. SEM micrographs of different whole slag fines particles: a)–c) at magnification=1600×; d)–e) at magnification=3000×. 5. Critical discussion It has been observed that the slag fines, generated in high amount in the gasifier (cf. Section 2), are very rich in carbon (cf. elemental and SEM-EDX analysis results). This is believed to be a clue in favor of the existence of a dense-dispersed phase (regime SC), Figures 1 and 2), which would be the source of high-C slag fines upon interaction with the quench bath. On the other hand, the slag waste presented a non-negligible content of carbon, mostly entrapped (in a segregated fashion) into the slag matrix. This would be consistent with the occasional establishment of regime E). It is interesting to analyze these findings in the light of results recently published by Li et al. [5,6]. These authors investigated the char-slag transition during entrained-flow oxidation of coal particles, observing a distinct transition from porous/non-sticky char to fluid/sticky slag occurring at temperatures above the ash flow temperature only when carbon conversion exceeded a certain threshold (around 90%). This means that the coexistence of regimes SC) –as the prevailing regime– and E) might be possible. If one assumes, based on Montagnaro and Salatino’s findings [7], that regime SC) is the dominant regime under typical operating conditions of entrained-flow gasifiers, then some char particles belonging to the densedispersed phase and in the late burn-off stage could well be permanently embodied into the slag layer, according to Li et al.’s findings [5,6]. Although the carbon content of these 8
Oviedo ICCS&T 2011. Extended Abstract
particles is likely to be modest, it is anyway not negligible and could possibly justify the estimated C-value of 3–4%. As far as fly ash is concerned, its carbon content (5%) was unexpectedly small. This finding can be interpreted by considering that bulk-to-wall transfer of char/ash particles is dominated by the inertial mechanism associated with turbophoresis and centrifugal forces due to swirl/tangential flow. This mechanism would make wall transfer of coarser particles more effective than transfer of fines [7]. Considering, as reported by Wu et al. [13], that the carbon content of coarser particles is generally higher, this mechanism would imply a more effective transfer of carbon to the wall, as compared with ash transfer. These two aspects (coexistence of regimes and preferential mass transfer) could explain why fly ash resulted selectively C-depleted and why slag fines did show such a large mass flow rate and C-content, contrary to the expectations. Moreover, it is recalled that part of the fly ash might derive by nucleation and growth of fine (carbon-free) inorganic particles from the gas phase under the extremely high temperature conditions experienced by the fuel in the flame region of the oxidizer. This would further justify the comparatively small carbon content of fly ash.
6. Conclusions The properties and partitioning of carbon among the three main sources (coarse slag, slag fines, fly ash) of solid residues coming from an industrial-scale entrained-flow coal gasifier have been characterized by a combination of experimental techniques. The carbon content of slag fines is very large, in the order of 60%. The carbon content of fly ash is around 5%. The carbon content of slag granules as assessed by standard elemental analysis techniques is negligible, but combined microscopy and EDX analysis of the granules’ cross-sections indicates that residual carbon is present in slag granules as segregated embedded carbon-rich patches, amounting to about 3% by mass of the sample. The properties of these three types of residues are consistent with a mechanistic framework of entrained-flow gasification of coal developed by the authors, which considers the bulk-to-wall transfer of solids and the establishment of segregated phases in the near-wall region of the gasifier.
Acknowledgement The authors wish to express their gratitude to Mr. F. García Peña, Dr. A. M. Mozos and Dr. P. Coca (ELCOGAS, Spain) for having supplied raw materials and for useful discussion. Dr. M. Urciuolo and Mr. S. Russo (IRC-CNR, Italy) are gratefully acknowledged for their support in solid characterization. 9
Oviedo ICCS&T 2011. Extended Abstract References [1] Walsh PM, Sarofim AF, Beér JM. Fouling of convection heat exchangers by lignitic coal ash. Energy Fuel 1992;6:709–15. [2] Shimizu T, Tominaga H. A model of char capture by molten slag surface under high-temperature gasification conditions. Fuel 2006;85:170–8. [3] Wang XH, Zhao DQ, He LB, Jiang LQ, He Q, Chen Y. Modeling of a coal-fired slagging combustor: development of a slag submodel. Combust Flame 2007;149:249–60. [4] Shannon GN, Rozelle PL, Pisupati SV, Sridhar S. Conditions for entrainment into a FeOX containing slag for a carbon-containing particle in an entrained coal gasifier. Fuel Process Technol 2008;89:1379–85. [5] Li S, Whitty KJ. Investigation of coal char-slag transition during oxidation: effect of temperature and residual carbon. Energy Fuel 2009;23:1998–2005. [6] Li S, Wu Y, Whitty KJ. Ash deposition behavior during char-slag transition under simulated gasification conditions. Energy Fuel 2010;24:1868–76. [7] Montagnaro F, Salatino P. Analysis of char-slag interaction and near-wall particle segregation in entrained-flow gasification of coal. Combust Flame 2010;157:874–83. [8] Seggiani M. Modelling and simulation of time varying slag flow in a Prenflo entrained-flow gasifier. Fuel 1998;77:1611–21. [9] Álvarez-Rodríguez R, Clemente-Jul C, Martín-Rubí JA. Behaviour of the elements introduced with the fuels in their distribution and immobilization between the coal-petroleum coke IGCC solid products. Fuel 2007;86:2081–9. [10] Aineto M, Acosta A, Rincón JM, Romero M. Thermal expansion of slag and fly ash from coal gasification in IGCC power plant. Fuel 2006;85:2352–8. [11] Font O, Moreno N, Díez S, Querol X, López-Soler A, Coca P, García Peña F. Differential behaviour of combustion and gasification fly ash from Puertollano Power Plants (Spain) for the synthesis of zeolites and silica extraction. J Hazard Mater 2009;166:94–102. [12] Acosta A, Aineto M, Iglesias I, Romero M, Rincón JM. Physico-chemical characterization of slag waste coming from GICC thermal power plant. Mater Lett 2001;50:246–50. [13] Wu T, Gong M, Lester E, Wang F, Zhou Z, Yu Z. Characterisation of residual carbon from entrainedbed coal water slurry gasifiers. Fuel 2007;86:972–82. [14] Xu S, Zhou Z, Gao X, Yu G, Gong X. The gasification reactivity of unburned carbon present in gasification slag from entrained-flow gasifier. Fuel Process Technol 2009;90:1062–70. [15] Zhao X, Zeng C, Mao Y, Li W, Peng Y, Wang T, Eiteneer B, Zamansky V, Fletcher T. The surface characteristics and reactivity of residual carbon in coal gasification slag. Energy Fuel 2010;24:91–4. [16] Song W, Tang L, Zhu X, Wu Y, Rong Y, Zhu Z, Koyama S. Fusibility and flow properties of coal ash and slag. Fuel 2009;88:297–304.
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Characterization of the hydrocarbon components from sequential flash pyrolysis for a vitrinite-rich and inertinite-rich coal Daniel van Niekerk , Carol du Sautoy and Johannes van Heerden Sasol Technology Research and Development, Coal Processing Technologies, 1 Klasie Havenga Road, Sasolburg, 1947, South Africa Corresponding author e-mail:
[email protected] Abstract Flash-pyrolysis is a potential process that can be utilized to produce hydrocarbon liquids directly from coal. During flash-pyrolysis coal is rapidly heated to high temperature (e.g. 600 °C within a few seconds) resulting in the release of gaseous and liquid hydrocarbons. Flash-pyrolysis gas chromatography mass spectrometry (GCMS) is an analytical method that can be used to investigate the pyrolysis hydrocarbon composition of a specific coal. In the current study the difference in pyrolysis hydrocarbon composition was determined for two maceral-different Permian-aged South African coals: A vitrinite-rich and an inertinite-rich coal. A weighed amount of coal was sequentially pyrolyzed from 300 °C to 1000 °C with 100 °C increments and analyzed using a GCMS. The chromatograms were analyzed and compounds were grouped according to representative functionalities (e.g. alkanes, alkenes, alkylbenzenes, phenolics, etc.). The effect of pyrolysis temperature versus the hydrocarbon composition for each coal was investigated. For both coals, tar evolution started between 400 and 500 °C and terminated at 700 °C. The maximum tar evolution was at 600 °C for both coals. The tar evolved consisted of various types of aromatic and aliphatic hydrocarbons. This work shows the effect of maceral composition (inertiniterich versus vitrinite-rich) and temperature on flash-pyrolysis hydrocarbon yield and composition. This hydrocarbon composition and temperature relation can fundamentally be related to the molecular structure of these two maceral-different coals.
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Oviedo ICCS&T 2011. Extended Abstract
1.
Introduction
Flash-pyrolysis is a potential process that can be utilized to produce hydrocarbon liquids directly from coal. During flash-pyrolysis coal is rapidly heated to high temperature (e.g. 600 °C within a few seconds) resulting in the release of gaseous and liquid hydrocarbons. Flash-pyrolysis gas chromatography mass spectrometry (GCMS) is a useful analytical technique for the chemical characterization of insoluble organic matter present in coals and other carbonaceous sediments [1]. Various studies have been conducted using flash-pyrolysis GCMS to elucidate the chemical structures of coals and coal macerals [2-4]. In this study flash-pyrolysis was used to investigate the pyrolysis hydrocarbon composition of two maceral-different Permian-aged South African coals: A vitrinite-rich and an inertinite-rich coal. The effect of pyrolysis temperature versus the hydrocarbon composition for each coal was investigated. 2. Experimental methods Coal samples: Two maceral-different South-African Permian-aged Gondwanaland coals were selected for this study: Vitrinite-rich coal and an inertinite-rich coal. Coal samples were analyzed using proximate, ultimate and petrographic analyses. The petrographic analyses included maceral composition and mean-random vitrinite reflectance. Pyrolysis gas chromatography mass spectrometry: Approximately 4-6 mg of each coal was placed in a quartz pyrolysis tube. The coal was sequentially pyrolyzed in 100 °C increments, starting at 300 °C and ending at 1000 °C using helium as flushing and carrier gas. Pyrograms were generated using a GCMS fitted with a 60m x 250 m x 0.5 µm DB Petro non-polar column. The oven temperature program was as follows: Initial temperature of 50 °C was held for 1 minute; thereafter three ramps were carried out: 5 °C/min to 100 °C; 7 °C/min to 180 °C; 10 °C/min to 320 °C; and this final temperature was held for 8 minutes. The split ratio for all runs was set at 50:1. The mass spectrometer was used in scan mode: 14 amu to 450 amu.
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Oviedo ICCS&T 2011. Extended Abstract
3. Results and discussion The maceral compositions of the two coals are summarized in Table 1.
Table 1: Maceral composition of the two coal samples Coal sample Vitrinite-rich coal Inertinite-rich coal
Vitrinite 82.5 13.9
Liptinite 4.5 1.5
Inertinite 13.0 84.6
Maceral composition in mineral matter free basis (mmf)
Two South-African Permian-aged coals were selected for this study: A vitrinite-rich coal (~82 % vitrinite, mmf) and an inertinite-rich coal (~85 % inertinite, mmf). Both coals had minor liptinite content (<5 %) which is typical of South African Permian-aged coals [5]. The dominating inertinite macerals in both coals was a combination of lowand high-reflecting semifusinite and inertodetrinite. Both coals were ranked as highvolatile bituminous by mean-random vitrinite reflectance. Proximate and ultimate analyses data are shown in Tables 2 and 3.
Table 2: Proximate analysis of the two coal samples (dry, ash-free) Coal sample Vitrinite-rich coal Inertinite-rich coal
Fixed carbon % 57.6 70.3
Volatile matter % 42.4 29.7
Table 3: Ultimate analysis of the two maceral-different coal samples (dry, ash-free) Coal sample Vitrinite-rich coal Inertinite-rich coal
%C 77.7 75.8
%H 5.5 4.2
%N 2.1 1.7
%S 1.7 1.0
%O 12.9 17.3
H/C 0.84 0.66
O/C 0.12 0.17
% Oxygen determined by difference
As expected, the inertinite-rich coal had a higher percentage fixed carbon and a lower percentage volatile matter content than the vitrinite-rich coal. The vitrinite-rich coal had a higher hydrogen content than the inertinite-rich coal (5.5 % for the vitrinite-rich coal and 4.2 % for the inertinite-rich coal, daf basis). The vitrinite-rich coal had a higher atomic H/C ratio than the inertinite-rich coal (0.84 % for the vitrinite-rich coal and 0.66 % for the inertinite-rich coal). The lower atomic H/C ratio of the inertinite-rich coal indicates a more aromatic hydrocarbon composition than the vitrinite-rich coal [6]. These results were consistent with previous vitrinite-rich and inertinite-rich results of South African Permian-aged coals [7]. Previous studies found that although these two coals differ in maceral composition they are similar in bulk structure [7]. The main structural differences between these coals are that the inertinite-rich is more aromatic,
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Oviedo ICCS&T 2011. Extended Abstract
having a more polycondenced aromatic structure and is more structurally ordered than the vitrinite-rich coal [7,8].
The two maceral-different coals were sequentially pyrolyzed in 100 °C increments (from 300 °C to 1000 °C) using a flash-pyrolysis coupled GCMS. Semi-quantification was conducted by determining the area percent of all the pyrolytes in the pyrograms. The amount of pyrolytes generated at each pyrolysis temperature for both coals are shown in Figure 1.
Figure 1: The amount of pyrolytes generated for both coals at various temperatures (normalized) In general, the vitrinite-rich coal had greater abundance of pyrolytes present than the inertinite-rich coal. Minor amount of pyrolytes were formed at 300 °C for both coal samples (less than 5 %). This mass loss was attributed to light organic hydrocarbons associated with the coal structure and the loss of inherent moisture. Pyrolysis started between 400 °C and 500 °C, with maximum pyrolyte evolution at 600 °C. Vitrinite-rich and inertinite-rich coals had similar devolatisation profiles, but differed in the amounts of mass loss. The various individual pyrolytes in the pyrograms were classified according to functionality: Alkanes, alkenes, dienes, alkylbenzenes, alkylphenols, alkylnaphthalenes,
alkylnaphthols,
polyaromatic
hydrocarbons
(PAH),
heteroatom
hydrocarbons and aromatic mixtures. The various hydrocarbon groups are summarized in Tables 4 and 5.
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Oviedo ICCS&T 2011. Extended Abstract
Table 4: Hydrocarbon distribution from vitrinite-rich coal Alkanes Alkenes Dienes Alkylbenzenes Alkylphenols alkylnaphthols alkylnaphthalenes PAH Heteroatom HC Aromatic mixtures
300 25.4 0.4 0.0 0.0 0.0 0.0 6.7 5.2 0.7 0.0
400 1.4 0.0 0.0 0.0 0.0 0.0 0.0 1.5 0.0 0.0
500 15.3 8.6 9.6 6.3 14.2 9.5 9.6 18.9 5.8 4.5
600 56.7 91.0 90.4 63.7 71.0 80.3 71.9 51.8 88.5 93.8
700 1.2 0.0 0.0 24.9 14.7 10.3 11.9 22.6 5.1 1.7
800 0.0 0.0 0.0 4.2 0.0 0.0 0.0 0.0 0.0 0.0
900 0.0 0.0 0.0 1.0 0.0 0.0 0.0 0.0 0.0 0.0
1000 0.0 0.0 0.0 0.0 0.0 0.0 0.0 0.0 0.0 0.0
Total 100.0 100.0 100.0 100.0 100.0 100.0 100.0 100.0 100.0 100.0
1000 0.0 0.0 0.0 0.7 0.0 0.0 0.0 0.0 0.0 0.0
Total 100.0 100.0 100.0 100.0 100.0 100.0 100.0 100.0 100.0 100.0
Table 5: Hydrocarbon distribution from inertinite-rich coal Alkanes Alkenes Dienes Alkylbenzenes Alkylphenols alkylnaphthols alkylnaphthalenes PAH Heteroatom HC Aromatic mixtures
300 27.7 0.0 0.0 0.0 0.0 0.0 5.4 0.8 1.2 0.0
400 1.8 0.0 0.0 0.0 0.0 0.0 0.6 1.1 0.1 0.9
500 13.9 8.4 12.7 5.2 7.0 8.8 12.1 22.0 8.6 26.9
600 51.1 85.2 78.3 55.9 64.3 75.8 58.3 40.2 68.5 56.9
700 5.5 6.5 9.0 27.8 28.2 15.4 21.0 33.5 21.1 15.3
800 0.0 0.0 0.0 8.3 0.5 0.0 2.6 2.3 0.6 0.0
900 0.0 0.0 0.0 2.1 0.0 0.0 0.0 0.2 0.0 0.0
The evolution of gas, aliphatic and aromatic species were determined and are shown in Figures 2 and 3.
Figure 2: Evolution of aliphatic, aromatic and gas species at various flash-pyrolysis temperatures from vitrinite-rich coal
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Oviedo ICCS&T 2011. Extended Abstract
Figure 3: Evolution of aliphatic, aromatic and gas species at various flash-pyrolysis temperatures from inertinite-rich coal Devolatisation profiles (including quantitative mass loss and product distribution) can be constructed for both coals using flash-pyrolysis GCMS. Pyrolysis started between 400-500 °C with the release of liquid hydrocarbons (similar amounts of aliphatic and aromatic compounds). The maximum pyrolytic evolution was at 600 °C with the majority of the pyrolytic products as liquid hydrocarbons (similar amounts of aliphatic and aromatic compounds). The main difference between vitrinite-rich and inertinite-rich coal are observed in the 700 °C devolatisation region. For both coals, at 700 °C mostly aromatic hydrocarbons and gaseous products were released. The inertinite-rich coal, however, released more aromatic hydrocarbons at this temperature than the vitrinite-rich coal (4.1 % for vitrinite-rich coal versus 8.4 % for the inertinite-rich coal). At 700 °C small amounts of aliphatic hydrocarbons (~ 2 %) was still released in the inertinite-rich coal. From 800 °C only gas was formed for both coals. Therefore, conducting pyrolysis above 700 °C for these two coals is a waste of energy if the goal is to produce liquid hydrocarbons. In general, the devolatisation behavior of these two maceral-different coals was similar.
The total hydrocarbons distribution evolved during flash-pyrolysis was determined and is shown in Figure 4. This total hydrocarbon distribution show the full liquid product produced up to 1000 °C.
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Oviedo ICCS&T 2011. Extended Abstract 30.0
25.0
Vitrinite-rich coal Inertinite-rich coal
% Abundance
20.0
15.0
10.0
5.0
0.0
Figure 4: Total tar hydrocarbon products from flash-pyrolysis (excluding gasses)
Flash-pyrolysis GCMS showed that the inertinite-rich coal in general was more aromatic; this was evident by the higher abundance of alkylbenzene, alkylnaphthalene and polyaromatic pyrolytes. The aromatic nature of inertinite was also confirmed by the elemental analysis (low atomic H/C ratio) and literature [7]. The abundance of aliphatic species (alkanes, alkenes and dienes) are assumed to be contributions from both the vitrinite and liptinite present in the coals [2,3]. Nip et al. and Hartgers et al. showed from pyrolysis of pure macerals that vitrinite typically has a high abundance of alkylphenols [2,3]. This trend was also observed in the South-African vitrinite-rich coal. The inertinite-rich coal resembles the pyrolyte profile of vitrinite, but with a greater aromatic character (greater abundance of alkylbenzenes, alkylnaphthalenes, and polyaromatic hydrocarbons). The inertinite-rich coal did not follow all the trends observed by Nip et al. and Hartgers et al. for semifusinite and fusinite [2,3]. For the inertinite-rich coal it was expected that the alkylnaphthalene and alkylbenzene hydrocarbon contribution would be greater [2,3]. The polyaromatic hydrocarbon contribution did follow expected trends: The inertinite-rich coal had a greater polyaromatic hydrocarbon distribution than the vitrinite-rich coal. The inertinite macerals present in the inertinite-rich coal was mostly low- and high-reflecting semifusinite and inertodetrinite. Low-reflecting inertinite is reactive and will have a similar composition to vitrinite [5,9]. Therefore, the similar pyrograms profiles for these 7
Oviedo ICCS&T 2011. Extended Abstract
two maceral-different coals may be due to the high abundance of reactive inertinites present in these samples (inertinites with similar reflectance than the corresponding vitrinite). Except for aromatic mixtures and polyaromatic hydrocarbons, the two coals were similar in pyrolytic composition. This similarity suggests a common origin of these groups of pyrolytes in these two maceral-different coals. This similarity in pyrolyte distribution was also observed in previous work [7]. 4. Conclusions In the current study the difference in pyrolysis hydrocarbon composition was determined for two maceral-different Permian-aged South African coals: A vitrinite-rich and an inertinite-rich coal. A weighed amount of coal was sequentially pyrolyzed from 300 °C to 1000 °C with 100 °C increments and analyzed using a GCMS. The chromatograms were analyzed and compounds were grouped according to representative functionalities. For both coals, tar evolution started at 400-500 °C and terminated at 700 °C. The majority of the tar evolved from both coals was at 600 °C. The tar evolved consisted of various types of aromatic and aliphatic hydrocarbons. This work showed the effect of maceral composition (inertinite-rich versus vitrinite-rich) and temperature on flash-pyrolysis hydrocarbon yield and composition. In general, the devolatisation behavior and hydrocarbon composition of these two maceral-different coals was similar. References [1] Tylor GH, Teichmuller M, Davis A, Diessel CFK, Littke R, Robert P, Organic petrology, Gebruder Borntraeger. 1st edition, Berlin, Germany; 1998. [2] Nip M, de Leeuw JW, Chemical structure of bituminous coal and its constituting maceral fractions as revealed by flash pyrolysis. Energy & Fuels 1992;6:125-136. [3] Hartgers WS, Sinninge Damste JS, de Leeuw JW, Molecular characterization of flash pyrolysates of two Carboniferous coals and their constituting maceral fractions. Energy & Fuels 1994;8:1055-1067. [4] Stout SA, Lasers in organic petrology and organic geochemistry, II. In-situ laser micropyrolysis-GCMS of coal macerals. International Journal of Coal Geology 1993;24:309-331. [5] Snyman CP, The role of coal petrography in understanding the properties of South African coal, International Journal of Coal Geology 1989;14:83-101.
8
Oviedo ICCS&T 2011. Extended Abstract [6] Maroto-Valer MM, Andresen JM, Snape CE, Verification of the linear relationship between carbon aromaticities and H/C ratios for bituminous coals. Fuel 1998;22:783-785. [7] Van Niekerk D, Pugmire RJ, Solum MS, Painter PC, Mathews JP, Structural characterization of vitrinite-rich and inertinite-rich Permian-aged South African bituminous coals. International Journal of Coal Geology 2008;76:290–300. [8] Van Niekerk D, Mathews JP, Molecular representations of Permian-aged vitrinite-rich and inertinite-rich South African coals. Fuel 2010;89:73-82. [9] Kershaw JR, Taylor GH, Properties of Gondwana coals with emphasis on the Permian coals of Australia and South Africa. Fuel Processing Technology 1992;31:127-168.
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Oviedo ICCS&T 2011. Extended Abstract
Wet oxidation of anthracene oil-based pitch – a way to porous carbons
H. Machnikowska, K. Torchała, G. Gryglewicz, J. Machnikowski Wrocław University of Technology, Division of Polymer and Carbonaceous Materials Gdanska 7/9, 50-344 Wroclaw, Poland E-mail:
[email protected] Abstract The oxidation of anthracene oil-based pitch with nitric acid, nitric/sulphuric acids mixture, potassium permanganate and sulphuric acid/hydrogen peroxide mixture has been studied as a possible alternative to the common process of stabilization in air. The effect of oxidation was monitored by weight uptake, elemental analysis and FTIR. The oxidized pitch was carbonized at 900oC and then activated with steam and CO2 up to 50 wt.% burnoff. The results show that all the oxidation agents are effective in suppressing pitch plasticity by the treatment for 0.5-1 h at room temperature. The wet oxidation results in different, compared to stabilization in air at 300oC, oxygen functionalities and somewhat lower coking yield (72-77 vs. 81 wt.%). Enhanced nitrogen content is characteristic of pitch treated with HNO3 and HNO3/H2SO4. The well developed ultramicroporosity of resultant cokes, determined by CO2 adsorption at 273K, proves that structural ordering has been completely destroyed as a result of the oxidative treatment of pitch. Wet and air oxidized pitches show comparable propensity to porosity development on activation. Clearly, CO2 is more effective in activating pitch derived chars than steam, BET surface area in the range of 1600 - 1800 m2/g can be measured from N2 adsorption isotherms.
1. Introduction Oxidative treatment to suppress pitch plasticity is a crucial step in the preparation of pitchbased carbon fibres [1]. It has been also reported that the process can be used for producing pitch-derived hard carbons with properties which are suitable for molecular sieves [2] and anodes of lithium-ion batteries [3]. The conventional air oxidation is a long lasting process and requires carefully adjusted conditions due to a narrow window between the temperature of sufficient reactivity with oxygen and the softening point of the precursor. Developing an alternative method allowing to perform the stabilization at a lower temperature and a shorter time is therefore of interest for both carbon fibres and carbon particulates. Earlier works using
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Oviedo ICCS&T 2011. Extended Abstract
low softening point petroleum pitch fibres [4] and coal-tar pitch spheres [5] showed the efficiency of nitric acid as a possible oxidizing agent. In this paper we present the study on the stabilization of anthracene oil-based pitch by the treatment with different oxidizing agents with respect to the suitability of resultant chars to the porosity development on subsequent activation with steam and CO2.
2. Experimental section The starting material for the study was anthracene oil-based pitch (AOP) of softening point (Mettler) of 244oC which was supplied by Industrial Quimica del Nalon, S.A. The sample was ground to pass the sieve 100 μm with a mean particle size of 40 μm. Oxidation agents used for AOP oxidation were: HNO3; H2SO4 + HNO3; H2SO4 + H2O2 and KMnO4 in 0.5M H2SO4 solution. The oxidants used were adapted from the studies on nanotubes purification/functionalization [6,7], however, the process variables had to be adjusted for this work. The AOP stabilization is very sensitive to the change of concentration, mixing ratio and reaction time as well as particle size distribution of pitch. The procedure that was developed to assess the effect of oxidation consisted of monitoring the weight uptake, oxygen content, surface chemistry (FTIR) and the carbonization behaviour. The latter included the yield, appearance and particle size distribution of cokes produced at 900oC. The oxidation conditions set for a given oxidizing agent to produce non-agglomerated char on heat-treatment and symbols of resultant stabilized pitches are given below: 38% HNO3, 1 h; 1g AOP / 50 ml solution
- AOP-N
60% H2SO4/38% HNO3 3:1, 0.5 h; 1g AOP / 50 ml solution - AOP-SN 74%H2SO4/30%H2O2 4:1, 1 h; 1g AOP / 100 ml solution
- AOP-SH
KMnO4, KMnO4/pitch 2:1, 0.5 h
- AOP-Mn
All treatments were performed at room temperature. The chars produced from successfully stabilized pitches were characterized by CO2 adsorption at 273K to measure ultramicroporosity development and next were activated with steam at 850oC and with CO2 at 950oC up to 50 wt.% burnoff. Porosity development in resultant activated carbons was assessed from N2 adsorption isotherms at 77K (ASAP 2020, Micromeritics). The sample AOP-ox, i.e. AOP which was stabilized in air according to the heating program from [1] was used as the reference material.
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Oviedo ICCS&T 2011. Extended Abstract
3. Results and Discussion The characteristics of AOP stabilized in different media under selected conditions are given in Table 1. Table 1. Effect of oxidation agent on pitch properties and carbonization behaviour Sample AOP-N AOP-SN AOP-SH AOP-Mn AOP-ox AOP
Weight uptake Δm [%] 11.0 13.3 5.0 -17.2 5.3
Oxygen content1 [%] 10.2 12.5 20.3 4.5 10.7 2.6
Nitrogen content1 [%] 3.8 4.5 1.4 1.3 1.4 1.5
Coking yield 2 [%] 77.4 77.7 70,7 72.6 81.2 65.2
VDR CO2 3
S0 CO2 4
[cm3/g] 0.21 0.19 0.22 0.25 0.27 0.02
[m2//g] 568 561 646 633 723 36
1
elemental analysis, direct determination 900oC for 0.5 h 3 ultramicropore (<0.7 nm) volume in coke from Dubinin-Radushkevich equation 4 surface area of ultramicropores (<0.7 nm) in coke. 2
The data show that under selected conditions the stabilization of AOP could be performed successfully with all studied oxidizing media. As evidenced by the appearance and particle size distribution of cokes, the stabilized materials practically neither agglomerate nor swell on the heat treatment. The mean particle size of AOP, about 40 μm, has been preserved in cokes. In contrast to the AOP-based coke, those from stabilized pitches are characterized by considerable ultramicroporosity development, corresponding to the ultramicropore volume VDR CO2 in the range of 0.19 – 0.27 cm3/g. However, different stabilization agents generate also some essential differences in the properties of stabilized pitches. Low weight uptake of AOP-SH and weight loss of AOP-Mn suggest that a partial burnt-off of pitch can occur on the respective treatment. A serious problems with filtration in the latter case is an evidence of partial disintegration of pitch particles on oxidation with KMnO4. Stabilization with nitric acid containing solution results in the noticeable nitrogen incorporation. In all cases the stabilization enhances carbonization yield compared to that from AOP, however the extent depends clearly on the oxidation agent, being the highest for AOP-ox. FTIR spectra (Fig. 1) reveal the significant effect of reactant on the mechanism of stabilization. Lack of clear differences between the spectra of AOP-N and AOP-SN proves that the stabilization with HNO3 and H2SO4 + HNO3 occurs in a similar way. Strong bands from
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Oviedo ICCS&T 2011. Extended Abstract
–NO2 groups at 1520 and 1320 cm-1 indicate the important role of nitration in the process. Very intensive signal from C=O group at ~1720 cm-1 and a broad band of C-O between 1200 and 1300 cm-1 in the AOP-SH spectrum can be attributed to the numerous carboxylic and etheric functionalities. AOP-Mn spectrum is distinguished by less pronounced absorption from oxygen functionalities. This suggests a weaker effectiveness of KMnO4 in oxygen groups generation.
AOP-N AOP-SN AOP-SH AOP-Mn AOP-ox
AOP 3500
3000
2500
2000
1500
1000
500
-1
Wavenumbers (cm )
Fig. 1. FTIR spectra of AOP stabilized using different oxidizing agents. Fig. 2 presents N2 adsorption isotherms of activated carbons produced by steam and carbon dioxide activation of stabilized pitches up to comparable burnoff of about 50 wt.%. Tables 2 and 3 give the porous texture parameters calculated from the isotherms. 600
600
b
400
400
3
V N2 (STP) [cm /g]
500
3
VN2 (STP)[cm /g]
a 500
300
AOP-N-S
200
300
AOP-N-C AOP-SN-C
200
AOP-SN-S
AOP-SH-C
AOP-SH-S 100
AOP-Mn-C
100
AOP-Mn-S
AOP-ox-C
AOP-ox-S 0
0
0
0.2
0.4
0.6
0.8
1
p/po
0
0.2
0.4
0.6
0.8
1
p/po
Fig. 2. N2 adsorption isotherms of steam (a) and CO2 (b) activated carbons from AOP powders stabilized using different activation agents. The analysis shows that the steam and CO2 activated carbons produced from all chars are Submit before May 15th to
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Oviedo ICCS&T 2011. Extended Abstract
typically microporous carbons. The chars from wet oxidized AOP have comparable propensity to porosity development as that from pitch oxidized in air. All steam activated carbons show similar porous texture characteristics. Activation using CO2 leads to a better developed porosity. In dependence on the stabilization method the BET surface area varies from 1400 to 1850 m2/g and the micropore volume from 0.5 to 0.7 cm3/g with the highest values for AAOP-SN-C and A-AOP-ox-C. However, it must be noted that the chars derived from oxidized AOP show much lower reactivity towards CO2 compared to conventional precursors. Table 2. Porous texture parameters of steam activated carbons from oxidized AOP pitches Sample A-AOP-N-S A-AOP-SN-S A-AOP-SH-S A-AOP-Mn-S A-AOP-ox-S
SBET
VT 3
VDR 3
L0
2
[m /g]
[cm /g]
[cm /g]
[nm]
1123 1101 1127 1162 1246
0.48 0.47 0.48 0.52 0.53
0.44 0.43 0.45 0.46 0.48
1.20 1.15 1.15 1.23 1.22
VDR/VT
VDR CO2 3
0.92 0.93 0.94 0.88 0.92
S0 CO2
[cm /g]
[m2/g]
0.19 0.17 0.17 0.24 0.18
491 436 452 610 467
SBET – BET surface area VT – total pore volume VDR – micropore volume from Dubinin Raduschkevitch equation L0 – mean micropore width
Table 3. Porous texture parameters of CO2 activated carbons from oxidized AOP pitches Sample A-AOP-N-C A-AOP-SN-C A-AOP-SH-C A-AOP-Mn-C A-AOP-ox-C
SBET
VT 3
VDR 3
L0
2
[m /g]
[cm /g]
[cm /g]
[nm]
1569 1861 1393 1597 1792
0.67 0.79 0,63 0.67 0.76
0.59 0.70 0,52 0.61 0.67
1.30 1.31 1,47 1.31 1.34
VDR/VT
VDR CO2 3
0.88 0.88 0,83 0.90 0.88
S0 CO2
[cm /g]
[m2/g]
0.23 0.21 0,16 0.21 0.22
567 511 382 509 521
4. Conclusions Anthracene oil-based pitch in the form of fine powder (< 100 μm) was successfully stabilized using oxidation with the solution of HNO3, H2SO4/HNO3 or H2SO4/H2O2 under individually selected conditions. Similar extent of oxidation but different oxygen functionalities were generated using a given reactant. Chars derived from wet oxidized pitches are characterized by ultramicroporous texture and high propensity to porosity development on physical activation. Microporous activated carbons of comparable pore volume, surface area and mean micropore width were produced using char from wet and air oxidized pitch. Clearly, CO2 is more
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Oviedo ICCS&T 2011. Extended Abstract
effective activating agent for pitch derived chars than steam. BET surface area as high as 1800 m2/g could be developed both from the air and wet stabilized pitch at the burnoff of about 50%. The wet oxidation with HNO3, H2SO4/HNO3 or H2SO4/H2O2 seems to be a possible alternative way of stabilization to prepare porous carbons in the form of fibres or powders from anthracene oil-based pitch. Acknowledgements The research leading to these results has received funding from the European Union’s Research Fund for Coal and Steel (RFCS) research programme under grant agreement No RFCR-CT-2009-0004.
References [1] Bahl OP, Shen Z, Lavin JG, Ross RA. Manufacture of carbon fibers. In: Donnet JB, Wang TK, Rebouillat S, Peng JCM, editors. Carbon Fibers, New York: Marcel Dekker, Inc; 1998, p. 1-83. [2] Vilaplana–Ortego E, Alcaniz-Monge J, Cazorla-Amoros D, Linares-Solano A. Effect of the stabilisation time of pitch fibres on the molecular sieve properties
of carbon fibres.
Microporous Mesoporous Mater 2008;109:21-7. [3] Fujimoto H, Tokumitsu K, Mabuchi A, Chinnasamy N, Kasuh T. The anode performance of the hard carbon for the lithium battery derived from the oxygen-containing aromatic precursors. J.Power Sources 2010;195:7452-6. [4] Vilaplana–Ortego E, Alcaniz-Monge J, Cazorla-Amoros D, Linares-Solano A. Stabilisation of low softening point petroleum pitch fibres by HNO3. Carbon 2003;41:1001-7. [5] Liu X, Liang X, Liu C, Zhang L, Qiao W, Ling L. Pitch spheres stabilized by HNO3 oxidation and their carbonization behavior. New Carbon Materials 2010;25:29-34. [6] Rasheed A, Howe JY, Dadmun MD, Britt PF. The efficiency of the oxidation of carbon nanofibers with various oxidizing agents. Carbon 2007;45:1072-80. [7] Wespasnick KA, Smith BA, Schrote KA, Wilson HK, Diegelmann SR, Fairbrother DH. Surface and structural characterization of multi-walled carbon nanotubes following different oxidative treatments. Carbon 2011;49:24-6.
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Oviedo ICCS&T 2011. Extended Abstract
Adsorption of phenol on nitrogen-enriched activated carbons prepared from coal-tar pitch and polymers E. Lorenc-Grabowska1, G. Gryglewicz1, M.A. Diez2, C. Barriocanal2 1
Wrocław University of Technology, Division of Polymer and Carbonaceous Materials, Gdańska 7/9, 50-344 Wrocław, Poland E-mail:
[email protected] 2 Instituto Nacional del Carbón, INCAR-CSIC, Apartado 73, 33080 Oviedo, Spain Abstract The adsorption of phenol on the activated carbon (AC) of enhanced nitrogen content was studied. The ACs were prepared from polyacrylonitrile (PAN) and a mixture of PAN and coal-tar pitch (CTP) by carbonization at 800 ºC followed by activation with steam at 850 ºC. For comparison purposes, the AC produced from a mixture of polyethylene terephthalate (PET) and CTP was used. The ACs were characterized by a similar porous texture but a different nitrogen content, in the range of 0.75-7.42 wt%, implying the differences in the basicity of the carbon surface. The phenol adsorption was carried out in static conditions at ambient temperature. The pseudo-second order kinetic model and Langmuir model were found to fit the experimental data very well. The adsorption of phenol on nitrogen-enriched ACs is characterized by a very high affinity of adsorbate towards the carbon surface. However, an enhanced content of nitrogen improves the adsorption capacity of AC only insignificantly. 1. Introduction The adsorption of phenols on ACs depends on many factors such as adsorbent properties (i.e. porous texture, surface chemistry), adsorbate characteristics (i.e. molecular size, solubility, pKa) and operational conditions (i.e. pH, temperature) [1-4]. Three mechanisms of phenolic compounds adsorption are considered, i.e. the π-π dispersion interaction, the electron-donor-acceptor complex formation and the hydrogen-bond formation. It is also postulated that ACs with basic surface characteristics are most suitable for the removal of weak acids such as phenol. However, most of the studies on the adsorption of phenols were carried out for non-modified, oxidized or heat treated ACs. There are only few reports concerning the adsorption on nitrogen enriched carbons which are characterized by basic surface properties [5]. One of the methods of preparing nitrogen-enriched porous materials is the carbonization followed by activation of
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Oviedo ICCS&T 2011. Extended Abstract
nitrogen-containing polymers as a precursor [5-6]. The aim of this work was to determine the mechanism and kinetics of adsorption of phenol in aqueous solution on nitrogen-enriched ACs. The role of nitrogen functionalities in the removal of phenols from water is discussed. 2. Experimental section 2.1. Materials The ACs were produced from polyacrylonitrile (PAN) and a mixture of PAN/coal tar pitch (CTP) at a ratio of 1:1 w/w by carbonization and subsequent steam activation at 850 oC [7]. For comparison purposes, an activated carbon prepared from a mixture of polyethylene terephthalate (PET) and CTP (1:1 w/w) under the same experimental conditions was used. In order to receive the ACs with a similar porous texture, the burnoff was 60, 50 and 40%, respectively. 2.2. Characterization of ACs The elemental analysis of C, H, N and S was performed using a Vario III Elemental Analyzer. The oxygen content was calculated by difference. The porous texture was determined from nitrogen adsorption isotherms measured at 77 K with a NOVA 2200 (Quantachrome) instrument. The specific surface area was calculated using the BET method. The amount of nitrogen adsorbed at relatively pressure of ppo-1 = 0.98 was employed to determine the total pore volume (VT). The micropore volume (VDR) was calculated applying the Dubinin-Radushkevich equation up to ppo-1 ≤ 0.15. The pore size distribution was determined by means of the DFT method. The pHPZC (point of zero charge) of the ACs was determined according to the procedure described by MorenoCastilla et al. [8]. Base-acid titration based on Boehm’s method was performed to measure the total amount of basic and acidic surface groups of carbons using solutions of NaOH and HCl. 2.3. Sorption of phenol The adsorption of phenol from aqueous solutions was carried out at 24oC in a static system. For the adsorption, 0.01-0.2 g of AC was placed into Erlenmeyer flasks and 0.10 dm3 of adsorbate solution (150 mg/dm3) was added to each flask. The stoppered flasks were kept in a thermostat shaker bath and were agitated to reach equilibrium. Each set of flasks included two flasks containing blank solutions to check for sorbate volatilization and adsorption on the glass walls. The adsorption isotherms were determined without
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Oviedo ICCS&T 2011. Extended Abstract
adding any buffer to control pH to avoid the presence of a new electrolyte in the system. The pH of the solution was measured by a digital pH-meter (Accumet Basic, Fisher Scientific) using a glass electrode. The concentration of each solute remaining in the water phase was quantified. The phenol concentration was determined using a HITACHI U-2800A UV-Vis spectrophotometer at the wavelength of 270 nm. 3. Results and discussion 3.1. Characteristics of ACs The elemental composition, porous texture parameters and surface properties of the ACs studied are given in Table 1. As can be seen the different extent of burn-off allowed to produce a series of ACs with comparable porous texture. The ACs are microporous in their nature as it is reflected by the 95% of the total pore volume. The BET surface area is between 720 and 760 m2/g. The pore size distribution is comparable for all the studied carbons as shown in Fig. 1. Table 1. Characteristics of ACs Parameter d
C Hd Nd Sd Od pHPZC basic acidic SBET, m2/g VT, cm3/g VDR, cm3/g VDR/VT
A-PAN
A-PAN/CTP
Ultimate analysis (wt%) 90.16 87.16 2.07 2.39 4.56 7.42 0.13 0.00 3.08 3.03 8.82 7.83 Boehm functional groups (mmol/g) 1.43 1.22 0.36 0.25 Porous texture parameters 740 730 0.307 0.302 0.292 0.286 0.95 0.95
A-PET/CTP 94.17 2.86 0.75 0.13 2.09 7.36 1.21 0.23 760 0.323 0.309 0.96
The nitrogen content of the ACs ranges from 0.75 wt% for A-PET/CTP to 7.42 wt% for A-PAN. The ACs enriched in nitrogen contain comparable amount of oxygen, i.e. around 3 wt%. All ACs studied are ash free therefore the surface properties are related to the chemical structure of the carbon surface. The pHPZC indicates the basic character of the surface, that is confirmed by the predominance of basic functional groups (Table 1).
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Oviedo ICCS&T 2011. Extended Abstract 0.07
A PAN
0.07
A PAN/CTP
0.06
0.05
0.05
0.05
0.04 0.03 0.02
0.04 0.03 0.02
0.01
0.01
0
0
0.2
1.0
1.8
2.6
3.4
4.2
5.0
A PET/CTP
3
3
pore volume, cm /g
0.06 pore volume, cm /g
0.06
3
pore volume, cm /g
0.07
0.04 0.03 0.02 0.01 0
0.2
pore width, nm
1.0
1.8
2.6
3.4
4.2
5.0
0.2
pore width, nm
1.0
1.8
2.6
3.4
4.2
5.0
pore width, nm
Fig. 1. Pore size distribution determined by DFT method. The basic character of A-PET/CTP arises mainly from the delocalized π electrons of the graphene layers [9], whereas in the case of nitrogen-enriched ACs the basicity is also due to the nitrogen functionalities such as pyridinic nitrogen [5]. 3.2. Adsorption of phenol from aqueous solution 3.2.1. Kinetics of adsorption The results of the adsorption of phenol on the ACs studied vs. time are shown in Fig. 2. The process is found to be rapid at the initial period and after 6 h over 50% of the equilibrium sorption capacity is obtained. The adsorption is the fastest for A-PET/CTP. The time needed to reach the equilibrium for this carbon is four-fold shorter (around 48 h) than for nitrogen-enriched ACs (90-120 h). The amount of phenol adsorbed at equilibrium increases slightly in the direction: A-PET/CTP
qt, mg/g
160 140 120 100 80 60 40
A-PAN
20
A-PET/CTP
A-PAN/CTP
0 0
50
100
150
200
time, h
Fig. 2. Effect of contact time on the extent of phenol adsorption. Several kinetics models are used to examine the kinetics of phenol adsorption. The pseudo-first (Lagergren) and pseudo-second order model are the mostly used equations [10-12]. The Lagergren equation is given by Eq. (1) whereas the second order model is given by Eq. (2):
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Oviedo ICCS&T 2011. Extended Abstract
log(qe exp-qt)=logqe-(k1 t)/2.303
(1)
t/qt=1/(k2 qe2)+1/qet
(2)
where t and qt are, respectively, time (min) and the amount of phenol adsorbed by carbon at time t (mg g-1); qe exp and qe are the amount of phenol adsorbed at equilibriumexperimental data and equilibrium-calculated data, respectively, expressed as mg g-1 sample; and k1 and k2 are the first (min-1) and second (g min-1 mg-1) order rate constant of adsorption. It was reported that the adsorption of phenols usually fits the pseudo-second kinetic model [10-12]. Both models have been applied in our work to describe the adsorption of phenol on nitrogen-enriched ACs. The results are given in Table 2. The correlation coefficient (R2) for the Lagergren equation appears to be low, ranging from 0.829 to 0.966. Furthermore, the experimental qe exp values do not agree well with the calculated ones. Whereas the linear plot of t/qt vs. time shows a very good agreement with experimental data with R2 close to 1 indicating that the pseudo-second order model is favorable for the phenol adsorption. As can be seen in Table 2, the calculated qe values agree with the experimental data very well. The k2 rate constant increases in the direction: A-PAN
First-order kinetic model qe k1 R2 qe exp (mg/g) (mg/g) (1/min) 158 113 0.0005 0.938 149 74 0.0004 0.966 148 47 0.0003 0.829
Second-order kinetic model qe k2 R2 (mg/g) (g/mg min) 1.41 10-5 161 0.997 2.15 10-5 149 0.996 -5 4.13 10 147 0.999
3.2.2. Equilibrium adsorption of phenol The adsorption isotherms for phenol on the ACs studied are shown in Fig. 3. According to Giles classification, the isotherm belongs to the type L for A-PET/CTP and type H for A-PAN and A-PAN/CTP carbons. The Langmuir class of isotherms is commonly reported for the adsorption of phenols from aqueous solution [1-5,10-12]. This is the case when no strong competition between the adsorbate and the solvent for occupying the adsorption sites is observed. The type H isotherm is attributed to an extremely strong adsorption at very low concentration that is unique in the case of adsorption from
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Oviedo ICCS&T 2011. Extended Abstract
aqueous solution. Langmuir (Eq. 3) and Freundlich (Eq. 4) models were used to interpret the equilibrium adsorption isotherms: qe=(b qmax ce)/(1+b ce)
(3)
qe=Kf ce1/n
(4)
where qe is the phenol amount adsorbed by carbon at equilibrium (mg/g), ce is the equilibrium phenol concentration in solution (mg/dm3), qmax is the monolayer capacity of the adsorbent (mg/g), b is the Langmuir adsorption constant (dm3/mg), Kf is the Freundlich constant (mg1-n dm3n/g), and 1/n is the heterogeneity factor. The calculated values of the Langmuir and Freundlich equation’s parameters are given in Table 3. 200
qe, mg/g
160 120 80 A-PAN
40
A-PAN/CTP A-PET/CTP
0 0
40
80
120
160
3
ce, mg/dm
Fig. 3. Equilibrium adsorption isotherms for phenol on ACs. Table 3 contains also the essential characteristics of the Langmuir equation that can be expressed in terms of a dimensionless separation factor RL, which is defined as: RL=1/(1+b co)
(5)
where c0 is the initial phenol concentration (mg/dm3) and b is the Langmuir constant (dm3/mg). The value of RL indicates the shape of the isotherm to be unfavourable (RL>1), linear (RL=1), favourable (0< RL<1) or irreversible (RL=0). The comparison of the R2 values of the linearized form of both equations indicates that the Langmuir model yields a better fit than the Freundlich model. The Langmuir monolayer capacity slightly increases as follows: A-PET/CTP<APAN/CTP
Submit before 31 May 2011 to [email protected]
6
Oviedo ICCS&T 2011. Extended Abstract
were between 0 and 1, which represents a favourable adsorption on all the ACs studied. It is worth noting that the RL of nitrogen-enriched ACs is significantly lower than that of A-PET/CTP, indicating a much stronger interaction between phenol and activated carbons of enhanced content of nitrogen. Table 3. Langmuir and Freundlich parameters for phenol adsorption on ACs Activated carbon A-PAN A-PAN/CTP A-PET/CTP
qmax (mg/g) 167 161 160
Langmuir b RL 3 (dm /g) 0.216 0.023 0.180 0.036 0.069 0.088
R2 0.998 0.999 0.999
Freundlich Kf n 1-n 3n (mg dm /g) 85 0.138 82 0.140 42 0.270
R2 0.922 0.969 0.901
4. Conclusions A considerable improvement of the affinity of phenol towards the carbon surface was observed for nitrogen-enriched activated carbons. With increasing the nitrogen content, the basicity of carbon surface increases. However, an enhanced content of nitrogen leads only to an insignificant increase in the adsorption capacity. This implies that the adsorption capacity of AC is mainly determined by its porous texture. The kinetics of phenol adsorption was very fast in the initial period of adsorption. The pseudo-second kinetic model describes the adsorption process of phenol on nitrogen-enriched activated carbons very well. The presence of nitrogen functionalities involves stronger interactions between the phenol molecule and the carbon surface. Acknowledgements This work was financially supported by Ministry of Science and Higher Education (N N523 557938) and Wrocław University of Technology (internal grant 344045/Z0309) and the CSIC Intramural Project (200480E616). The NOVAPET Company (Spain) which provides commercial PET is also acknowledged. References [1] Dąbrowski A., Podkościelny P., Hubicki Z., Barczak M. Adsorption of phenolic compounds by activated carbon-a critical review. Chemosphere 2005;58:1049-70. [2] Moreno-Castilla C. Adsorption of organic molecules from aqueous solution on carbon materials. Carbon 2004;42:83-94. [3] Salame I.I., Bandosz T.J., Role of surface chemistry in adsorption of phenol on
Submit before 31 May 2011 to [email protected]
7
Oviedo ICCS&T 2011. Extended Abstract
activated carbons. J. Colloid Interf. Sci. 2003;264:307-12. [4] Terzyk A.P., Further insights into the role of carbon surface functionalities in the mechanism of phenol adsorption. J. Colloid Interf. Sci. 2003;268: 301-29. [5] Lorenc-Grabowska E., Gryglewicz G., Machnikowski J. p-Chlorophenol adsorption on activated carbons with basic surface characteristics. Appl.Surf.Sci 2010;256:4480-87. [6] Lota G., Grzyb B., Machnikowski H., Machnikowski J., Frąckowiak E., Effect of nitrogen in carbon electrode on the supercapacitor performance. Chem.Physics Lett. 2007;404:53-8. [7] Lorenc-Grabowska E., Gryglewicz G., Machnikowski J., Diez M.A., Barriocanal C., Activated carbons from coal/pitch and polyethylene terephthalate blends for the removal of phenols from aqueous solutions. Energy Fuels 2009;23:2675-83. [8] Moreno-Castilla C., Lopez-Ramon M.V., Carrasco-Marin F. Changes in surface chemistry of activated carbon by wet oxidation. Carbon 2000;38:1995-2001. [9] Laszlo K., Szűsc A. Surface characterization of polyethyleneterephthalate (PET) based activated carbon and the effect of pH on its adsorption capacity from aqueous phenol and 2,3,4-trichlorophenol solution. Carbon 2001;39:1945-53. [10] Qian Q., Chen Q., Machida M., Tatsumoto H., Mochidzuki K., Sakoda A., Removal of organic contaminants from aqueous solution by cattle manure compost (CMC) derived activated carbons. Apply. Surf. Sci. 2009;255: 6107-14. [11] Mohanty K., Das D., Biswas M.N., Preparation and characterization of activated carbons from Sterculia alata nutshell by chemical activation with zinc chloride to remove phenol from wastewater. Adsorption 2006;12:119-32. [12] Hameed B.H., Chin L.H., Rengaraj S., Adsorption of 4-chlorophenol onto activated carbon prepared from rattan sawdust. Desalination 2008;225:185-98.
Submit before 31 May 2011 to [email protected]
8
Structural evolution upon thermal teatment of two South African anthracites as studied by XRD Marley Vanegas-Chamorro1,2,*, Katia Tamargo-Martínez1, Jorge Xiberta2, Amelia Martínez-Alonso1 and Juan M. D. Tascón1 1
Instituto Nacional del Carbón, INCAR-CSIC, Apartado 73, 33080 Oviedo, Spain
2
Departamento de Energía, Univ. de Oviedo, Independencia 13, 33004 Oviedo, Spain
* Corresponding author e-mail address: [email protected]
Abstract The effects of mineral matter and heat treatment temperature (HTT) on the graphitization degree of two South African anthracites (Hlobane and Somkhele Duff) were studied using X-ray diffraction (XRD). Raw and acid-demineralized anthracites were exposed to several different HTTs in the 900-2800 ºC interval under argon atmosphere. From the XRD patterns, the interlayer spacing (d002) and crystallite sizes along the c and a axes (Lc and La, respectively) were determined. The three parameters evolved in parallel with increasing HTT. Mineral matter was volatilized during the thermal treatment and had a clear catalytic effect on graphitization at temperatures below 2400 ºC, especially for crystallite growth along the a axis. Under equivalent conditions, Hlobane anthracite exhibited a higher degree of graphitization than Somkhele Duff anthracite, which can be attributed to a better degree of mixing between the mineral and organic constituents in the former coal.
1. Introduction The main objective of this work is to contribute to the development of reductants for Fe-Cr and Fe-Mn ores in submerged arc furnaces. It has been reported [1] that this process requires carbon materials with a high degree of graphitic order, and that these can be obtained from anthracites by high temperature treatments in an inert atmosphere. Accordingly, in order to broaden the scope of feedstock materials utilizable for this purpose, two South African anthracites have been submitted in this work to heat treatment at several temperatures in the 900-2800 ºC interval.
On the other hand, as it is known that the minerals present in anthracites generally have a strong catalytic effect on their graphitization [2,3], the inorganic components of the studied anthracites have been characterized and their effect has been investigated by comparing the structural evolution of raw and acid-demineralized anthracites as established by XRD.
2.
Experimental section
Materials. The starting materials were two anthracite samples from Holbane (H) and Somkhele Duff (S), both from South Africa, ground to <0.5 mm. Demineralization treatments with HCl and HF were performed following the ISO 602 standard. A batch of raw or demineralized sample was devolatilized at 900 ºC in Ar. Portions of it were heattreated at several temperatures in the 900-2800 ºC interval under an Ar atmosphere. Ssample identification codes used here contain the initial of the anthracite name (H or S), followed or not by D (i.e., demineralized or not). Methods. X-ray diffraction measurements were made in a Bruker D8 Advance diffractometer using Cu Kα radiation. The diffractograms were taken in the 5-95 º (2θ) interval with a step size of 0.02 º, step number of 4500 and step time of 2 s. The interlayer spacing, d002, was evaluated from the (002) peak position applying the Bragg equation. The average crystallite sizes Lc and La were calculated from the (002) and (100) peaks, respectively, by applying the Scherrer formula with K values of 0.9 for Lc and 1.84 for La.
3. Results and discussion Table 1 reports the results of proximate and elemental analyses of the starting anthracites, carried out following the corresponding ISO standards. Both anthracites, especially Hlobane, have a considerable ash content. Ash analyses, carried out by atomic absorption spectroscopy, indicated that Si, Ca and Na are more abundant in Hlobane than in Somkhele Duff, the opposite occurring with Fe. Mineral matter was characterized in the two starting anthracites by low-temperature ashing (LTA) followed by XRD, FT-IR spectroscopy and SEM-EDX. The mineral matter contents determined by LTA were 18.4 wt.% for Hlobane and 11.1 wt.% for Somkhele Duff. Minerals identified in these anthracites were principally quartz, illite, kaolinite, pyrite and calcite. Mineral matter is better crystallized and has a large particle size in Somkhele Duff anthracite than in Hlobane anthracite.
As can be deduced from Table 1, the demineralization treatments were effective in removing mineral matter since removal percentages of 97 wt.% (Hlobane) and 90.4 wt.% (Somkhele Duff) were attained. Control of mass loss during the thermal treatments and examination of mineral matter peaks in the X-ray diffractograms of samples heat-treated at progressively high temperatures evidenced a progressive loss of mineral matter by volatilization.
Table 1. Proximate and elemental analyses of the starting and demineralized anthracites.
H
HD
S
SD
Moisture (wt.%)
1.7
3.3
2.0
3.7
Vol. mat. (wt.%, db)
6.8
8.4
7.3
10.8
Ash (wt.%, db)
17.8
0.5
11.5
1.1
C (wt.%, daf)
92.22
90.29
93.41
89.48
H (wt.%, daf)
2.76
2.93
3.23
3.50
N (wt.%, daf)
2.26
2.17
1.91
1.92
S (wt.%, daf)
0.67
0.73
0.81
0.71
O (wt.%, daf)
3.17
4.04
1.37
3.73
Figure 1 shows the evolution of d002 as a function of HTT for the whole set of studied samples. As expected, the interlayer spacing decreases with increasing temperature. The blue line has been drawn just to guide the eye and to help one to differentiate two regimes: one below ~2000 ºC, where changes are bigger and strongly dependent on the presence of mineral matter, and another above this temperature, where changes in d002 are more moderate. The overall decrease of the interlayer spacing with increasing HTT is more continuous for the two demineralised anthracites than for their raw counterparts. The latter exhibit strong decreases in d002 at 1600-1800 ºC, whish are attributable to catalytic effects of mineral matter. Above 2600 ºC all d002 data converge to values that seem to be characteristic of the particular anthracite. Thus, at 2800 ºC, d002 takes values of 0.3370 (H) or 0.3369 (HD), versus 0.3399 (S) or 0.3391 nm (SD). In parallel with the improvement in the degree of graphene packing attested by d002 data, there is a progressive growth of graphite-like crystallites as HTT increases. Figures 2 and 3 show the evolution of the height and width of these crystallites, respectively, as a function
of HTT (for comparative purposes, the two plots are shown with the same vertical scale). It can be seen in Figure 2 that Lc is systematically higher for the raw anthracites than for their respective counterparts, and also systematically higher for Hlobane anthracite than for Somkhele Duff anthracite. 0.37
H
HD
S
SD
0.365
d(002) (nm)
0.36
0.355
0.35
0.345
0.34
0.335 700
900
1100 1300 1500 1700 1900 2100 2300 2500 2700 2900
Temperature (ºC)
Figure 1. Variation of the interlayer spacing, d002, with HTT. 60
60 HD
S
H
SD
50
50
40
40
La (nm)
Lc (nm)
H
30
HD
S
SD
30
20
20
10
10
0
0 800
1000
1200
1400
1600
1800
2000
2200
2400
2600
2800
800
Temperature (ºC)
1000
1200
1400
1600
1800
2000
2200
2400
2600
2800
Temperature (ºC)
Figure 2. Temperature dependence of the
Figure 3. Temperature dependence of the
graphitic crystallite height, Lc.
graphitic crystallite width, La.
However, in the case of La (Fig. 3) all values approximately coincide with each other up to 2000 ºC. Above this temperature, the crystallite width of raw Hlobane anthracite undergoes a strong increase, and in the 2000-2800 ºC interval it more than doubles the value corresponding to the demineralised counterpart. La values are systematically larger than those obtained for Lc, which reflects the well-known tendency of graphitic crystals to grow along the a axis in comparison with growth along the c axis [4].
As a comparison of Figures 2 and 3 clearly shows, under equivalent conditions Hlobane anthracite graphitizes more strongly than Somkhele Duff, even when it has been demineralized and particularly at temperatures above 2000 ºC.
4. Conclusions The two South African anthracites studied in this work exhibit similar rank characteristics. Their main mineral constituents are quartz, clays, calcite and pyrite, mineral matter being more abundant and with a les crystalline in Hlobane anthracite. Mineral matter present in the starting anthracites is lost by volatilization upon thermal treatment and exerts a notable catalytic effect on graphitization, especially at temperatures below 2400 ºC. The interlayer spacing and the crystallite sizes along the c and a axes evolve in parallel as temperature increases, crystal growth being systematically larger in width than in height. As a general rule, a higher graphitization degree is attained in the case of Hlobane anthracite, which can be attributed, among other factors, to the better degree of mixing between the mineral components and carbonaceous matrix in this coal.
Acknowledgements Colciencias and Universidad del Atlántico (Barranquilla, Colombia) are gratefully acknowledged for providing a doctoral research grant and a leave to M.V.C. We thank Dr. Henrique Pinheiro for providing the starting anthracite samples and for fruitful discussions.
References [1] Somma de Barros Pinheiro HJ. Springlake anthracite–Characterisation and potential industrial utilisation. PhD Thesis, University of Porto, Portugal, 2006. [2] Pappano PJ, Mathews JP, Schobert HH. Structural determination of Pennsylvanian anthracites. Prepr. ACS Div Fuel Chem 1999;44 (3):567-8. [3] González D, Montes-Morán M, García A. Influence of inherent coal mineral matter on the structural characteristics of graphite materials prepared from anthracites. Energy Fuels 2005;19:263-9. [4] Cuesta A. Factores determinantes de la reactividad de fibras de carbono y otros materiales carbonosos. Ph.D. Thesis, University of Oviedo, 1994.
Mineral matter and heat treatment temperature effects on the development of graphitic structure in two South African anthracites as studied by Raman spectroscopy and XRD Marley Vanegas-Chamorro1,2,*, Katia Tamargo-Martínez1, Jorge Xiberta2, Amelia Martínez-Alonso1 and Juan M. D. Tascón1 1
Instituto Nacional del Carbón, INCAR-CSIC, Apartado 73, 33080 Oviedo, Spain
2
Departamento de Energía, Univ. de Oviedo, Independencia 13, 33004 Oviedo, Spain
* Corresponding author; e-mail address: [email protected]
Abstract The effects of mineral matter on the graphitic structure attained by heat treatment at 900-2800 ºC of two South African anthracites (Hlobane and Somkhele Duff) were studied using Raman spectroscopy. To this end, both raw and acid-demineralized materials were studied. The Raman results were compared with those obtained previously by X-ray diffraction (XRD). A parallelism between both techniques was observed for demineralized samples. However, in the case of raw anthracites, mineral matter exerted strong catalytic effects on graphene layers stacking (three-dimensional order, as probed by XRD) but had a comparatively minor influence on the degree of graphene orientation (bidimensional order, to which Raman spectroscopy is sensitive).
1. Introduction In a previous work [1], the evolution in the structure of two anthracites submitted to heat treatment at several high temperatures was investigated using XRD. The aim of this work is to gain further insights into the type of thermal transformations that take place in these materials by means of Raman spectroscopy. Moreover, Raman and XRD results are compared with each other in an aim to exploit the complementary character of these two techniques [2]. It is expected that this understanding can contribute to selecting optimal conditions for the preparation of highly graphitic reductants for Fr-Cr and Fe-Mn minerals from South African anthracites [3]. As in the previous work, the effects of mineral matter are addressed by comparing the behaviors of raw and aciddemineralized anthracites.
2. Experimental section Materials. Two South African anthracites, Hlobane (H) and Somkhele Duff (S), were used as starting materials. Results from their analyses have been reported elsewhere [1] Removal
of
anthracite
inorganic
constituents
was
accomplished
by
acid
demineralization (using HCl and HF), according to the ISO 602:1983 International Standard for the Determination of Mineral Matter in Coals. All samples (demineralized or not) were devolatilized in a tube furnace under an Ar flow (350 cm3 min− 1), by heating several 32-g batches at 10 °C min− 1 to 900 ºC, an then keeping the temperature constant for 1 h. Portions of the batch decomposed at 900 ºC were exposed to higher heat treatment temperatures (HTTs) (1400-2800 ºC) for 1 h under Ar flow in a graphite furnace. Their identification codes contain the initial of the anthracite name, followed or not by D (i.e., demineralized or not), followed by the HTT in ºC. Thus, HD2400 stands for Hlobane anthracite, demineralized and then heat-treated at 2400 ºC. Methods. Raman spectra were obtained in the 800-3500 cm−1 spectral range with an HR 800 Jobin Yvon Horiba spectrometer equipped with a CCD detector. Each spectrum is the result of 4 accumulations of 15 seconds each using an incident power of 2-4 mW. In all cases, sampling was carried out on at least 8 different locations for the same sample. In the first order spectral region (800-2000 cm-1), deconvolution was carried out using the LabSpec® software. X-ray diffractograms were recorded in a Bruker D8 Advance diffractometer using Cu Kα radiation (λ = 1.5404 nm). Silicon was used as standard for peak position and broadening corrections. The interlayer spacing d002 was determined from the (002) peak applying Bragg’s law.
3. Results and discussion As illustrated in Figure 1 for Somkhele Duff anthracite, all samples carbonized at 900 ºC (either raw or previously demineralized) exhibit two main broad and overlapping bands in the 1st-order region of the Raman spectrum. They can be attributed [4-7] to defects (1350 cm−1, D band) and graphitic order (1580 cm−1, G band). According to carbon materials literature [8-10], three additional contributions located at about 1150 (I band), 1500 (D” band) and 1620 cm−1 (D’ band) must be considered. While the I band is attributed to C(sp3)–C(sp3) and/or H–free vibration modes, the D” and D’ bands have an origin that is still uncertain, but related to the graphitic disorder degree [4,5,11].
As the HTT increases above 900 ºC (Fig. 1), the bands in the 1st-order region progressively narrow, and the overlap of D and G bands practically disappears at 1600 ºC for non-demineralized samples and 1800 ºC for demineralized ones. In these Raman profiles, the D’ band contribution may be distinguished from a qualitative viewpoint as a shoulder is now observed on the G band (indicated by a black arrow in Fig. 1). In the 2nd-order spectrum, a new band at 2700 cm−1 named G’ is detected at 1600 ºC for the raw samples and at 1800 ºC for the demineralized ones. This band is related to stacking of graphene layers and evidences the incipient formation of a graphitic structure [7,9]. For the demineralized samples, the G and G’ band heights become comparable to each other at 2200 ºC, the D band height being smaller than the G band height (Fig. 1). For the non-demineralized materials, this trend is observed at significantly lower temperatures: 1800 ºC for Hlobane and 2000 ºC for Somkhele Duff.
a)
Demineralized
b)
Non-demineralized
G’
G
G’
G
D
2D’ D’
D
2D’
SD2800
D’
S2800
SD2600
S2600
SD2400
S2400
SD2200
S2200
SD2000
S2000
S1800
SD1800
S1600
SD1600
1000
1500
2000
2500 -1
Wavenumber (cm )
3000
SD1400
S1400
SD900
S900
3500
1000
1500
2000
2500
3000
3500
-1
Wavenumber (cm )
Figure 1. Raman spectra for heat-treated Somkhele Duff samples. The evolution of the structural disorder parameter, R = [ID/I(D+G)] for all samples is shown as a function of the temperature in Figure 2. For the two demineralized
anthracites, this intensity ratio decreases in an approximately linear way as the HTT increases. However, for the non-demineralized ones a considerable decrease in R (by ~40%) is detected between 1600 and 2200 ºC for Hlobane anthracite (Fig. 2a), and between 1800 and 2200 ºC for Somkhele Duff (Fig. 2b). Above 2200 ºC, R remains practically constant for both H and S series. The R value attained by H2800 sample is comparable to that of HD2800, but this is not so for the corresponding Somkhele Duff samples (compare Figs. 2a and 2b). One may therefore conclude that inherent mineral matter has a clear catalytic effect on the graphitization process, above 1600 ºC for Hlobane anthracite and 1800 ºC for Somkhele Duff. b)
a)
S
H
1
0
l
o
b
a
n
1
H
6
4
2
s
e
r
i
e
s
H
D
s
e
r
i
e
0
0
m
k
h
e
l
e
D
u
f
f
8
0
6
0
4
0
2
0
s
0
ID/I(D+G) (%)
ID /I(D+G) (%)
8
o
e
0
0
0
0
S
s
e
r
i
e
s
S
D
s
e
r
i
e
s
0
0
6
6
0
0
1
0
0
0
1
4
0
0
1
8
0
0
2
2
0
0
2
6
0
0
3
0
0
0
0
1
0
0
0
1
4
0
0
1
T
T
e
m
p
e
r
a
t
u
r
e
(
º
C
8
0
0
2
2
0
0
2
6
0
0
3
0
0
0
0
e
m
p
e
r
a
t
u
r
e
(
º
C
)
)
Figure 2. Relationship between the Raman intensity ratio disorder parameter and HTT for Hlobane (a) and Somkhele Duff anthracites (b). Working with the same anthracites, we have observed recently [1] that the graphitic interlayer spacing (d002) determined by XRD progressively decreased with increasing HTT. When d002 and ID/I(D+G. are plotted versus each other (Fig. 3), a reasonable correlation is observed for the two demineralized series over the entire range of graphitization degrees achieved by thermal treatment. However, for the two nondemineralized series, there is a “jump” at R ~58%; below this, d002 suddenly reaches its minimum value (0.337-0.340 nm). This occurs in the temperature range at which the catalytic effects became detectable (1600 ºC for H and 1800 ºC for S). For lower R values, d002 does not further decrease. A likely explanation to this finding lies in the different physical meanings of the information provided by Raman spectroscopy and XRD [11], as they are sensitive, respectively, to the degrees of bidimensional (2D) and
three-dimensional (3D) order. Accordingly, what Fig. 3 shows is that 2D order evolves in parallel with 3D order in the absence of catalytic effects. However, when these are operative, a good 2D stacking of graphene layers is attained as soon as the catalytic effect becomes detectable; at higher temperatures, there is a progressive improvement in 3D order, but the degree of 2D order remains virtually unmodified. a)
b) 0.37
0.37 H
HD
S
SD
0.36 d 00 2 (nm )
d0 02 (nm )
0.36
0.35
0.34
0.35
0.34
0.33
0.33 0
10 20
30
40
50 60
70
80 90 100
ID/(ID+IG ) %
0
10
20 30 40
50 60 70 80
90 100
ID/(ID+IG ) %
Figure 3. Relationship between d002 graphitic interlayer spacing and Raman intensity ratio for heat-treated Hlobane (a) and Somkhele Duff anthracites (b) 4. Conclusions Above 1600 ºC (Hlobane) or 1800 ºC (Somkhele Duff), the minerals associated with the two studied anthracites have a clear catalytic effect on their graphitization. However, in the case of demineralized samples, the graphitic order increases progressively with increasing temperature, there being a parallel development of 2D and 3D structural order. Unlike this, catalytic graphitization seems to affect selectively the stacking of graphene layers, but not their orientation degree.
Acknowledgements Colciencias and Universidad del Atlántico (Barranquilla, Colombia) are gratefully acknowledged for a doctoral research grant and a leave awarded to M.V.C. We thank Dr. Henrique Pinheiro for providing the starting anthracite samples and also for fruitful discussions.
References [1] Vanegas-Chamorro M, Tamargo-Martínez K, Xiberta J, Martínez-Alonso A, Tascón JMD. Structural evolution upon thermal treatment of two South African anthracites as studied by XRD. These ICCS&T 2011 Proceedings. [2] Pierson HO. Synthetic carbon and graphite: carbonization and graphitization. In: Bunshah RF, McGuire GE, Rossnagel SM, eds. Handbook of carbon, graphite, diamond and fullerenes: properties, processing and applications, New Jersey: Noyes Publications; 1993, p. 70-86. [3] Somma de Barros Pinheiro HJ. Springlake anthracite–Characterisation and potential industrial utilisation. PhD Thesis, University of Porto, Portugal, 2006. [4] Beyssac O, Goffé B, Petitet JP, Froigneux E, Moureau M, Rouzaud JN. On the characterization of disordered and heterogeneous carbonaceous materials by Raman spectroscopy. Spectrochim Acta Part A, 59, 2003;2267-76. [5] Swain G.M. Solid electrode materials: pretreatment and activation. In: Zoski CG, ed. Handbook of electrochemistry, 1st ed. Amsterdam: Elsevier; 2007, Ch. 5, p. 111-50. [6] Morishita K, Nara M, Matsuda J-I. Raman spectroscopic investigation of the isotopic effects in graphitic carbon and the interpretation of D-band. J Mass Spectrom Soc Jpn 2010;58(5):167-8. [7] Ferrari AC, Robertson J. Interpretation of Raman spectra of disordered and amorphous carbon. Phys Rev B 2000;61:14095-107. [8] Pimenta MA, Dresselhaus G, Dresselhaus MS, Cançado LG, Jorio A, Saito R. Studying disorder in graphite-based systems by Raman spectroscopy. Phys Chem Chem Phys 2007;9:1276-91. [9] Ferrari AC, J Robertson. Origin of the 1150-cm-1 Raman mode in nanocrystalline diamond. Phys Rev B 2001;63:121405–1-8. [10] Ferrari AC, J Robertson. Resonant Raman spectroscopy of disordered, amorphous, and diamondlike carbon. Phys Rev B 2001;64:075414–1-13. [11] Cuesta A, Dhamelincourt P, Laureyns J, Martínez-Alonso A, Tascón JMD. Raman microprobe studies on carbon materials Carbon 1994;32:1523-34. [12] Cuesta A, Dhamelincourt P, Laureyns J, Martínez-Alonso A, Tascón JMD. Comparative performance of X-ray diffraction and Raman microprobe techniques for the study of carbon materials. J Mater Chem 1998;8:2875-9.
SPECIATION AND FATE OF MERCURY IN OXY COAL COMBUSTION Font O1, Córdoba P1, Leiva C2, Romeo L M3, Bolea I3, Guedea I3, , Moreno N1, Querol X1, Fernandez-Pereira C2, Díez L I3. 1
Institute of Environmental Assessment and Water Research (IDÆA-CSIC), Jordi Girona 18-26, E-08034- Barcelona, Spain. E-mail: [email protected] 2
Escuela Superior de Ingenieros de Sevilla, Departamento de Ingeniería Química y Ambiental, Camino de los Descubrimientos, s/n. Isla de la Cartuja, 41092 Sevilla, Spain
3
CIRCE (Research Centre for Energy Resources and Consumption). University of Zaragoza. Mariano Esquillor, 15. Zaragoza, 50018. Spain Keywords: oxy-combustion, trace elements, mercury, fate ABSTRACT
The speciation and fate of Hg was evaluated in a 90KWth bubbling fluidised bed (BFB) oxy-combustion pilot plant, fed with anthracites and limestone (bed material). The Hg levels and speciation in the CO2-rich exhaust gas is of high relevance since this might affect subsequent purification, compression, transportation and storage of the CO2. Sampling of Hg in exhaust gases was undertaken through 3h at 70% load, 70:30 CO2/O2 ratio of raw gas and 800-820 oC. SOx, NOx, CO2, and CO in exhaust gases were measured simultaneously to Hg by monitor. Solid streams (coal, limestone, inert bed material, bottom ash, cyclones fly ash and bag filter fly ash) were also collected. Mercury concentrations of the exhaust gas attained 0.4 and 1.8 µg/m3N for Hg2+ and Hg0, respectively. Thus 82% of the emitted Hg is Hg0 and only 18% Hg2+. This speciation is the opposite to that of Hg in the raw gas escaping electrostatic precipitators in the conventional coal combustion (75-86 % Hg2+). Both, the low gas temperature (45 o
C) and high Ca content promotes condensation and preferential retention of Hg2+ in
bag filters fly ash. The exhaust gases are CO2-rich (95.8 %) with low levels of O2 (3.7%), CO (0.4%), NOx (0.06%) and SOx (4.2 ppm) indicating high retention efficiencies of SOx and NOx in oxy coal combustion. 1. INTRODUCTION The reduction of CO2 emissions from power plants is a crucial challenge to control and limit the global warming. The International Panel for Climate Change (IPCC) has
concluded that CO2 emissions must be reduced by 50 to 85% by 2050 if global warming is to be restricted to between 2 to 2.4 ºC [1]. The IEA Energy Technology Perspectives 2008 focus on returning emissions back to the level of 2005 and reduce CO2 emissions by 50% from the level of 2005 by 2050. The European Union (EU) is facing the challenge to reduce the European greenhouse gas emissions by 8% by 2012 compared to 1990 levels, while it is also dealing with security of energy supply, competitiveness and sustainable development in energy policy. About 30% of European power generating capacity is coal-fired and this percentage is much higher in certain EU countries (90% in Poland and over 50% in the Czech Republic, Greece and Germany).
The International Energy Agency (IEA) Energy Technology Perspectives 2008 shows possible routes to reach the challenging targets of the scenarios. Fuel oxy-firing is the most important new technology for the CO2 capture and storage (CCS) in power generation industry with an expected CO2 reduction of 14 to 19% (equivalent to CO2 savings up to 5000 Mtn/year by 2050). Nowadays oxy-firing technology is under development. There are 6 CCS demonstration projects in EU funded by the European Energy Programme for Recovery (EEPR).
Oxy-firing technology is based on using CO2 and O2 for combustion instead air to produce CO2-rich flue gases allowing an easy CO2 purification from other combustion gases (O2, NOx, SOx, HCl, HF,..) for subsequent geological storage. The CO2-rich gas is partially re-circulated to the boiler and mixed with O2 and H2O for combustion. The CO2-H2O rich atmosphere is considerably different to the atmosphere of conventional air combustion (78% N2 and 16 %O2). Both, the different composition of flue gas and the re-circulation of the CO2-rich exhaust gas may change the heat release and heat transfer patterns within the boiler, the behaviour of gases and particles, and gas/particle interactions, among others. Furthermore, the different gas composition leads to different operational conditions of particle and gas cleaning devices than in air-firing. Consequently the physical, mineral, chemical and leachable properties of slag/bottom ash, fly ash and FGD gypsum produced as well as the fate of trace elements in oxycombustion may be modified with respect to that produced in conventional air firing power plants. Furthermore, the levels of impurities (mainly SOx, NOx, HCl, HF, and H2O) in the resulting CO2-rich gas are crucial for the subsequent gas purification, compression, transportation and geological storage due to the high corrosion potential
of these gaseous species. It is well known the corrosion potential of H2SO4, HCl and HF. The occurrence of H2O, may increase the corrosion potential of H2SO4 and, in the presence of CO2, H2O may cause wet corrosion [2-3] and solid ice-like crystals [4]. Furthermore, due to its high volatility in combustion and its corrosion potential, Hg is regarded as an impurity of a high concern in the oxy-combustion CO2 rich exhaust gas. In view of the above issues, a simultaneous sampling and speciation of gaseous Hg and solid streams was devised and executed in a 90KWth bubbling fluidized bed (BFB) oxycombustion pilot plant. The main goal of this study if to determine the speciation and levels of Hg in the exhaust gas (including particulate Hg), and it’s partitioning among solid and gaseous streams in oxy-combustion conditions. Furthermore the partitioning of other major and trace elements among the global outputs and the physical, chemical, mineralogical and leachable properties of oxy-coal combustion wastes were also investigated.
2. EXPERIMENTAL 2.1. Oxy pilot plant description The oxy pilot plant was designed to operate with conventional air and oxy combustion conditions (using CO2 from a bundle of cylinders or from a flue gas recirculation) with high feed fuel and bed material flexibility and a wide range of fluidization conditions (700-900 oC and 0.8-1.4 m/s). The reactor is a bubbling fluidized bed combustor, of 90 kWth capacity (under oxy-fuel conditions), with a a total height of 2.5m and a bed diameter of 0.203 m. The heating-up of the bed material is achieved by an auxiliary propane burner. The bed is water-cooled (up to 3m3/h depending on fuel load) during stable operation, while the freeboard is refractory lined. Feed fuel system consists of two independent 0.2 m3 hoppers, located besides the reactor, for fuel and limestone, respectively. In the case of co-firing, a mixture of coal and limestone is fed by the same hopper while co-matter enters alone to prevent particle segregation. Each hopper discharges on a separate screw, controlled by a variable-frequency drive, therefore independent flow rates can be fed from every hopper. These screws discharge into a long mixing screw that introduces the powders in the reactor. The tip of the screw is water-cooled, and a temperature-based control prevents the back-propagation of the burning matter. Fuel and inert/ sorbent are entered 50 mm above the fluidizing plate, in order to reduce the initial elutriation. When operating oxy-fuel, the supply of O2/CO2
mixture is initially taken from two blocks of 12 cylinders each. Once the operation is stabilized, the consumption of CO2 from the cylinders can be replaced by flue gases recirculation, using the forced-draft fan. Oxidant entering the combustor is preheated by means of a 9 KW shell-and-tube heat exchanger, by using the sensible heat of flue gases leaving the reactor. A by-pass can be operated to accommodate the temperature levels.
Before entering the heat exchanger, gases are cleaned up of fly ash particles in a high efficiency Stairmand cyclone, designed for a cut size of 2lm (inlet velocity of about 16– 24 m/s). At the gas cold end, a fabric-filter bag is located for additional particle retention.
The rig is controlled by a Programmable Logic Controller (PLC), at which all the instruments (thermocouples, pressure gauges, and flow meters), control valves, motor starters and variable frequency drives are connected. Gas composition is on-line measured during the tests by a gas analyzer, determining CO2, O2, CO, SO2 and NOx concentrations.
2.2. Sampling Sampling of solid streams (coal, limestone, bottom ash, cyclones fly ash and bag filter fly ash) and exhaust gases (SOx, NOx, CO2, CO and Hg) was undertaken through around 4h operating (including start up and steady oxy-combustion conditions periods) at 70% load, 70:30 CO2/O2 ratio of raw gas and 800-820oC. The plant was fed with anthracitic coals as fuel and limestone as bed material. Limestone was fed during starting and stabilization periods of the plant up to reach the desired height of the bed. When reaching steady oxy-combustion conditions coal was fed to the reactor with a flow of 11.9 kg/h, attaining 87:13 coal/limestone ratio.
Isokinetic measurements of gaseous Hg were performed after gas passes the bag filters and heat exchanger. Sampling and speciation of gaseous Hg (Hg0 and Hg2+) run lasted 1.5h at steady oxy combustion conditions. This was devised according to EN 13211 and Meij and Winkel [5]. Accordingly, a sample of 75 mL was withdrawn from the flue gas stream through a filter system, maintained at the right temperature, followed by a train of dark glass bottles in an ice bath. The gas washing-vessel system consists of a washing-bottle filled with HCl to capture Hg2+ [5], a gas washing-bottle containing 3
w% H2O2 for SO2 removal, and two bottles containing 4 w% K2Cr2O7 with 20 w% of HNO3 for Hg0 trapping. Sampling of coal and limestone and bottom ash, cyclone and bag filter fly ash was carried out before and after oxy combustion test run, respectively. Table 1. Operational conditions of the oxy-combustion plant in sampling. Load 70% Exhaust gas flow 52.16 m3/h CO2/O2 70:30 Bed temperature 800-820 oC Coal input 11.9 kg/h Cyclone temperature 458-525 oC Limestone input 3.3 kg/h Bag filters temperature 45 oC 2.3. Analysis Mercury analysis were directly analysed on solid streams, solutions from flue Hg gas sampling and filter samples using a LECO AMA 254 gold amalgam atomic absorption spectrometer.
Fir solid samples, moisture and high temperature ash (HTA) were carried out at 105 and 750 oC, respectively according to ASTM recommendations. The contents of C, H and N were determined by means of conventional elemental analysers (Protein, NA 2100, Carlo Erba Instruments). The solid samples were acid-digested in duplicate by using a two-step digestion method devised by Querol et al [6]. The resulting solution was then analysed by Inductively-Coupled Plasma Atomic-Emission Spectrometry (ICP-AES) and by Inductively-Coupled Plasma Mass Spectrometry (ICP-MS) for 57 major and trace elements, respectively. The fly ash and coal international reference material (NBS1633b and SARM19) were also digested to determine the accuracy of the analytical and digestion methods. The Cl contents were determined by the Eschka method [7]. The procedure utilizes sample decomposition using Eschka mixture (two parts of MgO and one part of Na2CO3) at 800 °C for 2 h, in covered porcelain crucibles and subsequent determination by chloride selective electrode. The determination of F was performed following the method described by Sager [8]. This method uses a sample fusion with NaOH at 550 °C for 1 h, followed by residue dissolution by tiron (pyrocatechol-3, 5-disulfonic acid, disodium salt) which was subsequently analysed by fluoride sensitive electrode.
The mineralogy of solid samples was determined by X-ray powder diffraction (XRD) with a 218 Bruker D5005 diffractometer with monochromatic Cu Kα1,2 radiation operated at 40KV and 219 40mA., from 4 at 60° of 2 theta range, and a step size of 0.05° and 3s/step.
The European Standard leaching test EN-12457 (according to Council decision 2003/33/EC) was applied to the oxy coal combustion wastes (bottom ash, cyclone fly ash and bag filter fly ash) samples to determine the leaching potential of major, minor and trace elements The pH and ionic conductivity were determined by conventional methods. The content of major minor and trace elements of the leachates were determined by ICP-AES and ICP-MS. The content of Hg was determined directly on leachates by the same atomic absorption spectrometer described above, while leachable NH4+ and Cl by selective electrode.
3. RESULTS AND DISCUSSION 3.1. Mercury concentrations and enrichment factors in solid samples Coal arises from the anthracitic Bierzo coal basin (Northwest Spain) and it is a low moisture (4%), intermediate ash (17%) and low S (0.4 %) coal. Quartz, illite and traces of kaolinite and gypsum are the mineral phases detected by XRD.
Bed material is a limestone of 95% CaCO3 purity with minor concentrations of Mg (0.1%) and trace contents of Mn, Sr, and Ba (4 mg/kg) probably present as marginal carbonates or in the calcite structure (Table 2). Since quartz levels are under XRD detection limits (<1% wt) the remaining fraction of limestone (over 4%) has to be attributed to organic matter.
The concentrations of Hg in the feed fuel, bed material and ashes are depicted in Table 2. The Hg content of feed coal is in line with the relatively high Hg content of the Bierzo coals being higher than the mean values reported by Yudovich and Ketris [9] for worldwide coals. The Hg concentration in limestone is relatively low, attaining 0.005 mg/kg (Table 2). As regards the Hg concentrations in ashes, bag filters fly ash show the highest Hg concentrations, reaching 2.5 mg/kg while the content of Hg is lower in cyclone fly ash and bottom ash (Table 2).
The Al normalised enrichment factors (EFs) in bottom ash, cyclone fly ash and bag filter fly ash, calculated considering Al as a non-volatile element in oxy coal combustion by using the normalisation formulas by Gordon and Zoller [10], shows an increasing enrichment of Hg from bottom ash to bag filter fly ash, reaching EFs of 5.4 in this bag filter fly ash. Other elements, (S, F, Cl, B, As, Se and Sn) shows similar EFs trend among oxy-combustion ashes attaining EFs from 1.6 to 5.2 in bag filter fly ash. The high enrichment of Hg and the above elements is attributable to the low gas temperature (45oC) and high Ca content of bag filter fly ash. The low gas temperature promote condensation of volatile elements, while the high Ca content favours adsorption processes of Hg in Ca-bearing species surfaces [5-6].
It is worth noting that bag filter fly ash is characterized by a remarkable moisture (4% after air drying) and fine grain size (median of 8.6µm). Gypsum and calcite are the main crystalline phases present in this fly ash, with minor contents of illite and traces of bassanite, anhydrite, lime, halite, quartz, and hematite. The use of limestone as bed material favours the entraining of calcite particles by the gas, which are partially decomposed into lime and Ca-sulphate species in oxy-combustion process. The high levels of Ca particles not retained by the cyclones allows reaction with SO2 giving rise to the formation of Ca-sulphate species which are subsequently retained in the bag filter fly ash. Similarly As and F may react with Ca to form Ca arsenate and fluorite, respectively, while Cl mainly reacts with Na condensing in bag filters as halite.
Consequently it may be deduced that most of the Hg released from coal during oxycombustion is in oxidised form after cyclones being subsequently adsorbed in the surface of lime particles or other Ca-bearing species, most probably as HgCl2. Nevertheless, the high levels of SO2 after cyclones may favour the occurrence of condensed Hg-sulphate species in bag filter fly ash. Furthermore, high NH4+ leachable levels were determined in leachates from bag filter fly ash. This suggests high condensation of this N-bearing species in bag filters, most probably as NH4+ bi-sulphate due to the high SO2 content, but also may condense as nitrate/chloride NH4+species.
mg/kg Hg
Table2. Concentrations of Hg in feed fuel, limestone and ashes. Coal Limestone Bottom ash Cyclone fly ash Bag filter fly ash 0.12 0.005 0.032 0.067 2.8
3.2. Levels and speciation of Hg in the exhaust gas. As expected, the exhaust gas is CO2-rich (96%) with minor amounts of O2 (4%). The low levels of CO, SO2, NOx and Hg (Table 3) reveal the high abatement of these gaseous pollutants, attaining 99.9 % and 92.5 % retention for gaseous S and Hg, respectively. As regards Hg, this metal is mostly present in the exhaust gas as Hg0 (82%) while only 18% is Hg2+. This speciation is in line with the aforementioned prevalent adsorption/condensation of Hg2+ in bag filters. The gaseous Hg speciation in the oxycombustion exhaust gas is the opposite to that of Hg in the raw gas escaping electrostatic precipitators, but equivalent to the common Hg speciation after Flue gas desulphurization (FGD) in conventional pulverized coal combustion power plants [11]. The low PM levels of Hg (7 ng/m3) and S (<0.25mg/m3) and indicates high retention efficiencies (>99.9%) for particulate bound Hg and S species. Table 3. Composition of the exhaust gas and speciation of gaseous Hg. % % µg/m3 µg/m3 µg/m3 µg/m3 µg/m3 CO2 O2 CO SO2 NOx Hg total Hg2+ Exhaust gas 95.8 3.7 1.14 0.001 0.12 2.15 0.39
µg/m3 Hg0 1.76
3.3. Partitioning of Hg and other elements among global outputs.
The partitioning of major and trace elements among the global outputs (exhaust gas, bottom ash, cyclone fly ash, bag filter fly ash) was obtained by normalising the concentration of each element in each stream with the corresponding flow.
The partitioning of Hg among global outputs reveals that most of Hg (79%) is retained in bag filter fly ash and minor proportions in bottom ash and cyclone fly ash (8 and 5%, respectively). Nevertheless low fractions of Hg (7.5%) still present in the exhaust gas. Other elements such as S, Cl, F, B and As show similar partitioning than that of Hg, but with extremely low proportions or not present in the exhaust gas. The remaining elements are retained readily (>60%) in bottom ash due to their low volatile behaviour
and the higher bottom ash production (70%) with respect to that of cyclone (22%) and bag filter fly ash (0.8%).
4. CONCLUSIONS High retention efficiency for Hg (92.5) and major and trace elements (~99.9%) are reached by oxy-combustion of coal in a 90KW BFB reactor. Although the high retention of Hg, significant levels of Hg (2.15 µg/m3) still present in the exhaust gas, most of them as Hg0 (81%). The low temperature and the high occurrence of Ca-bearing species promote the prevalent condensation and capture of Hg2+ in the bag filters. Despite the high Hg abatement capacity the levels in the CO2-rich exhaust gas may be of concern for subsequent CO2 treatment and storage. Both, the high abatement of Hg and condensate bearing species of S, F, Cl and NH4+ and the speciation of the aforementioned pollutants in the bag filter fly ash allowed considering this clean particle device as desulphurisation-like system. Special attention should be paid to the leachable levels of these elements in this fly ash since may be of concern as regards the limits established by the European regulations for landfilling.
5. REFERENCES [1] Energy Technology Perspectives 2008. Scenarios & Strategies to 2050. OECD/IEA, 2008. [2] CaillyB, LeThiez P, Egermann P, Audibert A, Vidal-Gilbert S, Longaygue X. Geological storage of CO2: a state-of-the-art of injection processes and technologies. Oil Gas Sci Technol 2005; 60 (3): 517-e25. [3] GozalpourF,RenSR,TohidiB.CO2 EOR and storage in oil reservoirs. Oil Gas Sci Technol 2005; 60 (3):537-e46. [4] Andersson K, Johnsson F, Strömberg L. Large scale CO2 capture e applying the concept of O2/CO2 combustion to commercial process data. VGBPowerTech 2003; 83(10):1-e5. [5] Meij R. and Winkel. H.T. Mercury emissions from coal-fired power stations: The current state of the art in the Netherlands. Science of the total environment 2006; 368: 393-396. [6] Querol X, Fernandez-Turiel JL, López-Soler A. Trace elements in coal and their behaviour during combustion in a large power station. Fuel 1995; 74 ( 3): 331-343.
[7] Chakrabarti J N. Methods of determining chlorine in different states of combination in coal. In: Analytical methods for coal and coal products, 1978, Vol. I (Karr C. Jr, editor). Academic Press, New York, pp. 323-345. [8] Sager M.. Rapid determination of fluorine in solid samples. Monatshefte für Chemie 1987;118: 25-29 [9] Yudovich YE, Ketris MP. Valuable Trace Elements in Coal. Ekaterinburg 2006, p.538 [in Russian]. [10] Gordon GE, Zoller WH. Normalization and interpretation of atmospheric trace element concentration patterns. Proceeding of the 1st Annual NSF Trace Contaminants Conference. Oak Ridge, Tenn: Oak Ridge National Laboratory. CONF-73080; 1973; 314–25. [11] Córdoba P, Ochoa-Gonzalez R, Font O, Izquierdo M, Querol X, Leiva C, LópezAntón MA, Díaz-Somoano M , Martinez-Tarazona MR, Fernandez C, Tomás A. Partitioning of trace inorganic pollutants in a coal-fired power plant equipped with a wet flue gas desulphurisation system. Fuel, in press.
Oviedo ICCS&T 2011. Extended Abstract
Experimental Study on the Mechanism of SO2 Emission and Removal in the Coal Oxygen-Enriched Combustion Luning Tian1 Hanping Chen2 Haiping Yang3 Xianhua Wang4 Shihong Zhang5 Cheng Zeng6 State Key Laboratory of Coal Combustion Huazhong University of Science and Technology, Wuhan, 430074, P. R. China 1
4
[email protected], [email protected], [email protected], [email protected], [email protected], [email protected].
Abstract Oxygen-enriched fluidized bed combustion technology is a renewable technology, which combine the advantages of circulating fluidized bed (CFB) technology and oxygen-enriched combustion technology. Due to the presence of CO2, there are many differences in the SO2 emission and removal by limestone between oxygen-enriched combustion and conventional combustion. In this article, fundamental experiments have been done to investigate the behavior and mechanism of SO2 emission and removal in oxygen-enriched combustion. It was observed that the peak value of SO2 emission in O2/CO2 atmosphere is smaller than that in O2/N2 atmosphere, the total of SO2 emission in O2/N2 atmosphere is smaller than that in O2/CO2 atmosphere at the lower temperature, and opposite at the higher temperature. The rise in temperature and O2 concentration promote the SO2 emission, increase the peak vale of the SO2 emission curve, shorten the time length of SO2 emission. The total of SO2 emission increase with the temperature increasing in the O2/N2 and O2/CO2 atmosphere, except at the 950
in the
O2/CO2=20/80 atmosphere which the least total of SO2 release happen. The SO2 removal by limestone was affected not only by the reaction condition but also by the characteristics of SO2 release. The sulfation efficiency decreases remarkably with the particle size increases. The sulfation efficiency increase obviously with increasing of the Ca/S when the Ca/S is less than 2, and the increase become more gradual when the Ca/S is more than 2. Keywords: oxygen-enriched combustion, SO2, emission, removal, limestone 1.
Introduction
The CO2 and SO2 emission pose a severe threat to the climate change, ecological environment and human health. The emission of CO2 and SO2 are mainly from the combustion of fossil fuel. Thus oxygen-enriched fluidized bed combustion technology was invented and researched for the renewable utilize of fossil fuel. This technology Submit before 31 May 2011 to [email protected]
1
Oviedo ICCS&T 2011. Extended Abstract
combines their advantages of CFB combustion and oxygen-enriched combustion. It can obtain purity CO2 for capture and storage, low pollutants emission and high combustion efficiency [1]. Because of the different atmosphere in combustion, the SO2 emission characteristics in the O2/CO2 atmosphere have many differences than that in conventional air. Several investigates have shown the characteristics of SO2 emission and removal by limestone in oxygen-enriched combustion. Croiset and Thambimuthu [2] found that the conversion is 91% for the air case, drops to 75% for combustion in O2/CO2 atmosphere and 64% for recycle combustion, respectively. His possible explanations are the sulfur retain in the ash or SO2 is oxidized to SO3. This conclusion was also verified by Duan et al. [3]. Liu et al. [4] concluded that the SO2 emissions are more or less independent of the combustion media. Chen et al. [5] shown that higher concentration of SO2 was released in O2/CO2=30/70 or O2/CO2=40/60 atmosphere. Zhou et al. [6] indicated that the CaO in coal ash plays a significant role in effecting desulfuration, and the presence of CO2 markedly contribute to the desulfuration process. Jia et al. [7] reported that the sulfur capture efficenct decreases from the 68.4% during air firing to 40.1% during the oxyfuel firing in a mini-CFB Combustion. Related studies of coal oxygen-enriched combustion are in process, but there is less study on the mechanism of SO2 emissions and removal, and more for pulverized coal combustion. In this article, fundamental experiments have been done to investigate the behavior and mechanism of SO2 emission and removal in oxygen-enriched combustion, including the influence of atmosphere, temperature, Ca/S and limestone particle size on the SO2 emission and removal. This study will be helpful to research and develop the oxygen-enriched fluidized bed combustion. 2.
Experimental
Table 1. Ultimate and proximate analyses of Huangling bituminous coal Ultimate analysis (wt. %)
Proximate analysis (wt. %)
Cad
Had
Oad
Nad
Sad
FCad
Aad
Vad
Mad
50.83
2.76
7.35
0.59
1.75
41.09
34.27
22.18
2.46
Huangling bituminous coal and Fuxin limestone were used as sample whose properties are shown in Table 1 and 2. Table 2: Chemical composition of limestone samples (wt. %) Na2O
Al2O3
SO3
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CaO
Fe2O3
2
Oviedo ICCS&T 2011. Extended Abstract
2.26
0.97
0.18
53.16
0.10
Fig. 1 shows the experimental apparatus. An electric heated horizontal tube furnace was used for the experiment of SO2 emission and removal. A quartz tube (I.D. 50mm) was used as a reactor at the middle of the furnace. In the experiment process, the furnace was heated to the desired temperature, gas mixtures was then introduced into the furnace to provide the reaction atmosphere, the porcelain boat with thin layer of sample(coal or mixtures of coal and limestone) was pushed into the middle of the reactor, finally. The gas products were analyzed by Gasmet portable FTIR (Gasmet DX4000, Finland). The weight of coal is 1(±0.0005)g, particle size of coal is 0.3～0.45mm, and the particle size of limestone is ≤0.063, 0.063～0.1, 0.1～0.15 and 0.3～0.45mm, respectively. Gas mixtures were provided with cylinders. Toal gas flow of 3l/min was selected.
Fig.1 Schematic diagram of the experimental apparatus 3.
Results and Discussion
Effect of atmosphere on SO2 emission The curves of SO2 emission in different atmospheres of O2/N2 and O2/CO2 at the temperature of 950
are presented in the Fig. 2. It can be seen that the SO2 emission
curves show the same tendency of two peaks with the reaction time. The first releasing peak is the quick release of the easily decomposed organic sulfur and sulphide iron ore, the emission of sulfate sulfur with increasing reaction time form the second releasing peak. The peak value of SO2 emission in O2/CO2 atmosphere is smaller than that in O2/N2 atmosphere. This indicates that the present of CO2 decreases the actual reaction temperature by the high heat capacity. Fig. 3 shows the total of SO2 emission under
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Oviedo ICCS&T 2011. Extended Abstract
different atmospheres. As observed from Fig. 3, the total of SO2 emission in O2/N2 atmosphere is smaller than that in O2/CO2 atmosphere at the lower temperature (≤850 ), while the opposite tendency was shown at the higher temperature (≥950 ). The explanation is that the sulfur retention rate of coal ash is slow in the high CO2 concentration atmosphere at lower temperature by direct sulfation. The high concentration of CO2 improves the sulfur retention ability of coal ash, and promotes the transfer of sulfur from SO2 to other sulfurous phase at the higher temperature. 11.5
1200
CO2:O2=80:20
1000
N2: O2= 80:20
11.0 Total of SO2 Emission (l)
SO2 Concentration (ppm vol)
1400
800 600 400 200 0
CO2:O2=80:20 N2: O2= 80:20
10.5 10.0 9.5 9.0 8.5
0
200
400 600 800 Reaction Time (Second)
700
1000
800 900 1000 Reaction Temperature (℃ )
1100
Fig.3 Total of SO2 emission in different atmospheres
Fig.2 Curves of SO2 emission in different atmospheres
Effect of O2/CO2 concentration on SO2 emission 9.9
Total of SO2 Emission (l)
SO2 Concentration (ppm vol)
1500 1200 900 600
O2:CO2=80:20 O2:CO2=60:40 O2:CO2=40:60 O2:CO2=20:80
300 0
0
300
O2:CO2=10:90
600 900 1200 1500 Reaction Time (Second)
9.6 9.3 9.0 8.7 8.4
1800
10/90
20/80
40/60 O 2/CO 2
60/40
80/20
Fig.4 Curves of SO2 emission in different Fig.5 Total of SO2 emission in different O2/CO2 atmospheres O2/CO2 atmospheres The variations of SO2 emission with the reaction time are presented in Fig. 4 under different O2/CO2 concentrations at the temperature of 950 . The SO2 emission curves also show two peaks from 10/90 to 60/40 of the O2/CO2 atmosphere, while there is only one peak under the O2/CO2=80/20 atmosphere. The high O2 concentration effectively accelerates the combustion reaction and promotes the SO2 emission. The higher of the O2 concentration, the larger of the peak value, and the shorter of the time length which the curves reached the second peak. Especially in the O2/CO2=80/20 atmosphere, the
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Oviedo ICCS&T 2011. Extended Abstract
two peaks releasing of SO2 emission evolve to one peak releasing. Fig. 5 shows the total of SO2 emission in different O2/CO2 concentrations. With O2/CO2 of 10/90, the slower combustion is beneficial to SO2 emission because of the slow sulfur retention rate of coal ash. In 20/80 of O2/CO2 atmosphere, the sulfur retention ability of coal ash increases significantly, the emission amount of SO2 reduces markedly. Vigorous combustion can enhance considerably SO2 emission when the O2 concentration are higher high than 40%, but the coke phenomenon becomes more serious in the layer combustion procession with the O2 concentration up; it will inhibit the SO2 emission, as the release in the O2/CO2 of 80/20. Effect of temperature on SO2 emission 1400
1000
750℃
800
850℃
SO2 Concentration (ppm vol)
SO2 Concentration (ppm vol)
1400
O 2/N2 =20/80
1200
950℃
600
1050℃
400 200 0
O2/CO2=20/80
1200 1000 800
1050℃ 950℃ 850℃ 750℃
600 400 200
0
200
400
600
800
1000
0
0
200
Reaction Time (Second)
400
600
800
1000
1200
Reaction Time (Second)
Fig.6 Effect of temperature on SO2 emission in different atmospheres (a: O2/N2=20/80 b: O2/CO2=20/80) The curves of SO2 emission with the reaction time under different temperature are presented in Fig. 6. The SO2 emission curves show one peak of sulphide iron ore and organic sulfur releasing at the lower temperature (≤850 ), and present second peaks of sulfate sulfur releasing at the higher temperature(≥950 ). The rise in temperature promotes the SO2 emission, increases the shorten the whole time of sulfur release. Fig. 7 shows the total of SO2 emission at different testing temperature. It can be seen
12 Total of SO2 Emission (l)
peak value of the SO2 emission curve, and
atmosphere, except at the 950
in the
850℃
950℃
1050℃
10 8 6 4 2
that the total of SO2 emission both have the trend of ascent in the O2/N2 and O2/CO2
750℃
0
O2/N2=20/80 O2/CO2=20/80 Reaction Atomsphere
Fig.7 Effect of temperature on the total of SO2 emission
O2/CO2=20/80 atmosphere. The possible reasons have been explained above that the ability of sulfur retention of coal ash and the transfer of sulfur from SO2 to other sulfurous phase get better development at the 950
in the O2/CO2=20/80 atmosphere.
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Oviedo ICCS&T 2011. Extended Abstract
Influence of the O2/CO2 concentration on the removal of SO2 by limestone The limestone desulfuration efficiency in the different O2/CO2 atmospheres is presented in Fig. 8. It can be seen that the Desulfurization Efficiency (%)
desulfuration efficiency in the O2/CO2 concentrations of 20/80 and 80/20 is higher than that in the O2/CO2 concentrations of 40/60 and 60/40. lthough the total of SO2 emssion is the least. This is due to
sintering and decrease the SO2 emission
33 30 27 24
that the high CO2 in the O2/CO2=20/80 atmosphere can mitigate the limestone
36
20/80
40/60
60/40
80/20
O 2/CO 2
Fig.8 Effect of the O2/CO2 concentration on the SO2 removal by limestone
rate. In the atmosphere of O2/CO2=80/20, the coke phenomenon by vigorous combustion can inhibit SO2 emission, it makes a good condition for sulfation. Influence of temperature on SO2 removal by limestone 33
40 35 30 25 O 2/N 2 =20/80 20
750
800
850
900
950
1000
1050
Desulfurization Efficiency (%)
Desulfurization Efficiency (%)
45
32 31 30 29 O 2/CO 2=20/80
28 750
Temperature (℃ )
800
850 900 950 Temperature (℃ )
1000
1050
Fig.9 Effect of temperature on the SO2 removal in different atmospheres (a: O2/ N2 =20/80 b: O2/ CO2=20/80) The limestone sulfation efficiency at the different temperatures in the O2/N2=20/80 atmosphere are displayed in Fig. 9a, it can be seen that the sulfation efficiency increases with the temperature rising. Fig. 9b displays the limestone sulfation efficiency at the different temperature in the O2/CO2=20/80 atmosphere, the sulfation efficiency at the 750, 850, 950 and 1050℃ is 27.9%, 32.0%, 30.0% and 32.4%, respectively. At the 750℃, the lower temperature is not favorable for the direct sulfation reaction. The sulfation efficiency increase obviously by the improving of direct sulfation reaction is improved at the 850℃. At the 950℃, the sulfation efficiency decreases because the smaller total of SO2 emission reduce the limestone sulfation opportunity. At the 1050℃,
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6
Oviedo ICCS&T 2011. Extended Abstract
the limestone sulfation opportunity is increased by the larger total of SO2 emission, but the high temperature accelerates the limestone sintering, the sulfation efficiency is not enhanced markedly. Effect of atmosphere on SO2 removal by limestone Comparing the Fig. 9a and 9b, it can be obtained that the sulfation efficiency in the O2/N2 atmosphere is smaller than that in O2/CO2 atmosphere at lower temperature (<850 ), and the opposite tendency was shown with the temperature further increasing. This phnomence due to that: in O2/N2=20/80 atmosphere, the smaller release amount of SO2 reduce the opportunity for sulfation reaction at the 750 , 850
is the best sulfation
temperature in the O2/N2=20/80 atmosphere from the result of previous studies [8], so the sulfation efficiency increases obviously. The larger total of SO2 emission in O2/N2=20/80 atmospheres promtote the sulfur retain at the temperature of 950
and
1050 . Effect of limestone particle on SO2 removal by limestone Fig. 10 shows the sulfation efficiency with show that the sulfation efficiency decrease remarkably with the particle size increasing. This is due to the smaller particle size, the
Desulfurization Efficiency (%)
different limestone particle sizes. The results
32
larger specific surface areas. The sulfation reaction probability improved, and the sulfation rate increased [9].
30 28 26 24 22 20 18 ≤
0.063
0.1 ～ 0.15 0.063 ～ 0.1 Particle Size (u m)
0.3 ～ 0.45
Fig.10 Effect of limestone particle on SO2 removal
Effect of Ca/S on SO2 removal by limestone The influence of Ca/S on the sulfation efficiency is presented in Fig. 11. The increasing of the Ca/S when the Ca/S is less than 2. The increase of sulfation efficiency becomes more gradual when the Ca/S is
Desulfurization Efficiency (%)
sulfation efficiency increase remarkably with
35 30 25 20 15 10
more than 2. These results are in according with the general conclusions in SO2 removal
5 0 0
1
Ca/S
2
3
Fig.11 Effect of Ca/S on SO2 removal
by limestone in the conventional combustion [10]. 4.
Conclusions
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7
Oviedo ICCS&T 2011. Extended Abstract
Using horizontal fixed bed, the behavior and mechanism of SO2 emission and removal fundamental experiments were investigated in oxygen-enriched combustion. The main conclusions are drawn: (1) The peak value of SO2 emission in O2/CO2 atmosphere is smaller than that in O2/N2 atmosphere because the high heat capacity of CO2, the total of SO2 emission in O2/N2 atmosphere is small than that in O2/CO2 atmosphere at the lower temperature (≤850 ), the opposite tendency was shown at the higher temperature (≥950 ). (2) The high O2 concentration in O2/CO2 atmosphere promote the SO2 emission. The higher of the O2 concentration, the larger of the peak value. The rise in temperature promote the SO2 emission, increase the peak value of the SO2 emission curve. The total of SO2 emission increase with the temperature increasing in the O2/N2 and O2/CO2 atmosphere, except at the 950℃ in the O2/CO2=20/80 atmosphere which the least total of SO2 release happen. (3) The SO2 removal by limestone was affected not only by the reaction condition but also by the characteristics of SO2 emission. The sulfation efficiency decrease remarkably with the particle size increasing. The sulfation efficiency increase obviously with increasing of the Ca/S when the Ca/S is less than 2, and the increase become more gradual when the Ca/S is more than 2. Acknowledgement The author wishes to express the great acknowledges of financial support from National Nature Science Foundation of China (No. 50876036 and No. 51021065) and “Key Projects of National Fundamental Research Planning” (National 973 project: 2010CB227003) References [1]. Czakiert T, Bis Z, Muskala W, Nowak W. Fuel conversion from oxy-fuel combustion in a circulating fluidized bed. Fuel Processing Technology 2006;87(6):531-8. [2]. Croiset E, Thambimuthu K V. NOx and SO2 emissions from O2/CO2 recycle coal combustion. Fuel 2001;80(14):2117-21. [3]. Duan L, Zhao C, Zhou W, Liang C, Chen X. Sulfur evolution from coal combustion in O2/CO2 mixture Journal of Analytical and Applied Pyrolysis 2009;86(2):269-73. [4]. Liu H, Zailani R, Gibbs BM. Comparisons of pulverized coal combustion in air and in mixtures of O2/CO2. Fuel 2005;84(7/8):833-40. [5]. Chen J, Liu Z, Huang J. Emission characteristics of coal combustion in different O2/N2, O2/CO2 and O2/RFG atmosphere. Journal of Hazardous Materials 2007;142(1/2):266-71.
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8
Oviedo ICCS&T 2011. Extended Abstract [6]. Zhou Y, Zheng Y, Zhang L, Zheng C. Calcium-based desulfuration agent under the gaseous condition of air and O2/CO2. Journal of Engineering for Thermal Energy Power 2001;16(4):40911. [7]. Jia L, Tan Y, Wang C, Anthony EJ. Experimental study of oxy-fuel combustion and sulfur capture in a mini-CFBC. Energy Fuels 2007;21(6):3160-4. [8]. Liu D, Yan W. Fluidized bed combustion technology. 1rd ed. Beijing: China Electric Dower Press; 1995. [9]. Hajaligol MR, Longwell JP, Sarofim AF. Analysis and modeling of the direct sulfation of calcium carbonate. Industrial & engineering chemistry research 1988;27(12):2203-10. [10]. Anthony EJ, Granatstein DL. Sulfation phenomena in fluidized bed combustion systems Progress in Energy and Combustion Science 2001;27(2):215-36.
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9
Oviedo ICCS&T 2011. Extended Abstract
Current status of the Chemical Looping Combustion Technology F. García-Labiano, L. F. de Diego, P. Gayán, A. Abad, J. Adánez Department of Energy and Environment, Instituto de Carboquímica (ICB-CSIC) Miguel Luesma Castán 4, 50018 Zaragoza, Spain Tel: +34 976 733 977 Fax:+34 976 733 318, e-mail:[email protected]
Abstract This work describes the current situation of the Chemical-Looping Combustion (CLC) and Chemical-Looping Reforming (CLR) Technologies. The process is based on the transfer of the oxygen from air to the fuel by means of a solid oxygen-carrier avoiding direct contact between fuel and air. CLC has arisen during last decade as a very promising combustion technology for power plants and industrial applications with inherent CO2 capture which avoids the energetic penalty present in other competing technologies. CLR uses the chemical looping cycles for H2 production with additional advantages if CO2 capture is also considered. Although it can be considered as a new technology, during last decade CLC/CLR have suffered a great development regarding the use of gaseous or solid fuels, the manufacture of new oxygen-carriers and the operational experience in continuous units. Up to the end of 2010, more than 700 different materials based on Ni, Cu, Fe, Mn, Co, as well as other mixed oxides and low cost materials, have been produced. Among them, 36 have been tested in pilot plants under continuous operation. The total time of operational experience (≈ 3500 h) in different CLC/CLR units in the size range 0.3-120 kWth has allowed to demonstrate the technology and to gain in maturity. Considering the great advances reached up to date in a very short period of time, it can be said that CLC/CLR are very promising technologies within the framework of the CO2 Capture and Storage (CCS) options.
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1
Oviedo ICCS&T 2011. Extended Abstract
1. Introduction According to the analysis made by the IPCC [1] and IEA[2], the CO2 Capture and Storage (CCS) could account for 19% of the total CO2 emission reductions needed for this century to stabilises climate change at a reasonable cost. The purpose of CCS technology is to produce a concentrated stream of CO2 from industrial and energyrelated sources, transport it to a suitable storage location, and then store it away from the atmosphere for a long period of time. The IPCC Special Report on Carbon Dioxide Capture and Storage [1] gives an overview of the different options available for the capture, transport and storage processes. Regarding CO2 capture, three main approaches were considered for industrial and power plants applications: post-combustion systems, oxy-fuel combustion, and pre-combustion systems. All theses technologies have undergone a great development during the last years and some of them are available at commercial scale [3]. However, although most of the technologies can reduce CO2 emissions, they also have a high energy penalty, which results in a lower overall energetic efficiency and an increase in the price of the energy. In this context, the Chemical-Looping Combustion (CLC) process has been suggested among the best alternatives to reduce the economic cost of CO2 capture [4]. The estimated cost of the capture per tonne of CO2 avoided was 6-13 € for CLC, much lower than the calculated for other technologies of CO2 capture that can reach values as high as 40 € for the same purpose. The main drawback attributed to CLC is a low confidence level as a consequence of the lack of maturity of the technology. It must be considered that this is an emerging technology although during the last 10 years it has experienced a great development.
2. Concept The process is based on the transfer of oxygen from air to the fuel by means of a solid oxygen-carrier avoiding direct contact between fuel and air. Fig. 1 shows a general scheme of this process and the reduction and oxidation reactions taking place in the system. Submit before May 31st to [email protected]
2
Oviedo ICCS&T 2011. Extended Abstract
N2, O2
Reduction reaction
Oxidation reaction Air
CO2, H2O
MexOy
MexOy-1
Fuel
Figure 1. Scheme of the Chemical-Looping Combustion (2n+m-p) MexOy + CnH2mOp → (2n+m-p) MexOy-1 + n CO2 +m H2O
ΔHr
(1)
(2n+m-p) MexOy-1 + (n+m/2-p/2) O2 → (2n+m-p) MexOy
ΔHo
(2)
ΔHc = ΔHr +ΔHo
(3)
CnH2mOp + (n+m/2-p/2) O2 → n CO2 + m H2O
In a first step, the fuel is oxidized to CO2 and H2O by a metal oxide (MexOy) that is reduced to a metal (Me) or a reduced form MexOy-1. If the composition of the fuel gas is expressed as CnH2mOp, the global reduction process is given by reaction (1). The gas produced in this first step contains primarily CO2 and H2O. After water condensation and purification, a highly concentrated stream of CO2 ready for transport and storage is achieved. This concept is the main advantage of the process in relation with other CO2 capture technologies. In this sense, CLC is a combustion process with inherent CO2 separation, i.e. avoiding the need of CO2 separation units and without any penalty in energy. The metal or reduced metal oxide is further oxidized with air in a second step, and the material regenerated is ready to start a new cycle (reaction 2). The flue gas obtained contains N2 and unreacted O2. The net chemical reaction over the two steps, and therefore the combustion enthalpy, is the same as in conventional combustion where the fuel is burned in direct contact with oxygen from air (reaction 3). Therefore, the total amount of heat evolved in the CLC process is the same as in conventional combustion.
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3
Oviedo ICCS&T 2011. Extended Abstract
3. Applications The chemical-looping concept can have different applications, being the combustion and the H2 production the most relevant ones. For combustion purposes, the oxygen depleted solid material must be regenerated by oxygen in air. In general, these processes are known with the general term “Chemical-Looping Combustion” (CLC). CLC processes can address gaseous, liquids, and solid materials as primary fuels. Among the technologies for solid fuels can be differentiated: (1) the Syngas-CLC process where the oxygen-carrier comes into contact with the gasification products (syngas) obtained in a gasifier; and (2) the in-situ Gasification CLC (iG-CLC) process where the solid fuel and oxygen-carrier is mixed in the same reactor. If the oxygen-carrier is able to release gaseous oxygen for the fuel conversion, the process is called Chemical-Looping with Oxygen Uncoupling (CLOU). Figure 2 shows an example of the different type of reactions in which the oxygen-carrier is involved depending on the process. For H2 production, the regeneration of the oxygen-carrier can be done by using oxygen e.g. using air- or steam. When air is used for regeneration, it can be differentiated: (1) the Steam Reforming integrated to Chemical-Looping Combustion process (SR-CLC), where CLC is used to give the energy required for usual catalytic Steam Reforming; and (2) the autothermal Chemical-Looping Reforming process (a-CLR) where primary products from the chemical-looping system are H2 and CO.
CO2 H2O
H2 O
CO2 H2O
H2O
CO2 H2O
CO2
Oxygen-Carrier
Char
Volatiles
CO H2
Volatiles
CO H2
Oxygen-Carrier
Syngas Syngas-CLC (gas fuel)
O2
Char
Coal
H2O
Coal
CO2
Oxygen-Carrier
H2O and/or CO2
CO2
iG-CLC (solid fuel)
CLOU (solid fuel)
Figure 2. Main processes involved in the FR for solid fuel processing in a CLC system. Submit before May 31st to [email protected]
4
Oviedo ICCS&T 2011. Extended Abstract
The applications of Chemical-Looping technologies for fossil energy conversion have been briefly overviewed by Fan and Li [5], and Fang et al. [6], and a deeper description of Chemical-Looping technologies can be found in the works of Fan [7], Brandvoll [8], and Adánez et al. [9].
4. Oxygen carrier materials One of the key issues in the system performance is the oxygen-carrier material, which must accomplish the following characteristics: sufficient oxygen transport capacity, favourable thermodynamic regarding the fuel conversion to CO2 and H2O in CLC, high reactivity and maintained during many successive redox cycles, resistance to attrition, negligible
carbon
deposition,
good
fluidization
properties
(no
presence
of
agglomeration), limited cost, and environmental friendly characteristics. Normally, the pure metal oxides do not fulfil the above characteristics and reaction rates quickly decreased in a few cycles, showing the need to use of a support, which increases the mechanical strength and attrition resistance, and improves reactivity. In this sense, the method used in the preparation of the materials strongly affects the properties of the oxygen-carrier. Several methods for preparation of oxygen-carriers have been used in the past, most of them for laboratory scale production [9,10]. At the moment, the preparation methods planned for large-scale production include spray-drying, spin flash and impregnation. An important background has been reached in the development of the oxygen-carriers, with the testing of more than 700 different materials, mainly based on nickel, copper and iron. Hossain and de Lasa [11] reviewed the progress reached in the development of oxygen-carrier materials. A compilation of all the oxygen-carriers prepared up to the end of 2010, including metal oxide content, support material, preparation method and facility used for testing can be found in the review of Adánez et al. [9]. After the initial screening, 36 of them have been tested in continuous units, some of them used in more than 1 application. This gives us a total of 44 combinations of oxygen carriers and applications for a total of ≈3500 hours of operational experience in continuous units. Table 1 shows a summary of the operational experience regarding oxygen carriers, which includes the metal oxide and application. Submit before May 31st to [email protected]
5
Oviedo ICCS&T 2011. Extended Abstract
Table. Summary of the operational experience on CLC technologies in continuous units. Number of oxygen carriers Nickel Copper Iron Manganese Cobalt Mixed oxides Low cost materials TOTAL
Operational experience (hours)
CLC
CLCs
CLR
Total
CLC
CLCs
CLR
Total
16 3 4 1 1 5 3 33
2
6
24 3 5 1 1 5 5 44
2114 391 97 70 25 82 111 2890
160
284
2558 391 127 70 25 82 199 3452
1
2 5
6
30
88 278
284
It can be observed that Ni-based oxygen carriers have been the most extensively materials analyzed. These materials have therefore the most operational experience in continuous units, with a total of 2558 hours. These materials have been used for all the possible applications including combustion and H2 production processes (CLC, CLC with solid fuels –CLCs-, and CLR). The catalytic activity of metallic Ni made of them the best materials for CLR applications. However, the use of Ni-based materials may require safety measures because of its toxicity and other alternatives are currently under development. Copper is the other material with high operational experience for a total of 391 h in continuous units, only based on combustion processes. In addition, the use of low-cost materials is increasing in order to decrease the cost of the oxygen carrier regarding raw materials, particle preparation and the most importantly, disposal and environmental costs.
5. CLC pilot plants The chemical-looping concept showed in Figure 1 can be accomplished in different type of reactors and configurations. At the present, all the CLC plants existing worldwide use the configuration composed by two interconnected fluidised-bed reactors working at atmospheric pressure, although different configurations can be used [9, 12]. Table 2 shows a summary of the existing chemical-looping units with power output higher than 10 kWth including the type of gas fuel and oxygen carrier tested, the configuration, and the operational experience. It must be remarked the 10 kWth pilot plant located at Chalmers University of Technology used for the first long-time demonstration of the CLC technology during 100 h of continuous operation with the same batch of Ni-based
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Oviedo ICCS&T 2011. Extended Abstract
oxygen carriers. At the moment, this plant accumulates more 1350 h of operational experience using different Ni-based and Fe-based oxygen carriers, including the long term test (>1000 h) carried out with the same batch of material [13]. Up to now, the highest scale corresponds to the 120 kWth dual circulating fluidized bed pilot plant located at Vienna University of Technology, which has been successfully operated during more than 90 h with Ni-based materials, both for CLC and CLR applications [14,15]. The next step in the development of the CLC technology is the 1 MWth coal fuelled CFB unit located at Darmstadt University of Technology, TUD [16,17], whose operation is planned to start in 2011. However, to achieve competitive energy efficiencies it is necessary to operate at high temperatures and high pressures (1-3 MPa). In this sense, operating pressurised CLC plants using interconnected fluid bed technology could have some technical difficulties to maintain stable circulation between the reactors. Therefore, dynamically operated packed-bed reactors have been proposed if pressurised operation is desired. The operational experience of this technology is limited although could represent an important challenge for natural gas combustion in the future.
Table 2. Summary of the chemical-looping units with power output >10 kWth. Research Center Location Gaseous Fuels Chalmers University of CHALMERS, Sweden Institute of Carboquímica, ICB-CSIC, Spain IFP-TOTAL France
Technology,
Xi’an Jiaotong University China ALSTOM Power Boilers France Korean Institute of Energy Research, KIER, Korea Technical University of Viena, TUWIEN, Austria Solid fuels Chamers University of Technology, CHALMERS, Sweden Southeast University, China Darmstadt University of Technology, TUD, Germany
Unit size kWth
Configuration
Oxygencarrier
10
Interconnected CFB-BFB Interconnected BFB-BFB Interconnected BFB-BFB-BFB
NiO, Fe2O3
1350
CuO
200
NiO
n.a.
Fe2O3/CuO
15
NiO
100
NiO, CoO NiO, CoO NiO, ilmenite NiO
28 300 >90 20
ilmenite
90
NiO, Fe2O3 ilmenite
130 --
10 10 10 15 50 50 120 (CLC) 140 (CLR) 10 10 1000
Interconnected Pressurised CFB-BFB Interconnected CFB-BFB Interconnected CFB-BFB (KIER-1) BFB-BFB (KIER-2) DCFB
Interconnected CFB-BFB CFB-spouted bed Interconnected CFB-CFB
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Operation experience (h)
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Oviedo ICCS&T 2011. Extended Abstract
6. Remarks on CLC development CLC and CLR have arisen during last years as very promising technologies for power plants and industrial applications with CO2 capture. The main advantages come from their inherent CO2 capture which avoids the energetic penalty of this process in other competing technologies. This technology has suffered a great advance in different aspects such as material development, operational experience in continuous units, and process design and scaleup. An important background has been reached during last years in the development of oxygen-carriers, with the testing of more than 700 different materials mainly based on nickel, copper and iron. CLC has been demonstrated at scales up to 150 kWth with natural gas and the operation with coal is currently starting in a 1 MWth CLC plant. More of the experience has been gained for the use of gaseous fuels including long term tests (up to 1000 h). The total time of operational experience in continuous units including all fuels and technologies is about 3500 h at the end of 2010. However, this technology has important challenges in the near future: - The use of CLC for coal and solid fuels. In situ gasification of coal in the fuelreactor using cheap oxygen-carriers as natural minerals or industrial waste products, or the use of materials with CLOU properties are two very promising possibilities. - The use of cheap and friendly environmental materials as oxygen carriers. - The development of pressurised operation. - The scale up of the technology is a very important issue that needs to be accomplished in the near future to finally demonstrate that CLC is an economical alternative to other CCS technologies. As a consequence, the future of the chemical-looping technologies is very promising during the next years and with real prospects to be implemented at commercial scale.
Acknowledgement The authors want to thanks to the European Commission, CO2 Capture Project (CCP), Spanish National R&D&I Plan, and Gobierno de Aragón for their financial support on projects that helped the investigation in chemical-looping technologies. Submit before May 31st to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
References [1] IPCC. IPCC special report on carbon dioxide capture and storage, Cambridge, UK: Cambridge University Press; 2005. [2] International Energy Agency (IEA). Energy technology perspectives: Scenarios and strategies to 2050. París, France: OECD/IEA; 2006. [3] Toftegaard MB, Brix J, Jensen PA, Glarborg P, Jensen AD. Oxy-fuel combustion of solid fuels. Prog Energy Combust Sci 2010;36:581-625. [4] Kerr HR. Capture and separation technology gaps and priority research needs. In: Thomas DC, Benson SM, editors. Carbon dioxide capture for storage in deep geologic formations– Results from the CO2 capture project, Oxford, UK: Elsevier; 2005, vol. 1, Chapter 38. [5] Fan & Li. Chemical Looping technology and its fossil energy conversion applications. Ind Eng Chem Res 2010; 49: 10200-11. [6] Fang H, Haibin L, Zengli Z. Advancements in development of chemical looping combustion: A review. Int J of Chem Eng 2009. doi:10.1155/2009/710515. [7] Fan L-S. Chemical looping systems for fossil energy conversions. Hoboken, New Jersey, USA: John Wiley & Sons, Inc; 2010. [8] Brandvoll Y. Chemical looping combustion. Lap Lambert Academic Publishing; 2010. [9] Adánez J, Abad A, García-Labiano F, Gayán P, de Diego LF. Progress in ChemicalLooping Combustion and Reforming Technologies. A review. Progress Energy Combust Sci 2011. [10] Adánez J, García-Labiano F, de Diego LF, Gayán P, Abad A, Celaya J. Development of oxygen carriers for chemical-looping combustion. In: Thomas DC, Benson SM, editors. Carbon dioxide capture for storage in deep geologic formations formations – Results from the CO2 capture project, Oxford, UK: Elsevier; 2005, vol. 1, Chapter 34. [11] Hossain MM, de Lasa HI. Chemical-looping combustion (CLC) for inherent CO2 separations-a review. Chem Eng Sci 2008;63:4433-51. [12] Lyngfelt A. Oxygen carriers for chemical looping combustion - 4000 h of operational experience. Oil & Gas Sci Technol 2011;66:161-72. [13] Linderholm C, Lyngfelt A, Béal C, Trikkel A, Kuusink R, Jerndal E, Mattisson T. Chemical-looping combustion with natural gas using spray-dried NiO-based Submit before May 31st to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
oxygen carriers. In: Eide LI, editor. Carbon dioxide capture for storage in deep geological formations– Results from the CO2 capture project, UK: CPL Press; 2009, vol. 3. Chapter 6. [14] Kolbitsch P, Pröll T, Bolhar-Nordemkampf J, Hofbauer H. Operating experience with chemical looping combustion in a 120 kW dual circulating fluidized bed (DCFB) unit. Int J Greenhouse Gas Control 2010;4:180-5. [15] Pröll T, Bolhàr-Nordenkampf J, Kolbitsch P, Hofbauer H. Syngas and a separate nitrogen/argon stream via chemical looping reforming- A 140 kW pilot plant study. Fuel 2010;89:1249-56. [16] Beal C, Epple B, Lyngfelt A, Adánez J, Larring Y, Guillemont A, Anheden M. Development of metal oxides chemical looping process for coal–fired power plants. Proc 1st Int Conf on Chemical Looping. Lyon, France; 2010. [17] Ströhle J, Orth M, Epple B. Simulation of the fuel reactor of a 1 MWth chemical looping plant for coal. Proc 1st Int Conf on Chemical Looping. Lyon, France; 2010.
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Oviedo ICCS&T 2011. Extended Abstract
Oxygen transfer from metal oxides during chemical looping combustion of Victorian brown coal – An experimental and modelling study Chiranjib Saha, Teck Xin Seng, Anthony Auxilio*, Sankar Bhattacharya Department of Chemical Engineering, Monash University, Clayton Campus, Victoria – 3800, Australia * Corresponding author: [email protected] Phone: +61 3 9905 9623 Fax: +61 3 9905 5686 Abstract Chemical looping combustion is a novel technique where an oxygen carrier is used to transfer oxygen from air to the fuel, thus avoiding direct contact between air and fuel, and therefore mixing of the combustion product CO2 with nitrogen in air. This developing process is an alternative to conventional combustion and gasification where oxygen needs to be supplied from either air or energy-intensive air separation plants. Chemical looping has been widely studied for combustion of natural gas; however its application to solid fuel, such as coal, is being studied only recently. Coal is considerably more abundant and geographically more well spread than natural gas. This provides an impetus to assess the adaptability of coal in a chemical looping process. However, in doing so it is important to examine the extent to which oxygen can be transferred from the oxygen carrier to coal. This paper presents the results from an experimental and modelling study on oxygen transfer from NiO as oxygen carrier during chemical looping combustion. The reduction of NiO has been carried out using CO which is one of the main products of coal pyrolysis and gasification. The experiments are conducted in a thermogravimetric analyser. The operating temperature is varied and it is observed that the reaction rate was dependent on the reaction temperature and the gas concentrations. A general structural model is also presented for the description of non-catalytic gassolid reactions. The reaction between NiO particles and reacting gases during reduction is described by an unreacted shrinking core model to predict the transfer of oxygen from the matrix of NiO. The model, which incorporates parameters such as concentration of solid and gaseous reactants, particle size, weight fraction of the main and supporting particles, porosity and pore diffusion coefficient is applied to interpret the experimental results. The model is found to predict general trends exhibited by experimental data.
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1
Oviedo ICCS&T 2011. Extended Abstract
1. Introduction It is widely that the anthropogenic emission of greenhouse gas CO2, mainly due to power generation from fossil fuels is the major contributor to this increase in temperature. Among the fossil fuels coal currently accounts for 40% of the world’s electric power generation and in Australia it is almost 80%, a role that is expected to continue in the foreseeable future. Victorian brown coal represents a very large resource, over 500 years at current consumption rate and responsible for the steady economic development of the state as well as the country. Being carbon-intensive, conventional coal-fired power generation results in large CO2 emissions as well. Therefore, there is a strong incentive to develop technologies that would reduce or allow easier capture of CO2 emissions from coal-fired power plants. Presently three major approaches are under investigation for easier CO2 separation or capture from coal-fired units. These are post-combustion capture, pre-combustion capture and CO2 capture after oxy-fuel combustion. But all these processes suffer from high energy consumption, either due to the need for handling large volume of flue gas or for relying on energy-intensive oxygen separation units. Chemical looping combustion is is another developing technology for easier CO2 separation. 1.1. Chemical Looping Combustion – General Description The CLC system is composed of two reactors: an air reactor and a fuel reactor. In the fuel reactor, the fuel particles are reduced by the lattice oxygen of the metal oxide and reduced metal oxide is re-oxidized in the air reactor to be used in the next cycle. The main advantage of this system is a concentrated stream of CO2, free from dilution with nitrogen, can be obtained from fuel reactor after condensing the water vapour. A generalized description of the overall system is given in Figure 1 [1]. Reaction stoichiometry is given by the equations below: (2n+m) MyOx + CnH2m → (2n+m) MyOx-1 + mH2O + nCO2
(1)
MyOx-1 + ½ O2 → MyOx
(2)
where M stands for metal and MyOx as metal oxide. While the reduction reaction is endothermic, the oxidation reaction is exothermic. As a result, the total heat evolved for the combined reduction and oxidation steps remains the same, as the one observed in a conventional combustion where the fuel is burned in Submit to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
direct contact with oxygen from air. The hot flue gas from air reactor is passed through the HRSG for steam generation to expand using steam turbine for electricity. If the hot depleted air from air reactor is at elevated pressure, it can be expanded using a gas turbine for power generation in combined cycle mode (Figure 2). Thus CLC does not bring any enthalpy gains given the overall heat generation is equal to the heat of combustion. Its main advantage, however, resides in the inherent separation of both CO2 and H2O from the flue gases. In addition, CLC also minimizes NOx formation since the fuel burns in the fuel reactor in an air free environment and the reduced oxygen carrier is re-oxidized in the air reactor in the absence of a fuel, at comparatively lower temperatures. NOx formation usually occurs well above 1200°C a potentially maximum temperature for CLC. 1.2. Chemical looping combustion of coal The CLC process has in a few years grown from a paper concept to a promising process for CO2-capture and a number of prototypes have been designed [2]. However, important progress of this technology has been made primarily with gaseous fuel as feedstock [3]. Its application with solid fuel, such as coal, has started only recently. Three different approaches are under investigation for coal combustion through chemical looping. N 2 , O2
CO2 + H2O MyOx
Air Reactor
Fuel Reactor
MyOx-1 Air
Fossil Fuel
Figure 1. Schematic of Chemical Looping Combustion.
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Oviedo ICCS&T 2011. Extended Abstract H2O
STEAM CYCLE
STACK
POWER GENERATION
COMPRESSOR
HRSG
ST
HP
CO2-T
COMPRESSOR
LIQUID CO2
GT
IP
COAL FEEDER
CONDENSER CYCLONE
N2/O2
PUMP CO2 CAPTURE
FR
PUMP AR
MyOX MyOX-1
AIR
AIR
COMPRESSOR COMBINED CYCLE CHEMICAL LOOPING COMBUSTION
Figure 2. Combined cycle chemical looping combustion with CO2 capture Chemical Looping Combustion for syngas derived from coal: In this process coal is first externally gasified and the product syngas is then fed into the fuel reactor for combustion through chemical looping. The schemetic of this process is shown in Figure 3. With this basic approach the development of syngas chemical looping technology has been started for hydrogen production as well CO2 capture during electricity generation [4].However, in this process an energy intensive air separation unit is required for syngas production through coal gasification. In-situ chemical looping gasification and combustion of coal: In case of in-situ chemical looping combustion the coal particles are directly fed into the fuel reactor. The coal is then converted in to syngas in presence of some gasification agents like steam or CO2 in-situ. The syngas or mixture of CO and H2 then reacts with metal oxide from air reactor to produce CO2 and H2O in the fuel reactor as shown in Figure 4. Chemical Looping with oxygen uncoupling (CLOU): Chemical-looping with oxygen uncoupling (CLOU) is a novel method to burn coal in gas-phase oxygen. The carbon dioxide from combustion is inherently separated from the rest of the flue gases. CLOU is based on chemical-looping combustion (CLC) and involves three steps in two reactors, one air reactor where a metal oxide captures oxygen from the combustion air (step 1), and a fuel reactor where the metal oxide releases oxygen in the gas-phase (step 2) and where this gas phase oxygen reacts with a
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Oviedo ICCS&T 2011. Extended Abstract
fuel (step 3). N 2 /O2 CO2 /H2 O My Ox
CO2 Condenser
Air Reactor
Fuel Reactor
H2 O (l)
My Ox-1
CO +H2
CO2 or H2 O
Air
Coal Gasification Air
O2 ASU
Figure 3. Ex – situ CLC of coal N 2 /O2 CO2/H2 O My Ox
CO2 Fuel Reactor
Condenser
CO2 + H2 O
Air Reactor
Fuel Reactor
H2O (l)
MyOx-1
Oxygen carrier
Coal Humidity Ash Volatiles
CO H2
Air
CO2 or H2 O (V)
Coal H2O
H2 O
Figure 4. In – situ CLC of coal N2/O2
CO2 +H2O
Coal, H2O, CO2
N2/O2
CO2 +H2O
Reactions in Fuel reactor Air Reactor
Gasification Reactions
MyOx MyOx + H2 – MyOx-1 + H2O
Air Reactor
My Ox
C+ H2 O – CO+ H2
MyOx-2 + O2 (g) - MyOx
C+ CO2 – 2CO
MyOx + CO – MyOx-1 + CO2
Fuel reactor MyOx - MyOx-2 + O2 (g)
MyOx-2 + O2 (g) - MyOx
My Ox-2
MyOx-2
C nH 2m + (n+m/2)O2 (g) – nCO2 + mH 2O
CO/H2
Coal Char (CO 2/H 2O) Air
In-situ CLC
(A)
Air
CLOU
(B)
Figure 5. Reactions in fuel reactor in case of (A) in-situ CLC and (B) CLOU In case of in-situ chemical looping combustion of solid fuels there is a need for an intermediate gasification step of the char with steam or carbon dioxide to form reactive gaseous compounds which then react with the oxygen carrier particles. But no such gasification step is needed in case of CLOU [5]. Figure 5 shows the difference between Submit to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
in-situ CLC and CLOU process. The oxygen carrier for CLOU needs to have special characteristics and needs to react reversibly with gas-phase oxygen at high temperature. This places special thermodynamic and kinetic demands on the oxygen carriers for CLOU process. However, these special requirements of oxygen carrier particles are not needed in CLC [6,7]. In-situ chemical looping gasification process is advantageous over the syngas chemical looping as it eliminates the need for energy exhaustive air separation unit. Process simulation of both the syngas chemical looping (SCL) and in-situ or coal direct chemical looping (CDCL) is investigated by Gnanapragasam et al. [8] for co-production of hydrogen and electricity from Pittsburgh #8 coal. It has been observed that CDCL has fewer units in operations and hence less energy intensive as compared to the SCL. The CDCL system has higher H2/CO2 ratio as compared to SCL with lesser amount of steam requirement for gasification. Similar findings are reported by Fan et al. [9]. Insitu gasification of a German lignite and CO2 separation using chemical looping with a Cu-based oxygen carrier was investigated by Dennis and Scott [10] in laboratory scale fluidized bed reactor. Later on the same reactor was used by Dennis et al. [11] for performance evaluation of bituminous coal. A number of solid fuels (Mexican petroleum coke, coals from South Africa, China, Indonesia, Taiwan and France) have been used in both lab scale and bench scale in-situ chemical looping combustion experiments. But the application of in-situ chemical looping gasification and combustion with lignite is very limited. This project is the first ever study pertaining to application of in-situ chemical looping combustion with Victorian Brown coal [12]. Oxygen carriers perform the most important function in CLC. The most desirable characteristic of oxygen carriers is to transfer the oxygen from its matrix to the coal. As part of the big Project, the specific objective of this work is to present the non-catalytic gas solid reactions through a structural model. The reaction between metal oxide particles and reacting gases during reduction and oxidation is described by an unreacted core model to predict the transfer of oxygen from the matrix of metal oxides. The prediction of the model is compared with experimental works. 2. Experimental section The reduction of NiO has been carried out using CO which is one of the main products of coal pyrolysis and gasification. The experiments are conducted in a thermoSubmit to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
gravimetric analyser. Scanning electron microscope (SEM) was used for solids characterization. The amount of NiO used was 20 mg and the particle size was 100-150 µm. Experiments are carried out in alumina cruciales and the nickel oxide is heated from ambient temperature to two different operating temperature at a heating rate of 10°C/min under reducing environment. 3. Unreacted shrinking core model The transfer of oxygen during reaction of NiO with CO and conversion of NiO is described by an unreacted shrinking core model. The predictions of this model is compared with the experimental results. The assumptions of this model are described in literature [13]. The concept of gas-solid reactions in Unreacted shrinking core model assumes that the reactions proceed in the following two ways:
a) Diffusion through the product layer b) Chemical reaction on the solid surface The gaseous reactant A will form gas film and react with the unreacted core to form product layer. Then the gaseous reactant will have to diffuse through the product layer to react with unreacted core for complete reaction. a) For diffusion through the product layer, Rate of reaction of A is given by rate of diffusion to the reaction surface, (3) Flux within ash layer can be expressed by Fick’s law, (4) By integration from radius of unreacted core solid (R) to radius of shrinking unreacted core solid (rc ),time is obtained in term of conversion (x). (5)
Figure 6: Unreacted shirinking core model
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Oviedo ICCS&T 2011. Extended Abstract
b) For chemical reaction to occur (6) By integration from radius of unreacted core solid (R) to radius of shrinking unreacted core solid (rc ) ,time is obtained in term of conversion (x). (7)
Where
ρ
= Molar Density of Metal Oxide (kmol/m3 )
C AO = Molar Concentration of Gaseous Reactant (kmol/m3 ) R k
= Radius of Metal Oxide (m) = Reaction Rate Constant (m/s)
D
= Diffusion Coefficient (m 2 /s)
4. Results and Discussions Figure 7 represents the comparison of experimental results and models prediction for NiO conversion with time at different temperatures. It is clearly seen that the unreacted core model predicts the conversion of NiO well compared to the experimental results. Figure 8 shows that the particle size remains approximately the same after reaction, is porous and its shape is spherical which is an assumption used in the model. The EDX spectrum only detects Ni and O in the fresh and used particles (Figure 9). There is no sign of carbon deposition in both the used particles. The theoretical percentage of Ni and O in NiO is 78.58% and 21.42% respectively. The percentages of Ni and O in fresh particles used in this study are quiet similar values of 80.21% and 19.79% respectively which are close to theoretical values. The same percentages in case of used particles at 800°C are 90.75% & 9.25% and at 900°C are 89.10% & 10.90% respectively. The decrease in O-percentage from fresh particles is obvious due to donation of oxygen during combustion. However, a slighly lesser value of O-percentage in case of used particles at 800°C indicate a more conversion of NiO as compared to combustion at 900°C. Figure 7 makes this observation clear in case of both experimental result and model prediction. The reduced particles are more porous as expected (Figure 8). As assumed in the model, the porous product layer provides less resistance for the gas difusion (Figure 6) and the reaction can proceed further to the core. This observation clearly supports the above assumption.
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Oviedo ICCS&T 2011. Extended Abstract 1 0.9 0.8
Conversion, x
0.7 8 0 0 d e g re e c e lc iu s 9 0 0 d e g re e c e lc iu s
0.6 0.5 0.4
Solid lines – Model Prediction
0.3 0.2 0.1 0
0
100
200
300
400 500 T im e , t (s )
600
700
800
900
Figure 7: Effect of temperature on NiO conversion
(A) (B) (C) Figure 8: SEM images of NiO (500X) - (A) fresh, used (B) at 800°C and (C) at 900°C
(A) (B) (C) Figure 9: EDX Spectra of (A) fresh, used at (B) 800°C and (C) at 900°C NiO 5. Concluding Comments Chemical looping combustion is a novel technique where a metal oxide oxygen carrier is used to transfer oxygen from air to the fuel. In this way direct contact between air and fuel is avoided which makes CO2 separation easier from resulting flue gases. Metal oxide performs the most important function in CLC as it donates the oxygen required for fuel combustion from its lattice. An unreacted shrinking core model is presented for non catalytic gas-solid reaction during reduction of NiO with CO in a CLC system. The model’s prediction matches the experimental results. The assumptions of this model were verified by characterizing the combustion residues. It was observed that the porosity in used particles increased, which helped the reaction to proceed further and the gaseous products were able to diffuse from solid reaction surface into the surroundings.
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Oviedo ICCS&T 2011. Extended Abstract
No carbon deposition on the used particles was observed. Change in operating temperature had a minimal effect on particles conversion. This basic model is being updated by incorporating the coal properties to predict the solid-solid reactions as well. Series of experiments are in process using two different oxides, NiO & CuO, and a suit of Victorian brown coals-both raw and acid-washed forms. Acknowledgement The authors acknowledge the support of Brown Coal Innovation Australia (BCIA) and Monash Centre for Electron Microscopy for this work. Chiranjib Saha acknowledges Monash University for his scholarships. References
[1] Hossain MM, de Lasa HI. Chemical-looping combustion (CLC) for inherent CO2 separations – a review. Chemical Engineering Science 2008;63:4433-51. [2]Leion H, Lyngfelt A, Johanssona M, Jerndala E, Mattisson T. The use of ilmenite as an oxygen carrier in chemical-looping combustion. Chemical Engineering Research and Design 2008; 86:1017–26. [3]Adanez J, Cuadrat A, Abad A, Gayan P, de Diego LF, Labiano FG. Ilmenite activation during consecutive redox cycles in chemical-looping combustion. Energy & Fuels 2010;24:1402-13. [4]Gupta P, Velazquez-Vargas LG, Fan L-S. Syngas Redox (SGR) Process to Produce Hydrogen from Coal Derived Syngas. Energy & Fuels 2007;21:2900-08. [5] Mattisson T, Lyngfelt A, Leion H. Chemical-looping with oxygen uncoupling for combustion of solid fuels. International Journal of Greenhouse gas control 2009;3:11-9. [6]Leion H, Mattisson T, Lyngfelt A. Using chemical-looping with oxygen uncoupling (CLOU) for combustion of six different solid fuels. Energy Procedia 2009;1:447-53. [7]Mattisson T, Leion H, Lyngfelt A. Chemical-looping with oxygen uncoupling using CuO/ZrO2 with petroleum coke. Fuel 2009;88:683–90. [8]Gnanapragasam NV, Reddy BV, Rosen MA. Hydrogen production from coal using coal direct chemical looping and syngas chemical looping combustion systems: Assessment of system operation and resource requirements. International Journal of Hydrogen Energy 2009;34:2606-15. [9]Fan L-S, Li F, Ramkumar S. Utilization of Chemical Looping Strategy in coal gasification processes. Particuology 2008;6:131-42. [10]Dennis JS, Scott SA. In situ gasification of a lignite and CO2 separation using chemical looping with a Cu-based oxygen carrier. Fuel 2010;89:1623-40. [11]Dennis JS, Muller CR, Scott SA. In situ gasification of a lignite and CO2 separation using chemical looping with a Cu-based oxygen carrier: Performance with bituminous coals. Fuel 2010;89:2353-64. [12]Saha C, Roy B, Bhattacharya S. Chemical looping combustion of Victorian brown coal using NiO oxygen carrier. International Journal of Hydrogen Energy 2011;36:3253-59.
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Oviedo ICCS&T 2011. Extended Abstract
[13]Ryu H-J, Bae D-H, Han K-H, Lee S-Y, Jin G-T, Choi J-H. Oxidation and Reduction Characteristics of Oxygen Carrier Particles Reaction Kinetics by Unreacted Core Model. Korean Journal of Chem. Eng.. 2001;18:831-837.
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Oviedo ICCS&T 2011. Extended Abstract
Mineral-char interaction during the gasification of high ash coals in a fluidised bed gasifier: Redistribution of mineral phases within the char matrix. B.O. Oboirien12*, A.D. Engelbrecht1, B.C. North 1and R. Falcon2 1 Coal and Carbon Research Group, School of Chemical and Metallurgical Engineering Engineering, University of the Witwatersrand, Private Mail Bag 3, Wits 2050, South Africa, 2 CSIR Materials Science and Manufacturing, PO Box, 395, Pretoria 0001, South Africa
Abstract Mineral-char interaction can occur during the gasification of high ash coals in fluidised bed gasifiers when the melted minerals cover the char surface by penetrating into the cracks and pores of the char. The factors that determine the proportion of melted minerals covering the char surface are carbon conversion, temperature, and properties of the mineral matter (melting point). In this study, the extent of the l transformation of mineral phase and their redistribution within the char matrix was investigated for two high ash South African coals. The coal samples were gasified in a fluidised bed gasifier under oxygen-enriched conditions. The results indicate that for coals with a similar extent of transformation of mineral phases, carbon conversion and temperature are key factors that affect the distribution of the mineral phases. It was seen that an increase in carbon conversion and combustion temperature lead to more melted mineral covering the pores and surface of the char. When there is a variation in the relative amount of mineral phases, the melting point is the key factors that affect the distribution of the minerals. Mineral phases with high a proportion of low melting minerals such as Fayalite result in the formation of more slag (liquid) and a less viscous melt leads to more melted mineral covering the pores and the surface of the char
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Oviedo ICCS&T 2011. Extended Abstract
1. Introduction Mineral-char interaction can occur during the gasification of high ash coals in fluidised bed gasifiers when the melted minerals cover the char surface by penetrating into the cracks and pores of the char. Such inert mineral boundaries may be expected to severely reduce the ability of the carbon to react further by preventing the reaction gas from coming in contact with the carbon. [ 1-3].The melted minerals are mainly silicates that are formed from the decomposition of clays such as kaolinite, montmorillonite, illite and chamosite [4]. The extent of the mineral transformation mineral phases and their redistribution within the char matrix at a given temperature can be determined by the properties of the mineral phases[5]. Mineral phases with high a proportion of low melting point minerals result in the formation of more slag (liquid) and a less viscous melt leads to more melted mineral covering the pores and the surface of the char. In other studies[6] it was reported that the amount of residual carbon in char plays a role in the proportion of included minerals that covers the surface of the char. The lower the residual carbon in the char, the more included minerals become exposed on the char surface. They suggested that at a higher carbon conversion, more residual carbon in the char was consumed and there was not sufficient residual carbon to encapsulate the minerals.
In this study, the extent of the transformation of the mineral phases and their redistribution within the char matrix was investigated for two high ash South African coals. The coal samples were gasified in a fluidised bed gasifier under oxyen-enriched conditions
2. Experimental section 2.1 Samples Two South African coals used as feed in power stations were chosen for the experimental work. They are New Vaal (high ash with high inertinite content) and Grootegeluk power station middlings (high ash with high vitrinite content). Their properties are presented in Table 1.
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Oviedo ICCS&T 2011. Extended Abstract
Table 1: Coal properties Coal
Grootegeluk
New Vaal
Standard
Proximate analysis (as determined on air-dried basis) Ash content (%)
ISO 1171
31.70
37.15
Inherent moisture (%)
SABS 925
1.90
5.84
Volatile matter (%)
ISO 562
28.30
22.24
Fixed carbon (%)
By diff.
38.10
34.77
ISO 1928
21.4
18.11
Calorific value Calorific value (MJ/kg)
2.2 Experimental Procedure The gasification experiments were performed at atmospheric pressure in a bubbling fluidised bed reactor. Several gasified char samples were prepared from the two coal samples, using the in-house pilot-scale fluidized gasifier at the CSIR (Engelbrecht et al. 2010). The gasifier was operated under oxygen enriched conditions at 34% oxygen (13% enrichment) and at 910o and 960oC 2.3 Analysis The char samples were first milled to a particle size of -1.0 mm. A petrographic block of each sample was then prepared and polished in accordance with the ISO Standard 7404 2, 1985. The microscopic constituents were examined using a petrographic microscope at a magnification of X500 with oil immersion The relative proportions of the carbonrich organic constituents to inorganic materials were established on the basis of 500 point-counts on each sample according to the method set out in the ISO Standard 7404 3, 1994. Melted mineral matter was also identified. Some of this melted “slag” had coated carbon fragments or had penetrated into the carbon matrix, while some occurred in discrete separate bodies. The mineral compoistion of the coal and char samples were analaysed by XRD (X-ray Diffraction) in a Siemens D5000 powder diffractometer using Cu K α radiation. The mineral phase was identified using the reference intensity ratio (RIR). RIR values were determined from the PDF-2 date base (JCPDS-ICDD PDF2, 2003).
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Oviedo ICCS&T 2011. Extended Abstract
3. Results and Discussion Carbon content and Conversion The results of carbon and ash analyses conducted on the chars generated for both coals at bed temperaturas are presented in Tables 2 and 3. The results indicate that an increase in bed temperature led to an increase in carbon conversion (i.e. lower organic carbon content and higher ash content in the char particles) for both coals. At a constant temperature and residence time, New Vaal char particles had a higher carbon conversion than Grootegeluk char particles. New Vaal char particles had a low organic matter content and a high proportion of ash, while Grootegeluk char particles had a higher proportion of organic matter and lower proportion of ash. The melting of the ash can result in the melted minerals covering the surface of th char particles and preventing further carbon conversion. Table 2: Organic/ Inorganic constituents of Grootegeluk chars Bed
Carbon in the
Ash in the
Fixed carbon
Temperature
bed char (%)
bed char
conversion
(%)
(%)
(oC) 910
38.64
61.56
40.32
960
28.24
71.76
57.44
Table 3: Organic/ Inorganic constituents of New Vaal chars Bed
Carbon in the
Ash in the
Fixed carbon
Temperature
bed char (%)
bed char (%)
conversion
(oC)
(%)
910
5.49
94.51
80.45
960
3.29
96.71
85.57
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Oviedo ICCS&T 2011. Extended Abstract
Char-Mineral interaction The proportion of the melted mineral matter making contact with the carbon phase in th char matrix was estimated from petrographic analysis. The results obtained at 910oC and 960oC for both coals are presented in Table 4 and the Photomicrographs of the melted “slag” minerals (penetrating or surrounding carbon and occurring as separate bodies) are presented in Figure 1. Table 4:Organic/ Inorganic constituents Grootegeluk Char Samples
New Vaal Char Samples
910oC
960oC
910oC
960oC
Total organic material %
59
58
16
6
Relatively unchanged
33
18
24
16
1
10
2
4
7
14
58
74
visible minerals % Melted “slag” minerals penetrating/ surrounding carbon % Melted “slag” minerals separate bodies %
Figure 1: (a) Melted “slag” minerals - penetrating/ surrounding carbon (b) Melted “slag” minerals as a separate body
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Oviedo ICCS&T 2011. Extended Abstract
In the New Vaal char particles the highest consumption of carbon was found to have occurred in the char samples obtained at 960oC with 6% of total carbon remaining in the sample. The melted slag minerals predominated at 960oC (nearly 80% of the whole sample, compared to about 60% at 910oC). Very few instances of slag penetration into the carbon matrix were observed in either sample. The proportion of slag penetration into the carbon matrix doubled from 2% to 4 % when the temperature was increased from 910oC to 960oC
In Grootegeluk char, the proportion of carbon rich material in the char particles obtained at 910oC and 960oC were similar. The major difference lays in the proportion of melted mineral covering the char surface The proportion of slag penetration into the carbon matrix increased from 1% to 10 % when the temperature was increased from 910oC to 960oC.
A comparison of the proportion of slag penetration into the carbon matrix in New Vaal char and in Grootegeluk char particles showed that New Vaal char particles had a lower values despite the having a higher proportion of melted mineral. Based on the high proportion of melted minerals obtained in the NewVaal one would have expected a higher proportion of slag penetrating/covering the carbon. This suggest that the redistribution of the melted minerals can be determined from the behaviour of the mineral phases Mineral Transformation The mineral phases of the transformed mineral in the char samples were analysed by XRD. The XRD spectrum of Grootegeluk char sample generated at 960oC is presented in Figure 2. Smiliar spectra were obtained for the other test conditions. The content of mineral matters in the various chars generated at different temperatures was determined with the RIR method [7]. The results are presented in Figures 3 and 4. The main minerals that occur in the mineral transformation in char samples between 910oC and 960oC are mullite, quartz, wustite and fayalite. Mullite is formed due to the decomposition of metakaolinite when the temperature is higher than 900oC [8]. In this study the presence of mullite was only observed at 960oC for both New Vaal and
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Oviedo ICCS&T 2011. Extended Abstract
Grootegeluk char samples.
The concentration of mullite for Grootegeluk and New Vaal were
6% and 22 %
recpetively Wustite is formed at around 900 oC mainly due to the decomposition of hematite and siderite [9]. Wustite was only present in the Grootegeluk char samples. The concentration of musite decreased from 4.77% to 3 % when the temperature was increased from 910oC to 960oC.The decrease in the concentration of wustite at 960oC may be due to the formation of Fayalite. At a higher temperature, wustite reacts with mullite and quartz to form fayalite. [3]. The formation of low melting point minerals such as fayalite can lead to the formation of liquid slag phase. The molten material in slag can be used to determine the proportion of melted mineral that covers the char surface . Particles with lower melting point minerals such as fayalite will cover more surface area than a high melting point mineral such as mullite. [5,9].. This is due to different thermal stability The chemical activity of the highest occupied molecular orbits (HOMO) in mullite cluster is stronger than that of the lowest occupied molecular orbit (LUMO) in the mullite cluster. Based on the frontier orbital theory (Fujimoto, 1977), the frontier orbits in the HOMO or LUMO have more chemical reactivity and less thermal stability than any other orbits and play a key role in the reaction activity and thermal activity of minerals at the molecular level. Hence the energy which the electron in the external shell of the molecule needs to transfer from the HOMO to LOMO ( Δ E) in the mineral molecule can be used to calculate the thermal reaction activity and thermal stability of minerals a smaller value of Δ E means higher reaction activity and lower thermal stability. The Δ E value of fayalite is 1.659 eV, and is lower than that of mullite which is 6.8 eV [9] Ou results showed the presence of fayalite, a high melting point mineral in the Grooteguluk char.and fayalite was not found in New Vaal char.
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Oviedo ICCS&T 2011. Extended Abstract
Quartz Low
Counts GG 960 35
15000
Quartz Low
Oldhamite, syn Quartz Low
Quartz Low
Quartz Low Quartz Low
Fayalite Quartz Low
Anatase
Fayalite
Fayalite
Fayalite
Mullite Fayalite
5000
Fayalite
Oldhamite, syn
Quartz Low
10000
0 20
30
40
50
60
Position [°2Theta] (Cobalt (Co))
Figure 2 XRD spectrum of Grootegeluk char obtained at 960oC
100
90
80
70
(%)
60
GG910oC
50 GG960oC
40
30
20
10
0 Anatase
Calcite
Corundum
Fayalite
Magnetite
Mullite
Oldhamite
Pyrite
Quartz
Rutile
Wuestite
Figure 3 Mineral matter in Grootegeluk char samples at 910oC and 960oC
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Oviedo ICCS&T 2011. Extended Abstract
100
90
80
70
(%)
60 NV 910oC 50 NV 960oC 40
30
20
10
0 Anatase
C2AS Gehlenite
CaO Lime
Mullite
Oldhamite
Pyrite
Quartz
Figure .4 Mineral matter in New Vaal Char at 910oC and 960oC 4. Conclusions The results indicate that for coals with a similar extent of transformation of mineral phases, carbon conversion and temperature are key factors that affect the distribution of the mineral phases. It was seen that an increase in carbon conversion and combustion temperature lead to more melted mineral covering the pores and surface of the char. When there is a variation in the relative amount of mineral phases, the melting point and the viscosity of melted mineral are the key factors that affect the distribution of the minerals. Mineral phases with high a proportion of low melting minerals such as Fayalite result in the formation of more slag (liquid) and a less viscous melt leads to more melted mineral covering the pores and the surface of the char
Acknowledgement. The authors gratefully acknowledge the financial and other support received for this research from the CSIR and from the University of the Witwatersrand. EXXARO is acknowledged for supplying the coal samples. B.O.O. thanks CSIR for a doctoral scholarship.
References [1] Lin SY, Harito M, Horio M. Characteristics of coal char gasification at around ash melting melting temperature. Energy & Fuels 8 (1994) 598-606
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Oviedo ICCS&T 2011. Extended Abstract [2] Sekine Y , Ishikawa K, Kikuchi E, Matsukata M, Akimoto A. Reactivity and structural change of coal Turing steam gasification Fuel 85 (2006) 122-126. [3] Bai J, Wen L, Chun-zhu L, Zong-ging B, Bao-qing L. Influences of mineral matter on high temperature gasification of coal char. Journal of Fuel Chemistry and Technology 37 (2009) 134138. [4] Kosminski A, Ross DP , Agnew JB. Reactions between sodium and kaolin during gasification of low rank coal Fuel Processing Technology 87(2006), 1051-1062. [5] Gornostayev S S , Harkki JJ Mechamis of physical transformation of mineral matter in the blast furnance coke with refernce to its reactivity and strength. Energy& Fuels 20(2006) 26322635 [6] Li S, Whitty KJ. Investigation of coal char-slag transition during oxidation: Effect of temperature and residual carbon. Energy & Fuels 23(2009) 1998-2005 [7]Schreiner N A standard test method for the determination of RIR values by X-ray diffraction Powder Diffraction 10 (1995) 25-33. [8] Van Dyk JC, Melzer S, Sobiecki A Mineral matter transformation during Sasol fixed bed dry bottom gasification-utilisation of HT-XRD and Factsage modelling . Mineral Engineering 19 (2006) 1126-1135. [9] Zhang Z, Wu X, Zhou T, Chen Y, Hou N, Piao G, et al.. The effect of iron-bearing mineral meeting behaviour on ash deposition during coal combustión. Proceedings of Combustion Institute,33 (2011) 2853-2861.
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Oviedo ICCS&T 2011. Extended Abstract
Introduction of a Ternary Diagram for Comprehensive Evaluation of Gasification Processes for High Ash Coals
Martin Gräbner, Bernd Meyer TU Bergakademie Freiberg, Department of Energy Process Engineering and Chemical Engineering, Fuchsmühlenweg 9, 09599 Freiberg, Germany [email protected] Abstract Due to midterm depletion of oil and gas, coal gains importance not only as energy carrier but as feedstock for various chemical syntheses. Available gasification processes must be evaluated if they are capable for such conversion strategies if the ash content of the coal is increased. The present paper addresses the development of a comprehensive thermodynamic approach for the evaluation of gasification processes. A ternary diagram is introduced for a South African coal with an elevated ash content of 25 wt%(wf). The ternary diagram allows the evaluation of most of the commercially applied gasification technologies depending on the three variables O2, H2O and coal mass flow. Cold gas efficiency, dry CH4 yield, specific syngas production, H2/CO ratio as well as temperature and carbon conversion were selected as performance measures of the individual technologies. The graphical approach indicates the existence of optimum configurations for the specific coal depending on gasification pressure. A simplified user diagram was derived in order to adjust the existing processes to the needs of a high-ash coal. At a typical gasification pressure of 30 bar, a maximum cold gas efficiency of 87.4 % was identified at a temperature of 980°C, whereas the maximum syngas yield of 2.09 m²(H2+CO STP)/kg(waf) was located at 1135°C. The study is concluded by the evaluation of several fixed-bed, fluid-bed and entrained-flow processes assessing their potentials regarding the optimum domains within the ternary diagram.
1. Introduction The comprehensive assessment of gasification processes is difficult due to lots of independent variables such as coal composition and reactivity, temperature, pressure, H2O supply and varying boundary conditions. A well-known approach is to split all feed streams into their molar C-H-O composition plotting e.g. the equilibrium temperature for
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Oviedo ICCS&T 2011. Extended Abstract
soot formation it into a ternary diagram [1-3]. However, regarding performance parameters, technology comparison and optimum identification, the C-H-O plot has not been used, although a significant potential to illustrate basic relations is obvious. The present paper presents a new approach of plotting O2-H2O-coal mass flow in a ternary plot which allows assessment of the resulting temperature, carbon conversion, cold gas efficiency, dry methane yield, specific synthesis gas production as well as H2/CO ratio. A high-ash South African coal, see Table 1, was selected for the investigation since elevated ash contents pose a challenge to most of the commercial gasification processes. Hence, a detailed comprehensive investigation is recommended. Table 1. Ultimate analysis of South African HV C bituminous coal wt% (waf)
wt% (wf)
wt% (ar)
C
H
O
N
S
Ash
Moisture
79.58
4.06
13.34
2.07
0.95
25.33
6.04
2. Theoretical and Technological Background The O2-H2O-coal ternary diagram is developed by means of equilibrium modelling in the software Aspen Plus. The mass flow rates into the gasification systems are normalized to unity and treated as mass fractions in wt% in order to setup the diagram. According to Higman [4], a pressure of 30 bar is selected due to the suitability for various chemical syntheses and integrated gasification combined cycle (IGCC) power generation. The ConocoPhillips, GE, Shell, Siemens, HTW (high-temperature Winkler) and Sasol-Lurgi gasification technologies are selected for the assessment. For each entrained-flow and fluid-bed process a thermodynamic Aspen Plus model is developed applying unified boundary conditions. The models are verified by data given by Woods et al. [5] for ConcoPhillips, by McDaniel [6] for GE, by Rich et al. [7] for Shell, by Babcock [8] for Siemens, and by Bellin et al. [9] for HTW. The data for the Sasol-Lurgi fixed-bed gasifiers was taken from Modde and Krzack [10].
3. Results and Discussion a. Ternary diagram for temperature and carbon conversion In Figure 1, the temperature and carbon conversion is given while the regions for singlestage processes are marked. It can be seen that the dry-fed entrained-flow systems group on the O2-coal axis above the ash fluid temperature of 1440 °C indicating that no
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Oviedo ICCS&T 2011. Extended Abstract
moderator steam is necessary due to the high ash content. The slurry-fed systems can be found again above the ash fluid temperature, but limited by the mixing lines between the O2 and slurry concentration which is assumed between 50 to 70 wt% solids. If like in case of the ConocoPhillips gasifier a second stage exists, mixing points outside the slurry-fed domain are possible. The fluidized bed region is located below the ash softening temperature of 1285 °C and can be operated in carbon-rich bed inventory mode (right of 100 % carbon conversion line) or ash-rich bed inventory mode (left of carbon conversion line). The broad extension of the domain indicates the high flexibility in H2O supply in the fluidized-bed region. At temperatures below 800 °C, the fixed-bed gasifiers are located showing the highest observed H2O consumption. In this domain no equilibrium can be assumed, since kinetic restrictions and tar formation dominate the process. However, due to completeness the fixed bed gasifiers are integrated according to their consumption parameters.
Figure 1. Ternary diagram for temperature and carbon conversion (30 bar)
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Oviedo ICCS&T 2011. Extended Abstract
A great advantage of the diagram is to asses directly the effect of additional O2, H2O or coal injection with the help of mixing lines. For example, if it is considered to upgrade the Sasol-Lurgi fixed-bed gas by oxygen injection, the mixing line between the gasifier point and the top corner (O2) of the diagram applies following the law of inverse proportion for the intercepts as in every phase diagram. It is obvious that there is only limited potential if one considers performance parameters like cold gas efficiency or syngas yield which are described in the following.
b. Ternary diagram for cold gas efficiency and dry methane yield Figure 2 shows the cold gas efficiency which is defined on lower heating value basis [4]. A maximum region can be observed surrounding the 100 % carbon conversion iso-line. Along this line, a maximum cold gas efficiency of 87.4% was identified at 980 °C being inside the fluid-bed gasifier domain. At lower temperatures again an increase in cold gas efficiency can be determined. However, kinetic limitations and tar formation might offset the higher theoretical equilibrium values. Besides cold gas efficiency, the dry methane content of the product gas in vol% provides a useful measure to assess the gas quality regarding fuel gas application, synthetic natural gas production or CO2 capture capability. Figure 2 shows that in the carbon-rich region (carbon conversion < 100 %) the dry methane yield is mostly between 1 to 20 vol%, whereas left of the 100% carbon conversion line high cold gas efficiencies (>80%) accompanied by methane contents of <0.1 vol% can be reached which is typical for entrained-flow technologies. The very high methane yield near the H2O-coal line depicts the high expectable hydrocarbon production.
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Oviedo ICCS&T 2011. Extended Abstract
Figure 2. Ternary diagram for cold gas efficiency and dry methane yield (30 bar)
c. Ternary diagram for syngas yield and H2/CO ratio As a further measure to assess gasification, synthesis gas yield expressed as volume of CO and H2 at standard temperature and pressure related to the maceral fraction of the coal in m³(H2+CO STP)/kg(waf) was selected. For chemical syntheses, the molar H2/CO ratio is decisive as well to determine the gas treatment steps. Figure 3 presents the according plot for syngas yield and H2/CO ratio. The highest syngas yield of 2.09 m³(H2+CO STP)/kg(waf) was identified at 1135 °C. The maximum is again located in the fluid-bed region in the middle of a comparatively wide domain >2.0 m³(H2+CO STP)/kg(waf). This indicates that not only dry-fed entrained flow-processes deliver a high syngas yield as broadly concluded. The same amount of syngas can be supplied by fluid-bed processes as well accompanied by much higher H2/CO ratios up to 1.5 at 2.0 m³(H2+CO STP)/kg(waf) due to increased conversion of H2O to H2. The CO shift
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Oviedo ICCS&T 2011. Extended Abstract
conversion effort to meet the requirements of syntheses can thus be significantly reduced in comparison to dry-fed entrained-flow processes with H2/CO ratios of 0.25 to 0.5.
Figure 3. Ternary diagram for syngas yield and H2/CO ratio (30 bar) d. Optimum user diagram In order to merge all information which can be extracted from the ternary diagram, a simplified user diagram was developed for the specific syngas yield as exhibited in Figure 4. The user diagram allows identification of the maximum possible syngas yield for pressures between 1 and 100 bar and temperatures between 800 and 1500°C for the selected South African coal. The diagram is derived from the ternary diagram by setting the temperature and pressure to constants values and finding the maximum syngas yield. For the identified point the steam/oxygen ratio in kg/m³(STP), the molar H2/CO ratio, the temperature in °C and the oxygen consumption O2/caol in m³(STP)/kg(ar) were extracted or calculated from the flow streams. The relations are plotted in a quad
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Oviedo ICCS&T 2011. Extended Abstract
diagram where monotonic correlations are used to link the individual curves in the diagrams. The pink curve represents the maximum syngas yield path at varying pressure. Hence, it is possible to draw connection lines amongst the curves (see red arrow in Figure 4). All information to adjust the maximum syngas yield for a given pressure can be read off easily from the plot. The red arrows demonstrate the utilization of the diagram for a 30 bar process. The maximum syngas yield being 2.09 m³(H2+CO STP)/kg(waf) is selected as start point which can be extended along H2O/O2-line of 0.5 kg/m³(STP) to a temperature of 1135 °C. At constant temperature a molar H2/CO ratio of 0.6 with an oxygen consumption of 0.385 m³(STP)/kg can be found. The information might be used to design a new system or to adjust an existing gasifier in order to increase the syngas yield.
Figure 1. User diagram for optimum syngas yield for South African coal
4. Conclusions A new type of ternary diagram for comprehensive evaluation of gasification processes for a high-ash South African coal was developed. Depending on the three independent variables O2, H2O, and coal mass flow, an assessment of temperature, carbon conversion, cold gas efficiency, dry methane content, syngas yield, and H2/CO ratio was
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Oviedo ICCS&T 2011. Extended Abstract
accomplished. The diagram allows distinguishing operating regions for single-stage entrained-flow, fluidized-bed and fixed-bed processes. It is capable to evaluate multistage processes by simple mixing lines. The ConocoPhillips, GE, Shell, Siemens, HTW and Sasol-Lurgi gasification technologies were incorporated in the diagram. The maxima for cold gas efficiency and syngas yield were identified in the fluidized-bed region. An additional quad plot user diagram was derived for easy adjustment and design of gasifiers to achieve the optimum performance.
References [1] Mountouris A, Voutsas E, Tassios D. Solid waste plasma gasification: Equilibrium model development and exergy analysis. Energy Conversion and Management 2006;47(13-14):172337. [2] Li X, Grace J, Lim C, Watkinson A, Chen H, Kim J. Biomass gasification in a circulating fluidized bed. Biomass and Bioenergy 2004;26(2):171-93. [3] Li X, Grace J, Watkinson A, Lim C, Ergdenler A. Equilibrium modelling of gasification: a free energy minimization approach and its application to a circulating fluidized bed coal gasifier. Fuel 2001;80(2):195-207. [4] Higman C, van der Burgt M. Gasification. Elsevier Science, New York; 2003. [5] Woods MC, Capicotto PJ, Haslbeck JL, Kuehn NJ, Matuszewski M, Pinkerton LL, et al. Coast and performance baseline for fossil energy plants, volume 1 bituminous coal and natural gas to electricity, DOE/NETL-2007/1281. Tech. Rep.; National Energy Technology Laboratory; 2007. [6] McDaniel J. Polk power station 250 MW IGCC. Compact Course Gasification, TU Bergakademie Freiberg; 10.-12. November 2008. [7] Rich JWJ, Hoppe R, Choi GN, Hennekes RJ, Heydenrich R, Hooper M, et al. WMPI - waste coal to clean liquid fuels. Gasification Technologies Conference; 2003. [8] Babcock . Kombikraftwerk mit GSP-Flugstromvergasung. Deutsche Babcock Werke AG Brochure; 1992. [9] Bellin A, Dehms G, Karkowski G, Nassenstein C, Schrader L, Schumacher HJ. Kohlevergasung im Hochtemperatur-Winkler-Vergaser. Tech. Rep. FK 03E1092C, ISBN 3926732-07-5; Rheinbraun AG; 1988. [10] Modde P, Krzack S. Die Veredlung und Umwandlung von Kohle, Technologien und Projekte 1970 bis 2000 in Deutschland; chap. Gaserzeuger mit Drehrost. DGMK Deutsche Wissenschaftliche Gesellschaft für Erdöl, Erdgas und Kohle e.V.; 2008, p. 307-9. doi: 10.1002/cite.200950041.
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The determination of trace element distributions in coals—a new approach to an old method. D.A.Spears University of Sheffield, UK. E-mail [email protected] Abstract Differential dissolution is a long-established method for the determination of the composition of specific fractions in the coal. A standard method (British Standard 1017:6) to determine forms of sulphur in the coal is one example of such an approach. In an international collaborative programme on modes of occurrence of trace elements in coals (Davidson, 2000), three of the participating laboratories used differential dissolution. Unfortunately, quantitative extraction of all fractions is difficult to achieve and further treatment of the data is rewarding. If the analyses are comprehensive, and include major elements, then additional information is obtained from a statistical analysis of the data. With this approach, failure to quantitatively extract specific fractions can be overcome. The problem is analogous to statistically analyzing a set of samples representing one or more coal seams.
Introduction If it is known where trace elements are located within the coal, that is which fraction of the coal they are associated with, this is of value in resource prediction and exploitation. Improved coal characterisation benefits coal cleaning and utilisation, particularly in assessing environmental impacts from coal and coal by products. A number of methods have been employed to determine trace element associations in coals. Direct microanalytical methods are of value but may be restricted because of either detection limits or spatial resolution. The problem is that fine-grained minerals in coals, in some cases sub-micron in size, are intimately associated with the macerals. Scanning electron microscopes have the spatial resolution but do not have the required trace element detection limits. Detection limits with electron microprobes fitted with wavelength spectrometers are better, but still fall short of most requirements. We achieved a detection limit of about 100 ppm for elements such as Pb, with consequently many analyses falling below detection limits. The synchrotron XRF has at least two orders of magnitude improvement in sensitivity. The spatial resolution was about 50 μm when we had access to such an instrument (White et al., 1989), but is now much improved. X-ray absorption near edge structure 1
(XANES) spectroscopy has also proved of value for certain elements (Huggins et al, 2000). However, these are restricted facilities and beyond what could be expected in a well-found laboratory. This does not apply to laser ablation ICP-MS, which also has both the required sensitivity and spatial resolution. This rapidly evolving method was applied to coals by Chenery et al. (1995) and more recently by Spears et al. ( 2007). In the latter work the focus was on the trace element composition of the main maceral groups rather than the minerals. Another approach to the determination of trace element associations is the analysis of physically separated fractions. Ideally the fractions should contain concentrates of the different coal constituents, but quantitative separation is difficult to achieve and mathematical deconvolution is required, as in the work of Querol et al. in Davidson (2000).
An alternative strategy to physically
separating fractions in the laboratory is to select coal samples, usually a seam profile, which shows natural fractionation. The suite of samples could be based on more than one seam, possibly a coalfield, and statistical techniques are required to draw out the association. These are indirect methods of analysis as opposed to the direct instrumental methods. Finally, there is the chemical approach which uses differential dissolution to preferential attack and extracts specific fractions of the coal. This method was used by three of the participating laboratories, including ourselves, in an international collaborative investigation of trace element distributions in coal (Davidson, 2000). Although the procedures differ in detail they are broadly based on the standard methods to determine forms of sulphur in coal, in our case BS 1017:6. However, there is a need to standardise procedures in order to make results more comparable. Another difficulty is that quantitative extraction of specific fractions is not readily achieved, essentially the same problem as in the analysis of physically fractionated coals, and to address this problem statistical analysis has been performed, which is the new approach to an old method. The approach is therefore similar to that adopted, for example, in the analysis of geochemical data generated from the analysis of a suite of coal samples. In the one case however, the fractionation is natural and in the other it results from the differential dissolution. Differential dissolution stages and target coal fractions. As noted above the procedures adopted in our laboratory are based on the standard method to determine forms of sulphur in coals (BS 1017:6). In this method different forms of sulphur are released in sequential digestion steps. HCl is used to digest the sulphate S, and then HNO3 is used to digest pyritic S, which is indirectly determined from the Fe concentration in the HNO3 acid digest 2
assuming a composition of FeS2. The organic S is also determined indirectly from the difference between the total S and the S released in the HCl and HNO3 digestion steps. The digestion steps, with the introduction of an initial water digestion, were used to determine trace elements distributions (Cavender and Spears, 1995). In our contribution to the collaborative work reported in Davidson (2000) these steps were modified with a two step HNO3 acid treatment and the introduction of two stages to digest organic matter and then silicate minerals. In the procedures described below we have reverted to a single acid treatment for the pyrite following BS 1017: 6. Stage 1:
Approximately 0.5 g of the coal is shaken with 15 cm of deionised water in a sealed
plastic sample tube for a prolonged period of time up to 12 hours. The resulting suspension is centrifuged at 4000 rpm for 10 minutes. The supernatant solution is decanted off and made up to 25 cm3 with deionised water for ICP analysis. Stage 2:
The solid remaining from Stage 1 is shaken with 15 cm3 of dilute Hydrochloric acid (3%,
aristar grade) for up to 12 hours. The suspension is centrifuged at 4000 rpm for 10 minutes and the clear solution produced is decanted off and made up to 25 cm3 with deionised water for subsequent ICP analysis. Stage 3:
The residual solid from Stage 2 is shaken with 15 cm3 of dilute nitric acid (5%, aristar
grade) and digested on a hot plate for 30 minutes. The suspension is centrifuged as in Stages 1 and 2 and another 25 cm3 is prepared for analysis after decantation and dilution. Stage 4:
The solid remaining from Stage 3 is digested with 15 cm3 of concentrated nitric acid
(Aristar grade) using the microwave digestion system and sealed Teflon vessels.
The resulting
solution is transferred back to its sample tube and centrifuged at 4000 rpm for 10 minutes to separate any remaining solid from the supernatant solution which is diluted to 50 cm3 with deionised water for ICP analysis. Stage 5:
Any remaining solid is digested using the microwave equipment with 10 cm3 of
concentrated hydrochloric acid (Aristar grade) and 5 cm3 of HF. The final solution is transferred to a Teflon crucible and evaporated to dryness on a hotplate. The solid residue is redissolved in 10 cm3 of concentrated nitric acid and made up to 50 cm3 with deionised water for analysis. The interpretation of these dissolution stages in terms of coal components is as follows:Stage 1.
Elements present in pore fluids and soluble minerals.
Stage 2.
Carbonates (mainly calcite), exchangeable cations and monosulphides. 3
Stage 3.
Pyrite and carbonates (mainly dolomite and ankerite),
Stage 4.
Organic matter.
Stage 5.
Silicates.
It should be stressed again at this point that quantitative extraction of a component in any one stage is difficult to achieve and hence the necessity to adopt a mathematical approach to the interpretation of the analytical results. Coal samples In the present work only one coal has been analysed. This coal was obtained from Eggborough power station in the North of England, which at the time of sampling was burning local coals typical of the Yorkshire-Nottinghamshire coalfield. This coal was chosen for analysis in this work because it has been well characterised in earlier work. The whole coal mineralogy and geochemistry were described in Spears and Martinez-Tarazona (1993). Pulverised fuel size-fractions were analysed in Spears and Booth (2002) and Spears (2002) and trace element distributions determined indirectly by a statistical analysis of the data. In more recent work (Spears et al., 2007); trace element distributions have been obtained for the macerals in this coal by a direct method using laser ablation ICP-MS. In the collaborative research programme on element distributions (Davidson, 2000) one of the selected coals was from the UK. This coal, referred to as Gascoigne Wood, also comes from the YorkshireNottinghamshire coalfield and is similar in many respects to the coal under study here Analytical methods Solutions were analysed using both
ICP-AES and ICP-MS. The first method lacks the
sensitivity and accuracy for all elements and these were analysed by ICP-MS. The analysis method is indicated on Table 1. A more detailed discussion of the analyses and element specific problems are to be found in Davidson (2000). Samples were run in triplicate to check the reproducibility and analyses were summed and compared with whole coal analyses where available. Results and discussion. The analyses of the solutions resulting from the differential dissolution stages are shown in Table 1 and illustrated as percentages of the total amount extracted on Figure 1.
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Pyrite is known to be an important host for a number of trace elements in coals and Stage 3, the dilute nitric treatment, is designed to target pyrite. There is major extraction of both Fe and S in this stage, but there is also extraction in other stages. This is to be expected as Fe will be present in the clays and S in the organic matter. There is also the point made above that quantitative extraction of any component is difficult and this includes pyrite. Trace elements showing major extraction in Stage 3 include Cu, Zn, As, Mo, Cd and Tl. The agreement with published information for these coals is good. For example, in a recent investigation of the Parkgate coal in the same coalfield based on a comprehensive suite of trace elements (Spears and Tewalt, 2009), it was concluded that pyrite contains all of the Hg, As, Se, Tl and Pb and much of the Mo, Ni, Cd, and Sb. In the present work Hg and Se were not determined. Much of the Pb is returned in Stage 2 rather than Stage 3 possibly due to a sulphide other than pyrite. In Stage 1 we would expect to see the contribution of the porewaters and the water soluble components. Chloride was not analysed but in these coals a high concentration in Stage 1 would be anticipated. Sodium would also be expected to be high, but this is not so, possibly because of the effectiveness of the washery plant. In Stage 1 Ca is a major ion in solution, due in all probability to the dissolution of calcite or other carbonates, although a greater return of Ca would have been expected in Stage 2. Cobalt and Ni are linked with Ca with significant returns in Stage 1 and 2. Stage 4, in which organic matter is attacked, clearly does rather more than that based on the behaviour of Al and K. These elements are almost exclusively present in the clay minerals and it is apparent that the clay minerals have been destroyed with the nitric acid digestion in the microwave. Similar in behaviour are V, Cr, Rb, Sr, Y, Ba, Th and Sn and U to a lesser extent. In other work (Spears and Tewalt, 2009) elements closely linked with the clays are Rb, U, Cr, V, Y, Cu, Nb, Sn, and Th. (plus Cs, Li, Ga, Sc, Bi, Te, but not analysed in the present work). Over 95% of the Al is extracted in Stage 4 and 90% of the K. The latter is present in interlayer sites and some loss would be expected in the earlier treatments. Nevertheless, what has been achieved in Stage 4 is an efficient extraction of the clay minerals, that is the hydrolyzate fraction. The organic matter is also destroyed, as intended, and some of the V in particular will have this source. In Stage 5 there is dissolution of the remaining minerals. Most of the Ti, Zr and Nb are returned in stage and can be attributed to the resistate fraction with rutile and zircon likely hosts. A major difficulty with differential dissolution is that individual treatments may not be specific to one component, that is elements associated with a specific fraction in the coal will probably be returned in 5
more than one fraction. The problems of determining the forms of S in coals are well documented in the literature on the standard methods. The silicate fraction would be expected to release elements in stages, with the interlayer sites in the clay minerals the first to be attacked. The nitric acid stage, designed with the organic matter in mind, proved to be surprisingly effective at dissolving the clay minerals. Unfortunately this stage will also incorporate the organic matter and it is difficult to see how only the organic matter can be determined. Reversing stages 4 and 5 in an attempt to leave an organic residue has some merit, but it is difficult to envisage how an HF treatment would not strip some of the trace elements associated with the organic matter. Although failing to determine separately an organic association the final stage did provide additional information on the silicate fraction. The “new approach” to differential dissolution. A logical step in dealing with the behaviour of elements in the different dissolution stages is to consider the chemical fractions as comparable with analyses of a set of coal samples which have been physically fractionated. The latter is also comparable with a set of samples in which there is a natural variation in the proportions of the component fractions. In a correlation matrix the significant (>95%) correlations with Fe are S, Cu, Tl, Zn, Mo and As, similar to the qualitative conclusions above. The overall relationship between Fe and S corresponds to a FeS2 composition with the major return for both in Stage 3 as noted above. However, there are significant returns for both in other stages, reflecting the non-specificity of the dilute HNO3 stage. Based on the correlations with Fe it is possible to calculate some of the trace element concentrations in pyrite. These results are shown in Table 2 with previously determined results obtained using different methods. In general there is good agreement.
A notable feature of the dissolution stages is the high return of clay associated elements in the conc. HNO3 digestion. This is reflected in correlations involving Al., which are K (1.0), Cr (1.0), Rb (1.0), Sr (0.99), Th (0.99), Y (0.98), V (0.94), Mg (0.93), Ba (0.86) and U (0.79). As noted in the earlier discussion this suite of elements would be predicted for this fraction of the coal. The high values are noteworthy and as values fall sources other than the clay fraction could be contributing, although analytical error must also be considered. Furthermore, the laser ablation work (Spears et al., 2007) demonstrates that some of the V is also present in the organic matter and this fraction has not been isolated in the current dissolution work.
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The correlation matrix confirms the visual interpretation of Figure 1 with significant relationships between the three elements Ti, Zr and Nb. Rutile/anatase and zircon are probably the control. Showing a weaker association, although significant at the 95% level, with one or more of these elements are Sn and Sb. Tin would be expected in a resistate fraction but not Sb, for which sulphide minerals are the more likely host. Although U and Th are dominantly clay related there are indications that some U is present to a lesser extent in the resistate fraction, although the relationships fall below the 95% significance level. Calcium can be attributed to a small amount of carbonate and not the clay minerals based on the inverse relationships with the clay related elements. The only element to correlate at the 95% confidence level is Co. Nickel shows some association, as noted from the visual inspection of Figure 1, but the relationship falls below the 95% confidence level. Some phosphate mineral may be present based on weak links with Ba and Pb, but there is not a relationship with Ca, suggesting phosphate minerals are only present in minor amounts at the most. Conclusions. A conventional differential dissolution investigation of a coal to determine the distribution of elements between the main components has been extended using a statistical approach. This is seen as a logical step, putting the essentially visual approach to the interpretation of the dissolution analyses on a firmer footing. In addition a statistical treatment is desirable as the dissolution stages are not necessarily quantitative for specific fractions of the coal. Furthermore, different laboratories do not use exactly the same dissolution steps and a statistical treatment potentially overcomes this problem. Although it should be possible to successfully target porewaters, carbonates and sulphides, it remains a challenge to separate and distinguish between clay minerals and organic matter. Possibly there is some merit in reversing Stages 4 and 5, but it is difficult to envisage how differential dissolution of organic matter can be achieved without some attack of the clay minerals and vice versa. The answer to this problem could be a prior treatment in which organic matter is physically separated, possibly with heavy liquids, and analysed. There still remains the problem of separating silicates intimately associated with the organic matter and non-uniform organic composition. A potentially better approach would be to remove the organic matter after Stage 2 in a furnace at 395ºC and then repeat Stage 2 before proceeding to the pyrite dissolution in Stage 3. Organic matter can also be very effectively removed using low-temperature plasma ashing. Such equipment should be available in laboratories investigating the minerals in coals using X ray diffraction, which is essential for the study 7
of clay minerals. Both methods should be tested and possibly one or other included in a standardised dissolution procedure for which there is a need. Including the approach suggested for the organic matter it should be possible to determine the distribution of trace elements between the main components in a coal including the organic matter. It should be noted that trace element concentrations in the organic matter in the UK coals are relatively low compared with the mineral matter, although there is proportionally more organic matter, and that if distributions between organic components are required then laser ablation ICP-MS is probably the way forward. Acknowledgements. The significant support provided by Rio Tinto over a period of time for methods development is gratefully acknowledged. The encouragement given by Professor Chris Cross and Dr. Mike Whateley at that time was very much appreciated. However, the views expressed in this work are those of the author, who alone is responsible for any shortcomings.
REFERENCES Cavender, P.F. and Spears, D.A., 1995. Analysis of forms of sulphur within coal, and minor and trace element associations with pyrite by ICP analysis of extraction solutions. In: Pajares, J.A. and Tascon, J.M.D. (Editors), Coal Science, Coal Science and Technology, 24: pp. 1653-1656. Chenery, S., Querol, X. and Fernandez-Turiel, J.L., 1995. Quantitative determination of trace element affinities in coal and combustion wastes by Laser Ablation Microprobe-Inductively Coupled PlasmaMass Spectrometry. In: Pajares, J.A., Tascon, J.M.D. (Editors), Coal Science, Coal Science and Technology. Vol. 24, pp. 327-334. Davidson, R.M., 2000. Modes of occurrence of trace elements in coals: Results from an international collaborative programme. IEA Coal Research, 36pp. Huggins, F.E., Shah, N., Huffman, G.P., Kolker, A., Crowley, S., Palmer, C.A. and Finkelman,R.B. 2000. Mode of occurrence of chromium in four US coals. Fuel Processing Technology, 63, 79-92. Spears, D.A., 2002. Trace elements in some UK coals. Quantitative distribution within the coal. CD ROM publication of the Proceedings of the Eighteenth Annual International Pittsburgh Coal Conference, Newcastle, New South Wales, Australia. Spears, D.A. and Martinez-Tarazona, M.R., 1993. Geochemical and mineralogical characteristics of a power station feed-coal, Eggborough, England. Int. J. Coal Geol., 22: 1-20.
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Spears, D.A. and Booth, C.A., 2002. The composition of size-fractionated pulverised coal and the trace element associations. Fuel, 81, 683-690. Spears, D.A., Borrego, A.G., Cox, A. and Martinez-Tarazona, R.M., 2007. Use of laser ablation ICPMS to determine trace element distributions in coal with special reference to V, Ge and Al. International Journal of Coal Geology, 72, 165-176. Spears, D.A. and Tewalt, S.J., 2009. The geochemistry of environmentally important trace elements in UK coals, with special reference to the Parkgate coal in the Yorkshire – Nottinghamshire Coalfield, UK. International Journal of Coal Geology, 80, 157-166. White, R.N., Smith, J.V., Spears, D.A., Rivers, M.L. and Sutton, S.R., 1989. Analysis of iron sulphides from UK coal by synchrotron radiation X-ray fluorescence. Fuel, 68: 1480-1486.
Table 2. Concentrations (μg/g) of selected trace elements in pyrite A B C Cu 1090 756 315 Mo 95 110 110 As 750 1145 1030 Zn 790 450 20 Tl 28 Pb 308 310 320 Sb 10 20 Ni 410 310 Se 34 100 Column A calculated in this work. Column B indirect, geochemical calculation (Spears, 2002) Column C direct, synchrotron determination (White et al., 1989).
Figure 1
9
EGGBOROUGH COAL - MODIFIED METHOD
100%
HF / HCl / MW
80%
% extraction
HNO3 / MW
60% 2N HNO3 / HP
40% dilute HCl
20% WATER
0% 11 Na B
Mg 27 Si Al
P
S
39 Ca K
Ti
V
Cr
Mn Fe
59 60 Cu Zn Co Ni
75 85 Sr As Rb
89 90 93 95 111 118 121 Ba Y Zr Nb Mo Cd Sn Sb
205 Pb Tl
232 238 Th U
elements
10
Oviedo ICCS&T 2011. Extended Abstract-Oral
Solid-State NMR Study on Mineral Structure and Transformation Behaviors of Coal Ash Xiongchao Lin1, Keiko Ideta2, Jin Miyawaki2, Seong-Ho Yoon1, 2* and Isao Mochida3 1
Interdisciplinary Graduate School of Engineering Sciences, Kyushu University, Fukuoka 816-8580, Japan 2 Institutes for Materials Chemistry and Engineering, Kyushu University, Fukuoka 816-8580, Japan 3 Research and Education Center of Carbon Resource, Kyushu University, Fukuoka 816-8580, Japan *Corresponding author: Tel: +81-92-583-7959; Fax: +81-92-583-7897; E-mail: [email protected] Abstract
Original coal ashes contain various kinds of minerals, which coexist as crystal and non-crystal minerals at different temperatures. In this study, ash compounds were specifically treated as simple molten mixtures, and the continuous transformation and combination were investigated dependence temperature changes. Primary minerals and elements were analyzed by means of
27
Al-,
29
Si-solid state nuclear magnetic
resonance (NMR) for re-solidified molten ashes and slags at room temperature. Results demonstrated that, changes of minerals at low temperature region were mainly decomposition (clay minerals etc) and crystal transition (quartz etc). Amorphous minerals were found as a transition phase prior to the melting. Further, slags with amorphous morphology showed more dispersive coordination distributions than fused-ash due to the rapid quenching treatment. It was considered that crystal minerals continuously combining with cations (such as Ca2+, Mg2+) and eutectics with silica solid solution above fusing temperature caused the broadening of NMR peaks together with shifting of 29Si Qn structure peaks to lower magnetic field. Keywords: gasification, coal ash, solid state NMR, mineral
1. Introduction Slagging of coal ash is one of the most important troubles in the entrained type coal gasification. It is known that the fluidizing properties are strongly sensitive to compositional and temperature changes of molten coal ashes [1]. However, it is still big challenge to solve the problems of ash coating and smooth slagging out with or without adding fluxing agents. 1
Oviedo ICCS&T 2011. Extended Abstract-Oral Since coal ash always mainly contains aluminum and silicon oxides, an understanding of the transitional behaviors of alumina-silicate melts is most necessary for predicting the molten properties of slag in the practical gasification system [2]. However, almost liquid phases of coal ash show amorphous phase glasses and their mixed minerals, which makes them difficult to identify by general techniques (such as XRD). From above reason, we attempted to analyze the structural transitions of coal ashes using multi-nuclear solid state NMR for the quantitative understanding of their morphological changes under gasification range of temperature. 2. Experimental 2.1 Samples The representative ash denoting as D derived and slag from lignite were used as samples for this study. Low temperature ash (LTA) was prepared at 300 oC under air flow for 60 days. The other ashes were prepared under the heat treatment conditions as following: 600 oC, 2h; 815 oC, 2h; 1000 oC, 30min; 1200 oC, 30min; 1400 oC, 30min, and the molten ash was prepared at 1600 oC with 10 min through the re-solidification by quenching. 2.2 Sample characterization The coal ashes was analyzed by using XRF (X-ray fluorescence), XRD (X-ray diffraction) [3]. High-resolution
29
Si and
27
Al NMR spectra were obtained by using a JEOL NMR
spectrometer. Chemical shift scale in ppm was referenced to 1.0M AlCl3 aqueous solution and PDMS (±0.1 ppm). DQF-STMAS (double-quantum filter satellite transition magic angle spinning) was carried out after a bottom part of a 3.2 mm zirconia (ZrO2) sample rotor filled with Na2SO4 for magic angle adjustment. 3. Results and Discussion 3.1 Basic composition of coal and ashes D coal has an ash amount of 11.4 wt %, which is composed of 55.56 wt % of SiO2 and 28.12 wt % of Al2O3 as shown in Table 2. Compositions were divided into two parts as basic and acidic: b = %Fe2O3 + %CaO +%MgO + %Na2O + %K2O; a = %SiO2 + %Al2O3 + %TiO2.
2
Oviedo ICCS&T 2011. Extended Abstract-Oral Table 2 Coal and coal ash analysis (as equivalent oxide) Coal wt % Ash a Proximate SiO2 Moisture 4.41 Al2O3 Ash 2.81 Fe2O3 Volatiles 34.39 CaO Fixed carbon 58.39 K2O Ultimate TiO2 Carbon 75.26 SO3 Hydrogen 5.57 P2O5 Nitrogen 1.39 MgO Oxygen 17.66 Total Sulfur 0.12 b/a
wt % 55.56 28.12 11.52 1.38 1.18 0.97 0.57 0.36 NDb 99.66 0.19
Where, aND not detected b
Trace elements (e.g. Sr, Mn, V, Zr, Y, Zn, Cu, Rb) were ignored.
3.2 Solid state NMR analyses 3.3.1 29Si spectroscopy Fig.1 (I) shows the 29Si solid state NMR spectra of ash prepared at different temperatures, and the deconvolution of peak was carried out to fit the patterns using Gaussian equation that based on the resonances of standard materials measured before this experiments. The tetrahedral SiO4 units are identified according to their mutual connectivity as Qn, where n indicates the number of the bridging oxygen within a SiO4 tetrahedron [4] . Due to the low cation concentration in D ash, Q4 and Q3 are the primary coordination of silicon. The resonance of ash after 815 oC heat treatment shows a broad resonance and a sharp peak at -107.8 ppm (Fig.1(I) a). Broad peaks were derived through the destructive transformation of kaolinite; Sharp peak is attributed to the quartz at low temperature. -103.1 ppm, -98.8 ppm, -94.3 ppm and -88.7 ppm were ascribed to Q4(1Al) Q4(2Al) Q3(0Al) and Q3(1Al), respectively. Similar peaks are found in the ash after heat treat up to 1000 oC (Fig.1(I) b). More obvious change appears at 1600 oC, detected Q4 site at -110.2 ppm that is attributed to the amorphous SiO2 (Fig. 1(I) c). Simultaneously, the resonances at -85.6 ppm and -92.4 ppm are Q3(1Al) and Q3(0Al) respectively, begin to emerge at 1600 oC which is assigned as a poor crystalline mullite with SiO2 eutectic. Moreover, the main structure of Q4(1Al) at -103.4 ppm indicates the disorder or non-crystalline Si-O-Al minerals, which has a majority fraction at 1600 oC. 3
Oviedo ICCS&T 2011. Extended Abstract-Oral Rapid quenching was considered to be able to maintain the local structure of high temperature, thus slag shows a broad peak with a center at -112 ppm in Fig. 1(I) d, which reveals the amorphous silica of Q4(0Al) structure. The moderate peak indicates the coexistence of Qn species. However, we cannot conclude that the broad distribution of Qn structure was caused by rapid quenching or the higher temperature of gasification, because the combustion
ssb
15.7 1.5 16.1 4.7
ssb
10.3 5.3
58.4
65.6
ssb
58.7 87.0
100 50 0 Chemical shift (ppm)
ssb
65.3
150
-80 -100 -120 -140 Chemical shift (ppm)
ssb
(a)
(a)
4.7
(b)
34.1
58.5
(b)
ssb
Intensity(a.u.)
-107.6
(c)
-107.8
-88.2 -94.7 -99.4 -104.0 -88.7 -94.3 -98.8 -103.1
88.5
(d)
(d)
(c)
-60
55.3
-110.2
(II)
-110.5
Intensity (a.u.)
-85.0 -93.0 -98.0 -103.4
(I)
-86.1 -92.1 -98.4 -104.7
temperature was not exactly known.
-50
Fig.1. Single pulse 29Si NMR and 27Al NMR spectra (II)obtained with same contact time at (a) 815 oC, (b) 1000 oC, (c) 1600 oC and (d) slag as received. ssb denoted as spin side bond. 3.3.2 27Al spectroscopy The
27
Al NMR spectrum of low temperature ash contains three resonances as Fig. 1(II)
shown. 5-coordinated Al is formed at 815 oC, as seen in Fig. 1(II) a. The Al (V) was derived from the decomposing of kaolinite with the the 6-coordinate alumina to the 4-coordinate. In Fig. 1(II) b, two shoulder peaks were found at 65.6 and 10.3 ppm in the ash prepared at 1000 o
C, which were considered to come from the decomposing of 5-coordinate to both 6 and 4
coordinate. It may also reveal that the reactions and/or the mutual substitution of tetrahedral Al and Si by Ca2+, Mg2+, even Fe2+ cations starting at 1000 oC. While, due to the low containing of such cations, there is some isolating alumina generated at this temperature. The increase of Al (IV) at higher temperatures is attributed to the forming of the stable 4
Oviedo ICCS&T 2011. Extended Abstract-Oral frameworks of tetrahedral Al. Based on the identification of standard materials, mullite has both 6-coordinate and 4-coordinate structures, therefore, ashes at 1600 oC with lower 6-coordinate indicates the disorder crystal structure of mullite with silica eutectic (Fig. 1(II) c). Alpha-alumina transferred from γ-alumina was found with a AlIV resonance at about 17.1 and 16.1 ppm at 1600 oC, which is the independently solid part even at high temperature. Slag with rapid quenched has a broad resonance at a center of 55.6 ppm indicating the 4-coordinate alumina (Fig. 1(II) d). Weak peak at 15.7 ppm is the α-alumina and peak at 1.5 ppm is the 6-coordinate represent as the poor crystalline mullite. 3.2.3DQF- STMAS analyses Fig. 2 shows the
27
Al DQF-STMAS spectroscopy. Fig. 2a is the 2-dimensional
spectroscopy of D ash prepared at 815 oC. Four chemical shifts were detected. δi is the 6-coordinate, δii is the 5-coordinate and δiii, δiv are 4-coordinate. δi indicates the existing of AlO6 structure from the incomplete decomposition of kaolinite, meanwhile δii also reveals the transition from 6- to 4-coordinate via the 5-coordinate. We observed two AlO4 sites with both chemical shift and quadrupolar distributions due to its amorphous state, which indicates two types of Al structures. Fig. 2b demonstrates the structure of D ash prepared at 1000 oC, δi and δii are two different 6-coordinates, the separation of them means the isolation of Al structure, in other words, some α-alumina were generated. Further, the 5-coordinate disappears and 4-coordinate (δiii) gradually inclines to diversification; more obvious changes can be seen in Fig. 2c with D ash after 1600 oC heat treatment. 6-coordination completely divide to two sites, δi represents the 6-coordinate combining in alumina-silicate structure (e.g. mullite) and δii probably is the isolating alumina; Owing to the high temperature, the structures were further destroyed to various 4-coordinate of Al, at the same time the resonance noises increased as seen in δiii. Fig. 2d is the spectroscopy of quenched slag, which shows only tetrahedral Al structure (δi) and similar with ash made at 1600 oC, the quadrupolar distributions are dispersive. The disappearing of 6-coordinate may be caused by the rapid quenching that led to the 6-coordinate totally transfer to amorphous 4-coordinate. The distinction between STMAS and single pulse of slag may be contributed to the higher precision of STMAS method. 5
Oviedo ICCS&T 2011. Extended Abstract-Oral
δi
(b)
δi F1 dimension (ppm)
F1 dimension (ppm)
(a) δii
δiii
δii
δiii
δiv F2 dimension (ppm)
(d)
δi
F1 dimension (ppm)
F1 dimension (ppm)
(c)
F2 dimension (ppm)
δii
δi
δiii F2 dimension (ppm)
F2 dimension (ppm)
Fig.2. Two-dimensional DQF-STMAS NMR analysis of D ash: (a) 815 oC, δi (x=12.26, y=13.99), δii (x=42.37, y=45.50), δiii (x=63.76, y=65.55) and δiv (x=71.37, y=74.48). (b)1000 o
C, δi(x=11.57, y=13.11); δii (x=18.66, y=23.28) and δiii(x=63.79, y=65.47). (c)1600oC,
δi(x=10.05;y=10.29), δii (x=20.04, y=22.25), δiii (x=50-64;y=50-78) and (d) slag as received, δi (x=50-64;y=50-78). δ denoted as site of chemical shift. 4. Conclusions Silica and alumina are important components that play key roles as a framework-former and solid solution at different temperatures in high rank coal derived coal ash. Liquid phase was mainly formed by the mullite and silica eutectics. Fine particles, such as poor crystal mullite, cristobalite and alumina, were found as solid phase. They are considered to be derived from the kaolinite and dispersed in the system. Hereby, both Si and Al structures at high temperature show broad resonance distributions of NMR spectra.
6
Oviedo ICCS&T 2011. Extended Abstract-Oral Acknowledgement The authors are grateful to the financial support from new energy development organization (NEDO) and globe center of excellence (GCOE). References [1] Hurst, H.J., Novak, F. Patterson, J.H. Viscosity measurements and empirical predictions for fluxed Australian bituminous coal ashes, Fuel 1999; 78: 1831-1840. [2] Vargas, S., Frandsen, F.J. Dam-Johansen, K., Rheological properties of high-temperature melts of coal ashes and other silicates. Progress in Energy and Combustion Science 2001; 27: 237-429. [3] O. Font, Moreno, X. Querol, X-ray powder differaction-based method for the determination of the glass content and mineralogy of coal (co)-combustion fly ashes. Fuel 2010; 89: 2971-2976. [4] P. Pena, J.M. Rivas Mercury, A.H.de Aza, X. Turrillas, Solid-state
27
Al and 29Si NMR
characterization of hydrates formed in calcium-aluminate-silica fume mixtures, Journal of solid state chemistry 2008; 181: 1744-1752.
7
Characterisation of Coal and Biomass based on Kinetic Parameter Distributions for Pyrolysis N. Sonoyama1, J. Hayashi2 1
Coal & Environment Research Laboratory, Coal Business Office, Petroleum & Coal Marketing Department, Idemitsu Kosan Co., Ltd., 3-1 Nakasode, Sodegaura, Chiba, 299-0267, Japan. [email protected] 2 Division of Advanced Device Material, Institute for Materials Chemistry and Engineering, Kyushu University, 6-1, Kasuga Koen, Kasuga, Fukuoka 816-8580, Japan. Abstract We conducted the pyrolysis of various biomass samples by thermogravimery and obtained kinetic parameters that were a frequency factor and an activation energy. The kinetic parameters were derived from the Miura method based on the distributed activation energy model. Coal and biomass were compared based on the kinetic parameters. In addition, we validated the prediction of thermogravimetric curves on the basis of the kinetic parameters in the case of rapid pyrolysis. The rapid pyrolysis of biomass was conducted by using a wire-mesh reactor. Unlike coal, the kinetic parameters of biomass drastically changed with progression of pyrolysis. The activation energies of sugarcane bagasse were less than 200 kJ/mol during the initial step in pyrolysis. Some biomass samples indicated the decrease of kinetic parameters during pyrolysis, although it has been reported that the kinetic parameters of coal gradually increased during pyrolysis. The kinetic parameters of all biomass samples increased in the final stage, a charification process. We estimated that the low activation energy in the initial step was caused by hydration and volatilisation of lighter components. The decrease of activation energies occurred during pyrolysis was related to the softening and/or melting of a solid phase. The drastic change of the kinetic parameters was affected by the structural change of a solid phase. Biomass tended to have high frequency factors when the activation energies of coal and biomass were the same. The thermogravimetric curves based on the kinetic parameters of some biomass samples under rapid pyrolysis were corresponding to our experimental results obtained with the wire-mesh reactor. Analysing the change of kinetic parameters in detail provided the information on the behaviour of volatilisation, the change of solid state, the extent of char structural development, and so on. Kinetic parameter distributions of pyrolysis were expected to be useful for the characterisation of coal and biomass for pyrolysis as well as
1
combustion.
1.
Introduction World energy consumption increases yearly. We need to meet the demand of the
increased energy and also cope with environmental problems such as SO2 and NOx emissions and global climate change. The global climate change requires to decrease fossil fuel consumption and to increase the use of renewable energy resources. Therefore, the use of the renewable energy resources of great variety proceeded for energy generation. The operating condition of an existing boiler is altered to match the combustion characteristics of the renewable energy fuels and a new boiler is designed for the fuels. Coal and biomass co-combustion technology is also an effective one. The combustion characteristics of biomass are different from those of coal. Therefore, the behaviour of combustion and ash deposition in a boiler is different between biomass and coal. Many studies for structure and analysis of coal and biomass have conducted in the past. However, there has been little work to establish a method of the comprehensive evaluation of widely varying fuels. Pyrolysis is essential for cracking, combustion and gasification because of occurring first in decomposition by heating. Therefore, the well understandings of pyrolysis help us develop new processes and resolve the mechanisms of combustion and gasification. Thermogravimetric analysis is widely used in physical chemistry, material research, and thermal analysis. The thermogravimetric analysis provides the findings of decomposition behaviour in the wide range of temperature. The pyrolysis of solid fuels is complex due to including not only volatilisation but also secondary reaction such as particle-volatile interaction and the structural change of solid softening and melting. Excluding such effects is difficult for thermogravimetry. We experimentally investigated the effects of a heating rate and a sample height. In the present study, kinetics for the pyrolysis of solid carbonaceous materials was investigated in detail by kinetic parameter distributions based on a distributed activation energy model.
2.
Experimental Xylan that was oat-spelt xylan washed by ion-exchanged water after stirred in a 0.1 N
hydrochloric acid solution over 15 hours at ambient temperatures, cellulose, lignin that
2
was hydrolytic lignin washed by ion-exchanged water after a 5 N hydrochloric acid solution over 15 hours at ambient temperatures, sugarcane trash, sugarcane bagasse, pine sawdust, cedar sawdust, coffee residue, and Loy Yang coal were used. Xylan and lignin, cellulose, and sugarcane trash, sugarcane bagasse, pine sawdust, cedar sawdust, coffee residue, and Loy Yang coal were sieved in the range of 125 to 250 µm, 74 to 105µm, and 125 to 210µm, respectively. Pyrolysis of the samples was performed in thermogravimetry (TG) and a wire-mesh reactor (WMR). Thermogrvimetric data was analysed on the basis of a ditributed activation energy model propsed by Miura [1].
3.
Results and Discussion The mass loss profiles of the samples during pyrolysis in TG and WMR was
compared. Coffee residue and cellulose was affected by a packed bed height. Accordingly, we considered that the effect of a packed bed height was caused by secondary interaction such as char forming by particle-volatile interaction. It was confirmed that pine sawdust, sugarcane bagasse, and Loy Yang coal had little effect of a packed bed height in TG. Figure 1 shows the distributions of activation energy with conversion during the pyrolysis of biomasses. An activation energy distribution depended on the biomasses. Each activation energy distribution reflected the effects of not only compositions in a sample but also the change of solid structure.
Activation Energy [ kJ/mol ]
400 350 300 250 200 150 100 50 0 0.0
Sugarcane trash Sugarcane bagasse Pine sawdust Cedar sawdust Coffee residue
0.5
1.0
Conversion [ - ]
Figure 1 Activation energy distribution with conversion during pyrolysis.
3
Miura and Maki [2] were reported that the distribution of kinetic parameters differed depending on coal rank. Kinetic parameters that are frequency factor and activation energy for pyrolysis of coal and biomass were shown in Fig. 2. The kinetic parameter distributions of coals were different from those of biomasses. We showed the characterisation of pyrolysis for biomass by kinetic parameter distribution and predicted combustion characteristics such as char-burn out and a temperature of starting emission of volatiles.
25
Frequency Factor [ 1/s ]
10
20
10
Coal(C<75%) Coal(75%85%) Sugarcane bagasse Cedar sawdust Coffee residue
15
10
10
10
100
150
200
250
300
350
Activation Energy [ kJ/mol ]
Figure 2 Comparison of distributions of kinetic parameters for coal and biomass. The coal data were quoted from Miura and Maki[2].
Calculation for each conversion during the pyrolysis of pine sawdust was conducted on the basis of the distributed activation energy model. In the case of fast pyrolysis, a conversion profile derived from the calculation was good agreement with that of WMR. A mass loss profile for fast pyrolysis was predicted from that for TG.
4.
Conclusions The characterisation of biomass and coal was investigated by Miura’s distributed
activation energy model. We showed pyrolysis behaviour by the distribution of kinetic parameters that were frequency factor and activation energy, and considered that the evaluation of combustion characteristics of coal and biomass was allowed from the characterisation of those by pyrolysis.
4
Heating rate 1 K/s 1000 K/s
1.4 Exp. Calc.
Conversion [ - ]
1.2 1.0 0.8 0.6 0.4 0.2 0.0 400
500
600
700
800
900
Temperature [ K ]
Figure 3 Prediction of a conversion of pine sawdust during slow and fast pyrolysis including a cooling process. Experimental data were collected by WMR.
References [1] Miura K, A new and simple method to estimate f(E) and k0(E) in the distributed activation energy model from three sets of experimental data, Energy & Fuels 9 (1995), pp. 302-307. [2] Miura K, Maki T, A simple method for estimating f(E) and k0(E) in the distributed activation energy model, Energy & Fuels 12 (1998), pp. 864-869.
5
Oviedo ICCS&T 2011. Extended Abstract
Effect of the pyrolysis conditions on the microstructure of anthracene oil-based cokes P. Alvarez, N. Díez, R. Santamaría, C. Blanco, R. Menéndez and M. Granda. Instituto Nacional del Carbón, CSIC. P.O. Box 73, 33080-Oviedo, Spain [email protected] Abstract Anthracene oil is the heaviest fraction of coal tar distillation and by oxidative thermal condensation and subsequent thermal treatment this product can be transformed into a highly useful pitch-like material. The total absence of solid particles (e.g., ashes and primary quinoline insolubles) and the high aromaticity of its components make anthracene oil-based derivatives suitable for preparing cokes with a highly oriented microstructure. Two anthracene oil-based derivatives (an oil fraction and a pitch with a softening point of 110 °C) were pyrolyzed, under different nitrogen pressures (0-5 bar), in the temperature range of 420-460 °C with residence times of 5-21 h. The resultant green cokes were then carbonized up to 1000 °C. A comparison of the results obtained with both samples shows that the oil sample needs to be pyrolyzed under a higher nitrogen pressure and for longer residence time than the pitch to produce green cokes with a similar microstructure. This is due to the lower degree of polymerization experienced by the oil fraction with respect to the anthracene oil-based pitch. In both cases longer residence times at lower temperatures led to cokes with highly oriented and crystalline ordered microstructures, similar to those observed in needle cokes.
1. Introduction Cokes with a preferential orientation in their microstructure (i.e. needle-like cokes) are in great demand because of their superb properties (e.g., thermal and electrical conductivities, stiffness, etc.). The preparation of cokes with an oriented microstructure requires the use of precursors with specific characteristics, which must be processed under specific conditions to achieve the required microcrystallite orientation. Anthracene oil is a coal conversion product that due to its composition (i.e., it is highly aromatic) and the total absence of solid particles (i.e., primary insoluble quinoline, metals, etc.) could be a suitable precursor for the preparation of this type of cokes. This is a coal-tar fraction that distils between 280-400 °C and its transformation into a carbon precursor necessarily requires previous treatment to promote the polymerization of its
Submit before 31 May 2011 to [email protected]
1
Oviedo ICCS&T 2011. Extended Abstract
components and hence its carbon yield. This is achieved by oxidative thermal treatment at moderate temperatures and subsequent thermal treatment under inert atmosphere [1, 2]. By this procedure a carbon precursor (anthracene oil-based pitch) with the desired characteristics (e.g., softening point, beta-resin content, etc.) can be obtained. In the present study, two anthracene oil derivatives (an oil fraction and a pitch with softening point of 110 °C) were pyrolyzed, under different nitrogen pressures (0-5 bar), in the temperature range of 420-460 °C for residence times between 5-20 h. Green cokes were observed by polarized light microscopy and the different microstructures present in the cokes quantified by this technique. A relationship between the degree of microstructural orientation and the composition and experimental conditions used for obtaining the cokes was established.
2. Experimental section Two anthracene oil derivatives, AO-A and AO-B, were used as precursors for the preparation of the cokes. AO-A was obtained by thermal oxidative treatment of an industrial anthracene oil. AO-B was prepared by thermal treatment of AO-A at temperatures near to 400 °C [1]. About 300 g of AO-A, or 200 g of AO-B sieved and ground to < 200 μm, were placed in an aluminium crucible (130 mm in height and 58 i.d.) and vertically positioned in a stainless steel autoclave and then carbonized following a two/three consecutive step procedure. In all the cases the heating rate was ~10 °C min-1. AO-A was initially carbonized under a nitrogen pressure of 5 bar (first carbonization step). This pressure was subsequently reduced to 1 bar (second carbonization step) which was kept constant until the sample had cooled down to room temperature. In the case of AO-B both carbonization steps were carried out under a nitrogen flow of 100 mL min-1. The resultant cokes were characterized in terms of polarized-light optical microscopy and quantified by using the Optical Texture Index (OTI) value according to the formula: OTI = Σ fi x (OTI)i where fi represents the fraction of the anisotropic unit in the optical texture and (OTI)i the assigned factor for each anisotropic unit in the optical texture according to values reported elsewhere [3].
3. Results and Discussion Green cokes with different optical textures and high degrees of microstructural Submit before 31 May 2011 to [email protected]
2
Oviedo ICCS&T 2011. Extended Abstract
orientation were obtained from the two anthracene oil derivatives AO-A and AO-B. As a consequence of the different compositions and hence different pyrolysis behaviours of the derivatives, even when they come from the same feedstock, different experimental conditions were required to obtain a high degree of microstructural orientation in the green cokes. The chemical composition and molecular structure of the components of the derivatives are clearly related to the processing applied to the parent anthracene oil (Table 1).
Table 1. Elemental composition and main characteristics of the anthracene oil derivatives Elemental Analysis (wt.%) Sample
C/H
IAr
SP
TI
NMPI
CY
1.5
1.41
6.08
oil
1
0
9
0.6
1.73
7.79
110
23
3
51
C
H
N
S
O
AO-A
91.6
5.4
1.0
0.5
AO-B
93.3
4.5
1.1
0.5
C/H, carbon/hydrogen atomic ratio
IAr, Aromaticity index determined by 1H NMR SP, Mettler softening point (°C) TI, toluene insoluble content (wt.%) NMPI, N-methyl-2- pyrrolidinone insoluble content (wt.%) CY, Alcan carbon yield (wt.%)
During the thermal oxidative treatment applied to produce AO-A, the oxygen causes an initial polymerization/condensation of the anthracene oil components, but at the same time, some oxygen is incorporated into the components (1.5 wt.% for AO-A). The subsequent thermal treatment applied to AO-A to produce AO-B causes the polymerization/condensation of the components, thereby producing a significant reduction in the oxygen content (0.6 wt.%). Additionally, the hydrogen content is reduced from 5.4 to 4.5 wt.% and the solubility parameters, determined as TI / NMPI, increase from 1 / 0 to 23 / 3 wt.% for AO-A and AO-B, respectively. The Har/Hal ratio of the samples suggests that the polymerization/condensation reactions involved in the thermal treatment occur with the consumption of aliphatic hydrogen. The transformation experienced by AO-B during the thermal treatment leads to molecular moieties able to produce a solid sample at room temperature (softening point of 110 °C), with a carbon yield much higher than that of AO-A (9 and 51 wt.% for AO-A and AO-B, respectively). This is clearly evident from an analysis of the thermogravimetric curves (Figure 1), in which the carbon residue at 1000 °C is also observed to be higher in AO-B than in AO-A (8 and 38 wt.% for AO-A and AO-B, respectively). Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
This different composition and pyrolysis behaviour determines the experimental conditions used for the preparation of the green cokes. 100
Weight (%)
80 60
c 40
a 20
b 0 0
200
400
600
800
1000
Temperature (°C)
Figure 1. Thermogravimetric curves of the (a) parent anthracene oil (AO) and anthracene oil derivatives (b) AO-A and (c) AO-B
Preparation of green cokes from AO-A The pyrolysis behaviour of AO-A, shows a significant removal of volatiles in a short range of temperature, the carbonaceous residue being < 10 wt.% at temperatures above 400 °C (Figure 1). This indicates the convenience of using pressure during carbonization in order to mitigate the elimination of gases and to increase the yield of the process.
a
b
c
A 50 μm
Figure 2. Polarized light micrographs of the pyrolysis products obtained from AO-A at (a) 460 °C / 3 h / 5 bar, (b) 460 °C / 5 h / 5 bar and from AO-B at (c) 430 °C / 4 h
The carbonization of AO-A at 460 °C / 3 h /5 bar gives rise to carbonized products in which some incipient mesophase, in the form of microspheres, can be observed (Figure 2a, position A). A slight increase in the carbonization temperature and/or residence time (e.g., 460 °C / 5 h /5 bar) results in a non-plastic carbonized product (Figure 2b) whose
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Oviedo ICCS&T 2011. Extended Abstract
microstructural orientation has already been defined. This suggests that once initiated the formation of mesophase, rapidly increases leading to the formation of a non-plastic material (e.g., green coke) with a poor microstructural orientation. This microstructural orientation can be improved by prolonging the period during which the pitch passes through a plastic and fluid stage (i.e. mesophase). In order to achieve this, AO-A was carbonized in two consecutive steps. The first stage, carried out under a nitrogen pressure of 5 bar, led to the formation of large macromolecules and/or incipient mesophase, especially in the form of small microspheres, whereas the second stage, under milder conditions (lower temperature and pressure), favoured the growth and orientation of the mesophase. Moreover, the use of pressure during carbonization possibly favours the formation of a moderate viscous medium. This medium is known to be essential for development of bulk mesophase of large isochromatic area [4, 5]. Bearing this in mind, a series of experiments were conducted according to the experimental conditions given in Table 2.
Table 2. Experimental conditions, yield of the process and optical texture index of the green cokes prepared from AO-A and AO-B Optical Texture (vol.%) Sample
Treatment
PY 1
2
3
4
5
6
7
8
9
Iso
OTI
AO-A1
460 °C / 3 h / 5 bar + 450 °C / 5 h / 1 bar
26
3
1
3
0
13
19
4
6
50
1
25.05
AO-A2
450 °C / 8 h / 5 bar + 450 °C / 12 h / 1 bar
32
1
0
6
0
7
25
4
4
53
0
26.31
AO-A3
450 °C / 3 h / 5 bar + 450 °C / 17 h / 1 bar
30
1
0
3
0
3
15
5
6
67
0
26.97
AO-A4
450 °C / 3 h / 5 bar + 450 °C / 6 h / 1 bar + 450 °C / 11 h
28
1
0
1
0
12
13
3
14
55
1
25.89
AO-B1
430 °C / 4 h + 420 °C / 9 h
67
0
4
2
0
17
12
5
0
32
28
17.21
AO-B2
440 °C / 4 h + 430 °C / 4 h
55
0
2
8
0
21
24
8
7
30
0
22.98
AO-B3
440 °C / 3 h + 430 °C / 9 h
65
1
4
3
0
20
18
0
6
43
7
23.74
The green cokes obtained were characterized in terms of optical texture (Figure 3) and their main microstructural features quantified (OTI) according to the experimental procedure previously described (Experimental section)[3]. The OTI values are known to be good indicators of the microstructural orientation of the cokes and they correlate directly with the thermal expansion coefficients of the resultant cokes [6]. Under the experimental conditions studied, the green cokes obtained give optical texture index values (OTI), varying from 25.05 and 26.97 (Table 3), well inside the range of those reported for other needle-cokes [3]. It should also be mentioned that prolonged residence times in the second step of carbonization (AO-A3, OTI of 26.97) give rise to a
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Oviedo ICCS&T 2011. Extended Abstract
green cokes with higher OTI values than when the prolonged residence time is applied in the first carbonization step (AO-A2, OT of 26.31). Additionally, total depressurization (AO-A4, OTI of 25.89) during the second carbonization step leads to a green coke with a lower OTI value than when 1 bar of pressure is applied (AO-A3, OTI of 26.97). This could be related with the fact that percolation of the gas bubbles, that deforms the planar liquid-crystal components and arranges them into a uniaxial geometry characteristic of a needle-coke precursor, is required just at the point of solidification [7]. Therefore, a premature evolution of gases or a late evolution was found to be detrimental for the formation of needle coke [8]. a
b
c
d
500 μm
50 μm
Figure 3. Polarized light micrographs of the green cokes (a) AO-A1, (b) AO-A2, (c) AO-A3 and (d) AO-A4 obtained from AO-A. Top, montage photograph; bottom, micrograph of a representative region.
Preparation of green cokes from AO-B The thermogravimetric analysis of AO-B, unlike that of AO-A, indicates that this sample is able to generate coke with a relatively high yield at atmospheric pressure (Table 1 and Figure 1). This is probably because the preparation of AO-B includes an additional processing step (thermal treatment). Moreover, preliminary studies on the pyrolysis behaviour of this sample indicated that the initiation of mesophase is achieved
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Oviedo ICCS&T 2011. Extended Abstract
at temperatures in the range of 430 °C and at a residence time of 4 h (Figure 2c). Taking this into account, a two consecutive carbonization step procedure was designed to prepare green cokes from AO-B (Table 2). As can be observed (Figure 4), temperatures between 420-440 and residence times shorter than 13 h are enough to form anisotropic green cokes with a process yield of 55-67 wt.% and OTIs of 17.21 to 23.74 (Table 2). a
b
c
500 μm
50 μm
Figure 4. Polarized light micrographs of the green cokes (a) AO-B1, (b) AO-B2 and (c) AO-B3. Top, montage photograph; bottom, micrograph of a representative region.
If we compare the pyrolysis conditions for both samples that led to the formation of highly oriented green cokes, it can be seen that AO-B requires less severe conditions (lower pressure, temperature and time of reaction) to form the cokes. However, despite the more rigorous experimental conditions required to transform AO-A into an anisotropic green coke, this sample is able to attain a higher degree of microstructural orientation than AO-B, under the experimental conditions studied (higher values of OTI, Table 2).
4. Conclusions This work has demonstrated that it is possible to transform two anthracene oil derivatives (AO-A and AO-B) into highly oriented green coke by pyrolysis in nitrogen.
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Oviedo ICCS&T 2011. Extended Abstract
The reaction conditions required for this were 450-460ºC, 8-20 h and 1-5 bar for AO-A. The use of pressure was required to avoid a massive release of volatiles (up to the 90 wt. %) that would considerably decrease the reaction yield. In contrast, AO-B required a lower temperature and pressure to form highly oriented green coke (420-440 ºC, 8-13h), probably as a consequence of the extra thermal treatment to which this pitch was subjected during its preparation. The OTI values of the cokes in both cases were found to be between 25.05 and 26.97 for AO-A and between 17.21 and 23.74 for AO-B, over the entire range of OTI values obtained for other coal tar or petroleum derivatives. The lower values obtained for AO-B could be due to the low viscosity of the reaction media during pyrolysis, which is a consequence of the different compositions of the samples.
Acknowledgement. The research leading to these results has received funding for the European Union’s Research Fund for Coal and Steel research programme under Grant Agreement number RFCR-CT-2009-00004. Dr. Alvarez also gratefully acknowledges help received from the Spanish Ministry of Science and Education in the form of a Ramon y Cajal contract.
References [1] Álvarez P, Granda M, Sutil J, Santamaría R, Blanco C, Menéndez R, Fernández JJ, Viña JA. Preparation of Low Toxicity Pitches by Thermal Oxidative Condensation of Anthracene Oil. Environmental Science & Technology 2009;43:8126-32. [2] Álvarez P, Granda M, Sutil J, Menendez R, Fernández JJ, Viña JA, Morgan TJ, Millan M, Herod AA, Kandiyoti R. Characterization and Pyrolysis Behaviour of Novel Anthracene Oil Derivatives. Energy & Fuels 2008;22:4077-86. [3] Asao O, Zhanfen Q, Marsh H. Structural study of cokes using optical microscopy and X-ray diffraction. Fuel 1983;62:274-78. [4] Brooks JD, Taylor GH. In: Walker Jr. PL, Thrower PA, editors. Cemistry and Physics of Carbon Vol.4, New York: Marcel Dekker; 1968, p.243. [5] Hoover DS, Davis A, Perrota AJ, Spackman W. Extended Abstracts 14th biennial conference on Carbon. American Carbon Society; 1979, p. 393. [6] Mochida I, Korai Y, Oyama T. Semi-quantitative correlation between optical anisotropy and CTE of needle-like coke grains. Carbon 1987;25:273-8. [7] Mochida I, Maeda K, Takeshita K. Structure of anisotropic spheres obtained in the Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
course of needle coke formation. Carbon 1977;15:17-23. [8] White JL. In: Deviney ML, O´Grady TM, editors. Petroleum Derived Carbons (ACS Symposium Series No.21). American Chemical Society; 1976, p.287.
Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011.
Optical and Scanning Electron Microscopy of Coke: Microstructure & Minerals S Gupta1, E Lester2, M Ismail2 and G O'Brien3 1
SMART@UNSW, School of Materials Science & Technology, The University of New South Wales NSW, Australia Phone: 61-2-9385-4433 E mail: [email protected]
2
Department of Chemical and Environmental Engineering, The University of Nottingham, Nottingham, NG7 2RD 3
CSIRO Earth Science and Resource Engineering, QCAT, QLD, Australia
Abstract Characterization of coke microstructure and mineralogy is critical to reliably predict the effect of coal properties on coke quality. This paper presents applications of some of the advanced microscopic approaches to characterize complex heterogeneous structure of coke matter. An analytical study was carried out to develop a suitable approach for simultaneous examination of microstructure and mineralogy of coke lumps using optical and scanning electron microscopy. Optical microscopy of the tested cokes showed that the total pore phase area was primarily dictated by larger pores, and was found to be higher for lower strength cokes. Microscopic porosity of cokes was invariably less but consistent with the trend indicated by bulk porosity measurements. The microscopic data was processed using an erosion parameter to estimate coke wall thickness, however the estimated thickness was found to be dependent on the frequency of the application of erosion parameter. SEM porosity values of cokes were consistently lower than the total porosity of cokes based on optical microscopic estimates but indicated a similar trend except one coke. Small pore (0.75 mm) phase area rather than total pore area indicated an inverse association with coke strength parameters. The study showed that total porosity of cokes could not reliably distinguish the differences of coke strength. The SEM analysis is shown to quantify coke mineral phases as well as their grain size but requires optimization of mineral database. High reactivity cokes were found to be accompanied by high mineral phase area. The SEM approach is promising option due to simultaneous characterization of micro-structure and mineralogy but requires validation using wider range of samples.
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Oviedo ICCS&T 2011.
1.
Introduction
Coke is one of the most expensive and critical raw materials for steelmaking via a blast furnace route. Coke quality is characterized by a set of physical and chemical attribute of cokes which include lump size, shape, mechanical strength, high temperature strength as well as chemical reactivity. Coke mechanical strength and resistance to chemical and thermal environments are considered to be the most important properties for stable and efficient blast furnace operation. Cold strength is generally estimated by a range of mechanical tests using a several versions of drum tumbling tests [1]. The optimum range of these indices is wide and depends upon the operating practice of individual blast furnace operator. Cold mechanical strength is the ability of coke to withstand breakage at room temperature and reflects coke behavior in the upper part as well as outside blast furnace. The NSC test is commonly used for assessing coke strength after reaction with carbon dioxide. Microscopic imaging techniques are commonly used for characterizing coal properties [2-9]. However, the imaging techniques have been less frequently used for distinguishing microstructure of different type of cokes [10-18]. Characterization of coke pore structure is important as the total porosity and pore size distribution will dictate the coke behavior under load as well as during gasification at high temperatures. Therefore, effect of micro-structure on coke strength has been extensively studied [1925]. Bulk porosity of any material including that of coke is well known to affect the mechanical strength [19]. The contiguity of the pores will influence the accessibility of oxidising gases while large pores may cause the defects in the coke structure. The ratio between the sizes of the pores and pore walls of coke was linked to coke strength [12]. Generally, porosity increase in bench-scale tests occurs throughout the particle [14] while in a blast furnace porosity increases preferentially at the edges [24]. Effect of pore structure on coke strength is shown to be greater than other modification of coke characteristics during gasification [25]. Conventional petrography distinguishes coal grains into vitrinite, liptinite and inertinite macerals. The relative amount and reflectance of the vitrinite maceral of coals is often related to coke strength but may not be applied universally due to the differences in the identification standards as well as in the reactivity of the inertinites. The inertinities present in Australian coals are often believed to be more reactive [26] such that average
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Oviedo ICCS&T 2011.
fused semifusinite matter in cokes made from Australian coals was reported to be about 84% [13]. Size distribution of the unfused inertinite material in the coke also influence the coke strength [27]. Preferential removal of microtexture is dependent on parent coal rank such that a particular microtexture in some pore walls could cause a greater reduction in coke strength. However, coke microtexture and pore structure are interrelated to some extent such that the nature of fusible and unfusible coal maceral would determine the characteristics large and small pores of coke and hence its strength. Coke minerals are known to influence coke behavior in a blast furnace and hence high temperature strength including CSR [28]. Therefore, coke minerals would also influence pore structure of coke during gasification. Recently, an automated coal reflectogram parameter (FMR) was developed and empirically related to coke strength and reactivity [6,9]. With better understanding of the association between coke-microstructure parameters and mineralogy to coke strength, it would be possible to improve the reliability of relating coke strength with coal imaging parameters. However, the main focus of this study was to establish a suitable methodology to characterize micro-structure of large coke lumps using optical and SEM techniques using the same specimen. SEM Approach was used to characterize mineralogy of cokes.
2.
Experimental section
Four coals were selected based on a current ACARP research project [15]. Coal A and Coal B are low vitrinite coals while Coal C and Coal D are high vitrinite coals. Approximately 20 mm in diameter of coke pieces were examined using oil immersion and SEM facilities at Nottingham University. Sample A and B are low strength cokes (CSR < 60) while other C and D are high strength (CSR > 60) cokes. The same specimen blocks were used to visualize under oil immersion microscope in order to obtain isotropic/anisotropic domain information. Coke images of 15 x 15 mosaics were acquired using a 20x objective which has a 440 micron size scale. The images were post processed using a image processing software. SEM imaging of coke lumps was carried out using the same specimen.
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Oviedo ICCS&T 2011.
3.
Results and Discussion
Coke mosaics-oil immersion Figure 1a illustrates typical image of coke based on oil immersion microscopy. The images were processed to get pore size information. Figure 1b compares the distribution of pore size among four cokes and shows that coke C and coke D contained the largest size pores compared to other cokes.
(a)
(b)
Fig. 1 (a) Optical mosaics of coke A and (b) comparison of % phase area of pores. Figure 2a compares the total porosity of four cokes. Porosity of cokes (C and D) made of high vitrinite coals is more than 40%. These measurements were consistent with the bulk porosity measurements such that high strength cokes displayed high porosity. In all the tested cokes, more than 60% of total pore phase area is contributed by the pores passing through 745 micron circular sieve (Figure 2b).
Fig. 2 (a) Comparison of total pore area of cokes; (b) phase area of 745 micron pores. An erosion function was used to characterize wall thickness information. Figure 3a compares the amount of char walls that have been covered by 3 iterations. The thicker the initial coke walls, the smaller the amount that will be lost post erosion. The higher the number, the more complete the covering. Coke D has the thinnest char walls and consistently gives the highest values for the percentage of material covered. By
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Oviedo ICCS&T 2011.
increasing the pore removal severity, the percentage covered values decreases dramatically from 60% down to 20-30% with 3 iterations (Figure 3a). With 10-15 iterations, all four cokes are almost fully eroded. However, if the small pores are removed, then the percentage of wall covered dropped to 60%. The erosion routine requires further optimization. Figure 3b compares how the coke walls of four cokes are covered progressively with each iteration of the erosion function. The sharper the incline towards 100% coverage, the thinner the char walls. Coke C is consistently the thickest coke wall structure while coke D is the thinnest, and more porous structure.
Fig.3 (a) Effect of pore removal severity on the percentage of material covered in four cokes after 3 iterations (b) Cumulative area of coke wall covered vs iteration number. Coke Mosaics - SEM Analysis Figure 4a illustrates typical SEM images of a coke lump while Figure 4b and Figure 4c shows the processed the images before and after cleaning of pores less than 2500 pixels. Figure 5a compares the pore size distribution among four cokes. Figure 5b shows that total porosity of coke C and coke D are comparable and higher than coke A and coke B. Total porosity from SEM data provides a similar trend of total porosity as indicated by optical imaging. Both were also seen to be consistent with the bulk porosity of cokes such that in both cases coke C and coke D indicated a higher porosity.
Fig. 4 (a) Illustration of typical SEM image of coke A; (b) image before and (c) after
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Oviedo ICCS&T 2011.
cleaning of less than 2500 pixel pores.
(a)
(b)
Fig. 5(a) Comparison of cutoff size (equivalent circular sieve) corresponding to 20%, 50% and 80% of pore (b) comparison of total SEM pore phase area of cokes. SEM analysis and coke mineralogy The mineral content of cokes from the SEM Mosaics can be determined using grey scale thresholding of the SEM images. This approach provides total mineral content of cokes without identifying the actual mineral or ash material. Majority of mineral grains in all cokes were found to be less than 100 μm in size. Figure 6 illustrates various minerals dispersed in coke carbon matrix. Quartz and decomposed clays were observed to be major mineral phases. It is possible that some of clays could not sufficiently decompose during coking but requires further improvement for phase identification.
Fig. 6 SEM image of coke A illustrating mineral distribution in coke matrix. Implications of Microscopic Characterisation of Coke Composition Total porosity based on optical and SEM data could not be related to the measured coke strengths. The percentage area of 745 micron diameter pore appear to show a better empirical correlation with M40 values of cokes (Figure 7a) but less apparently with the CSR (Figure 7b). This implies that the nature of pores could have a greater impact on
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Oviedo ICCS&T 2011.
coke strength. However, the exact contribution of various pore sizes on coke strength still needs to be studied.
Fig. 7 (a) M40 vs % of phase area of 745 micron pores (optical) and (b) CSR. SEM porosity data indicated a similar trend such that cokes containing high percentage of -0.1mm pore phase are indicated low M40 (Figure 8a) as well as CSR values (Figure 8b). The SEM data also indicated an inverse relationship between total mineral phase areas with both coke strengths. The SEM approach has additional advantage as it can simultaneous provide pore and mineral phase information.
(a)
(b)
Fig. 8 (a) M40 vs percentage of < 0.1 mm pore phase area based on SEM data (b) CSR. On the basis of coal reflectance data, a Full Maceral Reflectance parameter can be calculated [16]. The FMR parameter can be related to cold coke strength (Figure 9a) and CSR (Figure 9b) such that Coal A and B with extreme FMR values show lower strength.
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Oviedo ICCS&T 2011.
(a)
(b)
Fig. 9 (a) Coal reflectance parameter (FMR) vs M40 and (b) CSR vs FMR.
4.
Conclusions
A sample preparation scheme was established which enabled optical and scanning electron microscope examination of coke lumps using the same specimen such that both approaches successfully distinguished the total porosity of cokes as well as pore distribution in cokes. High strength cokes displayed higher amount of total porosity was mainly attributed to the largest size pores but did not show any meaningful correlation with coke strength indices. Coal FMR parameters was also related to coke strength values. The SEM approach seems to be promising for simultaneous measurements of the microstructure and mineralogy of cokes but requires further investigations.
Acknowledgement. Authors will like to express thanks to ACARP for funding this study and the industry monitors Mr Chris Dempsey and Mr Tim Manton for their support. Authors also appreciate the sample, technical support and suggestions provided by Mr Phil Bennett.
References [1]
Deiz MA, Alvarez R and Barriocanal C. Coal for metallurgical coke production:
predictions of coke quality and future requirements for cokemaking, Int. J. of Coal Geology, 2002;50(1-4): 389-412. [2]
Kojima K. Prediction of the coke strength coal by coal microscopic analysis, Journal of
Fuel Society of Japan, 1971;50(12): 894-901. [3]
Kojiro K. and Yoshihisa S. A Development of automatic coal petrographical analyses
for evaluating coking coals, Tetsu-to-Hagane, 1978; 64(12):1661-1670. [4]
Gransden JF, Jorgensen JG, Manery N, Price JT and Ramey NJ. Applications of
microscopy to coke making, Int. J. of Coal Geology, 1991;91:77-107.
[email protected]
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Oviedo ICCS&T 2011. [5]
Cloke M, Lester E, Allen M, Miles NJ. Repeatability of maceral analysis using image
analysis systems. Fuel. 995;74:654-8. [6]
O'Brien G and Jenkins B. Coal characterization by automated coal petrography. Fuel.
2003:1067-1073. [7]
O’Brien G, Jenkins B, Ofori P. and Ferguson K. Semi-automated petrographic
assessment of coal by coal grain analysis. Minerals Engineering. 2007. 20:428–434. [8]
Wu, T., Lester, E. and Cloke, M. A burnout prediction model based around char
morphology, Energy & Fuels. 2006. 20(3):1175-1183. [9]
Gupta S and Shen F. Advanced characterization of coal petrography for cokemaking,
Final POSCO Project Report, UNSW Sydney, Australia 2010. [10]
Patrick JW, Sims MJ, Stacey AE. The relation between the strength and structure of
metallurgical coke. J. Phys. D: Appl. Phys. 1980;13:937-51. [11]
Garboczi E, Bentz D, Martys N. Experimental methods in the physical science. Methods
in the Physics of porous Media. Academic press, San Diego.1999;35: Chapter 1. [12]
Andriopoulos N, Loo C, Dukino R, McGuire S. Micro-properties of Australia coking
coals. ISIJ International. 2003;43(10):1528-1537. [13]
Bennett P, Rigg D. Inertinite fusibility determination, Project C13066 Report, ACARP
Brisbane, 2008. [14]
Bennett P, Reifenstein A, O’Brien G. Coke reactivity and characterization, Project
C12057 Report, ACARP Brisbane, 2008. [15]
Bennett P, Gupta S, Advanced characterization of coke microstructure for use in
prediction of coke strength, 2011, ACARP Project C18043, Brisbane. [16]
Koba K, Sakata K and Ida S. Gasification studies of cokes from coals. The effects of
carbonization pressure on optical texture and porosity, Fuel;1981;60:499-506. [17]
Pusz S, B Kwiecinska, Koszorek A, Krzesinska M and Pilawa B. Relationships between
the optical reflectance of coal blends and microscopic characteristics of their cokes, Int. J. of Coal Geology, 2009;77(3-4):356-362. [18]
Jasienko S and Kidawa H. The behaviour of petrographic constituents of hard coals in
coking process and structure of the cokes obtained. Fuel Process. Technology.1987;17:155-167. [19]
Nishioka K, Yoshida S. Strength estimation of coke as porous material, ISIJ
1983;23:475-481. [20]
Yamaoka, H and Suyama S, Prediction model of coke strength after gasification
Reaction, ISIJ International, 2003: 43 (3):338-347 [21]
Yamaoka H, Suyama S., and Nakano K. Development of a size degradation model of
coke particles at the Drum Test and inside the Blast Furnace. ISIJ Int. 2003;43(1):44-53. [22]
Sharma R, Dash PS, Banerjee PK, Kumar D. Effect of coke Micro-Textural and coal
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Oviedo ICCS&T 2011. petrographic properties on coke strength characteristics. ISIJ Int..2005;45(12):1820-1827 [23]
Krzesinska M, Pusz S, Koszorek A. Elastic and optical anisotropy of the single-coal
monolithic high-temperature (HT) carbonization products obtained on a laboratory scale. Energy & Fuels 2005;19:1962-1970. [24]
Fukuda K, Kato K, Kunitomo K, Naito M. Coke degradation after reaction with carbon
dioxide in a blast Furnace, Proceedings of AISTech Pittsburgh Conference, 2008. [25]
Pusz S, Krzesinska M, Smędowski L, Majewska J, Pilawa B, and Kwiecinska B.
Changes in a coke structure due to reaction with carbon dioxide. Int. J. of Coal Geology. 2010;81(4):287-292. [26]
Diessel CFK. Carbonisation reaction of inertinite macerals in Australian coals.
Fuel.1983;62: 883-892. [27]
Kubota Y, Nomura S, Arima T and Kato. Effects of coal inertinite size on coke strength.
ISIJ International. 2008;48(5):563-571. [28]
Gupta S, French D, Sakurovs R et al., Coke minerals and ironmaking reactions in Blast
Furnaces. Progress in Energy & Combustion Science. 2008;34:155-197.
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Small Scale Determination of Metallurgical Coke CSR Tony MacPhee1, Louis Giroux1, Ka Wing Ng1, Ted Todoschuk2, Marcela Conejeros3 and Cornelis Kolijn3 1
CanmetENERGY, 1 Haanel Drive, Ottawa, ON, Canada, K1A 1M1, ArcelorMittal Dofasco Inc., 1390 Burlington Street East, Hamilton, ON, Canada, L8N 3J5 3 Teck Coal Ltd., Suite 1000, 205-9th Ave. S.E.,Calgary, AB, CANADA T2G 0R3 2
ABSTRACT
Coke Strength after Reaction (CSR) and the concomitant Coke Reactivity Index (CRI) are useful parameters for assessing the behaviour of coke in blast furnace. CanmetENERGY has developed a procedure for producing coke to allow CSR measurement involving relatively small amounts of coal sample (~15 kg). The procedure involves producing semi-coke using the Sole Heated Oven in accordance with ASTM standard D2014-97(2004). The resulting semi-coke is quenched (either wet or dry) and subsequently reheated to 11000C under nitrogen for 1 hour. CSR and CRI are measured using the coke produced using this procedure. To demonstrate the validity of CSR and CRI measured using this approach, identical coal blends were carbonized concurrently using the CanmetENERGY pilot scale moveable wall oven (460 mm wide, 350 kg capacity) and the procedures mentioned above. CSR and CRI of the cokes produced by these two approaches were compared. Statistical analysis and coke textual analysis were performed to demonstrate the applicability of this novel approach on CSR determination with limited amount of coke sample. Keywords: Coke, Cokemaking, CSR, CRI, Carbon Forms
Introduction Current world-wide ironmaking capacity is dominated by blast furnace technology. In 2009, this technology accounted for 94% of global hot metal production, 912.2 Mt [1]. In comparison, direct reduced iron (DRI) production was limited to only 53.1 Mt [2]. The performance and efficiency of blast furnace ironmaking technology is continuously improving. One of the factors contributing to the success of blast furnace ironmaking is the continued amelioration in coke quality. Metallurgical coke is the major source of carbon in the operation of the blast furnace. Besides being responsible for producing reducing gas, coke also supports the descending burden and provides passages/voids through it for distributing reducing gas in the furnace. Moreover, the combustion of coke in the lower hearth by the injected blast generates heat for melting of the hot metal. Because of the numerous functions of coke
in the blast furnace, stringent requirement on its physical and chemical properties are needed to ensure smooth operation of high productivity modern blast furnaces [3]. In the lower hearth of the blast furnace, coke is the only solid material present to support the entire weight of the burden above. Coke, possessing high mechanical strength in extremely hot and dynamic environments, is required to cope with the increase in throughput of large sized blast furnaces. Prior to its descent to the lower hearth, coke reacts with CO2 produced from the reduction of iron ore to generate CO to replenish the source of this reducing gas required for reduction of iron ore in the upper portion of the furnace. Therefore, the coke charged into the blast furnace must be capable of generating CO by reacting with CO2 while simultaneously maintaining its physical strength after reaction. Among the important indicators for assessing the quality of coke for blast furnace application are Coke Reactivity Index (CRI) and Coke Strength after Reaction (CSR) developed by Nippon Steel Corporation in the early 1970’s [4]. To estimate the quality of coke produced from specific coal blends in terms of CSR and CRI, the most economical way is by carrying out carbonization in a pilot-scale coke oven having known and proven capability of producing industrial grade coke. CanmetENERGY currently operates two pilot-scale slot-type coke ovens (460 mm wide and 350 kg capacity). The pilot-scale coke oven used in this investigation is depicted in Figure 1.
Figure 1. CanmetENERGY Pilot-Scale Coke Oven The cokes produced in either of these ovens have been shown over time to be of similar quality to industrial coke via benchmarking against industrial ovens [5]. The pilot scale oven test requires a specific amount of coal to ensure the coke produced is appropriate for CSR evaluation. In such instances where the amount of coal available is insufficient to perform a pilot-scale test, as for the case of an exploration bore hole sample, CSR and CRI normally cannot be measured. In an effort to broaden the ability to measure CSR and CRI with limited amounts of coal sample, a novel procedure for producing coke involving carbonization using a soleheated oven was developed at CanmetENERGY. The sole-heated oven used in this study is shown in Figure 2.
Figure 2. CanmetENERGY Sole-Heated Oven To validate this new approach for producing coke for CSR and CRI evaluation, concurrent carbonization tests of identical coal blends using sole-heated oven and pilotscale moveable wall oven were performed. CSR and CRI of the cokes produced using both carbonization routes were compared. Besides comparing sole-heated oven and pilot-scale oven cokes for their CSR and CRI, these were also evaluated for their Apparent Specific Gravity (ASG), the ratio of the mass of a volume of dry coke to the mass of an equal volume of water. Coke ASG varies with the rank and ash content of the coal carbonized, the bulk density of the coal charge in the oven, the carbonization temperature and the coking time [6]. Furthermore, microscopic analysis of the textures was performed on the sole-heated and pilot-scale oven cokes to compare them for their carbon forms. As discussed later, this technique is, among numerous advantages, extremely useful for understanding the behaviour of coal during coking and for interpreting pilot movable wall oven results including pressure generation and coke quality results. Experimental Coke Preparation by Sole-Heated Oven The preparation of a coke sample for CSR evaluation using a coal sample of limited quantity involves two steps, namely semi-coke preparation and heat treatment. For preparation of semi-coke, the sole-heated oven was employed in accordance with the ASTM D2014-97(2010) Standard for expansion or contraction of coal. A total of 12 kg of sample (coal or blends) is divided equally and each half-charged into chambers approximately 280 mm in width, length and depth of a double-chambered oven. A weighted piston applies a constant load of 15.17 ± 0.35 kPa on the surface of the coal in each of the chambers throughout the test. The coal bed is heated from below by the sole plate initially set at 554 °C and gradually ramped up to 950 °C according to a prescribed temperature program. The test is considered complete when the temperature at the top of the coal bed reaches 500 °C after a period of 6-7 hours. Following completion of the test, the coke is pushed from the oven, quenched in water, drained and oven dried
overnight at 120 °C. In cases where expansion/contraction values are reported, the measured expansion or contraction of the sample is converted to a reference base of 833 kg/m3 and 2% moisture. For heat treatment, the dried semi-coke (8-9 kg) is subsequently introduced in pieces (50 x 175 mm) into a stainless steel holding box hermetically sealed on top with 3 mm thick section of stainless steel with a 1 cm exit hole in the centre for venting the hot coke gases. The holding box is connected to N2 gas for continually flushing the semi-coke (5-10 L/min flowrate) to prevent its combustion. The holding box with the semi-coke inside is then heated in a Muffle Furnace from ambient to 1100 °C in 2-3 h at the rate of 5-10 °C/min. Upon attaining 1100 °C, the coke is soaked for an additional hour. Then, cooling is allowed to take place to approximately 100 °C. The entire heating and cooling cycle is carried out in a continuous flow of N2 and requires about 15 hrs to complete. The average weight loss of coke during the heat-treatment process has been measured to be 5 ± 2%. CSR, CRI and ASG Determination After the heat treatment, the coke sample produced was prepared and tested for CSR and CRI measurement as per specifications in the ASTM D5341-99(2010) Standard. By definition, the CRI is the percent weight loss of the coke sample after reaction in CO2 at 1100oC for 2 hours. The cooled, reacted coke is then tumbled in an I-drum for 600 revolutions at 20 rpm. The cumulative percent of +9.5 mm coke after tumbling is denoted as the CSR. The repeatability limit (r) of this method for CRI is 2.4 and for CSR is 5.4; the reproducibility limit (R) of this method for cokes in CRI range 20-28 is between 5 and 7 and that for cokes in CSR range 55-70 (acceptable grade for ironmaking) is also between 5 and 7. ASG of cokes were determined following a method developed at CanmetENERGY and related to the ASTM D167-93(2004) and ISO 1014:1985 Standards. Coke Texture Analysis Carbon form analysis in this study was carried out as per a combination of the US Steel method [7] and the CanmetENERGY method. A single point count is made for each measured field of view. For each field, the stage is rotated in order to determine the possible highest rank carbon form. Normally 500 point counts are performed on a sample. Each carbon form is derived from an assumed parent coal V-type. From the coke texture analysis, one can determine the ‘effective coal reflectance (%Ro)’ and also the percentages of low-, medium- and high-volatile parent coals used in the blend. The CMSI [8] is the coke mosaic size index defined as:
CMSI= {(%Incipient)+2(%CF+%CM)+3(%CC+%LF+%LM)+4(%LC+%RF)+5(%RM+%RC)} (i) (100 - %Isotropic - %Total Inerts)
This is a mathematical method to summarize the carbon form analysis. The formula used in this work is an adaptation of Coin’s method. The higher the CMSI, the higher the rank based on carbon forms measured.
Results The primary objective of this work is to validate and compare the measurement of CSR and CRI of cokes produced in the sole-heated oven to those generated in a pilot-scale oven. To achieve this goal, nineteen (19) coal blends were thus carbonized concurrently in both types of ovens at CanmetENERGY. As described earlier, the CSR test has repeatability and reproducibility limits of 5 and 7, respectively (ASTM D5341-99(2010). The sole heated oven procedure produces coke for CSR testing that is within these limits for comparing to pilot-scale oven results. Table 1 compares CSR’s and CRI’s of cokes produced by the two types of ovens. CSR range of the cokes examined was between 42 and 65. CRI range of the cokes examined was between 23 and 37. The difference in CSR and CRI, expressed as percentage, between cokes produced by the two ovens are also listed in this table. Table 1. Comparison of CSR and CRI of Pilot Scale Oven and Sole Heated Oven Cokes Pilot Scale Oven
CSR Sole Heated Oven
% Difference
Pilot Scale Oven
CRI Sole Heated Oven
% Difference
61.9 60.5 60.3 60.3 58.0 65.0 51.3 60.5 60.6 62.7 59.9 42.8 60.9 50.6 60.3 57.1 54.5 52.1 57.9
62.4 60.8 63.9 61.8 54.7 63.2 50.9 51.4 61.5 51.4 56.7 41.5 60.2 51.8 61.7 55.6 59.0 52.4 55.6
0.8 0.5 6.0 2.5 -5.7 -2.8 -0.8 1.5 1.5 -2.1 -5.3 -3.0 -1.2 2.4 2.3 -2.6 8.3 0.6 -4.0
22.7 25.1 26.9 25.3 28.1 22.5 31.0 25.3 25.6 25.9 26.6 35.9 26.0 32.0 27.6 28.8 30.1 31.5 26.1
23.9 25.1 24.9 25.0 29.1 24.2 29.9 25.8 26.2 25.8 28.0 37.2 26.9 30.8 25.9 30.3 27.0 32.8 29.0
5.3 0.0 -7.4 -1.2 3.6 7.6 -3.5 2.0 2.3 -0.4 5.3 3.6 3.5 -3.8 -6.2 5.2 -10.3 4.1 11.1
Figure 3 shows the scatter plot between the CSR obtained for the sole heated oven coke and that for the pilot scale oven coke. A strong linear relationship was observed. In the same graph, a straight line passing through origin with unit slope (i.e. y=x plot) was superimposed for comparison. As can be seen, the data points are evenly distributed on both sides of the superimposed line, which indicating that there is no bias due to the coke preparation method with respect to the y=x plot. Furthermore, the deviations of the experimental data from the superimposed y=x plot were within the expected repeatability and reproducibility of CSR measurement. This indicates that CSR’s of the coke samples produced using both methods are statistically identical. The observed differences in the data points arise from the accumulated random error in coal handling, coke preparation and CSR measurement, etc. A more detailed analysis of
the difference in CSR measured between the two preparation methods reveals that the average of the difference in CSR among the measurements is -0.06% and the standard deviation is 3.6%. The calculated 95% confidence interval is 1.6%. As shown in this error analysis, the CSR measured using sole-heated coke is within ± 2% of the pilot oven CSR value.
Figure 3. CSR Relationship between Sole-Heated Oven and Pilot Scale Oven
Similar to the CSR analysis, Figure 4 shows the scatter plot between the CRI obtained for the sole heated oven coke and that for the pilot scale oven coke. A strong linear relationship was observed. As can be seen, the data points are evenly distributed on both sides of the superimposed line y=x plot, which indicating that there is no bias due to the coke preparation method. Furthermore, the deviations of the experimental data from the superimposed y=x plot were within the expected repeatability and reproducibility of CRI measurement. This indicates that CRI’s of the coke samples produced using both methods are statistically identical. The average of the difference in CRI between the two coke preparation methods is 1.1% and the standard deviation is 5.4%. The calculated 95% confidence interval is 2.4%. Hence, the CRI measured using sole-heated coke is within ± 3% of the pilot oven CRI value. To develop a better understanding of the influence of different carbonization methods on the properties of the resulting coke, an analysis of coke ASG was also performed. As indicated in Table 2, ASG of sole-heated oven cokes are consistently higher than that of pilot oven cokes. In fact for the nineteen (19) cokes examined, ASG of sole-heated oven coke is, on average, about 6% higher than ASG of pilot oven cokes. This finding is mainly attributed to the fact that sole-heated oven cokes are produced under a constant load of 15.2 kPa, which is higher than the pressure exerted on the coal during carbonization in the pilot scale oven, ~8 kPa. Figure 5 shows the scatter plot between ASG measured on sole-heated oven coke and pilot oven coke and compared with the superimposed y=x plot.
Figure 4. CRI Relationship between Sole-Heated Oven and Pilot Scale Oven
Table 2. Comparison of ASG in Pilot Scale and Sole-Heated Ovens Pilot Scale Oven 0.935 0.921 0.944 0.921 0.930 0.920 0.918 0.954 0.953 0.953 0.931 0.951 0.930 0.922 0.942 0.956 0.925 0.946 0.927
Sole Heated Oven 0.991 0.977 0.994 0.979 0.966 0.967 0.987 1.018 1.001 1.009 0.957 1.016 0.971 0.994 1.012 1.009 0.989 1.002 0.987
% Difference 6.00% 6.11% 5.21% 6.28% 3.88% 5.11% 7.54% 6.70% 4.96% 5.87% 2.81% 6.75% 4.42% 7.89% 7.39% 5.59% 6.95% 5.93% 6.45%
Figure 5. ASG Relationship between Sole-Heated Oven and Pilot Scale Oven Carbon form analysis was performed on seven coke samples from both the pilot-scale movable wall oven (A-G) and the corresponding sole-heated oven (A’-G’). A summary of the carbon form analysis results is given in Table 3. The data in this table shows the total of each category based on size of the carbon form itself, i.e., fine, medium and coarse. Inerts (filler phase), are also classified by subgroups such as fusinite, semifusinite, and unidentified inerts along with other inerts, which are summed up in the table. As shown, similar results were obtained from both the pilot-scale movable wall test oven and the sole heated oven in terms of percent binder phase (reactive) and filler (inert) phase.
Table 3 Carbon Form Analysis Results on Selected Cokes from the Pilot Scale Oven and SoleHeated Oven Total Binder%
Total Filler%
CMSI
Blend Ro
Pilot Scale Oven
Sole Heated Oven
Pilot Scale Oven
Sole Heated Oven
Pilot Scale Oven
Sole Heated Oven
Pilot Scale Oven
Sole Heated Oven
81.79 76.12 80.45 67.13 75.21 72.66 72.70
82.06 78.48 82.43 73.15 74.23 69.25 72.06
18.21 23.88 19.55 32.87 24.79 27.34 27.30
17.94 21.52 17.57 26.85 25.77 30.75 27.94
2.77 2.87 2.71 2.73 2.95 2.87 2.80
2.71 2.84 2.69 2.70 2.90 2.90 2.76
1.12 1.15 1.12 1.11 1.17 1.16 1.13
1.13 1.14 1.12 1.11 1.17 1.16 1.14
A good linear relationship between percent binder phase present in sole-heated and pilot scale oven cokes was observed, Figure 6. Also, the sole-heated and pilot scale oven cokes show excellent linear relationship for ‘effective coal blend reflectance, Ro’ as presented in Figure 7. This further denotes the similarity of the cokes produced in both types of ovens.
Figure 6. Binder Phase (Reactive) Relationship between Sole-Heated Oven and Pilot Scale Oven Cokes
Figure 7. Effective Coal Blend Ro Relationship between Sole-Heated Oven and Pilot Scale Oven Cokes Conclusions This paper describes the development of a novel procedure at CanmetENERGY for the evaluation of coke CSR using a small-scale carbonization oven – the sole-heated oven. A comparison between cokes produced in a sole-heated oven using the method developed in this work and those formed in a pilot-scale coke oven found the following: 1. CSR’s and CRI’s determined are statistically identical. 2. ASG’s for sole-heated oven cokes are higher on account of the higher pressure (load) applied on the coal bed.
2. Carbon forms expressed as binder phase (reactive) and filler phase (inert) are similar. 3. This procedure finds useful applications as a preliminary evaluation method and a reliable screening tool in both the cokemaking and coal mining industry as it provides relevant information on (i) Coking potential/ability of coal/coal blends (ii) CSR evaluation (iii) Carbon form development prior to designing and running pilot oven trials. Acknowledgements The authors would like to thank the Canadian Carbonization Research Association and CanmetENERGY personnel for their excellent contribution to the experimental work and for generating the data presented in this paper. References
[1] Statistic archive, World Steel Association, http://worldsteel.org/?action=stats_search&keuze=iron&country=all&from=2009&to=20 09, retrieved on February 21, 2011. [2] Statistic archive, World Steel Association, http://worldsteel.org/?action=stats_search&keuze=irondr&country=all&from=2009&to= 2009, retrieved on February 21, 2011. [3] Modern Blast Furnace Ironmaking – An Introduction, Geerdes, M., Toxopeus, H., van der Vliet, C., Verlag Stahleisen GmbH, Chapter 4, 2004. [4] Ida, S., Nishi, T., Nakama, H., Behaviour of burden in the Higashida No.5 blastfurnace. J Fuel Soc of Japan, Vol. 50, 645-654 (1971). [5] Leeder, W.R., Price, J.T., Gransden, J.F., ISS-AIME Ironmaking Conference Proceedings, Vol. 59, 55-65, (2000). [6] Price, J.T., Gransden, J.F., Metallurgical Coals in Canada: Resources, Research, and Utilization, Canmet Report 87-2E (1987). [7]Gray, R.J., DeVanney, K.F., “Coke Carbon Forms: Microscopic Classification and Industrial Applications”, International Journal of Coal Geology, Vol. 6, 277-297 (1986). [8] Coin, C.D.A., Microtextural Assessment of Metallurgical Coke, B.H.P. Central Res. Lab., CRL/TC/41, 1982.
Oviedo ICCS&T 2011. Extended Abstract
Chemical Looping Combustion of coal using a residue from alumina production T. Mendiara, G. Ferrer, P. Gayán, A. Abad, F. García-Labiano, L. F. de Diego, J. Adánez Department of Energy and Environment, Instituto de Carboquímica Miguel Luesma Castán 4, 50018 Zaragoza, Spain Tel: +34 976 733 977 Fax:+34 976 733 318 e-mail: [email protected]
Abstract In the recently developed Chemical Looping Combustion (CLC) technology, the oxygen needed for the oxidation of the fuel is provided by an oxygen carrier. The oxygen carrier circulates between two interconnected fluidized-bed reactors. In the fuel reactor the oxidation of the fuel takes place while the carrier is reduced. A product stream mainly consisting of CO2 and H2O leaves this reactor and therefore CO2 can be easily captured, once the water has been condensed. The reduced oxygen carrier is then transported to the air reactor, where it is re-oxidized in air before a new cycle begins. Today, there is an increasing interest on CLC application to solid fuels, especially coal. One of the possibilities to process coal in a CLC system is the in situ gasification and subsequent combustion of the product gases. Potential CLC oxygen carriers should comply with some chemical and mechanical requirements but it would be interesting that the carrier is as inexpensive as possible, as some losses are expected accompanying the coal ashes. In the present work, a residue from alumina production (red mud) mainly constituted by Fe2O3 has been tested as oxygen carrier in CLC of coal. Batch experiments were carried out in a fluidized bed reactor. Red mud performance was studied through several reduction/oxidation cycles using anthracite coal, sub-bituminous coal and lignite as fuels. The effect of operating conditions, such as temperature and gasification agent on the char conversion and combustion efficiency of gasification products were evaluated. The carrier showed high combustion efficiencies at all temperatures tested. Besides, independently of the fuel used, red mud was capable of burning the gases generated during the gasification of the corresponding char. The present results allow to consider red mud as a promising oxygen carrier for the iG-CLC of coal.
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Oviedo ICCS&T 2011. Extended Abstract
1. Introduction Carbon dioxide capture and storage (CCS) is currently considered as a mid-term solution to stabilize CO2 atmospheric concentration in an effort to mitigate the effects of global warming. Chemical Looping Combustion (CLC) technology is one of the technologies which allow combustion with inherent CO2 separation at low cost [1]. Two separate reactors, one for air and one for fuel are used. A solid carrier, normally a metal oxide (MxOy), transports the oxygen between the reactors. In the fuel reactor, the oxygen carrier is reduced while the fuel is oxidized. The product gas from the fuel reactor consists of CO2 and H2O. Water can be easily separated by condensation leading to a high CO2 concentrated stream, ready for compression and sequestration, without additional costs or energy penalties for gas separation. In the air reactor, the oxygen carrier is oxidized again with air to its original state. The net chemical reaction and combustion enthalpy is the same as in a conventional combustion. Several authors have successfully demonstrated the feasibility of this process in different CLC prototypes in the 10-140 kWth range using gaseous fuels and oxygen carriers based on nickel, cobalt and copper oxides. But increasing interest is developing on CLC using coal as fuel [2-4], regarding its intensive use. One of the possibilities for using coal in a CLC is the in situ Gasification-Chemical Looping Combustion (iG-CLC). Coal is physically mixed with the oxygen carrier in the fuel reactor and the carrier reacts with the gas products of coal gasification, where H2 and CO are main components. The efficiency of char gasification in the fuel reactor and the separation of ash from the oxygen carrier seem to be key factors for the development of this process. Reactions (1)(4) summarize the processes taking place in the fuel reactor: Coal → volatiles + char
(1)
Char + H2O → H2 + CO
(2)
Char + CO2 → 2 CO
(3)
H2, CO, volatiles (CH4) + n MxOy → CO2 + H2O + n MxOy-1
(4)
In the air reactor, the oxygen carrier is regenerated following reaction (5): MxOy-1 + ½ O2 → MxOy
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(5)
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Oviedo ICCS&T 2011. Extended Abstract
The development of appropriate oxygen carriers are one of the cornerstones in the evolution of the CLC technology. Besides the common chemical and mechanical properties demanded for an oxygen carrier, in CLC of coal it is especially interesting to find inexpensive oxygen carriers, as there might be losses of material during ash separation. Recently, several studies have pointed to the use of natural minerals [5] or waste materials as oxygen carriers [6, 7]. Red mud is an industrial residue generated in the alumina production via the Bayer process with a high amount of iron. In red mud, the active material Fe2O3 is present mainly together with Al2O3. The disposal of red mud constitutes a significant part of the overall production cost of alumina and is an area of ongoing concern to alumina refiners, as it is generated in a very large quantity. In previous studies by the authors [7], the reactivity of red mud to the main gases present in coal combustion was evaluated. Red mud showed high reactivity to H2 and CO and lower to CH4 in the temperature range 1100-1223 K. This work evaluates the behaviour of red mud as oxygen carrier in CLC of coal using different types of coal (anthracite, sub-bituminous and lignite). The experiments were conducted in a batch fluidized bed reactor. The objective of this work was to evaluate the effect of operating conditions, such as temperature (1173-1253 K) and gasification agent (H2O, CO2 or H2O/CO2 mixtures) on the char conversion and combustion efficiency of gasification products.
2. Experimental section
2.1 Oxygen carrier: Red mud The red mud sample used in this work was supplied by Alcoa Europe-Alúmina Española S.A. It was dried at room temperature for 72 hours and then sieved to the desired size (150-300 µm). Previously to its use, the dried sample was calcined at 1473 K during 18 hours to ensure complete oxidation of the sample and increase its mechanical strength. The main chemical and physical properties of the calcined material are shortly summarized in Table 1. The value of the oxygen transport capacity, RO, depends on the final oxidation state after reduction (Fe2O3–Fe3O4–FeO–Fe). Only the transformation from hematite (Fe2O3) to magnetite (Fe3O4) may be applicable for industrial CLC. Therefore, the value of the oxygen transport capacity, RO corresponds to the transformation to Fe3O4. Submit before May 31st to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
Table 1. Characterization of the red mud calcined sample
a
b
Fe2O3 (% wt)
71a
XRD main phases
Fe2O3, β-Al2O3
Crushing strength (N)
2.8
Oxygen transport capacity, RO, (%)b
2.4
Porosity (%)
3.7
Real density (kg/m3)
4500
Specific surface area, BET (m2/g)
0.1
Determined by TGA RO =
mox − mred mox
2.2 Solid fuel: coal char The fuel used was char from devolatilized Spanish anthracite, sub-bituminous South African coal and Spanish lignite. To produce the char, a batch of 300 g of coal particles (200-300 µm) was devolatilizated in a fluidized-bed reactor. The reactor was fluidized by N2 and it was heated up from room temperature to 1173 K with a temperature ramp of 20ºC/min and afterwards cooled down. The proximate and ultimate analyses of the obtained chars are shown in Table 2. The particle size of char was in the same range as the original coal. Table 2. Proximate and ultimate analyses of the different chars used Proximate analysis Sub-bituminous coal Water content 0.9 Ash 20.0 Volatile matter 1.1 Fixed carbon 78.0 Ultimate analysis Carbon Hydrogen Nitrogen Sulphur
76.5 0.2 1.6 0.8
Lignite 0.52 43.48 1.6 54.4
Anthracite 0.75 31.58 0.94 66.73
50.6 0.40 0.48 5.03
65.55 0.63 0.77 1.24
2.3 Experimental setup and procedure: fluidized-bed reactor Figure 1 shows the scheme of the experimental setup. The reactor was loaded with 200 g of red mud. The fuel was fed through a fuel chute which ends 3 cm above the distributor Submit before May 31st to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
plate and about 5-6 cm below the upper level of the fluidizing particles, so char particles are fed inside the fluidized bed. The upper part of the chute has a reservoir where the fuel is placed and later pressurized with nitrogen. v1
v2
N2
N2
P Solids feeding system
Filter v3 Gas analysis
Furnace H2O (l) Thermocouple
ΔP
Distributor plate
P
steam Air CO2 N2
Figure 1. Experimental setup: batch fluidized bed reactor The gas feeding system allowed feeding alternatively air, nitrogen or a mixture of steam/CO2. The total fluidizing flow was 200 LN/h, which corresponds to a gas velocity of 0.1 m/s at 1173 K in the reactor. Different gas analyzers continuously registered the CO, CO2, CH4, H2 and O2 concentrations in dry basis. CH4, CO and CO2 concentrations were measured by a non-dispersive infrared (NDIR) analyzer and H2 measurement was based on thermal conductivity. Oxygen concentration was determined using a paramagnetic analyzer. A downstream N2 flow of 90 LN/h was introduced to ensure a continuous dry gas flow feeding the analyzers as in most of cases the product gas is mainly composed by steam. During reduction periods, char was used as fuel whereas the reactor was fluidized with steam, CO2 or steam/CO2 mixtures (H2O:CO2 molar ratios of 90:10, 70:30, 50:50), which also acted as gasification agent. The reaction temperature was varied between 1173-1253 K. After every reducing period, red mud particles were fully re-oxidized with air before starting a new cycle after 2 min N2 purge. Loads of 1.5 g of char particles were fed in all cases in the upper part of the fuel chute (Figure 1).
2.4. Data evaluation The dry basis outlet flow (Fout) was calculated using the downstream introduced N2 flow. The rate of char conversion, rC(t), was calculated from a mass balance to carbon in
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Oviedo ICCS&T 2011. Extended Abstract
gaseous form in the reactor (no methane was detected). rC (t ) = ( yCO2 + yCO ) ⋅ Fout − FCO2 ,in ;
yi=component molar fraction
(6)
The evolution of char conversion, Xchar, with time can be calculated as: X char (t ) =
1
t
N C ,char
∫ r (t )dt
; Nc= carbon mol fed
C
0
(7)
The instantaneous rate of conversion of the char, rC,inst, is calculated as the rate of gasification per the amount of non-gasified carbon that is still in the reactor. rC ,inst (t ) =
rC (t )
∫
(8)
t
N C ,char − rC (t )dt 0
The rate of oxygen transferred from red mud to the fuel gas, rO(t), can also be calculated. The amount of hydrogen and oxygen in the fuel was not considered in the mass balances. The outlet water flow was calculated considering that the flow of hydrogen either in H2 or H2O comes only from introduced steam. The hydrogen coming from char moisture was considered to be negligible.
rO (t ) = FO ,out − FO ,in = Fout ⋅ ( 2 yCO2 + y CO − y H 2 ) − 2 FCO2 ,in
(9)
The conversion of red mud, Xred, in the fluidized bed for reduction reaction can be calculated from the integration of rO(t) with time: X red (t ) =
1 N O , rm
t
∫r 0
O
(t )dt
; NO,rm= oxygen mol in red mud
(10)
Finally, the combustion efficiency is defined as the oxygen gained by the fuel for its oxidation divided per the oxygen needed to fully oxidize the fuel.
ηc (t ) =
rO (t ) 2rC (t )
(11)
3. Results and Discussion
3.1 Effect of temperature
Figure 1 shows the instantaneous rate of char conversion (rC,inst) and the combustion efficiency (ηC) for the experiments performed with sub-bituminous char at different temperatures and using water as fluidizing agent. At higher temperatures, the char conversion rate is higher. There is an increase in the rate of char conversion with temperature, due to the increase in the char gasification rate, from around 10%/min to 40%/min at char conversion close to 0.75. Most of the CO and H2 generated during
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Oviedo ICCS&T 2011. Extended Abstract
gasification was consumed by the carrier an only small concentrations of these gases were observed. Therefore, the combustion efficiency was higher than 97% at all the temperatures tested. Moreover, the combustion efficiency was high and similar up to red mud conversions around 0.3 for the different temperatures tested, indicating that during a great part of the experiment red mud reacts fast with the gasification products at all the temperatures tested. 80
0.95 0.90
40
ηC
rC,inst (%/min)
60
1.00
1173 K 1223 K 1253 K
0.85 0.80
20 0 0.0
0.2
0.4
0.6
0.8
1.0
0.75
1173 K 1223 K 1253 K
0.70 0.0
0.2
0.4
Xchar
0.6
0.8
1.0
Xred
Figure 1. Instantaneous rate of sub-bituminous char conversion (rC,inst) as function of the
char conversion (Xchar) and combustion efficiency (ηC)) versus red mud conversion (Xred) at different temperatures and water as gasifying agent
3.2 Effect of the gasifying agent
Figure 2 shows the instantaneous rate of char conversion and the combustion efficiency for experiments at 1173 K using water and CO2 as gasifiying agents. 20
1.00
H2O CO2
15
CO2
0.90
10
ηC
rC,inst (%/min)
H2O
0.95
0.85 0.80
5
0.75
0 0.0
0.2
0.4
0.6
0.8
1.0
0.70 0.0
Xchar
0.2
0.4
0.6
0.8
1.0
Xred
Figure 2. Instantaneous rate of sub-bituminous char conversion (rC,inst) as function of the
char conversion (Xchar) and combustion efficiency (ηC)) versus red mud conversion (Xred) at 1173 K and different gasifying agents Submit before May 31st to [email protected]
7
Oviedo ICCS&T 2011. Extended Abstract
The gasification step is considered to be the limiting step in the iG-CLC process. Therefore, it should be optimized. In the experiments, the char conversion rate was lower when CO2 was used as the gasification agent instead of water. In this case, the main gasification gas is CO and its reaction rate with red mud is slower than in the case of H2. This would also explain the decrease in the combustion efficiency observed in the experiments with CO2. The results for the experiments with different H2O:CO2 molar ratios gave values between those presented in Figure 2. According to these results, the amount of CO2 in the gasifying stream should be limited to avoid low char gasification rates.
3.3 Effect of the coal type
Figure 3 shows the instantaneous rate of char conversion and the combustion efficiency for the experiments performed with different coal chars at 1173 K and using water as fluidizing agent. The rate of char conversion increases following the sequence anthracite, sub-bituminous and lignite, as it was expected considering the reactivity of these coals. Lignite char presents the highest char conversion rate, with values up to 9 times those observed for anthracite. The combustion efficiency is high in all the cases. These results indicate that even using high reactivity fuels, this oxygen carrier is capable of burning the gases that are generated during the gasification of the corresponding char.
60
1.00 0.95
40
Anthracite Lignite
Sub-bituminous
0.90 ηC
rC,inst (%/min)
Lignite
20
0.85 0.80
Sub-bituminous
0.75
0 0.0
Anthracite
0.2
0.4
0.6
0.8
1.0
0.70 0.0
Xchar
0.2
0.4
0.6
0.8
1.0
Xred
Figure 3. Instantaneous rate of char conversion (rC,inst) as function of the char
conversion (Xchar) and combustion efficiency (ηC) versus red mud conversion (Xred) at 1173 K and water as gasifying agent for anthracite, sub-bituminous and lignite chars
Submit before May 31st to [email protected]
8
Oviedo ICCS&T 2011. Extended Abstract
4. Conclusions
Red mud was evaluated as oxygen carrier for the in situ Gasification-Chemical Looping Combustion (iG-CLC) of coal. Red mud is a residue from alumina production mainly consisting of Fe2O3. In the iG-CLC, the coal is first gasified and the gasification products react with the oxygen carrier. For a better understanding of the process, the experiments performed in this work used char from the corresponding coal. Several variables influencing the process were analyzed. First, the temperature was varied between 1173 and 1253 K in experiments with water as gasifying agent. The carrier showed high combustion efficiencies at all temperatures tested. Another variable analyzed was the gasifying agent used. The values of the combustion efficiency were higher when water was used. Finally, different coals were tested: anthracite, subbituminous and lignite. The combustion efficiency in all cases was high. Independently of the fuel used, red mud was capable of burning the gases generated during the gasification of the corresponding char which leads to consider the further use of this carrier in the iG-CLC of coal.
Acknowledgement
The authors thank the Spanish Ministry for Science and Innovation for the financial support via the ENE2010-19550 project. T. Mendiara thanks for the “Juan de la Cierva” post-doctoral contract awarded by this Ministry. The authors also thank to Alcoa Europe-Alúmina Española S.A. for providing the solid material used in this work.
References [1] Eide LI, Anheden M, Lyngfelt A, Abanades C, Younes M, Clodic D. Novel capture processes. Oil Gas Sci Technol 2005;60:497–508. [2] Berguerand N, Lyngfelt A. Chemical-Looping combustion of petroleum coke using ilmenite in a 10 kWth unit – High temperature operation. Energy Fuels 2009;23: 5257-68. [3] Berguerand N, Lyngfelt A. Batch testing of solid fuels with ilmenite in a 10 kWth chemicallooping combustor. Fuel 2010;89:1749-62. [4] Linderholm C, Cuadrat A, Lyngfelt A. Chemical-looping combustion of solid fuels in a 10 kWth pilot-batch tests with five fuels. Energy Procedia 2011;4:385-92.
Submit before May 31st to [email protected]
9
Oviedo ICCS&T 2011. Extended Abstract [5] Cuadrat A, Abad A, Adánez J, de Diego LF, García-Labiano F, Gayán P. Design considerations for the in-situ Gasification Chemical-Looping Combustion process (iG-CLC) – Part 1. Experimental tests using ilmenite as oxygen carrier. Fuel Proc Tech 2011 (submitted for publication). [6] Ortiz M, Gayán P, de Diego LF, García-Labiano F, Abad A, Pans MA, Adánez, J. Hydrogen production with CO2 capture by coupling steam reforming of methane and chemical-looping combustion: Use of an iron-based waste product as oxygen carrier burning a PSA tail gas. J Power Sources 2011;196:4370-81. [7] Mendiara T, Abad A, de Diego LF, García-Labiano F, Gayán P, Adánez J. Red mud as oxygen carrier in chemical looping combustion of coal. 5th International Conference on Clean Coal Technologies. Zaragoza, Spain (2011).
Submit before May 31st to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
Chemical looping combustion of char with a Cu-based carrier Antonio Coppola1, Osvalda Senneca2, Roberto Solimene2, Riccardo Chirone2, Luciano Cortese2, Piero Salatino1 1 Dipartimento di Ingegneria Chimica, Università degli Studi di Napoli Federico II Piazzale Vincenzo Tecchio 80, 80125 Napoli (Italy) T: +39 081 7682258; F: +39 081 5936936; E: [email protected] 2 Istituto di Ricerche sulla Combustione – Consiglio Nazionale delle Ricerche Piazzale Vincenzo Tecchio 80, 80125 Napoli (Italy)
Abstract Chemical looping combustion with oxygen uncoupling of char from a bituminous coal with a Cu-based carrier is characterized by a combination of experimental techniques. The carrier consists of commercial reagent-grade CuO-dispersed regular alumina particles in the 850-1000µm size range. The inherent extent of oxygen uncoupling of the carrier has been characterized by oxygen switching in a thermogravimetric analyzer. The extent and rate of oxygen uptake by the carrier has been characterized over repeated oxidation/reduction cycles, and compared with those of pure copper oxide. Combustion of char in a bed of oxidized carrier has been characterized in a bench scale fluidized bed quartz reactor in oxygen-free atmosphere and monitored by continuous analysis of gases at the exhaust. Regeneration of the carrier has been carried out in the same reactor at constant temperature in air. The reactivity of the carrier has been determined after multiple oxidation/reduction cycles to assess the occurrence of deactivation.
1. Introduction The increase of CO2 concentration in the atmosphere is recognised as the main responsible for global warming [1]. Power plants firing fossil fuels represent the most important source of CO2 emissions. A promising way for reducing such emissions is to separate CO2 from flue gas to produce a concentrated CO2 stream ready for sequestration, along three alternative pathways: post-combustion, pre-combustion and oxy-fuel [2-5]. Chemical looping combustion [6-7] is an innovative combustion technology with inherent CO2 sequestration which does not belong to the categories mentioned above. The concept is based on utilization of solid oxygen carriers (OC), most typically metal
1
Oviedo ICCS&T 2011. Extended Abstract
oxides, which promote fuel oxidation without direct contact with atmospheric oxygen [8-10]. The process involves two reaction steps: 1. CnH2m + (n+½m)MexOy = nCO2 + mH2O + (n+½m)MexOy-2 2. MexOy-2 + O2 = MexOy where CnHm is a generic fuel, while MexOy and MexOy-2 are the metal oxide in oxidized and reduced form, respectively. The process involves cyclic steps of OC reduction by the fuel and oxidation by atmospheric air, which take place iteratively in the Fuel Reactor (reaction 1) and in the Air Reactor (reaction 2) (fig. 1). The oxidation of the carrier is exothermic, whereas the reduction step is endothermic, with the remarkable exception of CuO [11] which is characterized by an exothermic reduction stage. The Fuel and the Air reactors most typically consist of dual interconnected fluidized beds [12, 13]. Implementation of chemical looping combustion is conceptually fairly simple as far as gaseous fuels are concerned: contacting of the fuel with the carrier can be easily and effectively accomplished in this case. On the contrary, application of the CLC concept is far more difficult for solid fuels, as the solid-solid reaction between fuel and OC is not likely to occur with appreciable rate. On the other hand, CLC of solid fuels is a challenging goal, because of the large carbon content typical of solid fuels. A way to overcome fuel-OC contacting issues is represented by in situ or ex situ gasification of the solid fuel [14-17]. Ex-situ gasification involves the utilization of a pre-gasification reactor which produces syngas which is introduced in the Fuel reactor. The in situ gasification concept is based on promoting simultaneous gasification of the fuel and looping combustion of the produced syngas in the Fuel Reactor. In situ gasification has the advantage to be less expensive then the ex-situ gasification because the utilization of a gasification reactor is avoided. Moreover studies of syngas oxidation by metal oxides have demonstrated that this reaction is generally very fast at high temperatures, and that gasification is enhanced by the continuous removal of CO and H2, which are gasification inhibitors [18]. A major drawback of in situ gasification is represented by the large differences between the time scales of heterogeneous fuel gasification and of the syngas/OC reaction. This feature constrains the process to the establishment of large carbon inventories in the fuel reactor, which may be responsible for extensive carbon losses both in the bed drain and in the carryover [14]. A novel technique for the direct application of chemical looping combustion to solid
2
Oviedo ICCS&T 2011. Extended Abstract
fuels is based on the exploitation of chemical looping with oxygen uncoupling (CLOU) [19-24]. CLOU is based on the property of selected metal oxides to release gaseous oxygen at certain temperatures and oxygen partial pressures. This technique entails limited energy penalty and thus CO2 capture costs. The reaction scheme for CLOU is: 3. MexOy = MexOy-2 + O2 4. C + O2 = CO2 Both reactions take place simultaneously in the Fuel Reactor. The reduced OC is regenerated in the air reactor, much like conventional CLC, where it comes in contact with atmospheric oxygen and the reverse reaction 3 takes place. OC suitable for CLOU must meet specific requirements, including ability to reversibly uptake oxygen in the temperature range of interest for combustion, i.e. 800-1200°C. Possible OCs for CLOU are the systems: CuO/Cu2O, Mn2O3/ Mn3O4 and Co3O4/CoO [19]. Among them, CuO shows interesting properties for CLOU, as the equilibrium O2 concentration is about 0.045 at 950°C. The reactions in the fuel reactor are: 5. 4CuO = 2Cu2O + O2 6. C + O2 = CO2 Oxygen release (reaction 5) is endothermic but the global reaction is exothermic due to carbon combustion (reaction 6). The reduced OC is regenerated in the air reactor by means of the reverse reaction 5. According to reactions 5 and 6, the molar ratio R=C/Cu=0.25 fixes the maximum amount of carbon that can be oxidized per mole of copper during each cycle. The aim of the present study is the development of a protocol for the characterization of oxygen carriers suitable for application to CLOU of solid fuels. A combination of TG analysis and experiments in a bench scale fluidized bed reactor is directed to the characterization of the extent and rate of oxygen uptake in the regeneration stage, of the rate of reversible oxygen release during the desorption stage, of the influence of iterated reduction/oxidation cycles on the reactivity of the OC. The experimental protocol has been set up and tuned by applying it to a commercial reagent-grade oxygen scavenger. Table 1. Proximate and ultimate analysis of the coal char Ultimate Analysis Proximate %w %w (on dry basis) Analysis Moisture Volatile Matter Ash Fixed Carbon
3.34 7.04 20.45 69.17
Carbon Hydrogen Nitrogen Sulphur Ash Oxygen
71.92 1.37 1.48 0.52 20.45 4.26
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Oviedo ICCS&T 2011. Extended Abstract
Char 14 14
13 8 3
3
9
2
10
11
5
CO, CO2, O2
PID 15
4 12 1
6 7
N2
7
Air
Figure 1. The fluidized bed reactor. (1) gas preheating/premixing section; (2) fluidization column; (3) electrical furnaces; (4) porous plate; (5) thermocouple; (6) manometer; (7) mass flow meter/controller; (8) stack; (9) cellulose filter; (10) membrane pump; (11) gas analyzer; (12) personal computer; (13) hopper; (14) valve; (15) PID temperature controller.
2. Experimental section The material used for the tests was a commercial (Sigma-Aldrich) oxygen scavenger consisting of copper (II) oxide (13% by mass) supported on porous alumina. The carrier was supplied as regular particles in the size range 850-1000 μm. A few reference experiments were also carried out on pure CuO obtained by thermal decomposition of copper nitrate. The char used for the tests was obtained by devolatilization in a fluidized bed at 850°C for 5 min of mm-sized South African bituminous coal particles. Char was eventually ground and sieved in the size range 400-1000 μm . Char properties are given in Table 1. Experiments were carried out both in a Thermogravimetric analyzer and in a lab-scale fluidized bed reactor. Thermogravimetric analysis: The thermogravimetric analyzer NETSCH STA 409C/CD was equipped with on-line NDIR gas analyzer (ABB AO2020 Uras 14). Alternated oxidation/desorption conditions representative of those experienced in a CLOU cycle were simulated by the following sequence of steps: Step 1 (sample heating up and desorption) - Approximately 30mg of carrier were loaded into the TGA and heated in a flow of 100% nitrogen (200ml/min) up to the desired desorption temperature T at the heating rate of 10-50°C/min. The sample was then kept in nitrogen for a time td. Step 2 (oxidation) - The gas was switched over from nitrogen to a mixture of 21% oxygen in
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Oviedo ICCS&T 2011. Extended Abstract
nitrogen. The sample was held in oxidizing atmosphere for time to. Step 3 (desorption) The gas feeding to the TG was again switched over from 21% oxygen in nitrogen to 100% nitrogen and the sample was held in inert atmosphere for time td. Steps 2 and 3 were iterated several times. During the experiments the weight loss and uptake as well as the CO and CO2 concentrations at the exhaust were continuously monitored. The temperature in the experimental campaign was varied in the range 600-1000°C while the duration of the oxidation/desorption steps varied in the range 10-100min. Experiments in the fluidized bed reactor: CLOU was simulated in a bubbling fluidized bed reactor made of quartz, 30 mm ID operated at atmospheric pressure (Fig. 1). The reactor consists of two sections: a) the preheater/premixer of the fluidizing gas, 0.36 m high; b) the fluidization column, 0.42 m high. The reactor is equipped with a porous plate and is electrically heated with two semi-cylindrical ovens placed. A hopper located at the top of the fluidizing column is used to feed char to the reactor. A type-K thermocouple located 10 mm above the gas distributor is used to measure the temperature and to drive the control system of the oven electrical supply. Fluidizing gas is metered to the reactor via two mass flowmeters/controllers. Downstream of the reactor, a fraction of the exhaust gas is continuously sampled to measure CO, CO2 and O2 concentrations by analyzers in order to monitor the progress of reactions. Concentration signals are logged on a PC at a sampling rate of 1 Hz. FB experiments consisted of the following steps: Step 1 (sample heating up) Approximately 29 g of carrier were initially loaded into the reactor and heated in a flow of air (200 Nl/h) up to the desired desorption temperature T. The sample was then kept in nitrogen for a time td. Step 2 (char conversion) - The fluidizing gas was switched over from air to nitrogen. 30s after the switch, a batch of 0.35 g of char was loaded into the reactor via the hopper. The loading delay of 30 s was needed to avoid partial combustion of the char with the residual oxygen in the reactor after switch-over. The duration of the delay was selected on the basis of the analysis of the transient response of the reactor to step-wise switch of the fluidizing gas. The bed was kept in nitrogen for a time td. Step 3 (regeneration of the carrier) - The gas feeding to the FB was again switched over from nitrogen to air. The carrier was oxidized for time to. Steps 2 and 3 were iterated several times. The concentrations of CO and CO2 at the exhaust were continuously monitored during the experiments. Notably, the amount of
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Oviedo ICCS&T 2011. Extended Abstract
char loaded at the beginning of each char conversion stage (Step 2) corresponded to a value of R=C/Cu≅0.5, i.e. nearly twice the stoichiometric ratio consistent with reactions 5 and 6. The choice of operating with excess char combined with the fast rate of the char-oxygen reaction 6 were considered a good basis for assuming that all the released oxygen during step 2 reacted with the char and could be detected as CO2 and, to a lesser extent, CO at the exhaust. 102 N2
N2
Air
Air
Air
N2
100
w/wCuO, %/
98
96
94
92
90
88 0
50
100
150
200
250
t, min
Figure 2: Results of thermogravimetric analysis of pure CuO obtained from thermal decomposition of copper nitrate during simulated CLOU at 900°C.
3. Results and Discussion Experiments in the TG analyzer: Figure 2 reports the results of a reference TG experiment carried out with pure CuO obtained by in situ thermal decomposition of copper nitrate. The experiment was carried out at T=900°C. A nicely reversible behaviour is observed over three cycles as far as oxygen uptake/release is concerned. 105 Air
N2
Air
N2
Air
N2
104
w/wCuO, %
103 102 101 100 99 98 97 0
100
200
300
400
t, min
Figure 3: Results of thermogravimetric analysis of the commercial Cu-alumina oxygen carrier during simulated CLOU at 1000°C
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Oviedo ICCS&T 2011. Extended Abstract
The extent of oxygen uptake/release, nearly 10% of the sample mass, is fully consistent with the stoichiometry of reaction 5, confirming that all the copper participates in the reaction during each cycle. Oxygen release in the desorption stage is nearly complete over 20 min. Oxygen uptake in the oxidation stage is faster, being nearly complete over less than 10 min. Figure 3 reports results of TG analysis of the commercial carrier, expressed as weight of the sample on a CuO (alumina-free) basis during iterated oxidation/desorption at the temperature of 1000°C. The cyclic uptake and release of oxygen are clearly appreciated by the mass gain/loss profiles. Notably, the cyclic pattern appears to be superimposed to a permanent weight loss of about 2.5% over 400min. The reason of this weight loss is not certain, but it is possible that it is related to incomplete curing of the scavenger, which was used “as is”. When referred to the baseline, weight gains of 4.3, 3.7 and 3.2% are observed over the 45 min duration of the first, second and third oxidation cycles, respectively (to be compared with the theoretical maximum uptake of 10% on a CuO basis). Uptaken oxygen is released during the subsequent 60 min desorption stages. Experiments in the FB reactor: Figure 4 reports the time series of the CO2 concentration measured at the exhaust of the fluidized bed reactor during iterated char conversion (in nitrogen) and OC regeneration (in air) cycles at the temperature of 950°C.
7000 N2
Air
N2
Air
N2
Air
N2
6000
CO2, ppm
5000
4000
3000
2000
1000
0 5000
10000
15000
20000
25000
t, s
Figure 4: Time series of the CO2 concentration at the exhaust during simulated CLOU of coal char in the fluidized bed reactor at 950°C.
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Oviedo ICCS&T 2011. Extended Abstract
Concentrations of CO and O2 were also monitored, much smaller than those of CO2, and are not reported. The observed CO2 patterns are characterized by the following features: a) rapid CO2 emission in the early stage of the char conversion step due to char oxidation by the carrier and emission of residual volatile matter and surface oxides, followed by a more gradual decay; b) a pronounced CO2 peak during early regeneration of the carrier, due to rapid burn off of the unconverted char left after the previous char conversion stage (it is recalled that the C/Cu=R loading ratio at the beginning of the char conversion stage exceeded the stoichiometric value of 0.25); c) stable emission of CO2 during OC regeneration at levels corresponding to the baseline concentration in ambient air. The CO2 and CO profiles have been worked out to calculate the conversion degree of carbon batches loaded in each cycle as a function of time, reported in Fig. 5.
0.4 N2
Air
N2
Air
N2
Air
N2
0.35
0.3
X, ‐
0.25
0.2
0.15
0.1
0.05
0 5000
10000
15000
20000
25000
t, s Figure 5: Carbon conversion degree during each cycle of the simulated CLOU of bituminous coal char in the fluidized bed reactor at 950°C.
It is recalled that, due to the excess char loading (R=0.5), the maximum possible carbon conversion degree in each char conversion stage is 0.5. A carbon conversion degree of 15%, (corresponding to reduction of 30% of the stoichiometric Cu) is approached over nearly 18min during the first cycles, over about 45min in the subsequent cycles. Altogether, alumina-supported copper turns out to be far less reactive than pure copper oxide obtained from decomposition of nitrate and undergoes extensive deactivation over iterated cycles. This feature is consistent with results by Adánez-Rubio et al. [23, 24]. These authors have shown that CuO on γ-alumina is less reactive than CuO on different
8
Oviedo ICCS&T 2011. Extended Abstract
support materials and undergoes pronounced deactivation upon cycling. Literature on copper-based catalysis of oxidation reactions suggests than this might be related to solidsolid interaction between CuO and Al2O3 [25] with formation of a spinel (CuAl2O4). Moreover CuO may catalyze the phase change of γ-alumina into another phase, with smaller porosity at temperatures lower then pure Al2O3.
4. Conclusions An experimental protocol for the characterization of oxygen carriers for application to Chemical Looping combustion with Oxygen Uncoupling (CLOU) of solid fuels has been developed, based on the combined use of thermogravimetric analysis and simulated CLOU of coal char in a bench scale fluidized bed reactor. The protocol has been applied to a commercial Cu-based carrier supported on alumina. A benchmark was provided by reference data obtained using pure CuO from decomposition of nitrate as carrier. The commercial oxygen carrier displayed moderate-to-low reactivity and a marked propensity to undergo deactivation over iterated oxidation/reduction cycles, in accordance with previous studies on alumina-supported Cu carriers and catalysts.
References [1] IPCC. Climate Change 2007. http://www.ipcc.ch/ [2] IPCC. Carbon Dioxide Capture and Storage. http://www.ipcc.ch/ [3] Figueroa, J.D., Fout, T., Plasynski, S., McIlvried, H. and Srivastava, R.D., (2008). Advances in CO2 capture technology–The U.S. Department of Energy’s Carbon Sequestration Program, International Journal of Greenhouse Gas Control, 2, 9–20. [4] Kanniche, M., Gros-Bonnivard, R., Jaud, P., Valle-Marcos, J., Amann, J.M. and Bouallou, C., (2010). Pre-combustion, post-combustion and oxy-combustion in thermal power plant for CO2 capture, Applied Thermal Engineering, 30, 53–62. [5] D. Aaron, C. Tsouris. Separation of CO2 from flue gas: a review. Separation Science and Technology 2005; 40; 321-348. [6] A. Lyngfelt, M. Johansson, T. Mattison. Chemical-looping combustion-status of development. 9th International Conference on Circulating Fluidized Beds (CFB-9), May 13-16, 2008, Hamburg, Germany. [7] T. Mattison, A. Lyngfelt. Capture of CO2 using chemical-looping combustion. In Scandinavian-Nordic Section of Combustion Institute 2001. Göteborg. [8] M. Johansson. Screening of oxygen-carrier particles based on iron-, manganese, copper- and nickel oxides for use in chemical-looping technologies. Department of Chemical and Biological Engineering Environmental Inorganic Chemistry Chalmers University of Technology Göteborg, Sweden 2007. [9] J. Wolf, M. Anheden, J. Yan. Comparison of nickel- and iron-based oxygen carriers in chemical looping combustion for CO2 capture in power generation. Fuel 2005; 84; 993-1006. [10] J. Adánez, L.F. de Diego, F. García-Labiano, P. Gayán, A. Abad. Selection of
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Oxygen Carriers for Chemical-Looping Combustion. Energy & Fuels 2004; 18; 371377. [11] L. F. de Diego, F. García-Labiano, J. Adánez, P. Gayán, A. Abad, Beatriz, M. Corbella, J. M. Palacios. Development of Cu-based oxygen carriers for chemicallooping combustion. Fuel 2004; 83; 1749–1757. [12] P. Kolbitsch, T. Pröll, J. Bolhar-Nordenkampf, H. Hofbauer. Design of a Chemical Looping Combustor using a Dual Circulating Fluidized Bed Reactor System. Chem. Eng. Technol. 2009; 3; 398-403. [13] M.K. Chandel, A. Hoteit, A. Delebarre. Experimental investigation of some metal oxides for chemical looping combustion in a fluidized bed reactor. Fuel 2009; 88; 898908. [14] J. Wang, E.J. Anthony. Clean combustion of solid fuels. Applied Energy 2007; 85; 73-79. [15] T. A. Brown, J. S. Dennis, S. A. Scott, J. F. Davidson, A. N. Hayhurst. Gasification and Chemical-Looping Combustion of a Lignite Char in a Fluidized Bed of Iron Oxide. Energy Fuels 2010; 24; 3034–3048. [16] J. S. Dennis, S. A. Scott. In situ gasification of a lignite coal and CO2 separation using chemical looping with a Cu-based oxygen carrier. Fuel 2010; 89; 1623–1640. [17] J. S. Dennis, C. R. Müller, S. A. Scott. In situ gasification and CO2 separation using chemical looping with a Cu-based oxygen carrier: Performance with bituminous coals. Fuel 2010; 89; 2353–2364. [18] H. Leion, T. Mattisson, A. Lyngfelt. The use of petroleum coke as fuel in chemical-looping combustion. Fuel 2007; 86; 1947–1958. [19] T. Mattisson, A. Lyngfelt, H. Leion. Chemical looping with oxygen uncoupling for combustion for combustion of solid fuels. International Journal of Greenhouse Gas Control 2009, 3, 11–19. [20] H. Leion, T. Mattisson, A. Lyngfelt. Using chemical-looping with oxygen uncoupling (CLOU) for combustion of six different solid fuels. Energy Procedia 2009; 1; 447–453. [21] T. Mattisson, H. Leion, A. Lyngfelt. Chemical-looping with oxygen uncoupling using CuO/ZrO2 with petroleum coke. Fuel 2009; 80; 683–690. [22] A. Shulman, E. Cleverstam, T. Mattisson, A. Lyngfelt. Chemical-Looping with oxygen uncoupling using Mn/Mg-based oxygen carriers– Oxygen release and reactivity with methane. Fuel 2011; 90; 941-950. [23] I. Adánez -Rubio, P. Gayán, F. García-Labiano, L. F. de Diego, J. Adánez, A. Abad. Development of Cu-based oxygen carrier materials suitable for ChemicalLooping with Oxygen Uncoupling (CLOU) process. Energy Procedia 2011; 4; 417-424. [24] I. Adánez -Rubio, P. Gayán, A. Abad, F. García-Labiano, L. F. de Diego, J. Adánez. CO2 Capture in Coal Combustion by Chemical-Looping with Oxygen Uncoupling (CLOU) with a Cu-based Oxygen-Carrier. Clean Coal Technologies – CCT 2011; 8-12 May 2011; Zaragoza-Spain. [25] G.A. El-Shobaky, G.A. Fagal. Thermal solid-solid interaction between CuO and pure A12O3 solids. Thermochimica Acta 1989; 141; 205-216.
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Theoretical approach on the CLC performance with solid fuels: optimizing the solids inventory Ana Cuadrat, Alberto Abad, Pilar Gayán, Luis F. de Diego, Francisco García-Labiano, Juan Adánez Instituto de Carboquímica (ICB-CSIC), Dept. of Energy & Environment, Miguel Luesma Castán, 4, Zaragoza, 50018, Spain Email of corresponding author: [email protected] Abstract Chemical Looping Combustion (CLC) is a combustion technology with inherent separation of the greenhouse gas CO2. An option for CLC with solid fuels is the in situ Gasification-CLC (iG-CLC) process, where simultaneously the fuel gasification and the oxidation of the products of devolatilization and gasification take place. The objective of this work was to optimize the operating conditions for iG-CLC using ilmenite as oxygen-carrier. A theoretical simplified model for the fuel-reactor has been done. The modeled mass balances were compared to experimental results from tests performed in a 500Wth facility fuelled with “El Cerrejón” bituminous coal. Char gasification and reaction of ilmenite with H2, CO and CH4 kinetics are incorporated to the model. A contact efficiency for the volatile matter with the oxygen-carrier in the reactor was included. Simulations were done to evaluate the effect of the fuel-reactor solids inventory on the combustion and carbon capture efficiencies, the carbon stripper efficiency and fuel-reactor temperature. It is highly beneficial to increase the solids inventory up to 1000 kg/MWth, but the carbon capture has minor improvement with further increase and it is better to raise the carbon stripper efficiency. The carbon capture is promoted by increasing the temperature or residence time of char in the fuel-reactor. 1.
Introduction
Chemical-Looping Combustion (CLC) is a novel combustion technology with inherent separation of the greenhouse gas CO2 that involves the use of an oxygen-carrier, which transfers oxygen from air to the fuel avoiding the direct contact between them. Commonly, the CLC system is made of two interconnected reactors, the air- and fuelreactor, and an oxygen-carrier that circulates between them. In the fuel-reactor, the fuel is oxidized to CO2 and H2O by the oxygen-carrier that is reduced. The reduced oxygencarrier is further transferred into the air-reactor where it is re-oxidized with air. The flue
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Oviedo ICCS&T 2011. Extended Abstract
gas leaving the air-reactor contains N2 and unreacted O2. The exit gas from the fuelreactor contains only CO2 and H2O. After water condensation, almost pure CO2 can be obtained with little energy lost and low cost. One of the options to use the CLC technology with solid fuels is the so-called in-situ gasification Chemical-Looping Combustion (iG-CLC). In this technology the solid fuel is introduced directly in the CLC system. Fig. 1 shows the simplified reactor scheme of the iG-CLC process.
N2, O2
CO2
CO2, H2O
MexOy Air reactor
C MexOy-1
H2O(l) Fuel reactor
Carbon stripper Air
CO2
Coal Ash CO2
H2O(v)
Fig. 1. Reactor scheme of the CLC process using solid fuels (- - - optional stream). In the iG-CLC process the fuel-reactor is fluidized by a gasifying agent, e.g. H2O or CO2. Solid fuel devolatilization (1) and gasification (2) take place in the fuel-reactor and the resulting gases and volatiles are oxidized through reduction of the oxidized oxygencarrier, MexOy, by reaction (3). The oxygen-carrier reduced in the fuel-reactor, MexOy-1, is subsequently led to the air-reactor where it is re-oxidized with air by reaction (4). The net chemical reaction as well as the heat involved in the global process is the same as for usual combustion. Coal → Volatile matter + Char
(1)
Char + H2O / CO2 → H2 / 2CO
(2)
H2, CO, Volatile matter + n MexOy → CO2 + H2O + n MexOy-1
(3)
MexOy-1 + ½ O2 → MexOy
(4)
Experimentally it was found that working at high temperatures is required to reach high process performance [1,2]. The gasification step was found to be the limiting step in the fuel-reactor and the solids stream exiting the fuel-reactor can contain some unconverted char. Thus, implementing a carbon stripper is fundamental to ensure high extent of gasification [3]. The carbon stripper would increase residence time of char in the fuelreactor by separating and reintroducing ungasified char particles exiting the fuel-reactor. The objective of this work was to determine the operating conditions that optimize the
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Oviedo ICCS&T 2011. Extended Abstract
carbon capture and combustion efficiencies in a CLC system fuelled with coal. For that, a simplified model based on mass balances of a CLC system has been developed. For the simulations, a bituminous Colombian coal “El Cerrejón” was used as solid fuel and ilmenite, a natural ore, was selected as it has proven to be an adequate oxygen-carrier for iG-CLC [1-3]. To prove the validity of the model, the predicted results were compared to experimental tests done in a 500 Wth unit with this type of fuel and this oxygencarrier. The effect of the main operating parameters on the combustion and carbon capture efficiency of the CLC system was analyzed. The parameters considered were the solids inventory, the efficiency of a carbon stripper and the fuel-reactor temperature.
2.
Modeling of the fuel-reactor
A theoretical model is a very useful tool to understand the operation of the system in a general way, as well as to predict the influence of the different variables and optimize their values. A simplified theoretical model for the fuel-reactor has been developed, but describing with high accuracy the complex processes happening in the fuel-reactor. The model was developed and applied to El Cerrejón coal as fuel, whose properties can be found elsewhere [1]. The calculations were made with the basis of 1 MWth fuel power. The fuel-reactor is a fluidized bed, so the mixing of solids and gases in the reactor is determined by its fluid dynamics. To simplify, perfect mixing of the solids was assumed, gas plug flow in the bed, isothermal reactor and that there is no resistance to the gas exchange between bubble and emulsion phases. Besides, volatile matter is evolved in plumes, and no gas exchange between volatile plume and emulsion was considered. In this simplified model it was assumed that the fuel devolatilization was instantaneous and it took place in the bottom part of the bed. The calculation of the generated gases coming from the volatile matter is based in the distribution obtained by Matthesius et al. [4]. Furthermore, it is supposed that the volatile matter is released in form of H2, CO, CH4, H2O and CO2 as gaseous species. Furthermore, to adjust the theoretical to the experimentally measured species released by the volatiles, it was considered that 67% of the theoretically generated CH4 was reformed by the steam used as gasification agent. Thus, the calculated quantities of gases generated from the devolatilization of 100 g of El Cerrejón coal was 5.7g CO, 42.8g CO2, 7.1g CH4 and 5.5g H2. The kinetic parameters of char gasification for El Cerrejón were obtained by TGA analysis when using H2O and CO2 as gasification agents and taking into account the
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Oviedo ICCS&T 2011. Extended Abstract
inhibition effect of the gasification products H2 and CO. The oxygen-carrier considered in this work is ilmenite. Reaction kinetics for ilmenite were found by Abad et al. [5]. They found that ilmenite reacts with H2, CO, CH4 and O2 following a model of homogeneous reaction in the particle and reaction in the grains following a shrinking core model, for which there is control of the chemical reaction and with no resistance to mass transfer in the gas film and to diffusion inside the particle. Fig. 2 represents mass flow changes of the gases involved in the process for a differential mass inventory dmOC. Each dmOC is considered to have two zones: (1) the emulsion, where gasification takes place and (2) the volatile matter plume. The volatiles are released in a plume and they have poorer contact with the oxygen-carrier, which is taken into account with the parameter contact efficiency, Egs.
Emulsion mOC+dmOC
Fi,e,2
Volatile plume (-rilm)·Egs
Fi,v,2
rgasif + (-rilm) dmOC mOC
Fi,e,1
Fi,v,1
Fig. 2. Mass flow changes of the gases involved in the process –products of gasification and devolatilization- for a differential mass inventory dmOC. Eqs. (5) and (6) are the mass balances of gases in the emulsion in a differential mass inventory dmOC. For the volatile matter released reducing species, i.e., CH4, H2 and CO, the mass balances are Eqs. (7)-(9). These equations show the variation of the molar flows of H2, CO and CH4, i.e. FH 2 , FCO and FCH 4 , respectively. ∂FH 2 ∂mOC
+
f 1 Φ 1 (−rilm , H 2 ) + C (−rgasif ) H 2O =0 MC 2d ·M O 3 1 − fC
∂FCO f f 1 Φ 1 1 ( − rilm ,CO ) + C (− rgasif ) H 2O + + 2 C (− rgasif )CO2 =0 ∂mOC 2d ·M O 3 1 − fC 1 − fC MC MC ∂FCH 4 ∂mOC
∂FH 2 ∂mOC
(5) (6)
+
1 Φ (− rilm ,CH 4 )·Egs = 0 2 d ·M O 3
(7)
+
∂FCH 4 1 Φ (−rilm, H 2 )·Egs + 2 =0 2 d ·M O 3 ∂mOC
(8)
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Oviedo ICCS&T 2011. Extended Abstract
∂FCH 4 ∂FCO 1 Φ + (− rilm ,CO )·Egs + =0 ∂mOC 2d ·M O 3 ∂mOC
(9)
(−rgasif ) H 2O and (− rgasif )CO2 are the char gasification rates with H2O and CO2, respectively. fC is the carbon concentration in the fuel-reactor bed. Char is supposed to be uniformly distributed throughout the bed in perfect mixing and therefore the gasification is equal throughout the bed and is function of the char concentration.
(−rilm, H 2 ) , (−rilm ,CO ) and (−rilm ,CH 4 ) are the reaction rates of ilmenite with H2, CO and CH4, respectively. Ф is the characteristic reactivity in the reactor, which can be obtained from Fig.7 in the work done by Abad et al. [6] as a function of the conversion of solids at the reactor inlet and the variation of the solids conversion in the reactor, ΔXr. The molar flow variation of generated H2O corresponds to the molar flow variation of H2 disappeared and the molar flow variation of generated CO2 corresponds to the molar flow variation of CO disappeared. The H2O and CO2 in the emulsion and plume phases are also considered separately. The global balance to the whole reactor is done by integrating the differentials over the total mass inventory.
3.
Performance evaluation
The capture efficiency, ηCC, is the physical removal of carbon dioxide that would otherwise be emitted into the atmosphere and it is defined as the fraction of the carbon introduced with the coal feed flow, Ffuel,in, that is converted to gas in the fuel-reactor.
ηCC =
[C ] fuel ·Ffuel ,in / M C − FCO2 , AR [C ] fuel ·Ffuel ,in / M C
(10)
[C]fuel is the carbon fraction in the fuel. ηCC depends on the char conversion, Xchar, which is the fraction of the carbon flow from the char fed with the fuel that has been gasified. The implementation of a carbon stripper is therefore fundamental to ensure high carbon capture in iG-CLC. The carbon stripper efficiency, ηCS, is defined as the fraction that is separated and recirculated with respect to the ungasified char that exits the fuel-reactor.
ηCS = 1 −
FCO2 , AR m& OC · fC (1 − f C )
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(11)
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Oviedo ICCS&T 2011. Extended Abstract
where ( FCO + FCO + FCH )out is the sum fuel-reactor outlet flows of CO2, CO and CH4 2
4
obtained from the model. m& OC is the oxygen-carrier recirculation rate. The simulation results are obtained after adjusting the carbon fraction in the fuel-reactor bed so that the carbon balance is accomplished in the fuel-reactor:
FC , fuel =
[C ] fuel ·Ffuel ,in MC
(12)
= ( FCO2 + FCO + FCH 4 )out + m& OC · f C (1 − f C )
The combustion efficiency in the fuel-reactor, ηcomb
FR,
is the fraction of the oxygen
demanded by the volatiles and gasification products that is supplied by the oxygencarrier. It is calculated as the oxygen flow contained in the product gases from the fuelreactor that was supplied by ilmenite divided by the oxygen demanded by the coal fed in the fuel-reactor minus the ungasified char that escapes to the air-reactor, FCO2 , AR . [ H ] fuel and [O] fuel are the hydrogen and oxygen fractions in the fuel:
ηcomb FR =
4.
( FH 2O + 2 FCO2 + FCO ),out − ( FH 2O + 2 FCO2 + FCO ),in (2[C ] fuel / M C + 0.5[ H ] fuel − [O] fuel / M O )·F fuel ,in − FCO2 , AR
(13)
Results and discussion
A model has been developed to predict both the carbon capture and combustion efficiencies of the CLC process with coal as fuel as a function of various operational parameters. The model calculates the gas flow for CO2, CO, H2, H2O and CH4 exiting from the fuel-reactor, as well as the fraction of unconverted char passing to the airreactor. Firstly, the predictions of the model were checked with experimental results in a 500 Wth CLC plant. After that, simulations were performed using the model in order to evaluate the effect of operational parameters on the performance of iG-CLC systems. The operational parameters analyzed were the fuel-reactor solids inventory, carbon stripper efficiency and fuel-reactor temperature. All simulations performed are done for a corresponding thermal power of 1 MWth, which is a coal feeding rate of 0.0386 kg/s. 4.1. Comparison with experimental results
Previous to the modeling of the process its validation was carried out with experimental
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Oviedo ICCS&T 2011. Extended Abstract
data available from the experiments in the 500Wth CLC rig operated with El Cerrejón bituminous coal and using ilmenite as oxygen-carrier. The facility has no carbon stripper. The experiments for comparison and corresponding simulations were done at TFR of 900ºC, the oxygen-carrier to fuel ratio φ was around 2, steam was used as gasification agent with steam to fixed carbon ratio H2O/C=0.7 and the inventory was 3100 kg/MWth. In addition, ilmenite enters the fuel-reactor fully oxidized, i.e. Xox,AR=1. The experimental results showed some incomplete combustion, which was found to be due to volatile matter that was not fully burnt due to poor contact of the volatile matter with the oxygen-carrier [1]. A contact efficiency of volatile matter, Egs, was introduced to predict the experimental fuel-reactor combustion efficiencies obtained. It was obtained that the Egs that predicts the combustion efficiencies obtained experimentally is 0.53%. Besides, the char conversion and carbon capture efficiency simulated are very close to the values obtained experimentally. The simulations showed that in the emulsion the gasification products are fully oxidized by the oxygen-carrier and unconverted CH4, H2 and CO coming from volatiles. The resulting H2/CO and CH4/CO relative fractions in the fuel-reactor outlet obtained are also very similar to the experimental values. The model validity was thereby checked. After, simulation and process optimization proceed and thereby solutions for the modeling and scale-up purposes. 4.2. Influence of carbon stripper efficiency, ηCS
In order to get higher extent of gasification, there is a need in iG-CLC of separating the char that has not been gasified and recirculate it back to the fuel-reactor, in order to increase the residence time of char and thereby raise the carbon capture. This can be made by a carbon stripper, whose beneficial effect has been experimentally confirmed [2]. Fig. 3 shows the simulated char conversion, carbon capture and combustion efficiencies with increasing solids inventory for ηCS of 0%, 80%, 90% and 100%. Higher ηCS leads to higher ηCC. For example, with 1000 kg/MWth at 950ºC the resulting
simulated ηCC increase from 45.2% with ηCS=0% to 69.4% with ηCS=80%, and 79.8% with ηCS=90%. Thus, the iG-CLC process needs a char recirculation system to achieve satisfactory values of ηCC. For ηCS=100% both char conversion and carbon capture efficiencies are 100%, since there is no char escaping to the air-reactor. The process performance for all ηCS simulated has high improvements up to inventories of 1000-
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Oviedo ICCS&T 2011. Extended Abstract
2000 kg/MWth, but from an inventory of 2000 kg/MWth an increase in ηCS has minor influence. For adequate values of inventory and with the presence of a carbon stripper, there are still some unburnt gases at the fuel-reactor outlet. This is because the unburnt species come from released volatile matter, due to their insufficient contact with the oxygen-carrier. This was experimentally seen by Cuadrat et al. [1]. Thus, a subsequent polishing step with pure oxygen would be needed to fully oxidize the unburnt gases in the fuel-reactor outlet.
100 80 Efficiency(%)
ηCS=100% 90% 80%
ηCS=100% 90% 80%
60
ηCS
0% 0%
40 20
ηCC
Xchar
a)
ηcomb FR
b)
c)
0 0
1000 2000 3000 4000
Inventory(kg/MWth)
0
1000 2000 3000 4000
0
Inventory(kg/MWth)
1000 2000 3000 4000 5000
Inventory(kg/MWth)
Fig. 3. Variation of a) carbon capture, b) char conversion and c) combustion efficiency
with increasing solids inventory for ηCS=0, 80%, 90% and 100%. TFR=950ºC. H2O/C=0.7. φ=2. ηCS=0. Xo,AR=1. Egs=0.0053.
4.3. Influence of the fuel-reactor temperature
Fig. 4.a) represents the carbon capture of the process, where it can be seen that to get high performance of the process, an inventory between 1000 and 2000 kg/MWth seem to be necessary and reasonable at all temperatures tested. To get high gasification and oxidation rates, it is necessary to work at high temperatures. With an inventory of 2000 kg/MWth and ηCS = 90%, the simulated ηCC was 79.3% at 900ºC and 90.7% at 1000ºC.
Efficiency(%)
80 Temperature
60
Temperature
40 ηCC
20
ηcomb FR
a)
b)
Total oxygen demand (%)
100
100
80 60 40
Temperature
20
Ox. demand
c)
0
0 0
1000 2000 3000 4000
Inventory(kg/MWth)
0
1000 2000 3000 4000 5000
Inventory(kg/MWth)
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0
1000 2000 3000 4000 5000
Inventory(kg/MWth)
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Oviedo ICCS&T 2011. Extended Abstract
Fig. 4. Variation of a) carbon capture, b) combustion efficiency and c) total oxygen
demand with increasing solids inventory for several fuel-reactor temperatures. 900ºC,
950ºC and
1000ºC. H2O/C=0.7. φ=2. ηCS=90%. Xo,AR=1. Egs=0.0053.
Regarding the oxidation of volatiles and gasification, in Fig. 4.b) it can be seen that the fuel-reactor combustion efficiency grows with the temperature because of the increase in the oxidation reaction of ilmenite with the products of gasification and devolatilization. The minimum solids inventory needed to obtain ηcomb FR above 85% is 1000 kg/MWth. With an exemplar inventory of 2000 kg/MWth, the simulated ηcomb
FR
was 86.0% at
900ºC, 90.7% at 950ºC and 94.2% at 1000ºC. As there is some H2, CO and CH4 at the fuel-reactor outlet, there must be a subsequent polishing step with pure oxygen to fully burn the fuel-reactor stream. Fig. 4.c) represents the total oxygen demand, which is the fraction of the oxygen demanded by the total coal fed that would need a later oxygen polishing step to ensure final full combustion of the fuel. The total oxygen demand decreases with an increase of the fuel-reactor temperature. With an inventory of 2000 kg/MWth the total oxygen demand is 11.5 % at 900ºC, 8.2% at 950ºC and 5.3% at 1000ºC. Thus, for an optimum performance, it would be best working at temperatures above 950ºC.
5.
Conclusions
In this work a simplified model of in situ Gasification Chemical-Looping Combustion with solid fuels was developed. The model includes the kinetics of coal char gasification and reaction of the products of gasification and devolatilization with the oxygen-carrier. The model was validated with experimental results from tests performed in a 500Wth facility fuelled with “El Cerrejón” bituminous coal and ilmenite as oxygen-carrier. The impediment to obtain full combustion in the fuel-reactor was assigned to low contact efficiency of the volatile matter with the oxygen-carrier, calculated to be 0.53% for the simulated facility. After that, the operating conditions for iG-CLC with solid fuels were optimized by analyzing the effect of relevant operational conditions on the performance of the system. It was found that it is essential to have enough inventory in the fuelreactor to oxidize the fuel and to enhance the carbon capture. The carbon capture was directly related to the extent of gasification, which is promoted by increasing the temperature or the residence time of char particles in the fuel-reactor. It is highly
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Oviedo ICCS&T 2011. Extended Abstract
beneficial to increase the solids inventory up to 1000 kg/MWth, but further increase does not give a relevant improvement. As an example, with inventory in the fuel-reactor of 1000 kg/MWth, at 1000ºC and a carbon stripper efficiency of 90% the carbon capture obtained is 86.0%. To further enhance the carbon capture efficiency, it is preferred to increase the carbon stripper efficiency rather than increase the solids inventory. As there is some H2, CO and CH4 at the fuel-reactor outlet, there must be a subsequent oxidation step to fully burn the fuel-reactor stream. The corresponding calculated oxygen demand of this subsequent oxygen polishing step was 12.6% at 1000ºC.
Acknowledgments
This work was partially supported by the Spanish Ministry of Science and Innovation (Project ENE2010-19550). A. Cuadrat thanks CSIC for the JAE Pre. fellowship. References [1] Cuadrat A, Abad A, García-Labiano F, Gayán P, de Diego LF, Adánez J. The use of ilmenite as oxygen-carrier in a 500 Wth Chemical Looping Coal Combustion unit. Submitted for publication.
[2] Berguerand N, Lyngfelt A. Chemical-looping combustion of petroleum coke using ilmenite in a 10 kWth unit-high-temperature operation. Energy & Fuels 2009;23(10):5257-5268. [3] Berguerand N, Lyngfelt A. The use of petroleum coke as fuel in a 10 kW chemical looping combustor. Int J Greenhouse Gas Control 2008;2:169-179. [4] Matthesius GA, Morris RM, Desai MJ. Prediction of the volatile matter in coal from ultimate and proximate analyses. J. S. Afr. Inst. Min. Metal. 1987;87(6):157-161. [5] Abad A, Adánez J, Cuadrat A, García-Labiano F, Gayán P, de Diego LF. Reaction kinetics of ilmenite for Chemical-looping Combustion. Chem Eng Sci 2011;66(4):689-702. [6] Abad A, Adánez J, García-Labiano F, de Diego LF, Gayán P, Celaya J. Mapping of the range of operational conditions for Cu-, Fe-, and Ni-based oxygen carriers in chemical-looping combustion. Chemical Engineering Science 2007;62:533-549.
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10
Puertollano IGCC: Towards zero emissions power plants F. García Peña and P. Coca Llano ELCOGAS, S. A., Carretera de Calzada de Calatrava km27, 13500 Puertollano (Ciudad Real, Spain) [email protected] Abstract
ELCOGAS main current activities include both CO2 capture and co-gasification using biomass. CO2 capture and H2 co-production are being undertaken through its 14 MWth pilot plant – installed and integrated in the IGCC plant-, which is treating 3,600 Nm3/h of the syngas generated in the IGCC. CO2 first tonne was captured on 13th September 2010. Main tests and targets up to completion of the project are comparison of sweet & sour capture (using different catalysts in the shifting unit), sweet tests up to Feb 2011 and sour tests up to June 2011, optimization of steam/gas ratio at the shifting unit, optimization of energy balance and to obtain real costs of CO2 capture and H2 production. Details and results of this R&D project and others are summarised in this extended abstract.
1. Introduction
ELCOGAS S.A. is a Spanish company established in 1992 and shared by European electrical companies and equipment suppliers. It operates the Puertollano 335 MWeISO IGCC demonstration power plant (Integrated Gasification Combined Cycle). This IGCC plant is the largest IGCC plant in the world using solid fuel in a single pressurised entrained flow gasifier, being in commercial operation since 1998 with synthetic gas. Its design fuel is a mixture 50:50 of coal (high content of ash) and pet-coke (high content of sulphur). The total power production up to Dec 2010 is 21,052GWh (mainly using syngas).
As a demonstration plant, ELCOGAS has obtained important achievements showing the potential of IGCC technology, including its advantages and disadvantage and identifying its main improvement and optimisation lines. Regarding this point, the ELCOGAS IGCC R&D activities are based on the opportunity that the IGCC technology offers related to fuel flexibility (test with different coals, pet-coke, biomass, wastes, …), multi-production (electricity, hydrogen, synthetic gasoline, biodiesel, …) and zero emissions production (reduction of emissions, CO2 capture …).
2. Towards zero emissions power plants Diversification of raw fuels The activities related to the diversification of raw fuels and products line are mainly linked to the project PIIBE (CENIT Programme -Spanish initiative). ELCOGAS coordinated the subproject about biodiesel from gasification by real co-gasification up to 10% of biomass and syngas characterisation (F-T process in laboratory). The selected biomass to be tested in the IGCC was olive waste (orujillo). Next table shows the average composition of the received orujillo and the common fuel (coal, pet coke and limestone) used in the power plant. Table 1. Orujillo and ELCOGAS common fuel average composition analysed by ELCOGAS laboratory
Parameter Moisture (%) Ash (%) Volatiles (%) Cfixed (%) LHV (kcal/kg) C (%) H (%) N (%) S (%) Cl- (mg/kg)
Orujillo average composition (Laboratory of ELCOGAS) 13.13 8.51 68.89 22.52 3,695 49.40 5.96 144 0.14 2,735
ELCOGAS common fuel. 2010 average composition 0.81 20.97 15.99 63.08 6,170 65.57 3.21 1.40 4.13 205.25
Next table shows the different tests that were carried out using orujillo as fuel, including the duration as well as the operating hours of them. Table 2. Battery of co-gasification tests in ELCOGAS (2007-2009)
Test Month/Year 2007- 2009 2008 March 2009 June 2009 Sept. 2009
Orujillo dosage ratio in weight% 1 -2 % 4% 6% 8% 10 % TOTAL
Orujillo tonnes (t)
Test duration (h)
1,572.84 652.14 395.86 383.90 656.68 3,661.42
800.3 154 64.4 46 62 1,126.7
Main conclusions from the co-gasification tests can be summarised as follows: ¾ The technical viability of co-gasification up to 10% by weight has been demonstrated ¾ The operation was within design ranges
¾ Biomass handling: - Orujillo should not be stored for a long time, since the biomass absorbs humidity - Orujillo goes easily stodgy if a large quantity is stored in the feed hopper before its consumption. ¾ Grinding system: during the 8% and 10% tests the increase of the mills consumption and the DP were detected. ¾ Gasifier load: no influence on the gasifier load arises from the orujillo co-gasification when 1%, 2%, 4% & 6% tests were carried out. More difficult to maintain it in 8%-10% tests due to the mills load. ¾ Clean gas: Orujillo co-gasification has no impact on the clean gas quality; its characterisation is similar to those relating to ELCOGAS common operation. ¾ Emissions: the 8% and 10% addition of orujillo seems to have an influence on the SO2 emissions (although orujillo has no content in sulphur), but always within limits. Currently, ELCOGAS is involved in other project called FECUNDUS (“Advanced concepts and process schemes for CO2 free fluidised and entrained bed co-gasification of coals” (RFCR-CT2010-00009) undertaking biomass selection as well as co-gasification tests with the selected biomasses. These tests will demonstration the technical viability of co-gasification and its influence in CO2 capture process.
Carbon Capture Pilot Plant description
Currently, the ELCOGAS largest investment in R&D is focused on carbon capture topic, covered by the National Singular and Strategic Projects Initiative, called PSE-CO2. The main milestone of the PSE-CO2 project has been the construction of a 14 MWth pilot plant fed by a 2% slipstream of the Puertollano IGCC power plant and able to capture 100 t/d of CO2, while producing 2 t/d of high purity H2 and using proven and commercial technology. Pilot plant aims are to demonstrate the feasibility of CO2 capture and H2 production in an IGCC that uses solid fossil fuels and wastes as main feedstock as well as to obtain economic data enough to scale it to the full Puertollano IGCC capacity in synthetic gas production.
The participants in the project are ELCOGAS (coordinator), University of Castilla-La Mancha (UCLM), CIEMAT (Spanish research centre) and INCAR (coal Spanish research centre) being the original budget 18.5 M€, currently is €14.7 million.
Both Spanish Government (Spanish Science & Research Minister) and Regional Government are funding the project through the Strategic and Singular Project Programme (PSE), being their contribution approximately 50% of the total budget.
The following figure shows a general view of the IGCC plant and the pilot plant. PRENFLO Gasifier
Coal preparation
CO2 capture pilot plant
ASU
Sulphur Recovery
Combined Cycle
Figure 1. CO2 capture & H2 co-production pilot plant: location
The pilot plant is fed with syngas –approximately 3,600 Nm3/h, dry base- from the IGCC power plant that can be desulphusired, i.e., it comes downstream the IGCC desulphurisation unit (called sweet gas) or the syngas can be fed sulphurised, i.e. upstream the desulphurisation unit (called sour gas).
The process of the 14 MWth pilot plant (see Figure 2) consists on a two step shifting unit to convert CO into CO2, a CO2 separation unit -based on absorption processes with amines (for both sweet/sour capture)- and a H2 purification unit (Pressure Swing Absorption unit), being all of them commercial processes. Auxiliary systems and full control have been integrated in the existing IGCC. All processes used in the pilot plant are being utilised at this moment by the chemical industry, so its innovation, summarizing, is their integration and use in the power industry.
Figure 2. Flow diagram of the CO2 capture & H2 production pilot plant
Main features of the pilot plant are summarised in next table:
Table 3. Main features of the pilot plant (design - SWEET CASE)
Characteristics of clean gas (pilot plant inlet) Flow
3,600 Nm3/h (d.b.)
Pressure:
19.8 bar
CO (% volume)
60.45
Temperature:
126 ºC
H2 (% volume)
21.95
Characteristics of main outlet streams (pilot plant outlet) CO2 flow
100 tonne/day
CO2 capture rate
> 90%
Pure H2 flow
2 tonne/day
Purity of pure H2
99.99 %
Raw H2 flow
5 tonne/day
Purity of raw H2
77.4 %
The pilot plant consists mainly of the three following steps: 1st step – CO conversion with water steam (shifting unit)
The aim of this phase is to modify the clean gas composition in order to move the carbon contained in the CO to CO2, while maximizing the H2 content.
The syngas from the existing IGCC is fed in a sulphur removal reactor (Zn oxide based adsorber,
used only in sweet conditions) and mixed with saturated medium pressure water steam. The mixture is then heated up to 310 ºC to guarantee water-gas shift reaction conditions in a first catalytic reactor (the catalyst is been supplied by Johnson-Matthey) where the main conversion from CO to CO2 and H2 is produced. Since the reaction is exothermic, the first reactor gas outlet temperature is quite high (480ºC). An intermediate cooling stage is required (where IP steam is produced), after which the gas is sent to a second reactor where the final conversion is achieved. The outlet high temperature with which the gas comes out of the second reactor is used in a regenerative heat exchanger to heat up the first reactor inlet gas. Following to this, there is the gas three step cooling process to 45ºC (low pressure steam generator, air cooler and water cooler). 2nd step – CO2 and H2 separation unit (CO2 capture)
The target of this step is to separate CO2 and hydrogen, obtaining a hydrogen enriched gas and a CO2 product stream.
To this purpose, an amine solution (concretely aMDEA - active Methyl DiEthanolAmine) is used to capture the CO2 contained in the gas coming from the shifting unit. The CO2 capture rate is higher than 90%. The CO2 captured is recycled back to the IGCC process. Downstream the CO2 absorption, the resulting gas is a hydrogen enriched flow called raw hydrogen (77.4% of purity). This stream is split into two: 40% is sent to the H2 purification unit and the rest, approximately 5t/d, is recycled back to the IGCC plant.
The aMDEA is regenerated by means of temperature increase and pressure reduction. During this process the CO2 is desorbed, obtaining the CO2 pure product stream. Once the aMDEA is regenerated, it is conditioned (pressure increase and temperature decrease) to be re-used. 3rd step – Hydrogen purification unit
Pure hydrogen (99.99% purity,) is obtained in this step from the raw hydrogen coming from the previous step.
For this purpose, 40% of the hydrogen enriched flow, called raw hydrogen, is purified by means of a PSA unit (Pressure Swing Adsorption) supplied by LINDE. Impurities such as CO2, CO, N2,
Ar are trapped in an adsorption multi-bed system whilst the hydrogen passes through it. This purification unit consists of four stages: adsorption, decompression, regeneration and compression. This process is carried out in four adsorption beds, consisting each of them in activated carbon, alumina and molecular sieve. The pure hydrogen obtained is recycled back to the IGCC, but can be used for different applications in the future.
The capacity of this unit is 2 tonnes of hydrogen per day with 99.99% of purity, being the nominal hydrogen recovery 70%. The tail gas generated in this step is also recycled back to the IGCC, but could be used as heat source in other processes due to its high hydrogen content (>50%).
3. Results and Discussion
The first tonne of CO2 was captured on 13th September 2010 (thus becoming the first installation of this kind in the world) and the commissioning was accomplished by October 2010. Characterization tests are being carried out since November 2010 until June 2011, covering the two different feeding syngas conditions.
As a brief description of the main learning in project phase can be mentioned: the finance delay due to funding calls, delay in the main equipment supply -more than 12-14 months-, the detailed engineering was conditioned by suppliers and the pilot plant construction was also delayed due to internal safety working permits since it is installed in an operating plant, finally lack of experimented personnel implied a delay on the commissioning.
The first battery tests using the sweet catalyst were undertaken from Oct 2010 to Feb 2011. The table below shows the composition of the main streams, comparing the expected values and the analysed at ELCOGAS laboratory: Table 4. Main results obtained from pilot plant in sweet operation (dry base). ELCOGAS laboratory
Main learning in the sweet characterisation tests has been the high reactivity achieved in the first reactor of the shifting unit, near to 95% CO-CO2 conversion, what would make possible to consider a shifting process with only one step using the sweet catalyst.
So, the figure below shows how the CO concentration varies with the temperature in the shifting unit, both theoretically (blue line) and practice (pink line). 20,0
20/01/11 17:24-28:35
CO (%)
15,0
10,0
Theoretical equilibrium curve Practice equilibrium curve
Practice Theoretically
5,0
0,0 300,0
350,0
400,0
450,0
500,0
550,0
600,0
Temperatura (ºC) Johnson Matthey
De alta Tª
Practice
Theoretically
Theoretically equilibrium curve
Practice equilibrium curve
Figure 3. Shifting unit conversion – diagram of temperature influence
In the figure above, the conversion in the first shifting reactor is represented, reaching near 95% in the practical case. It is seen that the temperature achieved at this point (511ºC) is higher than expected (498ºC). The next step is the cooling stage (straight line), where the outlet theoretical temperature (347 ºC) is not achieved in practice (383ºC). The final conversion obtained after
second reactor is 98%, near to the expected one. In addition, the equilibrium curves have been represented for both cases.
In addition, auxiliary consumption was lower than estimated in design, being the integration of O&M in the existing IGCC very easy, the rate of CO2 captured is 91.7% and the cold gas efficiency is 89.5%. Finally, the pilot plant has provided very useful data to develop some calculations about the CO2 capture costs. For these studies, captured CO2 costs have been estimated as a quotient between capture plant costs -CAPEX (investment costs) and OPEX (operational costs)- and captured CO2 tonnes.
CO 2 capture cost, €/t CO 2 =
CAPEX + OPEX Captured CO 2 tonnes
[eq.1]
This point of view is singular and in addition, data to calculate the cost of captured CO2 in the established conventional way for new plants can be offered. Table 5 shows the values that have been fixed to define the base case for calculations.
Table 5. Base case data
Variables
Expected life
Bank interest
Bank fee
Scale factor
Operating hours (IGCC mode)
Data
25 years
3.0 %
0.5 %
0.75
6,500 h
Average load factor
0.92
Net efficiency of Electricity power plant price with CO2 capture
40 €/MWh
Treated gas
33%
100%
According to these data, the first estimation cost of avoided CO2 is approximately 25-30 €/t for the existing IGCC (retrofitting) with sweet catalyst processes, which has been obtained from the pilot plant data.
Moreover, using these results as a reference, some sensitivity analysis have been carried out to get how some parameters influence on the captured CO2 cost. Figure 4 shows the influence of the operating hours and IGCC plant efficiency with CO2 capture on the captured CO2 cost:
60
3.000 h 3.500 h
CO2 cost (€/t CO2)
50
4.000 h 4.500 h
40
5.000 h
30
5.500 h 6.000 h
20
6.500 h 7.000 h
10 27
28
29
30
31
32
33
34
35
36
37
7.500 h
IGCC plant efficiency w ith CO2 capture (%)
Figure 4. Influence of the IGCC plant efficiency on the CO2 capture cost depending on the operating hours
It can be observed that the slope is the same in all cases, which means that the influence of the efficiency on the CO2 capture cost is independent of the operating hours. Small efficiency decreases cause big variations on the the CO2 capture cost.
For the second battery tests (sour capture), which will take place from May to June 2011, the sour catalyst will be tested, expecting to get final results by the end of July 2011. These final results will include comparison of the pilot plant’s behaviour under the two different operation conditions, optimization of steam/gas ratio at shifting unit for the correct operation of the plant, optimization of energy balance and real costs obtaining of CO2 capture and H2 co-production.
4. Conclusions
With the aforementioned and taking into account the results to be obtained from the pilot plant, ELCOGAS has the opportunity to contribute to the optimisation of IGCC technology subsequently to optimisation of the clean coal technologies. So, improvements and processes, which are being set out for the design of new plants, can be tested and developed even at commercial scale, leading to ultra-efficient and zero-emissions energy plants based on gasification of low cost fuels.
Once the PSE project is finished, ELCOGAS proposal is to use the pilot plant as an R&D platform in order to develop new projects related to these research areas: optimisation of catalysts for shift reaction (including tests on a variety of different catalysts), development and
demonstration of new processes for CO2-H2 separation, demonstration of processes for CO2 treatment and the improvement of integration between the CO2 capture facility and the IGCC power plant to increase efficiency
Acknowledgement
It is worth to have an special mention to both Spanish Government, through the Spanish Science & Research Ministry and Regional Government, because of their contribution in the project funding it through the Strategic and Singular Project Programme (PSE).
The effect of minerals on the moisture adsorption and desorption properties of South African fine coal S.M. du Preez, Q.P. Campbell* School of Chemical and Minerals Engineering, North-West University, 2520, Potchefstroom, South Africa. *Corresponding author. Tel.:+27(0)18 299 1993, E-mail address [email protected]
The aim of this research was to study the influence of mineral matter content and type on the equilibrium moisture levels of South African coals under different environmental conditions.
Some of the moisture sensitive conditions to which
beneficiated and dried coals are exposed to, include open air stock piles subjected to heavy rainfall and varying climatic conditions. Under these conditions coal can experience an increase in moisture content which can result in extensive handling problems, the plugging of belt conveyors or even moisture contract penalties. A series of adsorption experiments were conducted to correlate the equilibrium moisture levels and different physical properties of different coal samples.
A climate chamber
varying the temperature and humidity of the environment in which the coal particles were placed produced the equilibrium sorption data. The equilibrium data obtained were then correlated with coal properties such as porosity, coal rank and more importantly mineral matter content.
1
1. Introduction Coal plays a key role in the South African economy and is a commodity responsible for approximately 77% of its primary energy production
[1]
. In 2007, South Africa
produced 247 million tons of coal, utilizing 182 million tons locally and exporting 68 million tons [1]. The response of coal moisture to changing environmental moisture levels are influenced by factors such as clay content and the percentage fines. Coal containing a significant amount of clay will potentially hold more moisture. This is particularly important as the mineral matter found in South African coals is predominantly clay minerals, largely in the form of kaolinite and illite.
The influence of varying
conditions can become particularly prominent when large quantities of coal are stored or transported over great distances. Various publications were found in the literature concerning the subject of water adsorption on coals as a function of vapour pressure [2, 3, 4]. Due to the heterogeneous nature of coal, it follows that several individual coal properties will influence the adsorption behaviour of coal. The extent of moisture adsorption is influenced by coal rank, mineral matter content, porosity, and that the specific adsorption sites are determined by oxygen functional groups. This study was done to determine the influence of the mineral matter content on the moisture adsorption properties of a specific coal. There is almost always a significant amount of mineral matter intimately associated with the coal, also referred to as inherent mineral matter. Inherent mineral matter cannot be effectively removed by the beneficiation process and is ever present in even the cleanest coal products [5]. It is imperative to take into account the effect of this mineral matter when assessing the coal’s behaviour during handling and storage. To better understand the mechanism of moisture adsorption and the subsequent influence of the mineral matter content on this mechanism it is also vital to understand the physical and chemical characteristics of the specific coal. Isotherms like the Dubinin-Astakhov equations are regularly used to characterize moisture adsorption on
2
carbonaceous materials. The results obtained from these isotherms can be used to characterize the chemical and physical properties of the coal surfaces, contributing to the understanding of the moisture adsorption mechanism on coal.
2. Experimental section Coal selection and preparation: Five coal samples from three different collieries in the major coalfields of South Africa were chosen for this study. These were the Waterberg (coal A), Witbank (coal B) and Free State (coal C) regions
[6]
. Coal samples from a Witbank - Highveld
colliery was taken from three different sampling points: the export product (B1), middling fraction (B2), and the discard stream (B3).
This was done to ensure a
variation in ash content while keeping most other coal properties constant, since that it was identified early in the study that coal rank is expected to have a major influence on the moisture adsorption properties of a specific coal. The export sample will also give better insight into the moisture adsorption properties of coal that will be transported to over 500 kilometres to the Richards Bay Coal Terminal, since the Witbank coalfield is a major source of steam coal for the export market [7]. The coal samples were milled and dry sieved in +1mm - 2mm size fractions. The main purpose of the milling was to reduce the particle size of the coal samples before sieving. A jaw crusher was specifically used to ensure that the particles break along their inherent lines of weakness, maximising the exposure of the finely dispersed mineral matter present in the coal particle. Climate chamber A climate chamber, where the relative humidity and temperature of the atmosphere can be independently controlled, was used to test the moisture adsorption and desorption properties of five selected coal samples. A load cell was installed inside the climate chamber to record the increase in mass due to moisture adsorption and any decrease in mass due to moisture desorption under varying climatic conditions.
3
The experimental procedure that was followed, involved the selection of about 70 grams of the coal sample. The sample was dried overnight in a vacuum oven at 105○C to ensure that all the moisture is excluded from the sample before each experiment. The sample was then placed into the climate chamber at a relative humidity (RH) of 20% at the specified temperature. The system was then allowed to reach equilibrium which was indicated by a constant mass reading. The humidity was the increased to 40%, 60% and 80% in turn, again allowing each step to reach equilibrium. After a maximum of 80% RH was reached the humidity was decreased in the same sequence as for the adsorption steps, from 80% to 60%, 40% and 20% RH while the temperature was kept constant. The overall time to complete an experiment was in the order of 70 hours to ensure complete equilibrium was reached in each step. Directly after each experiment was completed the sample was weighed and placed immediately in a vacuum oven for several hours after which the sample was weighed once again to determine the final moisture content of the coal sample using ISO and SABS standards. This data was used to back-calculate mass of adsorbed moisture per unit mass of coal. The humidity sequences were investigated at temperatures of 15◦C, 20◦C and 28◦C. A schematic representation of the experimental setup can be seen in Figure 2.1.
4
Figure 2-1: Schematic representation of the experimental setup The load cell continuously recorded the increase or decrease in mass during each experiment. These values and the relative humidity and temperature readings were continuously recorded by the computer.
3. Results and Discussion The proximate analysis of coal evaluates the moisture, ash, volatile matter and fixed carbon content was determined according to standard test methods.
The results
obtained from the proximate analysis can be seen in Table 3-1. Table 3-1: Proximate analysis*
Inherent moisture 5
%
A (ROM)
B1 (Export)
B2 (Middlings)
B3 (Discard)
C (ROM)
0.6
2.6
2.2
1.2
3.9
%
50.1
14.9
31.6
63.5
34.1
Volatile matter %
13.8
26.2
20.6
18.2
20.6
Fixed carbon
%
35.5
56.3
45.6
17.1
41.4
Total sulphur
%
0.42
0.33
0.48
6.99
0.14
Ash content
*Air dried basis, percentages reported as wt. %
According to the characterization results in Table 3-1 all the coal samples except for B1 export could be regarded as high ash coals. The percentage ash in coal A and coal B3 discard is noticeably higher than for the other coals. Coal C contains a distinctly higher amount of inherent moisture when compared to the other four coal samples, which is typical of Free State coals. According to the petrographic analysis conducted, coal A can be classified as a medium rank B coal where Coals B and C are reported to be medium rank C coals. All the coal samples are rich in inertinite except for coal A which is rich in vitrinite. In Table 3-2 the ash composition analysis results can be viewed for coal A, B1 Export, B2 middlings and B3 discard, as well as coal sample C. Table 3-2: Ash composition (XRF) analysis of coal samples* Inorganic species Al2O3
%
SiO2
%
CaO
%
6
A (ROM)
B1 (Export)
B2 (Middlings)
B3 (Discard)
C (ROM)
24.61
28.93
30.35
21.36
32.19
66.65
49.46
56.65
42.58
51.50
0.97
8.86
3.55
4.49
7.87
Fe2O3
%
2.71
1.81
2.61
22.38
2.37
K2O
%
2.45
0.49
0.54
0.7
0.48
MgO
%
0.53
1.32
0.75
1.44
0.71
Na2O
%
0.32
0.04
0.04
0.05
0.38
V2O5
%
0.16
1.61
0.28
0.22
0.1
TiO2
%
0.95
2.02
2.13
1.18
1.80
SO3
%
0.68
4.21
2.89
6.04
1.85
Other
%
0.25
1.25
0.27
0.36
0.39
100
100
100
100
100
Total
*All percentages are reported as wt. %
The results in Table 3-2 indicates that the ash is rich in Al2O3 and SiO2 which corresponds to high levels of quartz and kaolinite present in the coal. It can also be observed that the bulk of the ash for each of the five coal samples consists of SiO2 which is derived from quarts and clay minerals [8, 9]. The second largest contributor to the ash is Al2O3 which in turn corresponds to elevated amounts of clay minerals present. Fe2O3 can be related to the pyrite present in the coal samples. Coals A and B2 are the two coal samples containing the highest amount of clay minerals whereas B3 discard the least. The elevated levels of pyrite in the B3 discard sample are expected since pyrite is undesired in the final export product this also explains the low levels of pyrite present in the B1 export sample. A typical experiment conducted in the climate chamber with varying humidity at a constant temperature of 28○C can be seen in Figure 3-1.
7
Figure 3-1: A typical adsorption desorption experiment for coal C at 28○C. In Figure 3.1 the relative humidity was varied according to the sequence described in the experimental section.
The mass gain and loss for each step in the relative
humidity was recorded. A constant mass gain or loss indicated that equilibrium was achieved.
A temperature of 28○C was chosen as this is representative of the
environmental conditions that the coal will be exposed to in South Africa during summer. From the data obtained in Figure 3.1 the overall percentage moisture can be calculated. The moisture adsorbed at the equilibrium points can then be calculated as well as the amount of moisture adsorbed per amount of dry sample. Various models such as the Langmuir and Dubinin Astakhov can be fitted to the data. From these models, conclusions can be made about the maximum amount of moisture adsorbed as well as the degree of hysteresis that is present for each of the coal samples. Additionally, the dynamic part of each experiment can also be investigated and conclusions regarding the influence of mineral matter on the kinetic moisture adsorption can be drawn. The overall percentage moisture adsorbed is about 6.5% for coal sample C. It is also interesting to note that some form of hysteresis is present in this sample seeing that
8
the moisture adsorbed is not equal to the moisture desorbed. The hysteresis can clearly be seen in the adsorption and desorption isotherm illustrated in Figure 3-2.
Figure 3-2: Adsorption desorption isotherm of coal C A similar experiment was conducted on the B2 middlings sample and the percentage moisture adsorbed for this sample was 3.5%.
9
4. Conclusions •
The difference in the percentage moisture adsorbed between the B2 middlings and sample C can be attributed to the different minerals present in these samples since they are from the same rank the only other variable is the difference in mineral content. This statement will be further investigated and additional conclusions will be made regarding maceral content of each coal sample.
•
QEMSCAN results of the five coal samples will give further information regarding the amount and type of minerals present.
•
The hysteresis observed during the experiment can be attributed to capillary condensation and the ink bottle effect present during adsorption - desorption process.
In the very fine pores of the coal particle the mechanism of
adsorption is pore-filling rather than surface coverage. A possible explanation for low pressure hysteresis can be formulated in terms of swelling of the particles during adsorption. The swelling distorts the coal structure and opens up cracks which were previously inaccessible to adsorbate molecules. This can cause distortion that is not perfectly elastic and molecules can become trapped in the coal structure which cannot be released during desorption unless the temperature is elevated [10].
10
References [1] DEPARTMENT OF MINERALS AND ENERGY. 2008. South Africa’s Mineral Industry 2007/2008. Directorate: Mineral Economics Report, Pretoria, South Africa. [2] MAHAJAN, O.P. and WALKER, P.L. 1971. Water adsorption on coals. Fuel, Vol. 50, Issue 3. [3] UNSWORTH, J.F., FOWLER, C.S. and JONES, L.F. 1989. Moisture in coal: Maceral effects on pore structure. Fuel, Vol. 68, Issue 1. [4] MCCUTCHEON, A.L. and BARTON, W.A. 1999.
Contribution of mineral
matter to water associated with bituminous coals. Energy & Fuels 13, (1). [5] WARD, R.C. 2002. Analysis and significance of mineral matter in coal seams. International Journal of Coal Geology, Vol. 50: 135-168, 2 May. [6] JEFRFEY, L.S. 2005. Characterization of the coal resources of South Africa. The Journal of the South African Institute of Mining and Metallurgy. [7] MANGENA, S.J. and DE KORTE, G.J. 2004. Thermal drying of fine and ultrafine coal. [8] VAN DYK, J.C., and KEYSER, M.J.
2005.
Characterization of inorganic
material in Secunda coal and the effect of washing on coal properties. The Journal of the South African Institute of Mining and Metallurgy. Jan. [9] SPEARS, D.A. 2000. Role of clay mineral in UK coal combustion. Journal of Applied Clay Science, Vol. 16: 87-95. [10] GREGG, S.J & SING, K.S.W. 1982. Adsorption Surface Area and Porosity. 2nd ed. New York: Academic Press. 303p.
11
Structural Changes and Possible Modes of Interaction in Bituminous Coal Fly Ash Due to Treatments with Aqueous Solutions (Acidic and Neutral)
Roy Nir Lieberman,1, Roy Nitzsche,3, Haim Cohen1,2 123-
Department of Biological Chemistry, Ariel University Center at Samaria, Ariel, 40700 Israel, phone: 00972-52- 4306878, fax: 00972-8-9200749, email: [email protected]; [email protected] Chemistry Department, Ben-Gurion University of the Negev, Beer Sheva, Israel email: [email protected] TU Bergakademie Freiberg, Fakultät 4, Institut für Energieverfahrenstechnik und Chemieingenieurwesen 09599 Freiberg, Germany.
Abstract Coal fly ash is produced in Israel via the combustion of Class F bituminous coals. The bulk of coal fly ashes produced in Israel stems from South African and Columbian coals thus these ashes were the subject of the present study. It has been reported that the flyash can be used as a scrubber and fixation reagent for acidic wastes. Furthermore, the scrubbed product can serve as a partial substitute to sand and cement in concrete while the bricks have proved to be strong enough according to the concrete standards. Three possible modes of interaction were observed: cation exchange, chemical bonding and electrostatic adsorption of very fine precipitate at the flyash surface. In order to have a better understanding of the fixation mechanism we have decided to treat the flyashes with acidic (0.1M HCl) and neutral (DDW) solutions, thus changing the properties of the surface of the flyash particles. Surface analysis of the treated and untreated fly ashes have demonstrated that the treated flyashes have changed appreciably its' interactions with transition metal ions (e.g Cd2+, Cu2+). Introduction The major part of the electricity in Israel ( >60% in 20081) is produced by 4 bituminous coal fired power stations (the coals are imported mainly from South Africa Columbia, Indonesia, Russia and Australia2). Due to Israel strict environmental regulations the coal imported to Israel contains low content of sulfur and phosphorus. Thus, the fly ash produced in Israel has a highly basic reaction when exposed to water (pH >10), mainly because of the high content of CaO (and is defined as Class F). This means that the fly ash can act as a natural pozzolan. The present coal consumption is cal 13 MTons which yields ~1.3 MTons of Coal flyash. The fly ash particle size is in the range 3-250 μm and contains different types of glassy spheres. The first type are the Cenospheres (image 1A) which are glass bubbles and the second type are the Plerospheres (image 1B) which are also hollow glass bubbles filled with small particles inside. Both types are composed mainly of aluminates and silicates (>70%w). The fly ash has a large surface area which gives it the possibility to act as a potential fixation reagent. The chemical composition of the South African and Columbian fly ashes (SAFA, COFA respectively) is given in Table 1.
Table 1: the Major components and Minor elements in the SAFA and COFA of the South African and Columbian fly ashes COFA*
SAFA*
54.4 20.8 1.05 6.18 4.65 2.05 0.12 0.05 0.75 0.13 9-7
40.9 31.4 1.75 3.05 8.35 2.45 0.05 0.02 1.95 0.35 5-4
Component %Weight SiO2 Al2O3 TiO3 Fe2O3 CaO MgO K2O Na2O P2O5 SO3 C
Element COFA** SAFA** Ppm Ag 9.5 13.6 As <10 < 10 Ba 1,150 2,350 Be 5.07 9.43 Cd 2 <2 Co 27 40 Cr 133 150 Cu 60 77 Mn 375 360 Ni 70 68
It has been found that many of the trace elements in the fly ash have different washing potential which also is dependent on the nature and charge of the species. The leaching potential of divalent trace elements like copper, lead, barium and manganese are low, whereas trace elements such as selenium, molybdenum, boron and chromium having a very high leaching potential (as oxyanions). The trace elements with high washing potential are of course an environmental hazard. Modes of interaction of flyash with trace elements During recent studies 3 modes of fixation mechanism have been suggested: (1) Cation exchange The surface of the flyash particles contains several anionic functional groups, mainly aluminates AlO2- and silicates -SiO3- , thus it will behave as a cation exchange material (as shown in Scheme 1) Scheme 1: Cation exchange mechanism between the metal ions and the anionic groups on the fly ash surface
As already mentioned, the anionic groups are mainly aluminates and silicates:
Typical metal cations can be Cd2+, Cu2+ and Sr2+. (2) Coordinative bonding Coordinative bonding is a chemical bonding which is formed between the cation and the nonbonding electrons of functional groups located at the surface of the fly ash particle. The cation behaves as a Lewis acid and the fly ash as a Lewis base. Probably, the bond is formed with oxide ions (as shown in scheme 2): Scheme 2 – coordination bonding between the cations of the solution and the non-bonding electrons of the fly ash)
This is a mechanism in which a Lewis base donating the lone pair of electrons forms the bond with the metal cation which is equivalent to formation of a complex where the surface serves as the ligand. Energetically, it is very strong bond, up to 150kj/mole. Electrostatic interactions of fine precipitate– When a solution interacts with the fly ash (usually basic solution) a precipitate is usually formed. In a previous work, it was found that there is a very effective interaction between fine precipitate and the fly ash, probably by electrostatic interaction between the surface of the fly ash (which contains negative charge of anions) to the fine precipitates. (as shown in scheme 3). Scheme 3 – electrostatic interactions between the fine precipitate and the fly ash
Fly ash as a neutralization and fixation reagent for acidic waste Recently it was shown that the fly ash can act as an effective neutralization and fixation regent for dangerous acidic waste10. Two wastes have been tested: (i) Acidic Organic waste of regeneration processes of used motor oil. This waste contains more than 10M of acid waste per liter. The waste also contains high concentration of heavy and toxic metals. (ii) Acidic waste from the phosphate industry which is a byproduct of phosphate rock treatment with sulfuric acid. Results of these tests have shown that the fly ash is a very effective as a neutralization and fixation reagent for these wastes. The scrubbed product is a grey aggregate (sand like) which fixate the heavy metals effectively. The fixation quality was tested by two types of washing processes: TCLP13114 and CALWET5. The results showed that the concentration of leach heavy metals is well under the drinking limit (DL). These results, might indicate that the flyash can also serve as an efficient fixation reagent to low activity radioactive wastes.6 We have decided to check the effect of treatment of coal fly ash with water and hydrochloric acid on the surface properties of the flyashes which probably affects it's scrubbing and fixation quality of wastes.
Experimental section Materials and Equipment All water used in the study were DDW with a resistance > 10 MOhms/cm. Chemicals: The chemicals used were of analytical grade: Sigma/Aldrich, Riedel de Haen, B.D.H and Merck. Gases:
Nitrogen (99.95%, dry) and air (dry) were supplied by MAXIMA Ltd.
Equipment 1) The following instruments were used throughout this study: pH-meter – of El-Hama company model Cyberscan510. Calibration of this device was made by three buffer solutions: pH4; pH7; pH10. ICP-OES – Induced Chemical Plasma Atomic Emission spectroscopy – from Varian, model 710-ES. Some of the samples were analyzed in ICPOES from Varian model VISTAPRO. Orbital Shakers from Cocono, model TS-400 Teflon filers for PET syringe from Tamar LTD, model FP-030/0.2μ. Analytical scales from Mettler were used. The SEM – Scanning Electron Microscope used was a Joel JSM model-6510LV Fly ashes South Africa (SAFA) an Columbian (COFA) flyashes were supplied by the Utilities.. The ash content in Soth African coal is– 13.9% (% weight); and in Columbian coal - 8.7%. The density of SAFA – 0.98gr/cm3 and that of COFA – 0.85gr/cm3 Experimental method Treatments of the fly ash The treatment of the fly ash in DDW or 0.1N HCl is as shown in scheme 4. Scheme 4 –treatment of the fly ash with DDW or 0.1M HCl
The experiment was carried in 6 steps (as shown in scheme 5): 1) Preparing a simulant solutions (Cerium, Strontium, Cesium, Cadmium or Copper)) in a known volume 2) Mixing the solution with fly ash in a PET bottle. 3) Mixing the mixture in the orbital shakers at 250rpm for different periods of time. 4) Taking a sample and filter it. 5) Drying of the sample was carried out for 2 days under air in the hood at room temperature. 6) Measuring the pH and acidify the sample to prevent precipition. 7) Measuring the concentration in the ICP-OES Scheme 5 – work method
Results and Discussion SEM images (Scanning Electron Microscope) and weight loss SEM results show that the South African fly ash SAFA is different than the Columbian ash COFA (Figure 3). It is suggested that the source is the much higher concentration of lime which is relatively volatile thus it is present mainly at the outer surface of the ash particle. Indeed the pH of a DDW sample which is in contact with the fly ash (S/L ratio 1/20) is much more basic for SAFA, pH12.5 compared to COFA pH10. Figure 3 – SEM images taken for the SAFA (3a) and COFA (3b)
Treatment with DDW: During the washing process with the DDW, the major leaching is of the lime which is soluble and forms calcium hydroxide in solution, but also some compounds of sodium, magnesium and potassium are also soluble. The lime (CaO) content in SAFA and COFA is 8.35w% and 4.65w% respectively. However the weight loss during the DDW treatment of SAFA and COFA is only 3.45% and 2.65% respectively. This observation means that there is only partial dissolution of the lime, ~41% (SAFA) and ~57% (COFA) during the leaching process in DDW. In other words at least half of the lime in the fly ash particle is trapped in the fly ash structure and is not washed to the water. In the SEM images of the DDW washed ashes (Figure 4) it is clearly seen that the COFA particles are much smoother than the SAFA particles. Probably the dissolution of lime at the surface of the COFA by DDW is much more efficient than the case in the SAFA particles Figure 4 - SEM images taken for the SAFA (4a) and COFA (4b) treated with DDW
Treatment with 0.1N Hydrochloric acid: Treatment with 0.1M HCL solution of the fly ashes results in a much higher weight loss: ~16w% for the SAFA and ~13.7w% for the COFA. These weight losses are much higher than the lime content. If we assume that the iron oxides lime sulfates and phosphates are leached out, than one expects leaching out (table 1- Fly ashes composition) of 13.9% and 11.9% from SAFA and COFA respectively. The fact that it is cal 2% higher (within experimental error) means that indeed the acid penetrates to the inner parts of the particles and leach out all the lime and also there is some destruction of the matrix of the fly ash particles. Also, the SEM images (Figure 5) show again that the COFA particles are much smoother compared to the SAFA particles.
Figure 5 - SEM images taken for the SAFA (5a) and COFA (5b) treated with Hydrochloric acid Surface composition of SAFA and COFA- The major compounds of South African Fly Ash and Columbian Fly Ash are Al-,Ca-, Fe- and Si-oxides. The composition of the SAFA and COFA before treatment and after DDW and 0.1M HCl treatment is given (Figure 6). It is clear that DDW-treated and HCl-treated fly ash is very similar, however after the treatment with HCl, the percentage of Ca decreased by approximately 30-50 %. This is, probably, the result of the high solubility of Ca in aqueous solutions and the reaction between HCl and CaO causing dissolution of CaO.
Figure 6 – SEM results showing the % of elements on the fly ash surface (treated and untreated)
pH results – We have checked the amount of leached acid/base of untreated and treated SAFA and COFA to DDW (S/L 1/20 as a function of time (Figure 7). It is observed that SAFA and COFA treated with 0.1 M HCl have the lowest pH values. This can be attributed to the fact that 0.1M HCl treatment did neutralize most of the basic groups present in the fly ash particle. The DDW treatment of the fly ashes did neutralize only partially the basic functional groups in the ash ththus the pH is lowered compared to the untreated ashes. In all cases the resulting pH is higher for the SAFA compared to the COFA. This stems from the higher lime content in the SAFA (8.35w% compared to 4.65w% in the COFA). Furthermore it is definite that the leaching due to treatment is terminated within 1 hour as the pH is not changed for longer periods of shaking. Figure 7 –pH results of the SAFA and COFA before and after treatment as function of shaking period (S/L 1/20)
Conclusions (i) The treatment of the fly ashes particles with distilled water results mainly in partial dissolution of the lime content (ii)
The treatment of the fly ashes particles with acidic 0.1M HCl results mainly in total dissolution of the lime, iron phosphorous and sulfate content
(iii)
The leaching during treatment is relatively fast and terminates in cal 1 hour.
(iv)
The treatment results in formation of a cation exchange like moiety.
Acknowledgements We would like to thank the Israel Coal Ash Administration for funding the research, to Ariel University center and Dr. Hanan Teller for supplying us the equipment and work space. Also acknowledgement to the Israel Geological Survey for chemical analysis and Ms. Natali Litbek for the SEM analysis. References [1] Ravina, D.; Lulav, O.: Coal Ash in Israel: AshTech Conference in Birmingham: 2006. [2] Internet site: http://www.coal-ash.co.il. [3] H. A. Foner, T.L. Robl, J.C. Hower and M. U. Graham, Characterization of Fly Ash from Israel with Reference to Its Possiblre Utilization, FUEL, Vol. 78, 215-223, 1999. [4] U.S.-Environmental Protection Agency (EPA): Toxicity characteristic leaching procedure (TCLP). Part 261, Appendix II-method 1311, 40 CFR Ch. 1,1990 [5] California Waste Extraction Test (CAL WET), Environmental Health Standards - Hazardous Waste, 66261.126 Appendix II.
[6] T. Nokumura, The Material Flow of Radioactive Cesium 137 in the USA, 2000. [7] Internet site: www.epa.gov/rpdweb00/docs/source-nagement/csfinallongtakeshi.pdf. [8] Internet site: Strontium 90,k28 www.absoluteastronomy.com/topics/Strontium-90. [9] Cohen, H.: Fly Ash: A Potential Excellent Scrubber for Acidic Wastes in Israel: International Ash Utilization Symposium, Center for Applied Energy Research, University of Kentucky, paper #51; 2003. [10] Eli Lederman, "Coal Fly Ash as a Chemical Reagent for Scrubbing Industrial Acidic Waste", Ben Gurion University of the Negev, Israel, 2008.
Oviedo ICCS&T 2011. Extended Abstract
Prediction of Selective Trace Element Emissions During Oxy-CFB Combustion of Victorian Brown Coals Bithi Roy, Wei Lit Choo, Sankar Bhattacharya* Department of Chemical Engineering, Monash University, Clayton Campus, Wellington Road, Victoria 3800, Australia *E-mail address:[email protected] Abstract Victorian brown coal is an abundant resource with an estimated reserve of over 500 years at the current rate of consumption. Power generation from brown coals, however, results in higher CO2 emissions compared to that from high-rank coals. In reducing CO2 emissions, power generation using oxy-fuel based circulating fluidized bed (Oxy-CFB) combustion has emerged as a promising technology when utilising low-rank coals. Trace element emissions during Oxy-CFB combustion is an important issue to consider as the release of these elements in the gaseous effluent could present a serious health and environmental issue. Among the trace elements present in coal, Hg, Pb, Se, As and Cr6+ causes the greatest concern. Although some studies based on thermodynamic equilibrium calculations of trace element emissions during combustion of (mainly) bituminous coal have been carried out, the emission of trace elements during oxy-CFB combustion of Victorian brown coals is still unexplored. The FactSage thermodynamic package was used in this study to predict the concentration of trace elements (Cr, As, Se, Hg and Pb) species emitted during oxyCFB combustion of Victorian brown coal at different temperatures (800 - 1400oC). It was found that Ar, Se, Pb and Hg exist predominantly in the gas phase over the temperature range investigated. Chromium, however, was found to remain mainly as solid chromite in the reactor, over the entire temperature range. Toxic CrO3 was predicted in the gas phase, and in large concentrations above 1300oC. These results will be verified with experimental results in the future. 1. Introduction The growing energy demand for fossil fuels plays a key role in the upward trend in CO2 emissions. Hence, it is a formidable task to achieve a net reduction in CO2 emissions. CO2 capture and storage (CCS) can achieve a substantial reduction in greenhouse CO2 emissions from coal-fired plants. Oxy-fuel combustion has emerged as a promising
1
Oviedo ICCS&T 2011. Extended Abstract
technology for power generation from coal as it produces a concentrated CO2 stream which allows for easier carbon capture and storage [1]. On the other hand, circulating fluidized bed (CFB) combustion has already been established as a boiler technology suitable for utility-scale power generation for all types of fuels, predominantly low-rank coals. Advantages of the CFB reactor configuration include a homogeneous temperature distribution and the possibility of reducing SO2 emissions via addition of sorbents. As of 2008, almost 39000 MWe air-CFB steam generators are in operation or under construction [2].
Fig. 1. Simplified flowsheet of Oxy-CFB combustion
Combining these two technologies, oxy-fuel based circulating fluidized bed (Oxy-CFB) combustion (see figure 1) is emerging as a promising technique for coal combustion with CO2 capture. It combines the advantages of both oxy-fuel and CFB technologies and has the additional benefits of reduced combustor size, reduced operating costs and easier CO2 capture [3]. At present, Foster Wheeler, CIUDEN and to a lesser extent Alstom are the front-runners of this technology. The issue of trace element emissions during coal combustion has important implications on the operation of oxy-CFB combustors. The fate of trace elements (TEs), which are present in coal at very low concentrations (below 100 ppm) [4], are an important consideration as the inclusion of excessive amounts of these elements in the gas is harmful to the environment with additional implications for CO2 transport and storage. The combustion of coals containing only several parts per million of TEs could result in the release of several tons of pollutants into the environment [4]. Of the TEs present in coal, Hg, Pb, Se, As and Cr6+ are of greatest environmental and health concern [4]. In coal combustion, many TEs vaporise and then condense either homogeneously to form submicron ash particles or heterogeneously to adsorb on the surface of fine fly ash
2
Oviedo ICCS&T 2011. Extended Abstract
particles [4, 5]. However, significant amounts of a number of TEs leave the stack in gaseous forms, resulting in direct emission [5]. The emission and speciation of TEs depend on several factors including mineral contents, their distribution in the coal utilised, and the combustor temperature [4]. It is important to understand the effect of temperature on TE speciation as prediction of the possible TE species formed during coal combustion are important to control their emissions [5, 6]. Thermodynamic equilibrium calculations can be used to predict the possible phases formed and the concentrations of the compounds present in the equilibrium state. As experiments are costly to run, a pre-experiment forecast of the possible formation of these compounds may reduce the number of experiments required. Thus, many researchers have used thermodynamic equilibrium approaches to study the trace element distributions in coal combustion [5-12]. While many trace element studies have been carried out on (mainly) bituminous coals, the study of TE emissions from sub-bituminous coals, in particular Victorian brown coal, during oxy-CFB combustion has never been explored. Knowledge of trace element reactions and behaviour during oxy-CFB combustion is important for the control of pollutants and emissions. Hence, this study aims to predict the distribution of trace element (As, Cr, Hg, Pb and Se) species formed during the oxy-CFB combustion of Victorian brown coal using the FactSage thermodynamic modelling software. These predicted trace element emissions from oxy-CFB combustion will also be compared with that from air combustion. These results may help in the selection of suitable operating conditions such that the emitted trace elements are controlled within permissible limits. 2. Computational data and procedure The FactSage 6.2 software was employed to predict trace element speciation during combustion of Victorian brown coal under different conditions. Calculations have been carried out using equilibrium mode, which employ a Gibb’s energy minimization algorithm on the thermochemical function of ChemSage. Composition of one Victorian brown coal (Loy Yang) was used as input in this calculation. The trace elements considered in the study were As, Cr, Hg, Pb and Se. The composition of the coal samples is shown in table 1. The input used in the modelling work was 1 kg of dry coal.
3
Oviedo ICCS&T 2011. Extended Abstract
Table 1. Composition of the coal considered in the thermodynamic modelling Ultimate analysis (wt.% dry basis) Carbon 65.00
Minerals and inorganic (wt.% dry basis) SiO2 56.50
Trace elements (mg/kg dry basis) Cr 5-10
Hydrogen
4.60
Al2O3
20.50
Pb
5-10
Nitrogen
0.72
Fe2O3
4.60
As
2-4
Sulphur
0.50
TiO2
1.50
Se
2-4
Oxygen
25.37
K2O
1.30
Hg
0.1
Chlorine
0.11
MgO
3.60
Mo
10-20
Ash
3.70
Na2O
4.70
B
10-20
CaO
1.60
SO3
5.00
The oxy-fuel atmosphere used during combustion consisted of a mixed gas containing CO2, 2% excess O2, steam and N2. In case of air combustion, 6% excess oxygen assumed to be present in the system. The inputs were processed over the temperature range 800 to 1400oC, at 100oC intervals, and at atmospheric pressure. A key assumption to this modelling is that in the combustor, all phases are well mixed and reach thermal and chemical equilibrium. This temperature range covers both CFB and pulverised fuelfired (PF) systems, including any temperature excursion. 3.
Result and discussion
3.1.
Distribution of trace elements
The different species of trace elements considered in the analysis are listed in table 2. Table 2. List of species containing the five trace elements considered in this work1 Trace Element Cr As Se Hg Pb
1
Species CrO2 (g), CrS (g), CrO (g), Cr (g), CrO3 (g), CrO2Cl2 (g), CrN (g), FeCr2O4 (s). AsSe (g), AsN, AsS (g), As2 (g), AsH3 (g), As (g), As3 (g), As4 (g), As2Se2 (g), AsCl3 (g), As4O6 (g), As4Se3 (g), As4Se4 (g). H2Se (g), AsSe (g), SeS (g), Se (g), Se2 (g), SeO (g), PbSe (g), HgSe (g), SeO2 (g), As2Se2 (g), NSe (g), CSe (g), Se3 (g), SeCl2 (g), CSe2 (g), Pb2Se2 (g), (g), Se HgS (g), As (g),(g), Se HgO (g), BSe AlSe (g),2 (g), Hg (g),SeHgSe (g), (g), HgCl (g), (g), HgHSiSe (g),(g), Hg2As (g),SeHgCl (g), Hg (CH3)2 (g). PbS (g), Pb(g), PbCl (g), PbSe (g), PbO (g), PbCl2 (g), PbH (g), Pb2 (g), Pb2Se2 (g), PbCl4 (g).
g is the gaseous state; s is the solid state.
4
Oviedo ICCS&T 2011. Extended Abstract
The quantity of the major trace element species present in the vapour phase at equilibrium, within the 800-1400°C temperature range, is shown in table 3. For the sake of clarity, only those components present at quantity in excess of 1 x 10-12 moles are shown. Table 3. Major trace element species2 produced from the combustion of coal in the temperature range 800 to 1400°C. 1.9 Air nmolcombustion - 0.7 µmol
- 77 nmol Oxy-CFB0.19 combustion
47.5 - 96 µmol
47 - 96 µmol
CrO2 (g)
<1 pmol - 1 µmol
<1 pmol - 2 µmol
CrO (g)
<1 pmol - 6 nmol
<1 pmol - 14 nmol
<1 pmol - 0.16 µmol
<1 pmol - 0.3 µmol
<1 - 40 pmol
<1 - 78 pmol
0.12 - 47.6 µmol
0.23 - 47 µmol
AsN (g)
1.8 - 43 µmol
0.35 - 13.6 µmol
AsS (g)
0.3 nmol - 0.5 µmol
0.6 nmol - 0.7 µmol
As2 (g)
98 nmol - 1.2 µmol
0.66 - 6.4 µmol
As (g)
4.5 nmol - 9.6 µmol
11 nmol - 35 µmol
H2Se (g)
16 nmol - 1 µmol
44 nmol - 2.4 µmol
AsSe (g)
0.12 - 48 µmol
0.23 - 47 µmol
0.35 nmol - 0.9 µmol
0.2 nmol - 0.6 µmol
7 nmol - 10 µmol
6 nmol - 10 µmol
Se2 (g)
1.9 nmol - 0.77 µmol
1 nmol - 0.2 µmol
SeO (g)
0.88 nmol - 36 µmol
0.6 nmol - 35 µmol
PbSe (g)
4 nmol - 5.7 µmol
2.4 nmol - 3 µmol
SeO2 (g)
3 pmol - 7.4 µmol
1.7 pmol - 6.8 µmol
0.5 µmol
0.5 µmol
1 - 62 pmol
0.6 - 23 pmol
0.36 nmol - 39 µmol
0.2 nmol - 41 µmol
1.7 - 29 µmol
2.5 - 37 µmol
PbCl (g)
0.36 - 10 µmol
0.1 - 4 µmol
PbSe (g)
4 nmol - 5.7 µmol
2.4 nmol - 3 µmol
PbO (g)
24 nmol - 32 µmol
28 nmol - 31.5 µmol
Trace elements
PbCl Species 2 (g)
Cr
FeCr2O4 (s)
CrO3 (g) CrO2Cl2 (g) As
Se
AsSe (g)
SeS (g) Se (g)
Hg
Hg (g) HgSe (g)
Pb
PbS (g) Pb (g)
2
g refers to the gaseous state; s refers to the solid state.
5
Oviedo ICCS&T 2011. Extended Abstract
3.1.1. Chromium Table 1 shows the Loy Yang coal contains substantial quantities of chromium (5-10 mg/kg of coal). Figure 2 shows the quantity of chromium present in the vapour phase, as a function of temperature, during the combustion of these coals. The FactSage analysis carried out indicates that most of the chromium in the system would exist in the solid phase. Over the entire temperature range, almost all (>90%) of the chromium exists in the solid phase as chromite (FeCr2O4).
Fig. 2. Equilibrium distribution of chromium in the vapour phase during Oxy-CFB combustion.
It can be seen that above 1300oC, small amounts of chromite is vaporised to form predominantly CrO3, and small amounts of CrO2 and CrO2Cl2. The amount of chromium present in the vapour phase at equilibrium is predicted to be essentially the same for coals containing 5 and 10 mg of chromium per kg of coal. This suggests that the chromium concentration in the vapour phase is the result of vapour-solid equilibrium between the solid and gaseous compounds; it is expected that this equilibrium is dependent on temperature alone and not on the initial quantity of chromium present. As most of the gaseous chromium exists in the form of a toxic Cr(VI) compound, further investigation into the fate of the CrO3 is required to determine if effluent gas treatment is required when operating at temperatures above 1300oC, particularly for oxy-pf system. 3.1.2. Arsenic Arsenic is one of the more volatile trace elements present in coal. It is a toxic substance which affects the gastrointestinal tract, circulatory system, liver, kidney and skin. The toxicity of arsenic is dependent on the oxidation state of the arsenic species present As3+ is significantly more toxic than As5+ [13]. The equilibrium distribution of As at different temperatures is shown in figure 3. As Loy Yang coal could contain between 2
6
Oviedo ICCS&T 2011. Extended Abstract
to 4 mg arsenic per kg of coal, the solid and dotted lines represent the quantity ranges for arsenic species present in the vapour phase.
Fig. 3. Equilibrium distribution of arsenic in the vapour phase during Oxy-CFB combustion. The solid and dotted lines represent the result obtained for coal containing 4 mg As/kg coal and 2 mg As/kg coal, respectively.
The results indicate that at temperatures below 1200oC, AsSe is the main arsenic containing species in the vapour phase. However, at higher temperatures, elemental arsenic becomes the dominant form of arsenic in the gas phase. 3.1.3. Selenium As in the case of arsenic, the toxicity of selenium compounds depend on its oxidation state [13]. Se4+ is reported to be more toxic than Se6+. The quantity of selenium compounds present in the gas phase are presented in figure 4. Figure 4a shows the major selenium containing species present, whereas the species shown in figure 4b are present at lower quantities.
Fig. 4. Equilibrium distribution of selenium in the vapour phase during Oxy-CFB combustion. The solid and dotted lines represent the result obtained for coal containing 4 mg Se/kg coal and 2 mg Se/kg coal, respectively.
As shown in figure 3, the dominant species is AsSe at temperatures below 1200oC. However, at higher temperatures, substantial quantities of SeO and Se are formed. Figure 4b shows that small amounts of selenium dioxide (SeO2) are also formed at
7
Oviedo ICCS&T 2011. Extended Abstract
temperatures above 1300oC. This form of selenium is known to be toxic and its present could necessitate downstream gas treatment. 3.1.4. Mercury Mercury is considered to be one of the most volatile TEs in coal. Discharge of mercury in the gaseous effluent creates problem for both health and environment. Even when utilising coals with low mercury content, capture and removal of mercury by air pollution control systems is problematic [5].
Fig. 5. Equilibrium distribution of Hg (g) during Oxy-CFB combustion.
Figure 5 shows the equilibrium distribution of Hg as a function of CFB temperature. The results indicate that elemental mercury is the dominant mercury species over the entire temperature range. This could have important implications for downstream processes due to the corrosive nature of mercury [14]. At 1400oC, there is a slight drop (0.016%) in the amount of elemental mercury present due to the formation of HgO. 3.1.5. Lead The equilibrium distribution of Pb as a function of temperature is shown in figure 6.
Fig. 6. Equilibrium distribution of lead in the vapour phase during Oxy-CFB combustion. The solid and dotted lines represent the result obtained for coal containing 10 mg Pb/kg coal and 5 mg Pb/kg coal, respectively.
8
Oviedo ICCS&T 2011. Extended Abstract
The main lead species are PbS, Pb, PbO, PbCl and PbSe. Figure 6a shows that the dominant Pb species changes based on the temperature. At low temperatures (<900oC), lead sulphide is the dominant species. At very high temperatures (>1300oC), lead oxide becomes the dominant species. At temperatures between these extremes, most of the lead present in the gas phase exists as elemental lead. However, regardless of the lead species present, cold gas cleaning systems are very effective at removing traces of lead from the effluent gas [15]. 4. Conclusions This FactSage thermodynamic package was used to predict for the first time the behaviour of trace elements (Cr, As, Se, Hg and Pb) during Oxy-CFB combustion of Victorian brown coal a wide range of temperatures (800 - 1400oC). The dominant arsenic, selenium and lead species were found to be very dependent on operating temperature. Regardless, the results indicate that essentially 100% of these elements would exist in the gas phase, which could require gas clean-up prior to discharge. Almost all of the mercury also exists in the gas phase as elemental mercury, regardless of temperature. The toxic and corrosive nature of mercury means that a mercury removal step is required at the outlet of the furnace. Chromium was found to remain mainly as solid chromite in the reactor over the entire temperature range. Toxic CrO3 was predicted in the gas phase, and in large concentrations above 1300oC. These findings have important implication for setting parameters for oxy-CFB operation involving Victorian brown coal. Further modelling using a number of other brown coals and related experiments are in progress as part of a wider study. Acknowledgement The authors would like to acknowledge the Australian Department of Resource, Energy and Tourism (DRET), for their financial support of this work. References [1]
[2]
T. Wall, Y. Liu, C. Spero, L. Elliott, S. Khare, R. Rathnam, F. Zeenathal, B. Moghtaderi, B. Buhre, C. Sheng, R. Gupta, T. Yamada, K. Makino, J. Yu. An overview on oxyfuel coal combustion--State of the art research and technology development. Chemical Engineering Research and Design 2009;87:1003-1016. Platts. World Electric Power Plants Database. 2011.
9
Oviedo ICCS&T 2011. Extended Abstract
[3]
[4]
[5] [6] [7]
[8]
[9]
[10]
[11] [12] [13]
[14]
[15]
L. Duan, C. Zhao, W. Zhou, C. Qu, X. Chen. Effects of operation parameters on NO emission in an oxy-fired CFB combustor. Fuel Processing Technology 2011;92:379-384. F. Vejahati, Z. Xu, R. Gupta. Trace elements in coal: Associations with coal and minerals and their behavior during coal utilization - A review. Fuel 2010;89:904-911. R. Yan, D. Gauthier, G. Flamant. Possible interactions between As, Se, and Hg during coal combustion. Combustion and Flame 2000;120:49-60. R. Yan, D. Gauthier, G. Flamant. Partitioning of trace elements in the flue gas from coal combustion. Combustion and Flame 2001;125:942-954. K. Lundholm, A. Nordin, R. Backman. Trace element speciation in combustion processes--Review and compilations of thermodynamic data. Fuel Processing Technology 2007;88:1061-1070. J. R. Bunt, F. B. Waanders. Trace element behaviour in the Sasol-Lurgi MK IV FBDB gasifier. Part 1 - The volatile elements: Hg, As, Se, Cd and Pb. Fuel 2008;87:2374-2387. J. R. Bunt, F. B. Waanders. Trace element behaviour in the Sasol-Lurgi MK IV FBDB gasifier. Part 2 - The semi-volatile elements: Cu, Mo, Ni and Zn. Fuel 2009;88:961-969. J. R. Bunt, F. B. Waanders. Trace element behaviour in the Sasol-Lurgi fixedbed dry-bottom gasifier. Part 3 - The non-volatile elements: Ba, Co, Cr, Mn, and V. Fuel 2010;89:537-548. L. Zheng, E. Furimsky. Assessment of coal combustion in O2+CO2 by equilibrium calculations. Fuel Processing Technology 2003;81:23-34. E. Furimsky. Characterization of trace element emissions from coal combustion by equilibrium calculations. Fuel Processing Technology 2000;63:29-44. P. F. Nelson, P. Shah, V. Strezov, B. Halliburton, J. N. Carras. Environmental impacts of coal combustion: A risk approach to assessment of emissions. Fuel 2010;89:810-816. N. Imada, H. Kikkawa, K. Kobayashi, N. Oda. Study of mercury behaviour in flue gas of oxy-fuel combustion. In: The 35th International Technical Conference on Clean Coal & Fuel Systems, Clearwater, Florida, USA, 2010. S. Liu, Y. Wang, L. Yu, J. Oakey. Thermodynamic equilibrium study of trace element transformation during underground coal gasification. Fuel Processing Technology 2006;87:209-215.
10
Oviedo ICCS&T2011.Extended Abstract
Evolution of Victorian Brown Coal Char Structure during the Gasification in CO2 and Steam
Hui-Ling Tay, Shiro Kajitani and Chun-Zhu Li Fuels and Energy Technology Institute, Curtin University of Technology, GPO Box U1987, Perth, WA 6845, Australia. [email protected] Abstract Char structure is an important factor influencing its reactivity. Tracing the changes in char structure during gasification is indispensable to understanding the gasification mechanisms and pathways. This study investigated the changes in char structure during the gasification of Victorian brown coal with pure CO2, 15% H2O balanced with argon or 15% H2O balanced with CO2 at 800 oC in a novel fluidised-bed/fixed-bed reactor. The structural features of the resulting chars were examined using FT-Raman spectroscopy. Our results indicate that the reaction pathways for the char-CO2 and char-H2O reactions are different at 800 oC.
1. Introduction Char structure has been shown to affect its reactivity [1] during gasification. We have recently developed [2] an FT-Raman spectroscopic technique to investigate the changes in char structure. We have shown [3] that the presence of steam drastically decreased the char reactivity. However, the gasification in the mixed atmospheres containing both oxidising and reducing gasifying agents is not well understood. In this study, a Victorian brown coal was gasified in different gasifying atmospheres of steam, CO2 or CO2/steam mixture. The purpose of this study was to investigate the char structural changes in these atmospheres to understand the similarities and dissimilarities in the reaction mechanisms of char-CO2 and char-H2O reactions.
2. Experimental Section Victorian (Loy Yang) brown coal was used in this study. The coal was partially dried at low temperature (< 35 oC), before being crushed and sieved to get the particle sizes between 63 and 150 µm. The coal had a carbon content of 70.4 wt% and an ash yield of 1.1 wt%. Gasification experiments were carried out in a fluidised-bed/fixed-bed reactor [4]. Briefly, two frits were installed in the reactor body. The bottom frit was the support for the sand bed 1
Oviedo ICCS&T2011.Extended Abstract
as well as the gas distributor for the fluidising gas. The gasifying agents were 15% H2O balanced with argon, pure CO2 or 15% H2O balanced with CO2. Water was fed directly into the reactor using a HPLC pump to generate the required steam. At the required reaction temperature of 800 oC, coal was entrained into the reactor body via a water-cooled probe to be heated up rapidly for in situ gasification. The nominal feeding time for 1.5 g (accurately weighed) coal in each experiment was 18 min. The “holding time” commenced after the required amount of coal had been fed into the reactor. Most of char particles would be elutriated out of the sand bed and be held by the top frit. After the required holding time, the gasification was quenched by lifting the reactor out of the furnace. The FT-Raman spectra of chars were acquired [2] using a Perkin-Elmer Spectrum GX FT-IR/Raman spectrometer. The Raman spectra were deconvoluted into 10 Gaussian bands following the procedure outline previously [2].
3. Results and Discussion Figure 1 shows the changes in char yield with holding time in three different gasifying atmospheres at 800 oC. As expected, the gasification rate was the slowest in CO2 and the fastest in the H2O + CO2 mixture. 35 30
Char Yield, wt%(db)
25 20 15 10 5 0
0
5
10
15 20 25 Holding Time, min
30
35
40
15% H2O balanced with argon Pure CO2 15% H2O balanced with CO2
Figure 1 Char yields as a function of holding time and gasifying agent during gasification at 800 oC. 2
Oviedo ICCS&T2011.Extended Abstract
The Raman spectroscopic results are presented in Figures 2 and 3 as a function of char yield. The Gr, Vl and Vr are grouped together to represent the aromatic ring systems with less than 6 fused rings, whereas the D band represents the larger aromatic ring system (> 6 rings) as found in the amorphous structure. Therefore, the (Gr+Vl+Vr)/D ratio (Figure 2) reflects semi-quantitatively the relative ratio between the small and the large aromatic ring systems in char. The smaller aromatic ring systems in the char decreased with decreasing char yield for all the atmospheres investigated. The decrement was smaller for the chars produced in CO2 than for those produced in steam and H2O + CO2; the latter being similar. The steam has an almost decisive effect on the relative ratio of small and large aromatic ring systems in char.
1.6
(Gr+Vl+Vr)/D
1.4
1.2
1.0
0.8
0
5
10 15 20 25 30 Char Yield, wt% (db) 15% H2O balanced with argon
35
Pure CO2 15% H2O balanced with CO2
Figure 2 Raman band area ratio (Gr+Vl+Vr)/D of chars during gasification at 800 oC.
The Raman data in Figure 2 reflect the differences in the reaction mechanisms. The reaction between CO2 and char could produce intermediate such as C(O) [5]. On the other hand, steam would dissociate into H radicals and OH radicals on the char surface [6]. These radicals could penetrate into the char matrix to activate and grow the ring structures. It is speculated that the intermediates (e.g. H radicals) derived from steam have a higher chance to penetrate through the char matrix than those from CO2 and hence the Raman ratio decreased more rapidly for the chars produced in steam than those produced in CO2. 3
Oviedo ICCS&T2011.Extended Abstract
The total Raman areas (Figure 3) were similar for the chars produced from the gasification in steam and H2O + CO2. The presence of O-containing structures in char tends to increase the observed Raman intensity due to a resonance effect between O and the aromatic ring structure connected to the O-containing structure [2]. Therefore, it is believed that the oxygen-containing species from the dissociation of steam gave rise to the Ocontaining groups and hence increased the total Raman peak areas of chars produced from steam-containing atmospheres. The slow penetration of the CO2 intermediates into the char matrix would only show a slow increase in the total Raman area for the chars obtained from the gasification in pure CO2.
Total Raman Peak Area, arb. unit
3500
3000
2500
2000
1500
1000
0
5
10
15 20 25 Char Yield, wt% (db)
30
35
15% H2O balanced with argon Pure CO2 15% H2O balanced with CO2
Figure 3 Raman total area of char produced from the gasification at 800 oC in gasifying atmospheres as denoted by the legend.
4. Conclusions The FT-Raman spectral data show that the structure of chars changed with gasifying atmosphere in which the char was produced. The gasification in steam-containing atmospheres produced chars with very similar char structural features. Our data indicated that the char-H2O and char-CO2 reactions do not follow the same reaction pathways at least under our experimental conditions. 4
Oviedo ICCS&T2011.Extended Abstract
Acknowledgements The authors gratefully acknowledge the support of this study by the Australian Research Council (DP110105514). Some earlier part of this study was done in Monash University with support from the Victorian State Government under its Energy Technology Innovation Strategy program and the Latrobe Valley generators (International Power Hazelwood, International Power Loy Yang B, Loy Yang Power and TRUenergy).
References [1] Li X, Li C-Z. Volatilisation and catalytic effects of alkali and alkaline earth metallic species during the pyrolysis and gasification of Victorian brown coal. Part VIII. Catalysis and changes in char structure during gasification in steam. Fuel 2006;85:1518-25. [2] Li X, Hayashi J-i, Li C-Z. FT-Raman spectroscopic study of the evolution of char structure during the pyrolysis of a Victorian brown coal. Fuel 2006;85:1700-7. [3] Keown DM, Hayashi J-I, Li C-Z. Drastic changes in biomass char structure and reactivity upon contact with steam. Fuel 2008;87:1127-32. [4] Quyn DM, Wu H, Li C-Z. Volatilisation and catalytic effects of alkali and alkaline earth metallic species during the pyrolysis and gasification of Victorian brown coal. Part I. Volatilisation of Na and Cl from a set of NaCl-loaded samples. Fuel 2002;81:143-9. [5] Radovic LR, Walker PL, Jenkins RG. Importance of catalyst dispersion in the gasification of lignite chars. Journal of Catalysis 1983;82:382-94. [6] Hermann G, Hüttinger KJ. Mechanism of water vapour gasification of carbon--a new model. Carbon 1986;24:705-13.
5
3-D Structural analysis for metallurgical coke microstructure using micro X-ray CT Yoshiaki YAMAZAKI*, Kenichi HIRAKI*, Tetsuya KANAI*, Xiaoqing ZHANG *, Ataru UCHIDA*, Masakazu SHOJI*, Yohsuke MATSUSHITA*, Hideyuki AOKI*, Takatoshi MIURA*, Seiji NOMURA††, Hideyuki HAYASHIZAKI†† * Graduate School of Engineering, Tohoku University, 6-6-07, Aoba, Aramaki, Aoba-ku, Sendai, 980-8579 Japan, †† Environment & Process Technology Center, Nippon Steel Corp., 20-1, Shintomi, Futtsu, 293-8511 Japan.
Abstract Three dimensional (3-D) finite element model of coke microstructure was generated using micro X-ray CT and coke strength for the microstructure was numerically evaluated. First, 3-D stress analysis of a coke was compared with two-dimensional (2-D) stress analysis of the same one. Next, 3-D stress analyses of cokes, which were made from caking coal and slightly caking coal, were carried out. Furthermore, three dimensional wall thickness in coke microstructure was calculated to investigate the effect of the 3-D microstructure on coke strength. The results are summarized as follows: (1) Maximum principal stress magnitude and stress concentration area in 3-D stress analysis are smaller than the ones in 2-D stress analysis. (2) Stress concentration area in coke, which is made from slightly caking coal, is larger than one in coke which is made from caking coal. In coke made from slightly caking call, the wall is three dimensionally thinner than one in coke made from caking coal, so that the weak area develops. It is indicated that stress analysis based on images using micro X-ray CT is very useful to investigate relationship between microstructure of coke and coke strength. Key words; coke strength, 3-D microstructure, micro X-ray CT
1. Introduction
and estimated 3-D coke structure using fractal dimension.
Metallurgical coke’s roles are very important to achieve
However, there are few reports of 3-D coke analysis.
low reducing agent ratio and high productivity in the blast
In this study, microstructure of coke was scanned by
furnace operation. In particular, sustaining the flow passes
using micro X-ray CT and the resulting images were
of liquid metal and high temperature reducing gas in the
converted into finite element model. Then, stress analyses
blast furnace is the most important for stable operation of
were carried out to investigate relationship between
blast furnaces. Coke is porous material. It is known that
microstructure of coke and coke strength. Specifically,
microstructure of coke influences coke strength1,2).
3-D stress analysis of coke was compared with 2-D stress
The coke strength is recently estimated by numerical 3-5)
analysis
based on material and structural mechanics.
analysis of one. Next, 3-D stress analyses of cokes which were made from caking coal and slightly caking coal were
There are some reports from the standpoint of relationship
carried out to investigate the effect of 3-D microstructure
between microstructure and coke strength. However,
of these cokes on coke strength. Furthermore, we
these stress analyses are two-dimensional (2-D) analyses.
proposed the estimation method of coke strength based on
Actual cokes, in which matrix and pore connect,
3-D coke microstructure.
respectively, are three-dimensional (3-D) structures. In case of 2-D analysis, it seems that continuity of matrix is
2. Three-dimensional finite element analysis
underestimated. Therefore, it is necessary to investigate
2.1 Obtainment of three-dimensional structure
relationship between 3-D microstructure and coke strength based on actual coke structure. Micro X-ray CT is recently used to evaluate 3-D coke 6)
Table 1 shows coking condition and property of the specimens. The cokes were made from the caking coals (Goonyella coal and Peak Downs coal) and slightly
have observed 3-D coke
caking coal (NCBC coal) in an electrically heated pilot
microstructure before/after CO2 gasification reaction
coke oven (420 mm in width, 600 in length and 400 mm
structure. Yamamoto et al.
using micro X-ray CT. Tsafnat et al. have constructed
in height) 8). A coke lump was cut out as shown in Fig. 1.
3-D finite element model of coke using micro X-ray CT
Each specimen size was φ 19×30 mm. Coke specimens
7)
Table 1 Coking condition and property of the samples. Coal Peak Downs Goonyella NCBC
3
Size [mm]
Moisture [%]
Charging density [t/m ]
<3 (85%)
6～7
0.75
150
DI 15 83.9 82.8 69.9
MI0.212 MI0.6 44.2 13.1 45.2 12.0 38.6 2.5
Heated wall
Head
Body
Tail
60 mm
60 mm x3
φ 19 mm
θ
30 mm
X-ray
Body part
x1
x2
Sample
Fig. 1 Cut out and scanning procedure for coke sample.
Fig. 3 The procedure of 3-D model generation for coke microscopic structure.
were
scanned
by
using
micro
Fig. 2 The CT images of coke.
Fig. 4 Finite element model of Goonyella coke (Size: 5 mm×5 mm×5 mm).
X-ray
CT(TOSCANER-32250µhd, Toshiba IT & Control
2.2 Construction of finite element model
Systems Corporation, Japan). X-ray tube voltage was 180
Stress analyses were carried out using finite element
kV and X-ray tube current was 50 µA. The X-ray CT
analysis based on digital image modeling, which is widely
image is usually presented on a computer monitor using a
applied to composite material9) and cancellous bone10).
grey scale. Figure 2 shows cross-sectional images for each coke
Figure 3 shows the procedure of 3-D finite element model
generation
for
coke
microstructure.
First,
specimen. Black parts show pore, gray parts show coke
cross-sectional images of coke were obtained using micro
matrix and white parts show ash. Ash is displayed whitely
X-ray CT. Image analysis was done using Win Roof
since ash component is mainly silica and it is low
ver.6.0.1 (Mitani Corporation, Japan). Binary image for
radiolucency. The resolution was 32 µm/pixel for each
coke matrix and pore was generated based on histogram
coke specimen.
of brightness value. Discriminant analysis method which
automatically binarizes images was applied to decide threshold of binarization. In this study, ash was estimated
Analytical object
as coke matrix. Noise rejections were carried out by removing isolated coke matrices, which were about 0.14
x2
percent by volume. Cross-sectional images were stacked to form 3-D coke microstructure. Figure 4 shows finite
x1 x3
element model of coke. As shown in Fig. 4, each voxel,
Analytical conditions (a) Image of 3-D analysis Uniform y direction displacement x1 direction length 5.0 mm x2 direction length 5.0 mm x3 direction length 5.0 mm Bottom surface is restricted by x2 direction displacement
whose height was 32 mm/pixel, was finite element mesh. As a result, 3-D finite element model of coke was generated. x2
Analytical object
2.3 Analytical method Governing equations for this analysis are shown as: σ ij , j + Fi = 0 , (1)
1
ε ij = (ui, j + u j, i ), 2 σ ij = Eijkl ε kl .
(2) (3)
x1
(b) Image of 2-D analysis
Fig. 5 Computational domain and boundary condition.
Condition of static equilibrium Eq.(1) was used as governing equation. Relationship between strain and
3. Comparison of 2-D and 3-D elastic analyses
displacement was assumed to be infinitesimal strain Eq.
3.1 Analytical object and condition
(2). Linear constitutive law of stress and strain Eq. (3) was
2-D and 3-D stress analyses were carried out for
used. Stress analysis based on displacement method
Goonyella coke. In 3-D analysis, analytical area was 5
requires the formulation with explicit representation for
mm×5 mm×5 mm and resolution was 32 µm/pixel. In
displacement. Finite element formulation based on
2-D analysis, Computational objects were all images
displacement method is written as:
before stacking images for 3-D. Computational domain
n
ui ( x1 , x2 , x3 ) = ∑ N m ( x1 , x2 , x3 )u im ,
(4)
m =1
∫B Ω
T
DBdΩ ⋅ u = F,
B = LN,
(5) (6)
was 5×5 mm. Plane strain condition was assumed. 2-D numerical result was defined as average of maximum principal stresses which were obtained from all 2-D analytical results. Figure 5 shows the computational domains and boundary conditions for stress analyses
where Ω means whole region of continuum body.
assuming tensile tests. Constraint displacement of x2
Eight-node isoparametric hexahedral elements were
direction was uniformly applied to the top side nodes.
applied to this analysis. That is to say, the integers in Eq.
Constraint displacement corresponds to load of 1 MPa,
(4) is set n = 8 and i = 1, 2 and 3. In case that the number
which is calculated from reaction force on top side
of meshes is very large, it is impossible to calculate due to
boundary by cross-sectional area. Elastic modulus of coke
shortage of memory in one computer. Therefore, parallel
matrix was set 24 GPa based on result of nanoindentation
computations for stress analysis were carried out using
methods13). Poisson's ratio of coke matrix was set 0.3.
distributed memory type computer whose processors have
Elastic modulus of pore was 24×10-4 GPa3). In this
their own local memory.
numerical result, we confirmed that a stress distribution
In this study, each node displacement is calculated by the element-by-element conjugate gradient method11)
was the same as the one of case in which pore area mesh was not generated.
which is suitable to parallel computation and memory efficiency. It is known that coke is brittle material12), so analytical
3.2 Results and discussion Figure 6 (a) shows maximum principal stress
results were estimated by maximum principal stress
distributions of 3-D analyses and Fig. 6 (b) and (c) show
which is usually used to judge yield of brittle material.
maximum principal stress distribution in the same cross-sectional surfaces of 3-D and 2-D, respectively. In
Fig. 6, minus and plus values represent compressive and
4. The effect of coal type on coke strength
tensile stress, respectively. As shown in Fig. 6 (c), if coke
4.1 Analytical object and condition
matrix continuously connects from top side to bottom one
3-D stress analyses were carried out for three cokes
and wall thickness is thin, that area shows higher
which were made from Goonyella coal (caking coal),
maximum principal stress. This is because stress
Peak Downs coal (caking coal) and NCBC coal (slightly
concentration occurs since wall thickness is thin and
caking coal), respectively.
cross-sectional area of coke is small. In case of 3-D
resolution, material constants and boundary condition
analysis, similar tendency can be seen as well. However,
were the same as previous section. Uniaxial tensile tests
maximum principal stress magnitudes and stress
toward the x3 direction were assumed.
Computational domain,
concentration area in 3-D analysis are smaller than the ones in 2-D stress analysis.
4.2 Evaluation of 3-D wall thickness
Table 2 shows maximum and average values of
Actual coke is 3-D structure. Therefore, parameter
maximum principal stress. Both maximum and average
based on 3-D coke structure is necessary to investigate the
values in 3-D are lower than the ones in 2-D. It seems that
effect of coke microstructure on coke strength. We
continuity of matrix is underestimated in 2-D. In 3-D,
propose the method based on 3-D coke microstructure to
relaxation of stress concentration occurs since coke matrix
estimate coke strength.
connects in the depth (x3) direction. Therefore, numerical
Figure 7 shows a conceptual diagram of the method to
result of 3-D stress analysis is smaller than one of 2-D.
estimate 3-D wall thickness. We applied method which
We can estimate more accurate coke strength based on
was developed by Lindquist et al.15) to 3-D coke structure.
actual coke microstructure by 3-D stress analysis.
As shown in Fig. 7, each voxel in the pore is labeled with integer 0 and voxels in the coke matrix are initially
Fig. 6 Maximum principal stress distribution in same cross-sectional surfaces of 3-D and 2-D.
Table 2 Maximum and average values of maximum principal stress. Maximum value of maximum Average value of maximum principal stress[MPa] principal stress[MPa] 28.4 1.1 79.3 1.5
3D 2D
Coke matrix
Pore 0
0
0
0
0
0
0
0
1
1
2
3
4
4
4
4
0
0
0
0
0
0
0
0
1
1
2
3
3
3
3
3
0
0
0
0
0
0
0
0
1
2
2
2
2
2
2
2
0
0
0
0
0
0
0
1
1
2
2
2
1
1
1
1
0
0
0
0
0
0
0
1
2
2
1
1
1
0
0
0
0
0
0
0
0
0
1
1
1
1
1
0
0
0
0
0
0
0
0
0
1
1
1
1
1
0
0
0
0
0
0
0
0
0
1
1
1
2
2
1
0
0
0
0
0
0
0
0
1
1
1
2
2
2
2
1
0
0
0
0
0
0
0
0
2
2
2
2
3
3
2
1
1
0
0
0
0
0
0
0
3
3
3
3
3
3
2
2
1
1
0
0
0
0
0
0
4
4
4
4
4
3
3
2
2
1
1
0
0
0
0
0
Pore
Fig. 7 Image of the method to estimate 3-D wall thickness.
Fig. 8 Number distribution of Goonyella coke.
100
100
NCBC Peak Downs Goonyella
1
Goonyella
10
Frequency [%]
Frequency [%]
10
0.1 0.01 0.001
Peak Downs 1
NCBC
0.1 0.01 0.001
0.0001
0.0001
0.00001 -5
0
5
10
15
20
25
30
35
40
45
Maximum principal stress [MPa]
1 2 3 4 5 6 7 8 9 10 11 12 13
Number Fig. 9 Histogram of maximum principal stress for each coke.
Fig. 10 histogram of wall thickness number.
unlabeled. The algorithm proceeds by simultaneously
for each coke. In this study, 5 MPa14) was applied to the
assigning unlabeled voxels in the coke matrix with an
standard of tensile strength and we discussed coke
integer n + 1 for each voxel neighboring a pore-phase
strength based on it. As shown in Fig. 9, high stress can be
voxel, with n = 0 for this first iteration. This procedure
seen for coke, which is made from NCBC coal (NCBC
continues at a rate of one layer per iteration until all
coke). This indicates that NCBC coke includes the most
voxels in the coke matrix are labeled.
starting points of fracture and its strength is the lowest of
Figure 8 shows cross-sectional image of Goonyella
three cokes. Figure 10 shows histogram of wall thickness
coke. Red parts represent higher number voxels. This
number. Higher number can be seen for Goonyella coke.
indicates wall thicknesses are thicker. On the other hand,
This indicates wall thickness of Goonyella coke is thicker.
Blue parts show lower number voxels. This indicates wall
Lower number can be seen for NCBC coke. This
thickness is thinner. As shown in Fig. 8, wall thickness
indicates wall thickness of NCBC coke is thinner. In Fig.
appear to be thick, but this part is not red in the
10, higher stress can be seen for NCBC coke. This is
cross-sectional image. This is because coke matrix does
because stress concentration area of NCBC coke is larger
not connect in the depth (x3) direction.
since its wall thickness is thinner. This tendency agrees with result of Fig. 10. As a result, it is thought we can
4.3 Results and Discussion Figure 9 shows histogram of maximum principal stress
estimate 3-D coke microstructure using method which was developed by Lindquist et al.
5. Conclusion
εij
:
strain
[-]
σij
:
stress
[Pa]
Linear elastic stress analyses were carried out to investigate relationship between microstructure of coke
Reference
and coke strength using micro X-ray CT. First, 3-D stress
1) 2)
analysis of coke was compared with 2-D stress analysis of one. Next, 3-D stress analyses of cokes which were made from caking coal and slightly caking coal were carried out
3)
to investigate the effect of 3-D microstructure of these cokes in coke strength. Furthermore, we propose the method based on 3-D coke microstructure to estimate
4)
coke strength. The results are summarized as fo２llows: (1)
Average and maximum values of maximum principal stress in 3-D stress analysis were smaller
5)
than the ones in 2-D stress analysis. It is supposed that continuity of matrix is underestimated in 2-D. In
6)
3-D, relaxation of stress concentration occurs since 7)
coke matrix connects in the depth direction. (2)
Higher stress could be seen for NCBC coke. This indicates that NCBC coke includes the most starting points of fracture and its strength is the lowest of three cokes. This is because wall thickness of NCBC is thinner and its porosity is lower than the other
8) 9) 10)
cokes. (3)
It is suggested we can estimate 3-D coke microstructure using method which is developed by Lindquist et al.
It is indicated that stress analysis based on images using
11) 12)
micro X-ray CT is very useful to investigate relationship between microstructure of coke and coke strength. We
13)
can quantitatively estimate coke strength based on actual coke structure. 14)
Nomenclature
15)
B
:
B-matrix for finite element method
D Eijkl
:
elastic modulus matrix
:
elastic modulus tensor
[Pa]
Fi
:
body force
[N･m-3]
F
:
external force vector
[N]
L Nm
:
operator matrix
:
shape function
N ui
:
shape function matrix
:
displacement at (x1, x2, x3) in
[Pa]
[m]
element ui
m
:
nodal value of displacement
[m]
u xi
:
displacement vector
[m]
:
coordinate
[m]
x
:
coordinate vector
[m]
T. Arima: Tetsu-to-Hagané, 87(2001), 274. Y. Kubota, S. Nomura, T. Arima and K. Kato: ISIJ Int., 48(2008), 563. H. Hayashizaki, K. Ueoka, M. Kajiyama, Y. Yoshiaki, K. Hiraki, Y. Matsushita, H. Aoki, T. Miura, K. Fukuda and K. Matsudaira: Tetsu-to-Hagané, 95(2009), 593. Y. Yamazaki, H. Hayashizaki, K. Ueoka, K. Hiraki, Y. Matsushita, H. Aoki and T. Miura: Tetsu-to-Hagané, 96(2010), 536. K. Ueoka, T. Ogata, Y. Matsushita, Y. Morozumi, H. Aoki, T. Miura, K. Uebo and K. Fukuda: ISIJ Int., 47(2007), 1723. Y. Yamamoto, Y. Kashiwaya, M. Nishimura and M. Kubota: Tetsu-to-Hagané, 95(2009), 103. N. Tsafnat, G. Tsafnat and Allan S. Jones: Minerals Engineering, 22(2009), 149. S. Nomura, T. Arima and K. Kato: Fuel, 83 (2004), 1771. G. Nagai, T. Yamada and A. Wada: J. Struct. Constr. Eng., 509(1998), 77. H. Hirashita, N. Takano, M. Sugishita, S. Matsunaga and Y. Ide: The Computational Mechanics Conference, 20(2007), 385. K. Kashiyama and T. Tsukasa: Proc. Jpn. Soc. Civ. Eng. 668(2001), 43. T. Ogata, K. Ueoka, H. Hayashizaki, Y. Matsushita, H. Aoki, T. Miura, K. Fukuda and K. Matsudaira: Tetsu-to-Hagané, 93(2007), 736. H. Hayashizaki, K. Ueoka, Y. Yoshiaki, Y. Matsushita, H. Aoki, T. Miura, K. Fukuda and K. Matsudaira: Tetsu-to-Hagané, 95(2009), 460. M. Sakai, R. Nishimura, M. Nishimura and K. Fukuda: Tetsu-to-Hagané, 92(2006), 164. W. Brent Lindquist, Sang-Moon Lee, David A. Coker, Keith W. Jones and Per Spanne: Journal of Geophysical Research, 101(1996), 8297-8310
Preparation of High-strength Coke from Hot-briquetted Brown Coal A. Mori1, Y. Huang1, K. Norinaga1, S. Kudo1, T. Kanai2, H. Aoki2 and J.-i. Hayashi1,* 1
Institute for Materials Chemistry and Engineering, Kyushu University, Kasuga 816-0850, Japan, 2School of Engineering, Tohoku University, Sendai 980-8579, Japan *[email protected] Abstract Coke with tensile strength of 6–37 MPa was prepared through binderless briquetting of pulverized Loy Yang brown coal that showed no fluidity upon heating and subsequent heating of the briquette up to 900 °C. The briquetting at temperature of 130–200 °C and mechanical pressure of 64–128 MPa caused plasticization due to mobilization of low-mass components and deformation of the coal matrix, and thereby coalescence and adhesion among particles. The resulting coke had density as high as 1.1–1.3 g/cm3, which contributed to tensile strength of 28–37 MPa. 1. Introduction It is known that brown coal as well as lignite is not a suitable feedstock of metallurgical coke that is generally produced from caking coal or its blend with slightly caking or non-caking coals. Victorian brown coals have reflectance and Gieseler fluidity of 0.19–0.36% [1] and nearly zero, respectively, and these are both out of the ranges of MOF diagram [2] that gives composition of coal blend suitable for producing high-strength coke. Blending brown coal with caking coal may not be reasonable if its particular nature, i.e., high absorptivity of fluid material causing loss of fluidity of the blend upon heating [3,4], is taken into consideration. Formcoke technologies [5,6,7,8] may allow application of char/coke from brown coals to production of coke briquettes if appropriate binders such as raw or modified coal tars or asphalts are available. This technology requires the pyrolysis/carbonization for producing not only char/coke but also a binder (in application of coal tar) before briquetting and carbonizing them. There have been reports on binderless briquetting of a sub-bituminous coal and a lignite and successful production of coke from such briquettes [9,10]. Bayraktar and Lawson [10] briquetted an acid-demineralized Turkish lignite by applying 113–212 MPa mechanical pressure at ambient temperature and carbonized the briquette at 900 °C, thereby producing coke with a tensile strength of 6 MPa. This strength was as high as those of conventional blast furnace and foundry cokes, 2–6 MPa [11,12,13,14,15]. It is believed that Victorian brown coal is a most suitable feedstock of binderless briquettes
[16] by virtue of its great affinity with water (moisture) that may play roles of plasticizer, lubricant and/or adhesive, and, in addition, very low ash content. The present authors propose binderless briquetting of brown coal at temperature above 100 °C. Firstly, this method may enable to avoid formation of defects associated with the moisture release in early stage of heating (for producing coke) without mechanical pressure. Simultaneous moisture removal and mechanical pressurization are expected to be effective in maintenance of particles’ adhesion. Secondly, applying mechanical pressure at temperature of 100–200 °C may cause significant plasticization of brown coal over molecular to particle scales due to thermal breakage of hydrogen bonds [17] and mechanically induced extraction of low-mass components that are expected to cause significant particles’ deformation and adhesion by playing roles of self-plasticizer/binder [18]. In the present study, experimental investigation was made mainly on the combined effects of briquetting temperature and pressure on properties of resulting briquettes of Victorian brown coal and coke. 2. Experimental As-mined Loy Yang brown coal (Victoria, Australia) with a moisture content of more than 60 wt%-wet was dried in air at ambient temperature until its moisture content decreased to 10 wt%-wet, pulverized and then sieved for collecting a fraction that consisted of particles with sizes smaller than 106 µm, which was employed as the feedstock. According to the result from further sieving, the particles with sizes greater than 53 µm accounted for 61 wt% of the feedstock. The particle size range, i.e., –106 µm, was chosen according to the report by Bayraktar and Lawson [10]. The elemental composition of the feedstock was as follows: C = 69.2, H = 4.7, O = 25.2 (by difference), N = 0.6, S = 0.3 wt%-daf. The ash content was 0.80 wt%-dry. The feedstock of about 1 g was transferred into a 14.1-mm-diameter mold and heated to a prescribed temperature in a range from 25 to 230 °C. After confirming the temperature of the feedstock had reached the prescribed temperature, a mechanical pressure of 32–192 MPa was applied by hydraulic loading via the piston for 3–8 min. The pressure was then released, and the resulting briquette was recovered while the mold was cooled naturally to ambient temperature. The dimensions, mass and moisture content of the briquette in a shape of disc were measured. Typical diameter and thickness of the briquette were 14.0 and 5.0 mm, respectively. The briquette was heated in atmospheric flow of N2 (purity > 99.9999 vol%) at a rate of 3 °C/min up to 900 °C with a holding period of 10 min, and then cooled down to
ambient temperature at an average rate of 100–150 °C/min. The resulting coke was recovered and its dimensions and mass were measured. Mechanical strength of the coke was measured at ambient temperature by means of diametrical compression tests on a testing apparatus, Shimadzu EZ-L. The displacement and loading were measured during the compression at a displacement rate of 2.00 mm/min [19]. Assuming that the maximum loading at the breakage of the specimen corresponded to the maximum tensile stress, Pmax, it was determined based on an equation: Pmax = 2Lmax/πdl, where Lmax, d and l are the maximum loading, diameter and thickness of the specimen, respectively. Though not shown in detail, for all of the specimens tested, the loading (= stress) increased linearly with the displacement until its sudden drop. The coke samples prepared in the present study were thus broken obeying a brittle fracture mechanism. It was also confirmed that the coke samples cracked nearly equally into two semi-discs, which was a necessary condition in applying the above equation [20]. Fractured surfaces as well as top/bottom ones of some coke samples were observed by scanning electron microscopy (SEM) on a micrograph (Keyence, VE-9800). 3. Results and Discussion 3.1. Effects of briquetting temperature on properties of briquette and coke Effects of briquetting temperature, TB, were examined on the properties of resulting briquette and coke. The briquetting pressure, hereafter referred to as PB, was fixed at 128 MPa. The briquettes prepared at TB = 25, 70, 100 and 130 °C contained 6.6, 4.5, 2.4 and 0.5 wt% moisture, respectively, while less than 0.1 wt% for those at higher temperatures. Figure 1 shows the coke yield as a function of TB. The briquetting was effective for increasing the coke yield over the entire range of TB. Such increase was probably due to suppressed tar evolution caused by promoted cross-linking reactions and intra-particle charring of tar vapor. It was suggested that the briquetting at TB > 100 °C caused an additional positive effect on the suppression of tar evolution, implying changes in the coal structure at a molecular scale. The briquette density, ρB, was calculated on a moisture-free basis by an equation, ρB = (1 – w)ρ, where w and ρ were the residual moisture content and measured apparent density of the briquette, respectively. This equation is based on an assumption that the briquette undergoes little volumetric contraction in the final stage of moisture release [21]. As shown in Fig. 2, ρB greatly increases from 1.02 to 1.20 g/cm3 at 25–70 °C, and further but gradually to 1.26 g/cm3 at 230 °C. According to Higgins et al. [22] and Matsuo et al. [23], the true density and micropore volume of dry Loy Yang coal are 1.42 g/cm3 and 0.065 cm3/g, respectively. Assuming no meso-/macro-pores in the briquette, its density is 1.30 g/cm3. ρB for the briquettes prepared at 200–300 °C being as high as
1.24–1.26 cm3/g thus suggested that the briquetting at such temperature effectively eliminated inter-particle spaces by causing deformation of particles and their adhesion/coalescence and even loss of intraparticle macro/meso-pores. TB influenced the density of coke, ρC, in a way slightly different from ρB, i.e., there occured a minimum of ρC. The density of coke was a result from events causing increase or decrease in the density of carbonizing briquette. ρC had a minimum at TB = 130–160 °C. It was believed that during the carbonization ρC varied in three events; thermal relaxation of briquette, pore formation due to evolution of volatiles and densification of the nonporous part of the solid. No particular effects of applying TB = 130–160 °C was plausible on the second and third events so far as the result shown in Fig.1 was considered, and it was therefore suggested that the briquettes for TB = 130–160 °C underwent thermal relaxation that was associated with volumetric expansion in the course of reheating. In other words, the coal underwent a type of structural rearrangement during the briquetting at TB > 100 °C in a manner different from that at TB ≤ 100 °C. This hypothesis was consistent with the increase in the coke yield at TB > 100 °C, though the scale of such rearrangement is unknown. Another effect of applying TB > 100 °C on a coke property will be shown and discussed later. In Fig.3, the average tensile strength of coke, Pmax, is shown as a function of TB together with standard deviation, σ. Pmax is greater than those of conventional cokes, 2–6 MPa over the range of TB. It is also noted that Pmax steeply increases from 14 to 28 MPa at TB = 100–130 °C. Briquetting at TB > 100 °C was thus effective for preparing high strength coke with Pmax = 28–37 MPa. 3.2. Effects of briquetting pressure on properties of briquette and coke The effects of PB on ρB and ρC were investigated with TB = 200 °C. As seen in Fig.4, both ρB and ρC increase monotonously with PB in manners very similar to each other. ρB and ρC are also very close to each other. Figure 5 illustrates the effect of PB on the coke yield. Compared with the effect of applying TB = 200 °C on the coke yield (see Fig. 1), that of increasing PB was less significant. Figure 6 presents average Pmax as a function of PB. Pmax increases in a manner similar to that of ρC, and this indicates importance of ρC as a factor determining the tensile strength of coke. 3.3. SEM observation of coke samples Top/bottom and fractured surfaces of cokes prepared under different combinations of TB and PB were obserbed by SEM. Figure 7 displays SEM photographs of some selected coke samples. One of the features common among the coke samples was absence or
rarity of pores with sizes of 10 µm and greater, and this was a main reason why many of the present cokes had Pmax > 25 MPa. The fractured surfaces of a coke sample prepared with PB = 128 MPa and TB = 100 °C (see photos e and f) has morphology very similar to that of another coke from the same PB and TB = 200 °C (c and d) in terms of frequency of grain boundaries and size/frequency of pores. On the other hand, the fractured surface of a coke from TB = 200 °C and PB = 32 MPa (photos g and h) has clearly more grain boundaries and greater pores than that from PB = 128 MPa. The photographs c–f demonstrate that the briquetting leading to cokes with ρC > 1.2 g/cm3 successfully eliminated grain boundaries. It was found that the briquettes prepared at TB > 100 °C were tinged with black while the briquettes prepared at TB = 25 °C were dark-brown colored. This was an implication of that more or less amount of low-mass components of the coal was mechanically squeezed out of the macromolecular network, appeared on the particle surface and promoted its adhesion to another particle. It was also implied that such low-mass components, even inside the particle matrix, played a role of plasticizer, enhancing deformation of the particle. Some briquettes were crushed and then subjected to extraction with tetrahydrofuran (THF) at 30 °C and under ultrasonic irradiation. THF extracted material with 12.9 wt% of the briquette. This extraction yield was nearly equivalent with that from the starting coal, 13.6 wt%. The result was explained by that the mechanical extraction of low-mass components, if occurred, was not so extensive as the extraction with THF. 3.3. Factors crucial to mechanical strength of coke Figure 8 summarizes combined effects of TB and PB on the average Pmax assuming that ρC is a critical property for the mechanical strength of coke. Two different relationships between Pmax and ρC are seen in the figure. Pmax of the coke from briquetting at TB > 100 °C is a linear function of ρC over the ranges of ρC = 0.87–1.27 cm3/g and PB = 32–192 MPa. In case of TB ≤ 100 °C, Pmax is described by another linear function of ρC. Briquetting with TB > 100 °C was thus found to be necessary for preparing coke with Pmax > 15 MPa within the conditions examined so far. Each of the linear relationships are explained reasonably. Under an assumption that all of the cokes prepared in this work had similar true densities [24], difference in ρC among cokes from different conditions is attributed mainly to that in the porosity. Patrick and Stacey [24] showed that the tensile strength of coke was well described as a function of porosity. Arima [25] claimed that the strength of non-porous part of coke was not necessarily a function of coal rank/type and fusibility while frequencies of connected pores and non-adhesion grain boundaries were more crucial. ρC of the present
cokes, ranging from 0.87 to 1.27 cm3/g, are much higher than general blast furnace and foundry cokes. It was thus believed that such high density, in other words, low porosity was the primary reason of very high tensile strength of the cokes prepared in this work. Polished surface of a coke sample (TB = 200 °C, PB = 128 MPa) was observed by an optical polarized-light microscopy, which confirmed no progress of formation of optically anisotropic textures during the carbonization. A thermomechanical analysis of a briquette (TB = 200 °C, PB = 128 MPa) in a needle penetration mode revealed no softening/fusion of the briquette during the carbonization. These results are proofs of that high strength of the present cokes was not due to carbonization via mesophase formation. Here is discussed on the presence of two different ρC–Pmax relationships for the briquetting at TB > 100 °C and that at TB ≤ 100 °C. As shown in some photographs of Fig.7, it was difficult to explain the difference in the fractured surface between the cokes from briquettes with TB = 100 °C and 200 °C. The briquettes prepared at TB ≤ 100 °C contained 2.4 to 6.6 wt% moisture, which played a role of plasticizer and/or lubricant in the briquetting helping the formation of briquettes with ρB as high as 1.2 g/cm3. It was believed that low-mass components of the coal did not play important roles due to insufficiently high temperature for breaking hydrogen bonds and also the presence of water that was not necessarily a good solvent of such components. The briquetting at TB > 100 °C effectively removed water to a content below 0.5 wt% and also allowed mobilization of low-mass components. As shown in Fig.2, The densities of the briquettes with TB = 130 and 160 °C decreased upon the reheating (for the carbonization) to degrees much greater than those with TB ≤ 100 °C. This was probably arisen from thermal relaxation (volumetric expansion) of the briquettes with TB = 130 and 160 °C. Such thermal relaxation might release briquetting-induced mechanical stresses inside the briquettes, which could induce defects such as microcracks during the carbonization. Thus the thermal relaxation, if occurred, played a role of annealing in the early stage of the carbonization. For the briquetting at TB = 200–230 °C, the temperature would be high enough for the annealing within the period of briquetting. The resulting briquettes therefore underwent little or no density decrease (see Fig.2). 4. Conclusions Binderless briquetting of pulverized Loy Yang brown coal applying mechanical pressure of 32–192 MPa and temperature of 25–230 °C and subsequent carbonization with heating up to 900 °C successfully produced coke having tensile strength of 6–37 MPa. Such high mechanical strength of cokes was attributed mainly to low porosity. Briquetting at temperature over 100 °C caused plasticization/deformation of the coal
matrix due to low-mass components, and thereby promoting coalscence/adhesion of particles and eliminating inter-particle spaces and grain boundaries. Acknowledgement. This study was carried out as a part of a research project, “Scientific Platform of Innovative Technologies for Co-Upgrading of Brown Coal and Biomass”, which has been financially supported by MEXT, Japan in a Program of Strategic Funds for the Promotion of Science and Technology. The authors are also grateful to The Iron and Steel Institute, Japan (ISIJ), and Nippon Steel Corporation (NSC) for financial supports. Prof. Tsuyoshi Hirajima, School of Engineering, Kyushu University, and Dr. Seiji Nomura, NSC, are acknowledged for their technical advice. References [1] George AM, Mackay GH. Chapter 2: The Science of Victorian Brown Coal. Edited by Durie RA. 1991, Butterworth-Heinemann, Oxford. [2] Miyazu T. Proc. Int. Iron Steel Congress, Dusseldorf 1974;Paper 1.2.2.1. [3] Wornat MJ, Sakurovs R. Fuel 1996;75:867-871. [4] Sakurovs R. Fuel 2000;79:379-389. [5] Taylor JW, Coban A. Fuel 1987;66:1274-1280. [6] Rubio B, Izquierdo MT, Segura E. Carbon 1999; 37:1833–1841. [7] Paul SA, Hull AS, Plancher H, Agarwal PK. Fuel Processing Technology 2002;76:211-230. [8] Plancher H, Agarwal PK, Severns R. Fuel Processing Technology 2002;79:83-92. [9] Yip K. Wu H, Zhang Dk. Energy Fuels 2007;21:419-425. [10] Bayraktar KN, Lawson GJ. Fuel 1984:63;1221-1225. [11] Patrick JW, Stacey AE. Fuel 1972;51:81-87. [12] Patrick JW, Stacey AE, Wilkinson HC. Fuel 1972;51:174-179. [13] Patrick JW, Stacey AE. Fuel 1972;51:206-210. [14] Miyagawa T, Fujishima I. Nenryokyokai-shi 1975;54:983-993. [15] Miyagawa T, Fujishima I. Nenryokyokai-shi 1976;55:30-37. [16] Allardice, DJ. Chapter 3: The Science of Victorian Brown Coal. Edited by Durie RA. 1991, Butterworth-Heinemann, Oxford. [17] Miura K, Mae K, Sakurada K, Hashimoto K. Energy Fuels 1992;6:16-21. [18] Allardice DJ, Chaffee AL, Jackson WR, Marshall M. Chapter 3: Advances in Science of Victorian Brown Coal. Edited by Li CZ. 2004, Elsevier, Oxford. [19] Yamazaki Y, Hayashizaki H, Ueoka K, Hiraki K, Matsushita Y, Aoki H, Miura T. Tetsu-to-Hagané 2004;90:536-544. [20] Mellor M, Hawkes I. Engineering Geology 1971;5:173-225. [21] Evans DG. Fuel 1973;52:186-190. [22] Higgins RS, Kiss LT, Allardice DJ, George AM, King TNW. SECV Research and Development Department Report 1980, Report No. SC/80/17. [23] Matsuo Y, Hayashi Ji, Kusakabe K, Morooka S. Coal Science Technology 1995;24:929-932.
[24] Patrick JW, Stacey AE. Fuel 1978;57:258-264. [25] Arima T. Tetsu-to-Hagané 2001;87:274-281.
Coke yield, wt%-dry-coal
65 PB = 128 MPa 60
55
50
45
Char yield from pulverized coal
0
50
100
150
200
250
TB, °C
Figure 1. Effect of TB on the yield of resulting coke. 1.4 PB = 128 MPa
!B or !C, g/cm3
1.3 1.2 1.1 1
!Coke C
0.9 0.8
!Moisture-free briquette B
0
50
100
150
200
250
TB, °C
Figure 2. Effects of TB on ρB and ρC. 50
PB = 128 MPa
Pmax, MPa
40 30 20 10 0
2 sigma
0
50
100
150
200
TB, °C
Figure 3. Pmax as a function of TB.
250
1.4 TB = 200 °C
!B or !C, g/cm3
1.3 1.2 1.1 1
Coke
0.9 0.8
Moisture-free briquette 0
50
100
150
200
PB, MPa
Figure 4. Effects of PB on ρB and ρC.
Coke yield, wt%-dry-coal
65 TB = 200 °C 60
55
50
45
Char yield from pulverized coal
0
50
100
150
200
PB, MPa
Figure 5. Effect of PB on the yield of resulting coke. 50
TB = 200 °C
Pmax, MPa
40 30 20 10 0
2 sigma
0
50
100
150
PB, MPa
Figure 6. Pmax as a function of PB.
200
Figure 7. SEM images of fractures surfaces (a, c - h) and top/bottom surface (b) of coke samples. Images a, c and d: ΤΒ/PΒ; 200 °C/128 MPa. Images b, e and f: 100 °C/128 MPa. Images g and h: 200 °C/32 MPa. 50
Pmax, MPa
40 30
PB = 128 MPa TB = 200 °C
TB > 100 °C
20 10 0 0.8
TB ! 100 °C
0.9
1
1.1
1.2
3
!C, g/cm
Figure 8. Relationships between ρC and Pmax.
1.3
A Mechanism of improvement in coke strength by Adding a Solvent-Extracted Coal Noriyuki OKUYAMA, Takahiro SHISHIDO, Koji SAKAI, Maki HAMAGUCHI and Nobuyuki KOMATSU Coal & Energy Technology Dept., KOBE STEEL, Ltd, JAPAN [email protected] Abstract: A coal extract, produced by thermal extraction and solvent de-ashing in the coal derived methylnaphthalene solvent, has an excellent thermoplasticity even though the parent coal appears no thermoplasticity. We named it “HPC, High Performance Caking additive”, and have been developing to utilize as a thermoplasticity accelerator to make strong coke for blast furnace. Significant improvement in the thermoplasticity of coal blends are observed by HPC addition, especially with high blending ratio of slightly caking coals. It improves other dominant factors in coke strength, increasing in the dilatation of coals, increasing in the anisotropic texture and decrease in the inert structure in the coke. It is confirmed that the improvement of dilatation is the most important factor to perform filling the inter particle voids of coals in the coking reaction, that brings strong adhesive of coal particles to form a lump strong coke. The increasing in the charging bulk density of coal is also an important factor to decrease in the inter particle voids. The changing in the coke strength is quantitatively investigated as the functions of those factors in this study. 1. Introduction We have been developing a noble coal upgrading process to make an ash-less coal by applying the solvent de-ashing technology1). Coal is extracted in the coal-derived methylnaphthalene solvent at 350-400 oC. The solution is separated from the extraction residue by the gravity settling. Solvent is recovered and recycled in the process. HPC appeals an excellent thermoplasticity even though the parent coal appeals no thermoplasticity2). HPC softens at low temperature (less than 300ºC), keeps a highly fluidity in a wide temperature range and resolidifies at high temperature, nearly 500ºC. Therefore, HPC is available as a caking additive to make strong coke for the blast furnace. We participate in a Japanese national project called COURSE50 (CO2 Ultimate Reduction in Steelmaking process by innovative technology for cool Earth 50), in which we aim to apply HPC
as a caking additive to make strong coke for the iron blast furnace, which maximize the use of hydrogen as a reductant, in order to minimize coke addition to the furnace. Thermoplasticity is one of the most important characteristics for coke making. Japanese coke and steel companies make blast furnace coke by blending various types of coals, but the strongly caking coals have to be the main components in the coal blends. On the other hand, increasing in the use of slightly caking coals is an important subject to manage the shortage problem of strongly caking coals and to reduce accommodation cost of coal. This study investigates the effects of HPC as a caking additive to increase coke strength. Improvements in the caking properties by HPC addition, such as thermoplasticity, dilatation and anisotropic texture are examined, and coke strength is quantitatively understood with relation of those factors in this study. 2. Experimental section 2.1 Materials Three kinds Australian strongly caking coals (A, B, C) and a slightly caking coal (D) were used in this study. Table 1 shows the results of proximate, ultimate analysis and Gieseler plastometry (JIS M8801, similar to ASTM procedure) of the materials, and Table 2 shows the coal blending ratio of each blending base and total reflectance (Ro), total inert (TI) of each base. Base-1 consists with 75 wt% of strongly caking coals, and Base-2 consists with 50 wt% of strongly caking coals, coal-A : B: C : D = 15: 30 : 5: 50. When HPC was added to coal blend, coal-C was exchanged with HPC in the case of Base-1, and coal-B was exchanged in the case of Base-2. HPC was produced using a continuous bench scale unit (BSU), 0.1 t/d of coal consumption. The feedstock for HPC was an Australian steaming coal (MO), extracted at 400ºC. Table 3 shows the properties of MO-coal and HPC, and Table 4 shows the composition of the recycling solvent of BSU. Table 1 Properties of blending coals Proximate analysis name
ash [wt%]db
A B
10.3 10.5
VM
C
1)
21.2 21.8
Ultimate analysis H
N
S Odiff.
Maceral analysis4) Thermoplasticity 3) ST MFT RT Log (MFD) Ro Exinite Vitrinite Inertinite
[wt%] ( daf 2) ) 90.2 4.6 2.0 0.6 2.6 89.6 4.7 1.6 0.3 3.8
433 432
[ddpm] [ºC] 481 502 1.43 474 505 1.08
[%]
[-] 1.46 1.29
0.0 0.0
86.0 75.0
TI [%]
14.0 25.0
20.7 28.4
C
8.0
35.5
85.7 5.4 2.2 0.6 6.0
383
443 476
3.27
0.91
1.6
85.2
13.2
16.5
D
10.0
28.0
87.0 4.8 2.0 0.6 5.7
422
455 473
1.00
1.00
0.4
75.2
24.4
30.8
1) dry basis, 2) dry and ash free basis, 3) Gieseler plastometry (JIS M8801), ST: Softening temp., MFT: Maximum fluid temp., MFD: Maximum fluidity RT: Resolidification temp. 4) JIS M8816， Ro: Mean reflectance, TI: Total inert
Table 2 Coal blending conditions
Table 3 Properties of coal extraction solvent
Blending ratio [wt%] Base blend A B C D Ro (calc.) TI (calc.) 1.11 23.8 15 26 (34) 25 Base-1 1.16 27.9 Base-2 20 (30) 0 50
Ultimate analysis Atomic ratio H/C O/C H N S Odiff. [wt%] 91.45 6.9 0.02 0.7 0.93 0.905 0.008 [wt%] Component 1-methylnaphthalene 54.23 2-methylnaphthalene 37.46 naphthalene 2.10 dimethylnaphthalene 1.92 biphenyls 0.32 benzenes (1-ring compounds) 0.68 alkanes 0.48 thiol (S-containing aromatics) 2.04 others 0.76
( ) ：Exchanging coal with HPC
C
2.2 Coal carbonization experiment The coal blends were carbonized according to the conventional metallurgical coke making procedures. Coals were pulverized under 3 mm and HPC was pulverized under 0.15 mm. Coal blends were charged into an iron-made container of 0.02 m3, at 800 kg/m3 of bulk density, and were heated to 1000 ºC at a rate of 3 ºC/min under inert atmosphere. Duration was 30 min. 9 containers were filled in a coke oven, which has 300 kg of coal charging capacity. Drum Test (JIS K2151) was carried out to measure the coke strength. This is common way to evaluate coke strength in Japan. A 10kg of coke sample of the +50mm square hole was placed in the tumble drum and rotated for 30 revolutions, removed, screened and replaced in the drum and given impacts by further 150 revolutions. The drum index, DI15015, means the percentage of remaining +15mm square hole after 150 revolutions. The larger number indicates the stronger coke. 3. Results and Discussion 3.1 Changing in coke strength with fluidity of coal blends Thermoplasticity of coal is improved by HPC addition. That effect strongly appeals with lower grade coal, such as none or slightly caking coal3). Figure 1 shows Gieseler curves of the coal blends and HPC additional cases. Base-1 has sufficient fluidity, but base-2 has insufficient fluidity since blending ratio of the slightly caking coal (Coal-D) is 50 wt%. When HPC is added, Gieseler curves expanded to wider range and higher fluidity. And those effects appear more strongly on Base-2. Figure 2 shows the relation between MF value and coke strength (DI15015), and other dominant factors for coke strength, total reflectance (Ro), total dilatation (TD) of coal blends and anisotropic texture of coke. The DI value was increased more over to 88 by increase in MF value, and decreased in reversely. In the case of Base-2, DI value was too weak to use for
blast furnace without HPC addition and strongly increased with increase in MF value by HPC addition. Looking at other dominant factors in coke strength, it can be said as follows; (a) Total reflectance of coal blends are not influenced in this study (b) Total dilatation is strongly increased with increase in MF value especially in Base-2.
Log (Fluidity/ddpm)
(c) Optical anisotropic texture is developed with HPC addition, and inert texture, which can be regarded as the start point of crack, is decreased. Base blend + HPC 5% + HPC 10% + HPC 15%
3 Base-1
Base-2
+ HPC 20%
2
1
0 380
420
460
420 500 380 Temperature [ºC]
460
500
(5)
Mean reflex ratio (10)
(0)
85
(15)
(20)
(5)
80 (
): HPC amount
75 0
1 2 3 Log (MF/ddpm) Base-1 Base-2’
1.3
60
(Ro)
1.1
(25)
(10)
(0)
Ro
90
Optical texture [%] Total dilatation [vol%]
Coke strength DI 150 15
Fig.1 Changing in Gieseler curves by HPC addition
4
Total dilatation
(TD)
40 20 0 80
Optical texture Anisotropic texture
60 40 20
Inertﾄ texture
0 0
1
2
3
4
Log (MF/ddpm)
Fig.2 Relation between the maximum fluidity (MF) of coal blend and coke strength, and Relation between MF and other dominant factors in coke strength
3.2 Dilatation property in HPC addition As mentioned above, the dominant factors were improved accordingly with increase in the fluidity by HPC addition. Especially, increase in the total dilatation was strongly appeared by HPC addition. Figure 3 shows the behaviors of shrinkage and dilatation of coal blend in the dilatometer tests using Base-1 and Base-2. Base-1 shows sufficient behavior, it indicates nearly 40 % of dilatation after shrinkage. The shrinkage means the increasing in bulk density of coal by softening and fluidizing. On the other hand, base-2 shows insufficient behavior, dilatation is not appeared after shirincage without HPC. It can be seen that the large difference of the dilatation behaviors between Base-1 and Base-2 causes the large difference of DI values of coke between Base-1 and Base-2, about 87 and 78 respectively. When HPC is added to Base-2, dilatation is strongly increased, and both shrinkage and dilatation temperatures are shifted to lower
300
350
400
450
500
Displacement [%]
dilatation
Base-1 Base-1 + HPC 5% Base-1 + HPC 10%
Total dilatation (TD)
25 20 15 10 5 0 -5 -10 -15 -20 -25
shrinkage
Displacement [%]
temperatures according to the changing in Gieseler curves. 25 20 15 10 5 0 -5 -10 -15 -20 -25 300
Temperature [ºC]
Base-2 Base-2 + HPC 5% Base-2 + HPC 10% Base-2 + HPC 20%
350
400 450 Temperature [ºC]
500
Fig.3 Shrinkage and dilatation behaviors of coal blends and HPC addition in the dilatation tests
Nomura, et al.4) reported that the coke strength largely depended on the effect of filling the void of coal. In the view point of that, the dilatation property of coal and the bulk density at the charging should be equivalent effect, and he defined the void filling ability as following equation; Void filling ability [-] = Specific dilatation volume [cm3/g]･Charging bulk density [g/cm3] In that equation, specific dilatation volume expresses the volume of a unit mass of coal after expansion using the dilatation rate in the dilatometer test (Fig.3).
In reference to that, we examined the contribution to the coke strength of the “void filling ability” by HPC addition. Figure 4 shows the changing in the coke strength as a function of the dilatation rate. Coke strength increases with increase in the dilatation rate, and the series of Base1 and Base-2 can be plotted on a same line. Looking at the same dilatation rate, the higher charging bulk density is plotted at higher coke strength. Figure 5 shows relation between the “Void filling ability” and coke strength. The influences of the dilatation and the bulk density can be expressed on a same line. Therefore, the increasing mechanism of coke strength by HPC addition can be also explained as the void filling ability. That result conforms to the principle of a coal blending theory to satisfy the enough conditions of caking properties. The relation can be approximated following quadratic equation in this study; DI15015 = -31.92・(Void filling ability-1.2)2 + 88.8 For example, when the coal charging density is 0.85 g/cm3, the highest coke strength (DI15015=88.8) will be obtained by setting a blending condition of HPC to give the dilatation rate at around 30 % in this study.
Coke strength , DI15015
90
0.95 0.80
85
0.72 Coal charging density (g/cm 3 )
80
Base-2
Base-1
75 -30 -20 -10
0
10
20
30
Dilatation [vol%] Fig.4 Changing in the coke strength as a function of the dilatation rate
Coke strength , DI150 15
90 Coal charging density g/cm3
85
Base-1 Base-1 Base-1 Base-1 Base-1 Base-2 Base-2
80
1.01 0.95 0.90 0.82 0.80 0.85 0.80
75 0.4
0.6 0.8
1.0
1.2 1.4
1.6
Void filling ability [-] Fig.5 Relation between the “Void filling ability” and coke strength 4. Conclusions Significant improvement in the thermoplasticity of coal blends are observed by HPC addition, especially with high blending ratio of none or slightly caking coals. HPC gives a large effect to filling void of coal by expanding. The changing in the coke strength is quantitatively investigated as the functions of that factor in this study. HPC is expected to be a high performance caking additive to widen the coal resources to lower rank coals for blast furnace coke. Acknowledgement This study has been carried out as a part of Japanese national project called COURSE50 (CO2 Ultimate Reduction in Steelmaking process by innovative technology for cool Earth 50). We thank NEDO for the support of this study. References [1] Okuyama N, Komatsu N, Kaneko T, Shigehisa T, Tsuruya S,: Fuel Processing Technology 85 (2004) 947967 [2] Okuyama N, Komatsu N, Kaneko T, Shigehisa T, Tsuruya S,: Coal Preparation, 25 (2005) 295-311 [3] Okuyama N, et al., Proceedings in the 27th Annual International Pittsburgh Coal conference, (2010), Istanbul [4] Nomura S, Arima T, Kato K, Yamaguchi K: NIPPON STEEL TECHNICAL REPORT No. 94 (2006)
Oviedo ICCS&T 2011. Extended Abstract
Coal Drying and Dewatering for Power Generation – Current status, Research and Development Needs David Stokie1, Jianglong Yu2, Anthony Auxilio1 and Sankar Bhattacharya1 1
Department of Chemical Engineering, Monash University, Victoria, Australia [email protected] 2 Key Laboratory of Advanced Coking Technology, Liaoning Province, School of Chemical Engineering, University of Science and Technology Liaoning, Anshan (114051), P.R. China Abstract Around 45% of the world’s coal has either high-moisture or high-ash levels resulting often in inefficient power plants using these coals. There is a strong need to develop less energy-intensive coal drying technologies to improve the efficiency of plants using highmoisture coals. While some efforts in coal drying are in progress in Australia, Germany and the USA, accelerating these efforts into large-scale integrated demonstration is important. Success in developing more efficient coal drying and beneficiation technologies will promote the wider use of the high-moisture low-rank coals in efficient technologies such as either ultra-supercritical pulverised coal, and IGCC application for power and co-production of chemicals and fuels. This paper discusses the status of major drying technologies and identifies the key R&D needs for their wider commercial deployment. It does not deal with modelling of coal drying. 1. Introduction Low-rank coals containing high-moisture (30-70% on as-received weight basis) represent a significant resource worldwide. An estimated 45% of world’s coal reserves are lignites, some of which are referred to as brown coal. Most brown coals are cheap, are low in ash and sulphur contents, but have high moisture contents up to 70% on asreceived basis. Low-rank coals represent the major source for power generation in several countries – Australia (~50%), Germany (~75%), Greece (~90%), Poland (~55%), Russia (~45%), the USA (~10%) are examples. Figure 1 identifies the major countries that use high ash and/or high-moisture coals, and indicative ranges of moisture content, ash content and calorific values of lignites in those countries.
Coal pre-drying is an important step towards improving the efficiency of both existing and new pulverised coal-fired power plants using high-moisture coals. In general the efficiency of a unit using coal drops by about 4% point and 9% point when coal moisture content increases from 10% to 40% and 60% respectively. This is quite significant as 1%-point increase in efficiency can often result in up to 2.5%-point
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Oviedo ICCS&T 2011. Extended Abstract
reduction in CO2 emission. Apart from efficiency reduction, high moisture increases coal handling feed rate, auxiliary power in coal handling systems and pulverisers, and plant operating and maintenance costs.
70 6
1 moisture content of raw coal, %
60 4 Mj/kg
5
9
3
50
7 8
2
40
12 30
8 Mj/kg
10
4
12 Mj/kg
20
16 Mj/kg
11 10
20 Mj/kg
0 0
5
10
15
20
25
30
ash content of raw coal, %
Figure 1: High-ash and/or High-moisture containing coals often termed as lignites – their location, and calorific values (LHV, Mj/kg); country labels as follows: 1: Australia; 2: Indonesia; 3: India; 4: USA (Texas, North Dakota); 5: Germany; 6: Greece; 7: Spain; 8: Poland; 9: Czech Republic; 10: China; 11: Turkey; 12: Romania Compiled from various sources
Figure 2: Illustration of boiler size variation with moisture content in coal (Allardice, 1991) However, drying high-moisture coals increases the risk of spontaneous combustion as, due to their high oxygen content, they are usually more reactive than hard coals.
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Oviedo ICCS&T 2011. Extended Abstract
Therefore, in most power plants using high-moisture coals, drying has to be carried out immediately prior to combustion, i.e. in and around the mill, by recirculating some of the flue gases from the upper part of the boiler. This requires the boiler to be substantially larger to cope with the water vapour; As illustrated in Fugure 2, the higher the moisture content, the larger the boiler. In addition, to handle the additional volume of water vapour, the fan power requirement would be higher resulting in higher auxiliary power requirement and reduced efficiency. If high-moisture coal could be pre-dried, the boiler size could be smaller; and if low grade or waste heat could be used for drying, the boiler efficiency could be higher as well. The objective of this short paper is to summarise the status of different coal drying and dewatering technologies, identify the technical challenges and the immediate needs for research and development. Modelling of coal drying is not discussed in this paper. 2. Water in brown coal, physical nature of brown coal, and major variables affecting drying Moisture occurs in the forms of: surface water – which is weakly attracted to the particle surface, capillary water – found in capillaries and cracks in the coal structure, interparticle water – found in the space between two particles and interior water – dispersed through the coal structure. This is highlighted in figure 3, taken from (Muthusamy 2009).
Figure 3: Different forms of water associated with coal. (Karthikeyan, 2009)
Moisture in coal can also be described as freeze-able and non-freeze-able water. Freezeable water is weakly bound to the coal surface and is found in the bulk phase and the pore water (Allardice 2004). In contrast non-freeze-able water is molecularly bonded to the coal. Through hydrogen bonding (in conjunction with acidic functional groups, such
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as carboxylic acid) these bonds allow the moisture to be retained in the coal structure beyond the bulk water removal phase. While it varies from coal to coal, chemical bonding constitutes approximately 20% of the water in Victorian brown coal. Moisture removal from coal can be classified as evaporative (where the water comes out as vapour and can be condensed later) or dewatering, depending on how water is released from the coal matrix. Key parameters that affect drying rate and final moisture content of a dried brown coal include following: pressure and temperature of the drying process, flow rate of drying medium, relative humidity, residence time inside the dryer, size and physical natures of the particles to be dried. The kinetics of water removal depends on the moisture within the coal, as illustrated in figure 4. During the first phase, the surface and intermediate capillary water is evaporated. As moisture content decreases, the internal moisture has to be diffused out overcoming the capillary forces, and the drying rate decreases due to the increased binding forces.
Figure 4: Drying rate of coal with residual moisture content
Drying or dewatering rate and hence the design of a coal dryer is affected by several factors, apart from the the nature of the bonding between the organic matter and the water - the physical nature of coal and its porosity, particle size, pressure, temperature. Figure 5 shows the morphology of a brown coal particle in three micrographs; these clearly show the woody character as micrographs are taken at higher resolutions. Such particles are difficult to fluidize resulting in improper mixing in a steam fluidized bed dryer and therefore ineffective drying of the mass of coal inside the dryer. Woody particles constitute up to 5% of some Victorian brown coals.
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Oviedo ICCS&T 2011. Extended Abstract
Figure 5: Morphology of some of the brown coals – from left to right the electron micrographs show progressively woody character of the coal surface. Figure 6 represents the effect of pressure and particle size on heat transfer during drying of German brown coal in a steam fluidized bed dryer. Higher heat transfer rate is always desirable, as it results in smaller dryers. Smaller particles and higher pressure in general tend to increase heat transfer. However, smaller particles are not always amenable to easier fluidization. Additionally, higher pressure inside the dryer can make fluidization difficult; it also means extraction of high pressure steam, which will reduce the overall efficiency of a process in which the dryer is integrated. Therefore, the choice of particle size, pressure and temperature is always a balance between the degree of drying and the overall efficiency of the overall process.
Figure 6: Effect of particle size and pressure on heat transfer during steam fluidized bed drying of German brown coal (Hoehne et al, 2009) 3. Major technologies - brief description and their status The major technologies for coal drying and dewatering include the following: •
Fluidized bed dryer – steam or hot air fluidized, heating could be steam or hot gas
•
Entrained flow dryer – with high volumes of low-oxygen containg flue or fuel gas
•
Mechanical thermal expression dewatering
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Oviedo ICCS&T 2011. Extended Abstract
•
Hydrothermal dewatering
The subsequent sections present further discussions on the above technolgies and their current status. 3.1 Steam fluidized bed drying High-moisture coals are prone to spontaneous combustion when dried. They should preferably be dried in the absence of oxygen – or, alternatively, at a low oxygen level at lower temperatures - in an inert medium such as steam, which is readily available in a power station. Steam fluidized bed drying with in-bed heating was invented by Professor Owen Potter at Monash University. Steam drying underwent extensive testing and development in Germany, and to a lesser extent in Australia, between 1990 and 2002 (von Bargen, 2007). Figure 7 shows the schematic of a steam fluidized bed dryer. In a steam fluidized bed dryer, raw coal is fluidized by steam, and heat is supplied through immersed tubes using high temperature steam. Usually, a temperature gradient of around 50°C between the heating steam and the dryer bed is preferred to ensure an optimum level of drying and drying time. This means for drying to be accomplished at atmospheric pressure (around 100°C saturation temperature in the bed), the heating steam has to be around 5bar. This steam can potentially be supplied from low-pressure turbines in a coal-fired plant. Variations of the process shown in Figure 7 are possible, e.g. the vapour compressor may be completely dispensed with and the vapour either released into the atmosphere or used for thermal recuperation. In such a case, the heating steam, which is in the immersed coil, could be sourced from the steam cycle of the plant. The volume of the dryer and the level of drying that can be achieved in a steam fluidized bed dryer depend on a number of factors, including: • The conditions of the steam used for heating • The particle size of raw coal feed, which would, in turn, affect the drying time • The fluidisation velocity, which is important to ensure maximum contact between the heating steam and the particles. There are also issues that would be important in scaling up the capacity of the dryer. For example, though finer coal size is conducive to faster drying, effective fluidisation is made more difficult, which influences the level of drying that can be obtained within a reasonable time. RWE of Germany holds the state of the art in both brown coal combustion plant and brown coal drying in its technology (abbreviated in German as
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Oviedo ICCS&T 2011. Extended Abstract
BoA) with WTA steam fluidized bed pre-drying. In this process, raw coal is milled to a fine grain (0-2mm), which is then dried in a steam fluidized bed. Use of fine grain coal for drying reduces the size and, therefore, the cost of the dryer, it reduces the steam required to maintain fluidisation and it also requires slightly lower steam conditions than would be the case for drying coarser coal. The resulting dried coal contains 66% less than 90µm and less than 9% greater than 1mm (Klutz et al., 2006), while coal moisture content reduces from about 50% to between 10 and 18%, prior to feeding into the mills. Figure 7 shows the general schematic of the WTA steam fluidized bed dryer.
Large-scale demonstration of the WTA process is being carried out by RWE at one of their supercritical units, Niederaussem K, where 25% of Unit K’s input fuel is being treated; energy is saved by feeding only low-grade heat (120°C), in the form of low pressure steam, to fluidise and directly dry the coal. Much of the latent heat from the liberated water vapour is recovered in a heater and used in the system. Results indicate that the unit achieves a much better efficiency than previous lignite units because of the efficiency maximising measures of the plant’s ‘BoA’ or optimised efficiency technology system.
Figure 7: General schematic of WTA steam fluidized bed drying Courtesy RWE Power In addition to the demonstration at Niederaussem K, the WTA technology is also planned for demonstration at the Hazelwood Power Station in Victoria, Australia (Innocenzi, 2008). A WTA drier is planned for retrofitting to an existing 200MWe unit to dry 50% of the original feed of high-moisture coal, to reduce the moisture content 7
Oviedo ICCS&T 2011. Extended Abstract
from about 60% to 12%. The dried coal would then be co-fired with 50% high-moisture coal into the boiler. In addition to RWE, Vattenfal in conjunction with Brandenburgh Technical University is also known to develop this technology, although at a slightly positive pressure.
3.2 Hot Air (low temperature) fluidized bed drying Great River Energy (GRE) of the USA demonstrated the Lignite Fuel Enhancement System (LFES), which uses waste heat to dry fuel before being fed into the boiler. The LFES, illustrated in Figure 8, exploits the low-grade heat, which otherwise has little use. In the LFES, low temperature hot air (as opposed to steam in WTA) fluidises and heats the lignite to remove moisture from it. The air stream is cooled and humidified as it flows upward through the fluidized bed. The amount of moisture that can be removed is limited by the drying capacity of the air stream, which is supplemented by an in-bed hot water coil. GRE tested a number of lignites in a 2 tonnes/h pilot-scale dryer to evaluate the drying potential of different feedstocks. Tests confirmed the viability of coal drying and provided a basis for a larger-scale demonstration under the Clean Coal Power Initiative (CCPI). The CCPI-funded project (USDOE, 2008) progressed in phases. The full-scale integrated four-dryer system is designed to reduce the moisture content of all coal burned at the plant by 8.5%-points, from 38.5%.
Figure 8: A schematic of the Lignite Fuel Enhancement System, which uses waste heat from condenser water and flue gas Courtesy: Great River Energy 3.4 Entrained flow or flash drying In the Latrobe valley power stations, Victorian brown coals are dried by recirculating
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Oviedo ICCS&T 2011. Extended Abstract
hot flue gas through the itegrated mill/drying systems. HRL, as part of its IDGCC process, developed an entrained flow dryer using hot pressurised fuel gas. CRC for Lignite trialled a pressurised entrained flow drier, up to 10 bar, (Ross et al., 2005) using hot flue gas with low oxygen content. Developed properly, this has the potential to be used at least in a two-stage drying process wher the first stage could be entrained flow dryer while the finaly drying can be accomplished in a steam-fluidized bed dryer. White Energy in Australia is another developer of this technology, mainly for sub-bitumnous coals. 3.5
Mechanical Thermal Expression (MTE) dewatering
The MTE process removes moisture from coal without evaporation (dewaters), and is one which builds on work that was undertaken at the University of Dortmund and Diffenbacher (Bergins, 2003) in the late-nineties. It was demonstrated that if coal is heated to 150-200°C, at saturation pressure to prevent evaporation, the water in coal can then be ‘squeezed’ out by applying mechanical pressure. Raising the temperature makes the coal easier to deform under compression and makes the water more mobile by reducing its viscosity and surface tension. A substantial amount of research and development work was undertaken over 12 years in Australia on MTE dewatering by the Cooperative Research Centre for Clean Power from Lignite. This included work at bench scale and then successful development work at the 1 tonne/h scale (McIntosh and Huynh, 2005), wherein the effect of process variables (pressure, temperature, coal type, duration of heating and compression) on the extent of dewatering and throughput were established. Numerous tests were carried out in continuous and cycling batch mode. With Victorian brown coals, it was demonstrated that approximately 70% of the original coal water could be removed at around 200 C and at compression pressures of 60 to 110bar. A 15 tonne/h rig was designed, constructed and operated at the Loy Yang Power (a utility) site in Victoria. No development work is currently in progress.
3.6 Hydrothermal dewatering In hydrothermal dewatering, combination of high temperature and pressure (300 C and 100 bar) decarboxylates the coal. As CO2 is ejected the coal physically shrinks, expelling liquid water from its interstices in the process. The decarboxylation reaction is exothermic and contributes energy to the process, which uses about 2% of the wet brown
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Oviedo ICCS&T 2011. Extended Abstract
coal’s energy content for the overall process. Although hydrothermal dewatering has a long history of exploratory development, several key issues remain unknown. These include capital cost, scalability, treatment of discharged water. Australian company Exergen is currently the frontrunner in the development of this technology. Other companies working in related technologies are Ignite Energy Resources of Australia and Evergreen Energy of the USA. 3.7 Other drying technologies There are other drying technologies, which are under development or developed from small to medium scale. However, their application to large coal-fired plants for highmoisture coals is yet to be demonstrated. Some of these are: •
Rotary drier – It consists of a slightly tilted large rotating cylindrical shell which is slightly tilted from the feed to the discharge end. The drying medium can be superheated steam, inert gas or hot gas relatively free of oxygen. This is fed in either co-current or counter-current direction to the coal feed. Several companies such as Pinch Technologies and Keith Engineering from Australia and Tsukishima Kikai Ltd of Japan are front-runners of such developments.
•
Microwave drying – Most of the coals are usually transparent to microwave energy. As water is an absorbent of microwave energy, microwave energy if properly delivered can release both bound and free moisture in coal.
Table 1 shows the typical energy consumption and typical drying time for some of the major evaporative drying processes. Such information for the dewatering processes are not yet conclusively known. Table 1: Comparison of energy consumption and typical drying time for major drying processes (Wilson et al, 1992; Karthikeyan et al., 2009) Type of dryer
Steam fluidized dryer
Energy consumption, kJ/kg Drying time to water equilibrium moisture minutes 3100-400 30-40 Lower values achievable with process integration
Hot air or gas fluidized dryer Rotary dryer
3100 3700
1-5 15-40
4. Practical issues and R&D needs
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Oviedo ICCS&T 2011. Extended Abstract
Coal drying technologies, once commercialised, will make the vast resource of low-rank coals of varying moisture contents much more attractive for utilisation in new power generation units or for co-producton purposes. Significant changes to the structure and physical properties of coal do happen depending on the drying process; chemical properties also get upgraded. However, most high-oxygen containing low-rank coals will remain prone to spontaneous combustion after drying. It is, therefore, likely that pre-drying would still be carried out close to the power stations to avoid problems relating to spontaneous combustion. Disintegration of particles during drying is also an important practical issue. Reduction of particle size depends on the drying process, drying temperature and the type of coal. Experience with pilot-scale hot-gas drying (Ross et al, 2005) indicate that up to 20% finer particle size incerase takes place during drying of Victirian brown coal. Finer particles theoretically dry at a faster rate, but elutriate away from the dryer spending less residence time in the dryer. They also widen the size distribution of the particles being dried, making the design and operation of a dryer difficult. Finer particles are also known to be more hygroscopic, affecting their storage and handling properties. Generation of fines during drying, moisture readsorption and spontaneous combustion are all inter-related issues. Singnificant work have been done (Fei et al, 2009) on spontaneous combustion, but targeted experiments as a function of drying processes (in particular hot-gas drying and steam fluidized bed drying) are required to assess these issues, and the design of appropriate dryers. Other major and immediate needs for development of drying are: •
a reliable feeding system at high pressure, >25 bar, for high-moisture coals to used in gasification systems
•
drying kinetics of fine coal particles using waste heat, low-oxygen flue gas, fuel gas or low-grade steam and changes in their physical structure during the drying process
While all of the aforementioned issues are being addressed to varying extents, there is an urgent need to test the technologies at full-scale on coal-fired power plant.
5. Concluding Comments Drying is the key process before high moisture coals can be used in any applications. There are plenty of drying options for different range of applications. The type of application - whether the application is integrated power generation requiring operation at high availability, or fuels and chemical production through gasification requiring low 11
Oviedo ICCS&T 2011. Extended Abstract
availability - determines the type of coal drying technology needed. It also depends on whether partial drying is acceptable in situation such as retrofit for power generation. If partial drying is acceptable in the first instance, water can potentially be recovered from the flue gas through membranes, which are still in the developmental stage. For power generation purposes, steam fluidized bed drying is decidedly the front-runner and its demonstration for drying of high-moisture coal (~60% moisture) to commercial scale should be vigorously pursued. However, there are certain factors wich will influence commercial success of any technology (Godfrey, 2010). These, apart from maturity and scalability mentioned earlier, include footprint of the drying plant, capital and operating cost, abd capability of integrating with the overall coal conversion process.
Despite significant fundamental research in drying for all sorts of solid materials, coal drying has not yet been fully commercialized to a large scale for power industry application. Scaling-up of steam fluidized bed dryer to large commercial scale appears to be the major stumbling block before its wider application in the power industry. The application of computational fluid dynamics modeling along with targeted measurements in smaller rigs will significantly help in this matter. It is also necessary to generate fundamental data on drying kinetics from a range of coals under different conditions of drying temperature, particle size and heating medium, in particular steam. While modeling effort will shed insights into further development, one should also consider different types of fluidized bed dryer, and two-stage drying. In two-stage drying, first stage could be an entrained-flow dryer using hot fuel or flue gas, while final drying will be carried out in a smaller fluidized bed dryer with less scaling-up and fluidization issues. Incentives are required to accelerate the development and commercialisation of coal predrying technologies to full scale. International cooperation is also important among technology developers and utilities that use high-moisture coals.
Acknowledgement The authors acknowledge the Department of Resource, Energy and Tourism, Australia for their support of this project under its Joint Cordination Group (JCG) Clean Coal Technology program. References
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D. Allardice, The water in brown coal, Chapter 3 in RA Durie (ed.) The Science of Victorian Brown Coal, Butterworth-Heinemann, Oxford 1991, pages 102-150. C. Bergins, Kinetics and mechanism during mechanical/thermal dewatering of lignite, Fuel, V83, 2004, pages 355-364 Y. Fei, A. Abd Aziz, S. Nasir, W.R. Jackson, M. Marshall, J. Hulston, A.L. Chaffee, The spontaneous combustion behavior of some low rank coals and a range of dried products , Fuel, Volume 88, Issue 9,2009, pages 1650-1655 B. Godfrey, Recent Development s in innovative drying technologies, International Lowrank Coal Symposium, Melbourne, April, 2010 O. Hoehne, S. lechner, M. Schreiber, HJ Krautz, Drying of lignite in a pressurized steam fluidized bed – theory and experiments, Drying Technology, V28, Issue 1, 2010, pages 5-19 M. Karthikeyan, J. Kuma, C. Hoe, D. Ngo, Factors affecting quality of dried low-rank coals, Drying Technology, V25, Issue 10, 2007, pages 1601-1611 M. Karthikeyan, W. Zhonghua, A. Mujumdar, Low-rank coal drying technologies – Current Status and new Developments, Drying Technology, V27, Issue 3, 2009, pages 403-415 M. Mcintosh and D. Huynh, Pre-drying of high moisture content Australian brown coal for power generation, Proc 22nd Annual International Coal Preparation and Aggregae processing Conference, Lexington, USA, 2005 D. Ross, S. Doguparthy, D. Huynh, M. McIntosh, Pressurised flash drying of Yallourn lignite, Fuel, Volume 84, Issue 1, January 2005, pages 47-52.
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Pre-combustion CO2 capture: Laboratory- and Bench-scale Studies of a Sweet Water-Gas-Shift Catalyst for H2 and CO2 production J.M. Sánchez, M. Maroño, D. Cillero, L. Montenegro, E. Ruiz CIEMAT, Combustion & Gasification Division, Avenida Complutense, 22, 28040 Madrid (SPAIN; corresponding author: [email protected]
Abstract CIEMAT is currently engaged in several R&D projects dealing with pre-combustion CO2 capture and H2 production studies. In the pre-combustion CO2 capture approach, the carbon of the fuel is removed prior to combustion by partial oxidation or gasification, followed by steam reforming and water-gas shift (WGS) so that a CO2 and H2-rich gas is produced. Both components are subsequently separated usually by means of chemical or physical scrubbing. High purity H2 production often includes a PSA unit. This “conventional” concept is now being proved at a 14 MWth pilot plant in ELCOGAS IGCC. Nonetheless, a lot of R&D effort is still required, specifically focusing on the effect of process conditions on the activity of shift catalysts and looking into enhanced technologies for the reaction. In the first section of this paper, the activity of an iron-chromium-based catalyst for the water-gas-shift reaction is studied. To meet the goals set in the PSE-CO2 projects, the influence on the activity of the sweet shift catalyst of the main operating parameters temperature, space velocity, excess steam, and gas composition- is studied at laboratory scale. Results are presented and discussed. Best operating conditions to maximize selective shift conversion of CO to CO2 and H2 are determined. Lab-scale results also provide base-case data for the CAPHIGAS and FECUNDUS projects, in which a hybrid system for CO2 capture with H2 production is being studied, comprising the WGS catalyst, a CO2-selective sorbent, and a selective membrane for hydrogen removal in a single reactor. Based on the results obtained at lab scale, the second part of this paper deals with subsequent WGS testing at bench-scale, as part of the activities conducted by CIEMAT in the PSE-CO2 project. Preliminary results of bench-scale studies are presented.
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Oviedo ICCS&T 2011. Extended Abstract
1. Introduction In recent years, production of a hydrogen-rich gas from fossil fuels is attracting interest in power generation processes especially if coupled to CO2 capture. The approach is the so-called “pre-combustion fuel decarbonisation”, in which the carbon of the fuel is removed prior to combustion by partial oxidation or gasification, followed by steam reforming and water-gas shift (WGS) so that a CO2 and H2-rich gas is produced. Both components are subsequently separated and CO2 in the exhaust gas is captured using chemical or physical solvents (e.g. amine scrubbing). The water gas shift (WGS) reaction is, therefore, a key step on production of hydrogen and CO2-capture-ready streams from gasification. The water gas shift (WGS) reaction has been used industrially since the beginning of the 20th century in hydrogen production via coal gasification, as a part of ammonia synthesis by the Haber-Bosch process [1]. It is a mildly exothermic reaction, limited by chemical equilibrium, thermodynamically favoured at low temperatures, and kinetically favoured at high temperatures. To achieve a significant conversion at intermediate temperature, a catalyst is required. According to literature, WGS catalysts can be classified in three main groups: High temperature sweet shift catalysts, usually based on iron oxide with addition of chromium oxide, which operate at inlet temperatures in the range of approximately 320360ºC. Studies are being addressed to the identification of new promoters that are able to increase catalytic activity over larger range of temperature, such as copper [2], [3], rhodium, platinum, nickel, cobalt, manganese, palladium, [4] or to the formulation of Cr-free catalysts, replacing chromium e.g. by aluminium, manganese or cobalt [5] Low temperature sweet shift catalysts, typically based on a Cu/ZnO/Al2O3 or Cu/ZnO/Cr2O3 structure, which are used as second shift stage conducted in the temperature range from 150-250ºC for further decreasing of CO concentration and increasing hydrogen production. R&D efforts deal with the development of advanced catalysts, for instance transition metal supported-ceria (CeO2), including noble metal catalysts such as platinum, rhodium, palladium or gold [6-10]. Sulphur-tolerant WGS catalysts: During (co-)gasification of rich sulphur fuels, e.g. mixtures of coal and pet-coke, H2S concentration in the gas can be as high as 1% v/v. To prevent poisoning of Fe-Cr and especially Cu-Zn-Al catalysts, sulphur is removed before the water gas shift reactor. Shifting the sour gas, using sulphur-tolerant catalysts
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Oviedo ICCS&T 2011. Extended Abstract
can be an attractive option. Explored catalysts include those based on cobaltmolybdenum compositions [11], the use of Pt/ZrO2 materials [12], [13] or the use of carbides as promising candidates such as molybdenum carbide [14]. In this paper, the activity of an iron-chromium-based catalyst for the water-gas-shift reaction is studied, first at laboratory scale and then at bench scale level.
2. Experimental section 2.1 Catalyst A commercial Fe-Cr-Cu-based sweet WGS catalyst in pellet form was selected for water gas shift studies. The catalyst was used as received in cylindrical tablets of 6x3 mm. For testing at lab scale, 5 g of fresh catalyst was used in every test. For the bench-scale studies the amount of catalyst loaded in the reactor was 2000 g. 2.2 Test rigs Activity of the catalyst was examined firstly at lab-scale in a Microactivity Pro Unit. The unit can work at up to 700ºC and 30 bar. Maximum operating gas flow rate is 4.5 Nl/min. A detailed description of the rig can be found elsewhere [15].
Figure 1. Picture of the bench-scale test rig For testing at bench-scale an existing bench-scale test rig was adapted to WGS studies. This plant had been used extensively for hot gas desulphurisation. It is a high temperature, high-pressure (HTHP) bench scale facility for sorbent and catalyst testing. The plant, shown in Figure 1, can treat up to 20 Nm3/h of a gas mixture simulating the
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Oviedo ICCS&T 2011. Extended Abstract
composition of gasification gases. It is designed to operate at a maximum temperature of 700ºC and a pressure of 30 bar. Main modifications to the existing facility included a new pump with higher capacity to deliver the flow-rate of water/steam demanded by the WGS reaction and an automatic system for removal of the excess water, installed downstream the reactor. Complete description of the rig can be found in [16]. Gas composition at the reactor inlet and outlet was analysed by gas chromatography. 2.3 Experimental programme Studies at lab scale were focused on the optimisation of shift reactor operating conditions to maximise selective conversion of CO to CO2 and production of hydrogen while ensuring catalyst lifetime. The effect of main operating parameters on the activity and selectivity of the catalyst was determined in the following range: - Temperature, (T): From 250ºC to 380ºC - Gas hourly space velocity (GHSV): From 2500 h-1 to 5000 h-1 - Steam to CO molar ratio (R) between 1 and 3 Despite the fact that WGS is a moderately exothermic reaction, laboratory experiments presented in this work can be considered isothermal, because heat release was almost negligible, due to the small amount of catalyst and the low value of feed gas flow-rate. WGS tests were performed using different gas compositions as shown in Table 1. Table 1. WGS catalytic activity tests: Gas composition (% v/v dry basis) Component H2 CO CO2 N2
M1 23 60 4 13
M2 23 60 17 -
M3 60 40
M4 60 40 -
Under the scope of the PSE-CO2 and CAPHIGAS projects, M1 composition was set according to the gas composition expected for entrained flow oxygen gasification at ELCOGAS IGCC plant. M2 is similar to M1 but it contains more CO2 and is the composition expected for oxygen-carbon dioxide gasification in fluidised bed, which is the approach in the FECUNDUS project. For comparison purposes, M3 and M4 consist of binary mixtures of CO in N2 and CO2 respectively. Studies at bench scale focused on evaluating the performance and the activity of the catalyst at operating conditions which had proved to be successful in laboratory
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Oviedo ICCS&T 2011. Extended Abstract
experiments. The study has included the effect of steam to CO ratio and the effect of gas composition on the performance of the catalyst. In contrast to the laboratory scale studies, runs conducted at bench-scale level proceeded adiabatically.
3. Results and Discussion 3.1 Laboratory-scale studies The catalyst started to be active around 280-300 ºC, depending on operating conditions and it did not require any activation or became active on-stream. CO conversion increased on increasing temperature and reached a maximum between 350 ºC and 380 ºC, as Figure 2 shows.
100 90
CO Conversion (% mol)
80 70 60 50
20bar
40 30 20 10 0 300
320
340
360
380
Temperature (ºC)
Figure 2. Effect of temperature on CO conversion (SV=4715 h-1, P= 20 bar, Gas composition (%v/v): N2 13 %, CO 60 %, H2 23 %, and CO2 4 %, R=3)
In all tests, despite the high CO content in the feed gas -60 % v/v dry basis- the catalyst showed very good performance at intermediate temperatures 320 ºC-380 ºC rising hydrogen content in the gas from 23%v/v (dry basis) at the reactor inlet to around 4850 %v/v at the reactor outlet and reducing CO concentration at the reactor outlet to less than 3 % v/v. Regarding the influence of space velocity, higher GHSV values led to lower CO conversion and a higher temperature was required to achieve a given CO conversion as shown in Figure 3.
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Oviedo ICCS&T 2011. Extended Abstract
CO conversion for different GHSV values. Steam/CO=2 100 90
Equilibrium R=2
CO conversion (% mol)
80
-1
SV= 2885 h ; R= 2
70 60 -1
50
SV= 4715 h ; R= 2
40 30 20 10 0 -10 240
260
280
300
320
340
360
380
400
420
Temperature (ºC)
Figure 3. Effect of space velocity and on CO conversion Gas composition (%v/v): N2 13 %, CO 60 %, H2 23 %, and CO2 4 %
The effect of steam to carbon monoxide ratio was also evaluated. At least a steam to CO ratio of 1 is required to convert all CO into CO2. It is known, however, that running the WGS reaction at stoichiometric steam to CO ratio can promote undesirable secondary reactions, such as Boudouard reaction or methane formation. In order to prevent secondary reactions and to drive the reaction thermodynamically to the products side, excess steam is often used, though consequently, this poses a penalty on the energy efficiency of the process. In this study steam to CO ratio was varied from 1 to 3. Effect of excess steam on CO conversion 100
Equilib R=2
90
R=2 CO conversion (% mol)
80 70
R= 1
60 50 40 30 20
-1
Gas space velocity SV = 2885 h
10 0 240
260
280
300
320
340
360
380
400
Temperature (ºC)
Figure 4. Effect of steam to CO ratio on CO conversion Gas composition (%v/v): N2 13 %, CO 60 %, H2 23 %, and CO2 4 %, P=10 bar
At low space velocity a noticeable enhancement of catalytic activity is gained on doubling steam to CO ratio as Figure 4 shows. Secondary reactions such as methane formation was not found to be taking place, even when steam to CO ratio was set at the
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Oviedo ICCS&T 2011. Extended Abstract
stoichiometric value (R=1). However at higher space velocities (SV> 4715 h-1) and low steam to CO ratio, R ≤ 2, carbon built on the surface of the catalyst and methane formation was detected, so that a steam to CO ratio above 2 is required. A further increase of the steam to CO ratio (R=3) did not result in significantly higher CO conversion. The effect of CO2 content in the feed gas on the activity of the catalyst is shown in Figure 5. Above 330 ºC, increasing the content of CO2 resulted in a decrease of CO conversion, which was directly related to CO2 content in the feed gas. At 300 ºC, there was not a clear relationship between CO2 content in the feed gas and CO conversion, what might indicate that the presence of H2 was also influencing CO conversion.
Influence of feed gas composition 100 -1
SV= 4715 h , R H2O/CO=3, P= atmospheric 90
xCO (% mol/mol)
80
70
CO 60%; N2 40% CO 60%; H2 23%; CO2 4%; N2 13%
60
CO 60%; H2 23%; CO2 17% CO 60%; CO2 40%
50
40 300
320
340
360
380
Temperature (ºC)
Figure 5. Effect of gas composition and CO2 content on CO conversion SV=4715 h-1, R=3, P=1 bar
3.2 Bench scale studies Several studies were performed to examine the activity of the catalyst at bench-scale. The system showed an adiabatic behaviour, with temperature increasing due to the exothermic nature of the WGS reaction. In all tests, CO conversion increases quickly as soon as the reaction starts, what happens around 310 ºC, and so does the catalyst bed temperature, though never exceeding 560 ºC, what ensures that the catalyst is not damaged due to sintering. As an example, Figure 6 shows the test carried out at SV=4715 h-1, R=3. As can be seen, temperature and CO conversión is stabilized and maintained over time.
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Oviedo ICCS&T 2011. Extended Abstract
100
540 520
90
500 480
Temperature (ºC)
70
440 Sv = 4715 h-1 R=3
420 400
60
Bed temperature CO Conversion
380 360
50 40
340 30
320 300
20
CO Conversion (% mol)
80
460
280 10
260 240 0
30
60
90
120
150
180
210
240
0 300
270
Time (min)
Figure 6. Evolution of temperature profile and CO conversion during WGS study at bench scale. Gas composition (%v/v d.b.): N2 13 %, CO 60 %, H2 23 %, and CO2 4 %, P=10 bar, SV=4715 h-1, R=3, inlet T= 331 ºC
The effect of steam to CO ratio on CO conversion was determined for three R values (R = 1.61, 2 and 3). As shown on Figure 7, slightly higher CO conversion is achieved on increasing steam to carbon as predicted by the laboratory study. Secondary reactions were not found to be occurring. CO concentration at the reactor outlet decreased to less than 5% when using R=3, and to less than 9% for R=1.61. Average increase of bed temperature (ΔT) ranged between 210 and 230ºC. Average ΔT
280 260 240
100
Average CO Conversion Sv = 4715 h-1 R = 1,61
Sv = 4715 h-1 R=2
Sv = 4715 h-1 R=3
80
200
ΔT (ºC)
180
60
160 140 120
40
100 80 60
20
CO Conversion (% mol)
220
40 20 0
0
Figure 7. Bench scale -WGS study: Effect of steam to CO ratio on bed temperature and CO conversion Gas composition (%v/v d.b.): N2 13 %, CO 60 %, H2 23 %, and CO2 4 %, P=10 bar, SV=4715 h-1, R=1.61, 2 & 3, inlet T between 294 and 344 ºC
4. Conclusions The activity for the water gas shift reaction of an iron-chromium-based catalyst has been studied on laboratory and bench-scale. Despite the high CO content in the feed gas, by
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Oviedo ICCS&T 2011. Extended Abstract
choosing the right shift reactor conditions, CO concentration at the reactor outlet reached values below 3%, whereas H2 increased up to above 50% v/v (dry basis). Isothermal tests conducted at laboratory-scale experiments show that increasing space velocity leads to lower CO conversion and a higher temperature is required to achieve a given CO conversion. A higher H2O/CO ratio leads to higher carbon monoxide conversion, and without undesired secondary reactions. Higher CO2 content in the feed gas inhibits slightly the activity of the catalyst, especially above 300 ºC. Adiabatic tests carried out at bench-scale level shows that with a gas temperature at the reactor inlet of about 310 ºC, the WGS proceeds successfully. In the less efficient case, CO content at the reactor outlet reaches values below 9% v/v (dry basis), that is overall conversion is about 85 % (%mol/mol). Maximum gas temperature due to the heat release of the WGS reaction stays below 560 ºC and therefore loss of activity due to sintering is not expected. As happened for laboratory scale experiments, increasing steam to carbon monoxide ratio leads to slightly higher CO conversion which makes possible to achieve a content of CO at the reactor outlet as low as 5%.
Acknowledgement This research is financed by the European Union, Research Fund for Coal and Steel, FECUNDUS project (RFCS-CT-2010-00009), and by the Spanish Ministry of Science and Innovation through PSE-CO2 project, (PSE-120000-2008-29) and CAPHIGAS project, (ENE2009-08002).
References [1] Topham SA. The history of the catalytic synthesis of ammonia. In: Anderson JR, Boudart M, editors. Catalysis Science and Technology, New York: Springer Verlag; 1985, p. 1–50. [2] Andreev A, Idakiev V, Mihajlova D, Shopov D. Iron-based catalysts for the water gas shift reaction promoted by first-row transition metal oxides. Appl Catal 1986;22:385–7. [3] Edwards MA, Whittle DM, Rhodes C, Ward AM, Rohan D, Shannon MD et al. Microestructural studies of the copper promoted iron oxide/chromia water gas shift catalyst. Phys Chem Chem Phys 2002;4:3902–8. [4] Lei Y, Trimm DL, Cant NW, Tran T. Novel Fe-Cr Oxide catalyst for water gas shift reaction. In: Coombs S, Dicks A, editors. Proceedings of the First Nanomaterials Conference. NSW, Australia; 2004, p. 21–23.
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Oviedo ICCS&T 2011. Extended Abstract [5] Natesakhawat S, Wang X, Ozkan US. High-Temperature Water-Gas Shift Reaction over CrFree Fe-Al Catalysts Promoted with First Row Transition Metals. In: AIChE editors ISBN 08169-0965-2. Proceedings of the AIChE Annual Meeting 2004. Austin, Texas; 2004, paper 24d. [6] Jacobs G, Patterson PM, Williams L, Chenu E, Sparks D, Thomas G et al. Water gas shift: in situ spectroscopic studies of noble metal promoted ceria catalysts for CO removal in fuel cell reformers and mechanistic implications. Appl Catal A-Gen 2004;262:177–87. [7] Tibiletti D, Bart de Graaf EA, Phen The S, Rothenberg G, Farrusseg D, Mirodatos C. Selective CO oxidation in the presence of hydrogen: fast parallel screening and mechanistic studies on ceria-based catalysts. J Catal 2004;225:489–97. [8] Luengnaruemitchai A, Osuwan S, Gulari E. Comparative studies of low-temperature watergas shift reaction over Pt/CeO2, Au/CeO2 and Au/Fe2O3 catalysts. Catal Commun 2003;4:215– 21. [9] Fu Q, Kudriavtseva S, Saltsburg H, Flytzani-Stephanopoulos M. Gold-ceria catalysts for low-temperature water gas shift reaction. Chem Eng J 2003;93:41–53. [10] Panagiotopoulou P, Kondarides DI. Effect of morphological characteristics of TiO2supported noble metal catalysts on their activity for the water gas shift reaction. J Catal 2004; 225:327–36. [11] Song C. Overview of hydrogen production options for hydrogen energy development, fuelcell processing and mitigation of CO2 emissions. In: Proceedings of the 20th International Pittsburgh Coal Conference. Pittsburgh, PA; 2003, paper Nº 40-3. [12] Xue E, O`Keeffe M, Ross JRH. Water gas shift conversion using a feed with a low steam to carbon monoxide ratio and containing sulphur. Catal Today 1996; 30:107–118. [13] Maroño M, Sánchez JM, Ruiz E, Cabanillas A. Study of the suitability of a Pt based catalyst for the upgrading of a biomass gasification syngas stream via the WGS reaction. Catal Lett 2008;126:396–406. [14] Patt J, Moon DJ, Phillips C, Thompson L. Molybdenum carbide catalysts for water gas shift. Catal Lett 2000;65:193–5. [15] Maroño M, Sánchez JM, Ruiz E. Hydrogen-rich gas production from oxygen pressurized gasification of biomass using a Fe–Cr Water Gas Shift catalyst. Int J Hydrogen Energy 2010;35: 37–45. [16] Sánchez JM, Ruiz E, Otero J. Selective Removal of Hydrogen Sulfide from Gaseous Streams Using a Zinc-Based Sorbent. Chem Eng Sci 2005;60:2977–89.
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Oviedo ICCS&T 2011. Extended Abstract
Regeneration of used alkali carbonates for removal of gaseous sulfur compounds in gasification process Slamet Raharjo1, Yasuaki Ueki2, Ryo Yoshiie1 and Ichiro Naruse1 1
Department of Mechanical Science and Engineering, Nagoya University Furo-cho, Chikusa-ku, Nagoya 464-8603, JAPAN 2 Energy Science Division, EcoTopia Science Institute, Nagoya University Furo-cho, Chikusa-ku, Nagoya 464-8603, JAPAN [email protected]
Abstract Integrated gasification combined cycle system (IGCC) is expected to play an important role in one of high efficient coal conversion technologies. However, it is necessary to remove H2S and COS almost completely to supply the gasified gas into a gas turbine as a fuel. Although molten alkali carbonates of 43 mol-Na2CO3 + 57 molK2CO3 can completely absorb them even at high temperature, the regeneration of the used molten alkali carbonates is one of the important issues in enhancing the overall efficiency more in the gasification system. Therefore, the present work studies the regeneration of the used molten alkali carbonates by using CO2 and CO2 + steam as regeneration agents. As a result, only about 8 mass% and 5 mass% of sulfur were remained as solid sulfur after the regeneration at 900 K and 773 K by using CO2 and CO2 + steam as the regeneration agents, respectively. The steam addition contributed to decreasing the regeneration temperature. The dominant gaseous sulfur produced during the regeneration was COS and H2S for CO2 and CO2 + steam, respectively. Keyword: Molten alkali carbonates, Hot gas desulfurization, Coal gasification, Regeneration
1. Introduction Coal gasification processes emit pollutants of gaseous sulfur of H2S and COS, which are hazardous and corrosive [1]. H2S and COS contained in the gasified gas may cause some failures to the subsequent units such as a gas turbine and a fuel cell. Therefore, direct utilization of the gasified gas as a fuel or a chemical feed stock requires removal of H2S and COS in the gasified gas almost completely [2]. The integrated gasification fuel cell combined cycle (IGFC) operated under the EAGLE project in Japan currently uses the cold gas clean-up system to remove the gaseous
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Oviedo ICCS&T 2011. Extended Abstract
sulfurs [3]. However, integrating the hot gas desulphurization (HGD) system offers some potential benefits including improved thermal efficiency and environmental performance, reduced capital and operating costs, and the use of a more advanced, high efficiency gas turbine [2]. The hot gas desulfurization system by using molten alkali carbonates (43 mol-Na2CO3 + 57 mol-K2CO3) was studied in our previous work [4]. It suggested that the molten alkali carbonates could completely remove the gaseous sulfurs of H2S, COS and SO2 even at high temperature of 1173 K. Reactions between the molten alkali carbonates and the gaseous sulfurs result in formation of alkali sulfides (Na2S and K2S) mainly. However, they must be converted back to the alkali carbonates in order to recycle the used solvent into the gasification system. Therefore, the present work studies the regeneration of the used molten alkali carbonates by using CO2 and CO2 + steam as the regeneration agents.
2. Experimental First, the FactSage, which could estimate chemical equilibrium solutions, was used to determine the optimum regeneration condition for the regeneration experiments. The regeneration experiments were conducted by a tube reactor made of quartz. Figure 1 shows the schematic diagram of the tube reactor employed in this study. Steam generator Quartz tube reactor
Outlet
M2S sample N2 CO2
GC-FPD
Activated carbon
Waterpump EMIA-120
Residue
Ultrapure water
Fig. 1 Schematic of experimental apparatus Nitrogen gas was used as a purge gas prior to supplying the regeneration agent. The alkali sulfide sample set inside a ceramic boat was put into the quartz tube reactor. The first holding temperature was 383K in order to remove the moisture and the crystal
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Oviedo ICCS&T 2011. Extended Abstract
water in the sample. Then, the temperature was raised to the regeneration temperature, and was kept for 30 min under nitrogen atmosphere before switching to the regeneration agent of CO2 or CO2 + steam.
It kept 60 min for the regeneration. Solid sulfur
concentration remaining in the residue after the regeneration experiments was analyzed by using a sulfur analyzer, while the gaseous sulfur compositions in the exit gas were detected by a FPD gas chromatograph.
3. Results and Discussion 3.1 Chemical equilibrium calculations to estimate the optimum regeneration conditions Figure 2 shows the results of chemical equilibrium calculation for the regeneration process by CO2 as the regeneration agent. These graphs display the production fractions of (a) sodium carbonate and (b) potassium carbonate, which correspond to the regeneration efficiency. The higher amount of sodium or potassium carbonate means that the higher amount of sodium or potassium sulfide reacted with CO2. The variation in CO2 amount has minor effect to the production fraction of sodium carbonate, whereby, the relatively higher amount of CO2 would be favorable for potassium carbonate. Additionally, the simulations suggest that the optimum regeneration temperature lies between 600 K and 900 K. Figure 3 presents the results of chemical equilibrium calculation for the regeneration process by CO2 + steam under various ratios of CO2/steam. The various ratios of CO2/steam result in a little effect on sodium carbonate formation, while CO2/steam ratio of 1/2 exhibits the highest production fraction of potassium carbonate at low temperature. The optimum range of temperature would be between 500 K and 900 K. Based on those calculation results, the experimental conditions of regeneration experiments on the tube reactor for CO2 as the regeneration agent are shown in Table 1.
3.2 Regeneration experiments Figure 4 shows the remaining solid sulfur in the sample after the regeneration experiments for CO2 as the regeneration agent. As seen from the figure, around 900 K would be the optimum regeneration temperature. At this temperature, the remaining solid sulfur becomes only 8 mass% of the total sulfur content in the sample. Based on
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Oviedo ICCS&T 2011. Extended Abstract
this result, around 92 mass% of sulfur might have been vaporized as gaseous sulfurs during regeneration.
(a) Production fraction of sodium carbonate 100
Na2CO3 (s) [mole%]
80 60 40 Alkali sulfide/Carbon dioxide=1/4 (mole ratio)
20
Alkali sulfide/Carbon dioxide=1/50 (mole ratio) 0 470 510 550 590 630 670 710 750 790 830 870 910 950 990 Temperature [K]
(b) Production fraction of potassium carbonate 100 Alkali sulfide/Carbon dioxide=1/4 (mole ratio) Alkali sulfide/Carbon dioxide=1/50 (mole ratio)
K2CO3 (s) [mole%]
80
60
40
20
0 470 510 550 590 630 670 710 750 790 830 870 910 950 990 Temperature [K]
Fig. 2 Results of chemical equilibrium calculation of regeneration reactions for CO2 as a regeneration agent Further experiments for the mixture of CO2 and steam as the regeneration agent were also carried out by using the tube reactor. The experimental condition is provided
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Oviedo ICCS&T 2011. Extended Abstract
in Table 2. The ratio 1:2 of CO2/steam was chosen, based on the chemical equilibrium calculation result.
(a) Production fraction of sodium carbonate 100
Na2CO3 (s) [mole%]
80
60 Carbon dioxide/steam=1/1 (mole ratio) 40
Carbon dioxide/steam=1/2 (mole ratio) Carbon dioxide/steam=1/4 (mole ratio)
20
Carbon dioxide/steam=2/1 (mole ratio) Carbon dioxide/steam=4/1 (mole ratio)
0 450 490 530 570 610 650 690 730 770 810 850 890 930 970 1010 Temperature [K]
(b) Production fraction of potassium carbonate 100
K2CO3 (s) [mole%]
80
60
40
Carbon dioxide/steam=1/1 (mole ratio) Carbon dioxide/steam=1/2 (mole ratio) Carbon dioxide/steam=1/4 (mole ratio)
20
Carbon dioxide/steam=2/1 (mole ratio) Carbon dioxide/steam=4/1 (mole ratio)
0 450 490 530 570 610 650 690 730 770 810 850 890 930 970 1010 Temperature [K]
Fig. 3 Results of chemical equilibrium calculation of regeneration reactions for CO2 + steam as a regeneration agent
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Oviedo ICCS&T 2011. Extended Abstract
The remaining solid sulfur after the regeneration experiments by CO2 + steam is shown in Fig. 5. The result suggests that the lowest remaining solid sulfur of around 5 mass% of the total sulfur content in the sample can be achieved at the optimum regeneration temperature of 773 K. That means around 95 mass% of sulfur is expected to be vaporized during regeneration. Table 1 Experimental conditions of tube reactor for CO2 as the regeneration agent Sample
Na2S.9H2O + K2S
Na2S/K2S mole ratio
1/1
Atmosphere
N2, changed to the regeneration agent
Particle size
< 500 μm
N2 flow rate
1 L/min
CO2 flow rate
1 L/min
1st holding temperature
383 K for 8 min
2nd holding temperature
650 K or 925 K for 30 min
Regeneration temperature
650 K or 925 K for 60 min
Remaining solid sulfur [mass%]
50 45 40 35 30 25 20 15 10 5 0 650K
773K
873K
900K
925K
Temperature [K]
Fig. 4 Results of the remaining sulfur after the regeneration experiment by CO2 as the regeneration agent
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Oviedo ICCS&T 2011. Extended Abstract
Table 2 Experimental conditions of tube reactor for CO2 + steam as the regeneration agent Sample
Na2S.9H2O + K2S
Na2S/K2S mole ratio
1/1
CO2/steam mole ratio
½
Particle size
< 500 μm
N2 flow rate
1 L/min
H2O (l) flow rate
0.57 ml/min
CO2 flow rate
0.40 L/min
1st holding temperature
383 K for 8 min
2nd holding temperature
550 K or 873 K for 30 min
Regeneration temperature
550 K or 873 K for 60 min
Remaining solid sulfur [mass%]
20 17.5 15 12.5 10 7.5 5 2.5 0 550K
650K
773K
873K
Temperature [K]
Fig. 5 Results of the remaining sulfur after the regeneration experiment by CO2 + steam as the regeneration agent Comparing between Figs. 4 and 5, the regeneration process of the addition of steam results in lower regeneration temperature, compared to that without steam. It is obvious that the steam addition contributes to decreasing the optimum regeneration
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Oviedo ICCS&T 2011. Extended Abstract
temperature, which is favourable in enhancing the overall efficiency more in the gasification system.
(a) CO2 as the regeneration agent 60 Gas Concentration [ppmV]
Carbonyl sulfide 50 Hydrogen sulfide 40
Sulfur dioxide
30 20 10 0 15'
30'
45'
60'
Sample (minutes)
(b) CO2+steam as regeneration agents
Gas Concentration [ppmV]
110 100
Carbonyl sulfide
90
Hydrogen sulfide
80
Sulfur dioxide
70 60 50 40 30 20 10 0 15'
30'
45'
60'
Sample (minutes)
Fig. 6 Gaseous sulfur concentrations in the exhaust gas during regeneration
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Oviedo ICCS&T 2011. Extended Abstract
Figure 6 shows some gaseous sulfur concentrations including COS, H2S and SO2 in the exhaust gas during regeneration at 650 K under the conditions of (a) CO2 and (b) CO2 + steam as the regeneration agent. From this figure, COS is the dominant gaseous sulfur among others for CO2. Meanwhile, H2S is the dominant gaseous sulfur for CO2 + steam. Based on those results, the expected regeneration reaction scheme for both cases can be proposed as follows: With CO2 as the regeneration agent: Na2S[s] + K2S[s] + 4CO2[g] Æ Na2CO3[s] + K2CO3[s] + 2COS[g] With CO2 + steam as regeneration agents: Na2S[s] + K2S[s] + 2CO2[g] + 2H2O[g] Æ Na2CO3[s] + K2CO3[s] + 2H2S[g]
4. Conclusions The used alkali carbonates could be regenerated by using CO2 and CO2 + steam. The lowest remaining sulfur after regeneration of around 8 mass% of the total sulfur content in the sample was achieved at around 900 K by CO2 as the regeneration agent. Meanwhile, the regeneration by a mixture of CO2 and steam resulted in lower temperature between 650 K and 773 K. For the regeneration process by CO2, COS emitted as the dominant gaseous sulfur, while H2S was dominant for CO2 + steam as the regeneration agents.
References [1]
Lew S, Sarofim AF, Stephanopoulos MF. Sulfidation of zinc titanate and zinc oxide solids. Ind Eng Chem Res 1992;31:1890–1899 .
[2]
Mitchell SC. Hot Gas Cleanup of Sulphur, Nitrogen, Minor and Trace Elements. London: IEA Coal Research; 1998.
[3]
Kimura N. EAGLE project perspective on coal utilization technology. APEC Clean Fossil Energy Technical and Policy Seminar. Philippines; 2005, Jan.
[4]
Raharjo S, Ueki Y, Yoshiie R, Naruse I. Hot gas desulfurization and regeneration characteristics with molten alkali carbonates. International Journal of Chemical Engineering and Application 2010;1:96–102 .
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Oviedo ICCS&T 2011. Extended Abstract
Step Change Adsorbents and Processes for CO2 capture “STEPCAP”
T.C. Drage1 A.I Cooper3, R Dawson3, J Jones3, C Cazorla Silva4, C.E. Snape1, L. Stevens1, X. Guo4, J. Wood2, J. Wang2 1
Department of Chemical and Environmental Engineering, Faculty of Engineering,
University of Nottingham, Nottingham, NG7 2RD, UK. Fax: 44 (0)115 951 9514115; Tel: 44 (0)115 9514099; E-mail: [email protected]. 2
Chemical Engineering, The University of Birmingham, Edgbaston, Birmingham, B15
2TT, UK. 3
Department of Chemistry, Crown Street, The University of Liverpool, Liverpool, L69
7ZD, UK 4
Department of Chemistry, University College London, London, WC1H 0AJ, UK.
Abstract STEPCAP is a multipartner consortium project, the aim of which is to develop a targeted range of novel CO2 adsorbents for carbon capture. This research into materials and process development is essential to achieve the potential cost and efficiency benefits offered by solid sorbents capture technologies over the current state of the art processes. Firstly, this paper will discuss the key materials and process challenges associated with developing solid sorbents. This will lead into a discussion of materials development in STEPCAP which is based on a fundamental understanding of adsorption processes to design and optimise material properties and form. The development and performance of the three classes of materials under development in this study, microporous polymers, surface modified hydrotalcites, and co-doped sorbents, which offer potential for a step change increase in adsorption capacity and performance over previously developed materials will be discussed. Modified hydrotalcites such as, amine modified layered double hydroxides (LDH’s) have been synthesized via the exfoliation and grafting route. In addition, novel conjugated microporous polymers synthesized through SonogashiraHagihara coupling have also been investigated and have demonstrated similar capacities. Critically, due to the hydrophobic nature of some of these adsorbents, identical performance has been observed in the presence of moisture, an advantageous property for operation in the water saturated environment of flue gases. This presentation will
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Oviedo ICCS&T 2011. Extended Abstract
also present data on the performance of these materials in simulated flue gases as well after simulated temperature swing regeneration cycles to assess the stability and lifetime of the sorbents.
1. Introduction Recognising that fossil fuels will continue globally as part of a diverse energy mix for some time[1], targets and strategies have been developed to reduce greenhouse gas emissions, for example the European Unions Sustainable Energy Technology (SET) Plan[2]. Rapid development and implementation of these strategies will be required if the warnings of potentially damaging climate change reported by the Intergovernmental Panel on Climate Change (IPCC) are to be avoided[3], a task that is made more challenging when set against the significant global increase in energy demand[1]. Europe is committed to an 80% reduction in greenhouse gas emissions by 2050[4] and similar emissions reduction targets have been proposed and committed to on a global scale[5]. The current state of the art technologies for post-combustion capture, amine solvent scrubbing, uses aqueous solutions of alkanolamines to achieve CO2 separation from flue gas[6-8]. Whilst this technology is the current state of the art and will be used in the first generation of carbon capture plant, the technology has a number of drawbacks in terms of complexity in operation, high pH solvents leading to corrosion of metal piping, and the energy-intensive regeneration of the solvents[6]. This high energy usage of this process has led to the proposal of a range of potentially more efficient and less energy intensive second and third generation capture technologies[9]. The development of a solid adsorbent capture technology is one of the most promising alternative capture technologies[9]. A key motivation for the development of solid adsorbents for carbon capture is the potential energy saving shown by theoretical studies. These studies suggest that an adsorbent system with a cyclic capacity approaching or better than 3 mmol g-1 could significantly reduce the energy requirement of post-combustion capture by 30-50% compared with amine solvent systems[10]. A wide range of materials have been developed for this application[11] and include, supported amines and immobilized amines[12-16] activated carbons[17-19], Hydrotalcites[20], zeolites[21], inorganicorganic hybrid materials such as Metal Organic Frameworks (MOFs)[22]. Of all the materials developed and tested the challenge still remains to develop materials that
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2
Oviedo ICCS&T 2011. Extended Abstract
achieve these performance targets and are fully stable under the conditions of postcombustion flue gases[23].
2. Experimental Adsorbent materials have been characterised and tested using a range of techniques. Characterisation of materials has focussed on determining the physical and chemical properties of the solid sorbent materials. This has been conducted using a range of techniques, for example, elemental analysis, powder x-ray diffraction (XRD), diffuse reflectance infrared Fourier transform spectra (DRIFTS), textural properties have been determined by N2 adsorption analysis . Thermogravimetric analysis (TGA) has been used to determine the thermal stability of the materials as well as measure CO2 adsorption capacity and cyclic capacity[14].
3. Results and Discussion To realise the potential of solid sorbent for carbon capture two developments are required, new porous materials and plant integration processes. The key challenges for materials development and requirements in terms of: operating conditions, gas composition, stability and lifetime required to make solid sorbents a viable large scale CO2 capture process are described in this presentation[24]. To date a wide range of materials have been developed and tested as part of the STEPCAP project. The key materials under development are, microporous polymers, surface modified hydrotalcites, and co-doped sorbents. Performance of these materials has been assessed under a range of conditions and will be presented. Current key developments are summarised as follows: Hydrotalcites and conjugated microporous polymers have been studied as potential adsorbents for CO2 capture[25, 26]. Modified hydrotalcites such as, amine modified layered double hydroxides (LDH’s) were synthesized via exfoliation and grafting route. The influence of primary and secondary amines on carbon dioxide adsorption was investigated. One hydrotalcite with
3-[2-(2-Aminoethylamino) ethylamino]propyl-
trimethoxysilane, containing both a primary and secondary amine functional groups showed a steady increase in CO2 adsorption capacity of 0.74 - 1.76 mmol g-1 from 25 80oC through the flue gas temperature range. Synthetic microporous polymers possess some of the highest reported surface areas[27]
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3
Oviedo ICCS&T 2011. Extended Abstract
and some preparative routes might in principle be applicable to CCS applications[28]. A key benefit of porous organic chemistries is the very diverse synthetic organic chemistry which is available, both in terms of the wide range of monomers that can be exploited either by direct incorporation[29-31] or by the possibility of post-synthetic modification of networks to include functional groups reactive to CO2. These routes to materials synthesis are being explored as part of the STEPCAP project.
Incorporation of
functional monomers has been shown to be useful in tuning the isosteric heat of adsorption of CO2 by these materials[32]. A further advantage of organic polymeric networks over other highly porous synthetic materials such as hybrid inorganic-organic materials is their high moisture stablility together with high thermal stability[27]. However, despite recent reports of uptakes of around 3 mmol g-1 at ambient temperatures[33] microporous organic polymers have yet to achieve high enough CO2 loadings under the required conditions to be commercialised.
4. Conclusions The development of solid sorbents for CO2 capture is an area of significant academic and industrial interest. The composition of the flue gases in post-combustion capture and the requirements for material performance to minimise the energy penalty of the capture process present a significant challenge for materials development. To date, a wide range of functional materials have been and will continue to be developed with potential to make breakthrough. Whilst at present the required cyclic capture capacities can be achieved, one of the main challenges still remains to develop materials that can operate reliably and over a large number of cycles in a flue gas environment, a challenge which will form the future focus of the STEPCAP project.
Acknowledgement. The authors would like to thank E.ON-EPSRC strategic call on CCS for funding the Step Change Adsorbents and Processes for CO2 Capture research programme EP/G061785/1.
References [1] World Energy Outlook 2008 Edition, International Energy Agency, 2009. [2] P. Capros, L. Mantzos, N. Tasios, A. De Vita, N. Kouvaritakis, EU energy trends to 2030Update 2009, 2010.
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Oviedo ICCS&T 2011. Extended Abstract
[3] E. Rubin, L. Meyer, H. de Coninck, IPCC Special Report: Carbon Dioxide Capture and Storage, 2005. [4] Department of Trade and Industry. Meeting the Energy Challenge: A White Paper on Energy, 2007. [5] United Nations Framework Convention on Climate Change (UNFCCC). Report of the Conference of the Parties on its sixteenth session, held in Cancun from 29 November to 10 December 2010. Part one: Proceedings., 2011. [6] C.L. Leci, Financial implications on power generation costs resulting from the parasitic effect of CO2 capture using liquid scrubbing technology from power station flue gases, Energ Convers Manage, 37 (1996) 915-921. [7] H.J. Herzog, E.M. Drake, Greenhouse Gas R&D programme. IEA/93/oE6, (1993). [8] R.M. Davidson, Post-combustion carbon capture from coal fired plants - solvent scrubbing. IEA Clean Coal Centre, 2007. [9] J.D. Figueroa, T. Fout, S. Plasynski, H. McIlvried, R.D. Srivastava, Advancesn in CO2 capture technology - The US Department of Energy's Carbon Sequestration Program, Int J Greenh Gas Con, 2 (2008) 9-20. [10] M.L. Gray, K.J. Champagne, D. Fauth, J.P. Baltrus, H. Pennline, Performance of immobilized tertiary amine solid sorbents for the capture of carbon dioxide, Int J Greenh Gas Con, 2 (2008) 3-8. [11] R. Davidson, Post-combustion carbon capture – solid sorbents and membranes. IEA CLean Coal Centre, 2009. [12] X.C. Xu, C.S. Song, J.M. Andresen, B.G. Miller, A.W. Scaroni, Novel polyethylenimine-modified mesoporous molecular sieve of MCM-41 type as high-capacity adsorbent for CO2 capture, Energ Fuel, 16 (2002) 1463-1469. [13] X.C. Xu, C.S. Song, J.M. Andresen, B.G. Miller, A.W. Scaroni, Preparation and characterization of novel CO2 "molecular basket" adsorbents based on polymer-modified mesoporous molecular sieve MCM-41, Micropor Mesopor Mat, 62 (2003) 29-45. [14] T.C. Drage, A. Arenillas, K.M. Smith, C.E. Snape, Thermal stability of polyethylenimine based carbon dioxide adsorbents and its influence on selection of regeneration strategies, Micropor Mesopor Mat, 116 (2008) 504-512. [15] P.J.E. Harlick, A. Sayari, Applications of pore-expanded mesoporous silicas. 3. Triamine silane grafting for enhanced CO2 adsorption, Ind Eng Chem Res, 45 (2006) 32483255. [16] R. Serna-Guerrero, E. Da'na, A. Sayari, New Insights into the Interactions of CO2 with Amine-Functionalized Silica, Ind Eng Chem Res, 47 (2008) 9406-9412. [17] C. Pevida, T.C. Drage, C.E. Snape, Silica-templated melamine-formaldehyde resin derived adsorbents for CO2 capture, Carbon, 46 (2008) 1464-1474. [18] A. Arenillas, K.M. Smith, T.C. Drage, C.E. Snape, CO2 capture using some fly ashderived carbon materials, Fuel, 84 (2005) 2204-2210. [19] T.C. Drage, A. Arenillas, K.M. Smith, C. Pevida, S. Piippo, C.E. Snape, Preparation of carbon dioxide adsorbents from the chemical activation of urea-formaldehyde and melamineformaldehyde resins, Fuel, 86 (2007) 22-31. [20] S. Walspurger, L. Boels, P.D. Cobden, G.D. Elzinga, W.G. Haije, R.W. van den Brink, The Crucial Role of the K+-Aluminium Oxide Interaction in K+-Promoted Alumina- and Hydrotalcite-Based Materials for CO2 Sorption at High Temperatures, Chemsuschem, 1 (2008) 643-650. [21] P. Xiao, J. Zhang, P. Webley, G. Li, R. Singh, R. Todd, Capture of CO2 from flue gas streams with zeolite 13X by vacuum-pressure swing adsorption, Adsorption, 14 (2008) 575582.
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Oviedo ICCS&T 2011. Extended Abstract
[22] A. Torrisi, R.G. Bell, C. Mellot-Draznieks, Functionalized MOFs for Enhanced CO2 Capture, Cryst Growth Des, 10 (2010) 2839-2841. [23] S. Sjostrom, H. Krutka, Evaluation of solid sorbents as a retrofit technology for CO2 capture, Fuel, 89 (2010) 1298-1306. [24] T. Drage, C. Snape, L. Stevens, J. Wood, J. Wang, A. Cooper, R. Dawson, G. Guo, C. Satterley, R. Irons, Materials challenges for the development of solid sorbents for postcombustion carbon capture, Journal of Materials Chemistry, (In Prep). [25] J. Wang, L. Stevens, T. Drage, J. Wood, Preparation and CO2 adsorption of amine modified Mg-Al LDH via exfoliation route, Chem Eng Sci, (In Press). [26] J. Wang, L. Stevens, T. Drage, J. Wood, Preparation and CO2 adsorption of amine modified layered double hydroxide via anionic surfactant-mediated route (In Prep). [27] H. Ren, T. Ben, E.S. Wang, X.F. Jing, M. Xue, B.B. Liu, Y. Cui, S.L. Qiu, G.S. Zhu, Targeted synthesis of a 3D porous aromatic framework for selective sorption of benzene, Chem. Commun., 46 (2010) 291-293. [28] C.D. Wood, B. Tan, A. Trewin, H.J. Niu, D. Bradshaw, M.J. Rosseinsky, Y.Z. Khimyak, N.L. Campbell, R. Kirk, E. Stockel, A.I. Cooper, Hydrogen storage in microporous hypercrosslinked organic polymer networks, Chem Mater, 19 (2007) 2034-2048. [29] R. Dawson, A. Laybourn, R. Clowes, Y.Z. Khimyak, D.J. Adams, A.I. Cooper, Functionalized Conjugated Microporous Polymers, Macromolecules, 42 (2009) 8809-8816. [30] R. Dawson, A. Laybourn, Y.Z. Khimyak, D.J. Adams, A.I. Cooper, High Surface Area Conjugated Microporous Polymers: The Importance of Reaction Solvent Choice, Macromolecules, 43 (2010) 8524-8530. [31] J.-X. Jiang, A.I. Cooper, Microporous organic polymers: design, synthesis and function, Topics Curr. Chem., 293 (2009) 1-33. [32] R. Dawson, D.J. Adams, A.I. Cooper, Chemical tuning of CO2 sorption in robust nanoporous organic polymers, Chemical Science, (2011). [33] M.G. Rabbani, H.M. El-Kaderi, Template-Free Synthesis of a Highly Porous Benzimidazole-Linked Polymer for CO2 Capture and H2 Storage, Chemistry of Materials, 23 (2011) 1650-1653.
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6
Oviedo ICCS&T 2011. Extended Abstract
Development of a new synthesis gas production process from coal by catalytic gasification of HyperCoal using steam-CO2 as gasifying agent
Atul Sharma and Toshimasa Takanohashi Advanced Fuel Group, Energy Technology Research Institute, National Institute of Advanced Industrial Science and Technology, 16-1, Onogawa, Tsukuba, Ibaraki, JAPAN. Corresponding author: [email protected]
Abstract Growing energy costs and demand for fossil fuels because of rapid industrialization of emerging economies have revived the interest in coal gasification as a clean coal technology. Gasification can be used for both power generation and for production of liquid fuels/chemicals. To overcome the major drawbacks; high capital and/or operating cost, new advanced gasification processes have to be developed for more efficient and clean gasification of coal. Low temperature catalytic gasification of coal is a highly efficient way to convert coal to fuel gases but has never become a commercial process due to technical and economical factors. Primary reason is the loss of catalyst due to deactivation. In addition, due to the catalytic effect, the control of product gas composition is not easy. We developed a new process to clean coal before use, called HyperCoal process to produce ash less coal. With HyperCoal as feedstock, the loss of catalyst due to deactivation can be overcome and catalyst can be recycled and reused. For indirect production of liquids/ chemicals from coal, FT synthesis process is a primary process. However, a H2/CO ratio of 1 for direct DME production and 2 for methanol production is needed. From conventional high temperature coal gasification, synthesis gas with desired H2/CO can not be produced in a single step. The process is a two step process and efficiency is about 38~42%. In case of catalytic gasification, product gas is mainly H2 and CO2 at atmospheric pressure and methane rich at high pressure. The main highlight of the present study is the production of the synthesis gas from catalytic gasification at 650~700 oC temperature by gasifying HyperCoal in H2O/CO2 mixed environment and controlling its composition by adjusting the composition of the gasifying agent. Using this approach synthesis gas was produced with H2/CO ratio of 1~3 at 650~700 o
C in a single step which can be directly used as feedstock for DME and methanol production by FT
synthesis process.
1
Oviedo ICCS&T 2011. Extended Abstract
1. Introduction Growing understanding to reduce CO2 emissions and at the same time growing energy costs because of increased demand for fossil fuels due to rapid industrialization of emerging economies have revived the interest in coal gasification as a clean coal technology. Gasification can be used for both power generation and for production of liquid fuels/chemicals. However, for commercial acceptance, it has to economically compete with oil based industries. The major drawbacks such as high capital/operating cost and low efficiency have to be overcome. And for that, new advanced processes for more efficient and clean gasification of coal have to be developed. Low temperature catalytic gasification of coal is a highly efficient way to convert coal to fuel gases but has never become a commercial process due to technical and economical factors. Primary reason is the loss of catalyst due to deactivation. In addition, the primary gas at atmospheric pressure is hydrogen and at high pressure is methane. Therefore, the control of product gas composition is not easy. We developed a new process called HyperCoal process to produce ash less coal. With HyperCoal as feedstock, the loss of catalyst due to deactivation can be overcome and catalyst can be recycled and reused. For indirect production of liquids/ chemicals from coal, FT synthesis process is a primary process. However, a H2/CO ratio of 1 for direct DME production and 2 for methanol production is needed. From conventional high temperature coal gasification, synthesis gas with desired H2/CO can not be produced in a single step. The process is a two step process and efficiency is about 38~42%. In case of catalytic gasification, product gas is mainly H2 and CO2 at atmospheric pressure and methane rich at high pressure. The main highlight of the present study is the production of the synthesis gas from catalytic gasification at 600~700 oC temperature by gasifying HyperCoal in H2O/CO2 mixed environment and controlling its composition by adjusting the composition of the gasifying agent. Using this approach synthesis gas was produced with H2/CO ratio of 1~3 at 650~700 oC in a single step which can be directly used as feedstock for DME and methanol production by FT synthesis process.
2. Experimental A subbituminous coal, Pasir (PAS) from Indonesia was selected for the investigation. HyperCoal production method has been described in detail elsewhere. Briefly, HyperCoal (HPC) was produced by the solvent extraction of the coal with 1-methylnaphthalene at 360 C and subsequently separating the extract (HyperCoal) from the solvent. The extraction yield was 51 % for Pasir coal. The properties of the Pasir coal and HyperCoal produced from Pasir coal are shown in Table 1.
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Oviedo ICCS&T 2011. Extended Abstract
Table 1: Properties of coal and HyperCoal. Coal
Ash
Elemental analysis (wt % daf)
(wt %, db) C
H
N
S
O
Pasir coal
4.2
68.2
5.4
1.1
0.16
25.14
Pasir HyperCoal
0.01
79.5
5.9
1.1
0.29
13.21
HyperCoal has nearly no mineral matters. Because of almost no mineral matter, all the inorganically associated sulfur will be removed. The only sulfur in HyperCoal will be the organically associated sulfur. A detailed characterization of catalyst, HyperCoal, original coal, and chars using XRD, NMR, and SEM-EDX mapping techniques had been carried out and reported elsewhere. The experimental setup is shown in Figure 1.
Ar/O2 mixer
Purge gas /(Ar+O2) gas Steam generator 3-way valve
4-way valve
Sample
water TG carrier gas
Flow controller
TGA
Ice Trap
Ar
CO2
Ar
O2
HPLC pump
Gas out
Flow meter
CO2
µGC
Figure 1. Schematic of experimental set up for catalytic steam gasification of HyperCoal.
Samples for catalytic gasification experiments were prepared with 50 % catalyst loading. Catalyst loading was on dry and ash free wt % basis of coal and HPC. Catalyst mixing method has been described in detail elsewhere. Briefly, a desired amount of K2CO3 was added on the top of a measured sample already loaded into a test crucible as solid particles and stirred with a small spatula until white K2CO3 disappears by capturing moisture from the air. The particle size of coal and HyperCoal sample was under 75 µm. The gasification experiments were carried out with and without K2CO3 as a catalyst at 700, 650 and
3
Oviedo ICCS&T 2011. Extended Abstract
600 C with different steam to carbon dioxide (H2O/CO2) ratios as gasifying agent. Experiments were carried out in a thermogravimetric (TG-DTA 2020S, MAC) apparatus. A desired amount of water was pumped by a HPLC pump to a steam generator held at 250 C. CO2 was flowed to the steam generator as a steam carrier gas. By changing the amount of water pumped by the HPLC pump to the steam generator and the flow rate of CO2 as the carrier gas, different steam to CO2 ratio were achieved. For pure steam gasification, argon gas instead of CO2 was used as carrier gas. In case of steam and CO2 only conditions, partial pressure of steam and CO2 were kept at 0.5 atm by mixing argon gas. A 4-way valve at the inlet of the TG-DTA was used to change (Ar+O2) flow to (CO2+steam) flow. The flow lines were kept at 250 C by using ribbon heaters. First, a desired amount of sample was heated to 200 C and held in pure argon flow for 60 min to remove moisture and O2 from the reaction zone. After 60 min hold, the sample was heated to the desired temperature at 20 C /min in pure argon. When the desired temperature was reached without any hold time the pure argon gas was switched to the preset steam/CO2 gas mixture. The steam+CO2 mixture flowing from top comes into contact with the sample in the crucible. The evolved gases flow out together with the purge gas from the side into an ice cooled tar trap to remove tar before injecting to the micro gas chromatograph (Agilent 3000A). The total gas flow rate at the outlet was measured every 3 min by a film flow meter. 800
20
Weight [ mg]
16
700
Ar (100%) Steam + CO2 50 + 50
12
600 500 400
8
300
Temperature [C]
Ar + O2
200
4
100 0
0 0
30
60
90
120
150
Time [min]
Figure 2. Weight loss profile of HPC pyrolyzed in argon up to 700 C followed by steam+CO2 gasification.
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Oviedo ICCS&T 2011. Extended Abstract
3. Results and Discussion Figure 2 shows a typical weight loss curve for HPC+50 % K2CO3 sample pyrolyzed in argon up to 700 C followed by gasification with [50% steam+50% carbon dioxide (vol/vol)] as gasifying agent (from hereon in this manuscript steam to carbon dioxide ratio will be addressed as H2O/CO2). In a previous study gasification rate and gas composition was investigated at 10, 20, 40, and 50 % catalyst mixing ratio. It was reported that gasification rate was affected by the catalyst amount up to 50 % loading and above this catalyst loading rates were almost independent of catalyst amount. All experiments were carried out at atmospheric pressure. The weight loss curve can roughly be divided into three stages; moisture removal or drying stage, devolatilization stage and fixed-carbon gasification stage. The initial weight loss (up to 200 C) during heating from room temperature to 200 C in O2+Ar mixture is mainly due to moisture captured from air. Pre-oxidation was done to reduce the extremely high swelling propensity of HyperCoal. After pre-oxidation stage, sample was switched to 100 % argon for 60 min. The coal/HPC conversion on dry, ash, catalyst and volatile free basis (dacvf) (from hereon called char conversion) was calculated during the fixed-carbon gasification stage by the following equation: X (char conversion, % dacvf) =
W0 W 100 W0 (1 Wash Wcat )
[1]
where W0 is the weight when the gasification begins (db, mg) (weight at t= 45, 41 and 39 min for T= 700, 650 and 600 C, respectively), W is weight at any gasification time (db, mg, >39 min), Wash is weight fraction of ash content in coal or HPC, Wcat is weight fraction of catalyst content. Results of gasification rate and gas composition only in the char gasification stage will be discussed further. Figure 3 shows the gasification profile, gas composition and H2/CO ratio of produced gas from K2CO3 catalyzed coal and HPC at 700 C as a function of H2O/CO2 ratio. In general, rate decreased with increasing CO2 fraction in the gas mixture. Similarly, H2 decreased and CO increased with increasing CO2 fraction in the gas mixture. Under H2O/CO2 mixed gas environment, three reactions as shown below: C-H2O reaction (1), C-CO2 reaction (2) and water-gas shift (WGS) reaction (3) are expected to take place. C
+
H2O
→
CO
C
+
CO2
⇌
2CO
CO +
H2O
⇌
H2
+
H2
[1] [2]
+
CO2
[3] 5
Oviedo ICCS&T 2011. Extended Abstract
100
100/0
Char conversion [ %, dacvf]
Char conversion [ %, dacvf]
100 80 60
0/100
30/70 50/50
40
60/40
(a)
70/30
20
COAL＋K2CO3 T= 700
100/0 80 30/70
60
50/50
0 0
10
20
30
40
50
60
70
80
0
90
10
20
COAL＋K2CO3
125 100 75 50 25
Gas yield [mmol/g-char]
Gas yield [mmol/g-char]
175
CO H2
(b) T= 700
150
60
70
80
90
(e)
CO H2
T= 700
HPC＋K2CO3
100 75 50 25
50/50 30/70 0/100
100/0 70/30 60/40
Gasifying agent, H2O/CO2, (vol/vol)
50/50 30/70 0/100
Gasifying agent, H2O/CO2, (vol/vol) 30
29.5
(c)
25
27.1
15 10
10 4.4
2.8
2.1
1.0
0.1
HPC＋K2CO3 T= 700
20
H2/CO
15
(f)
25
COAL＋K2CO3 T= 700
20
H2/CO
50
0 100/0 70/30 60/40
5
40
125
0
30
30
Gasification time [min]
Gasification time [min]
150
(d) HPC＋K2CO3 T= 700
70/30 20
0
175
0/100
60/40
40
5
5.0
3.2
2.5
1.3
0.1
0
0 100/0 70/30 60/40 50/50 30/70 0/100 Gasifying agent, H2O/CO2, (vol/vol)
100/0 70/30 60/40 50/50 30/70 0/100 Gasifying agent, H2O/CO2, (vol/vol)
Figure 3. Effect of H2O/CO2 ratio of the gasifying agent on (a, b) gasification profiles, (c, d) gas yields and (e, f) H2/CO ratio of K2CO3 catalyzed coal and HyperCoal at 700 C.
Figure 3(a, d) shows the gasification profile of HPC and coal with 50 wt % catalyst in steam and carbon dioxide (H2O/CO2) mixed environment at 700 C. The H2O/CO2 ratios selected are H2O/CO2=100/0, 70/30, 60/40, 50/50, 30/70, and 0/100. H2O/CO2=70/30 means 70 % H2O and 30 % CO2 on volume basis. The gasification rates were high enough for commercial application. Figure 3(b, e) shows H2 and CO amount in the produced gas at different H2O/CO2 for coal and HPC at 700 C. At H2O/CO2=100/0, produced gas contained mainly H2 and very little CO. As H2O/CO2 changed to 70/30, 60/40, 50/50 and 30/70, H2 decreased and CO increased. At H2O/CO2=0/100, produced gas contained mainly CO and very little H2. Figure 3(c, f) shows change in H2/CO ratio with H2O/CO2 ratio of the 6
Oviedo ICCS&T 2011. Extended Abstract
gasifying gas. Results show that by changing H2O/CO2 ratio of the gasifying agent, H2/CO ratio of the produced gas can be controlled. These results show that synthesis gas with H2/CO ratio from 1~3 can be produced by gasification of coal in a single step by changing the H2O/CO2 ratio of the gasifying agent. Synthesis gas with H2/CO=1, 2 and 3 can be used as feedstock for FT synthesis process to produce DME, methanol, methane and other chemicals. For commercial application a high gasification rate or short gasification time is required. Considering the scale of operation of a coal gasifier typically about thousand tons per day, a long gasification time means large residence time leading to huge gasifier size in addition to additional energy requirements. Figure 4 shows a correlation between H2/CO ratio of synthesis gas, gasification time and H2O/CO2 ratio of the gasifying agent at 700, 650 and 600 C. A gasification rate of about 0.3 h-1 (18 min) at temperature 700 C and below would be necessary for commercial application. From Figure 4 it can be seen that synthesis gas with H2/CO=1~3 can be produced in a single step from catalytic coal gasification at 700 C with gasification time <20 min by changing H2O/CO2 ratio from 50/50~30/70.
6 H2O/CO2=70/30 5
H2/CO
4
700 650 600
H2O/CO2=50/50
3 2
H2O/CO2=30/70
1
H2O/CO2=0/100
0 0
50
100 150 200 Gasification time, [min]
250
Figure 4. Correlation between H2/CO ratio, gasification time and H2O/CO2 ratio of the gasifying agent as a function of temperature.
Figure 5 shows a schematic of such a process for producing synthesis gas from coal. HyperCoal is first produced from coal by HyperCoal production process not included in this process. The process flow shows that gasifier and combustor are separated. Combustor can be used to produce steam and supply necessary heat to gasifier by combusting HyperCoal residue. HyperCoal is mixed with catalyst and fed
7
Oviedo ICCS&T 2011. Extended Abstract
to gasifier where synthesis gas with desired H2/CO is produced in a single step by catalytic gasification of HyperCoal with H2O/CO2 as gasifying agent. Catalyst is recovered from the bottom and recycled. As HyperCoal is used, catalyst will not be deactivated. Coals especially with low ash contents such as Victorian brown coal can also be used instead of HyperCoal although their use will depend on the cost of the catalyst. The produced gas is then used as a feedstock for FT synthesis process to produce DME, methanol, methane and other chemicals. CO2 is produced as end product of FT synthesis process. CO2 is first separated and mixed with steam (H2O) before recycled to the gasifier as gasifying agent. The recycled CO2 can also act as fluidizing agent in case a fluidized bed gasifier is used. The proposed process is only a simplified overview of a new single step low temperature synthesis gas production and control process. There are several technical and operating challenges such a tar formation, heat supply to the gasifier, operating pressure of the gasifier, energy and commercial economics of the process, to list a few to be solved before a pilot scale process can be conceived. However, the results from the present study do suggest that there is a possibility to develop a new commercial process to produce synthesis gas as a feedstock for FT synthesis process from coal in single step at low temperature.
Synthesis gas H2/CO = 1, 2, 3
Coal/HPC
5
6
8
DME Methanol Methane
2
3 Coal/HPC residue Steam
1
4
CO2 7
1. Boiler 2. Feeder 3. Gasifier 4. Steam/CO2 mixer 5. Heat recovery unit 6. Gas cleaning 7. Catalyst recovery 8. Compressor 9. FT process
9
Recycled CO2
Catalyst recycle Water
Figure 5. Schematic process flow of a new low temperature single step coal to synthesis gas production process.
4. Conclusions Effect of gasifying agent on gasification rate and composition of product gas from catalytic gasification 8
Oviedo ICCS&T 2011. Extended Abstract
of coal and HyperCoal at 700, 650, and 600 C was investigated. Synthesis gas can be produced from low temperature catalytic coal gasification by using H2O/CO2 mixture as gasifying agent. Synthesis gas composition (H2/CO ratio) can be controlled in a single step by changing H2O/CO2 ratio of the gasifying agent. Gasification rate was affected by both temperature and CO2 fraction in the H2O+CO2 mixed gas. H2/CO ratio was only affected by CO2 fraction in the H2O+CO2 mixed and little or negligible with temperature. A new process to produce synthesis gas as a feedstock for FT synthesis process from coal in single step at low temperature was proposed.
References [1] Nahas, N. C. Fuel 1983, 62, 239-241. [2] Sharma, A.; Takanohashi, T.; Morishita, K.; Takarada, T.; Saito, I. Fuel 2008, 87, 491-497. [3] Sharma, A.; Takanohashi, T.; Saito, I. Fuel 2008, 87, 2866-2690. [4] Sharma, A.; Saito, I.; Takanohashi, T. Energy & Fuels 2008, 22(6), 3561-3565. [5] Sharma, A.; Kawashima, H.; Saito, I.; Takanohashi, T. Energy & Fuels 2009, 23(4), 1888-1895. [6] Sharma, A.; Saito, I.; Takanohashi, T. Energy & Fuels 2009, 23(10), 4887-4892. [7] Sharma, A.; Takanohashi, T. Energy & Fuels 2010, 24, 1745-1752.
9
A CeO2-La2O3-based Cu Catalyst for Application in HighTemperature Water Gas Shift Reactions L D. MORPETHa*, Y. SUNb, S.S. HLAa, G.J. DUFFYb, J.H. EDWARDSb, D.J.HARRISa and D.G. ROBERTSa * Corresponding Author; [email protected]. CSIRO Energy Technology, PO Box 883, Pullenvale, QLD 4069, Australia b CSIRO Energy Technology, PO Box 52, North Ryde, NSW 1670, Australia a
ABSTRACT Major hurdles to the widespread adoption of integrated gasification combined cycled (IGCC) with pre-combustion capture of CO2 include the comparatively high capital and operating costs of such systems. Improved operating efficiencies and reduced costs could potentially flow from elevated processing temperatures for the syngases leaving the gasifier and reduced steam injection rates. If the required water gas shift (WGS) process could be conducted at high temperatures and then integrated with a suitable membrane in the one reactor to separate the product hydrogen and CO2, high reaction rates and conversions of CO to CO2 and H2 approaching 100% could potentially be achieved. However, there are no commercial WGS catalysts suitable for continuous operation at temperatures higher than 450°C.
In this study, the performance of a CeO2-La2O3-based Cu catalyst for the WGS reaction was investigated using simulated coal-derived syngas in the temperature range of 450-600°C at approximately atmospheric pressure. This paper includes the results of a kinetics study of a CeO2La2O3-based Cu catalyst at atmospheric pressure. The effects of CO, CO2, H2O and H2 concentration on the WGS reaction rate were determined over the catalyst using selected gas compositions that might be encountered in coal-based gasification systems. The Ce-based Cu catalyst developed in this study and tested in our system performed strongly at temperatures up to 600°C, well beyond the maximum operating temperature of commercially available WGS catalysts. The catalyst was shown to work well in relatively low steam-to-carbon ratio syngas which is significant as it could lead to lower costs and higher thermal efficiencies through a reduced requirement for steam.
Keywords: 1. Water Gas Shift Reaction 2. Kinetics 3. Steam-to-Carbon Ratio 4. Coal Derived Syngas
2011 ICCS&T Oviedo, Spain A CeO2-La2O3 based Cu catalyst for application in high temperature water gas shift reaction.
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1.
INTRODUCTION
The water gas shift (WGS) reaction is one of the key processes in next generation IGCC plants that will incorporate CO2 capture. It involves the reaction of the CO leaving the gasifier with steam to produce CO2 and more H2 CO (g) + H2O (g)
CO2 (g) + H2 (g)
H298K = -41.1kJ
(1)
This is an equilibrium reaction that favours the formation of CO2 and H2 at low temperatures; and in practice catalysts are required to increase the reaction rate to ensure that the equilibrium conversion of CO is approached under the selected reaction conditions. Commercially available WGS catalysts developed over the last 40-60 years are not capable of operating for extended periods above 450oC. CSIRO has previously investigated the performance of these catalysts under simulated coal-derived syngas conditions where there are high CO concentrations, and has also characterised the performance under the low H2/high CO2 environments characteristic of the backend of catalytic membrane reactor systems. CSIRO has developed a range of ceria-based catalysts that are capable of operation at 600°C. We have previously investigated both CeO2-La2O3based and CeO2-La2O3-Al2O3-based Cu catalysts as well as La0.9-xCexFeO3 perovskite-like catalysts for the HT-WGS reaction in the temperature range 450 to 600°C and at a pressure of 1 atmosphere, with H2O:CO = 3.1/1 and a dry-gas hourly space velocity = 239,000 ml/hr.gcat [1-3]. The results show that these catalysts have high activity and stability, decreasing in the following order: La0.9xCexFeO3
perovskite-like catalysts > CeO2-La2O3-based Cu catalysts > CeO2-La2O3-Al2O3-based
Cu catalysts. Whilst it is possible to operate at reduced steam to carbon ratios, additional water (taking the steam to carbon ratio up towards a value of 3:1) is often required to suppress catalyst deactivation due to metal sintering and carbon formation on the catalyst surface. Operation at low steam to carbon ratios would be attractive as it would reduce steam requirements and lead to potentially improved operating efficiencies through reduction the energy penalties associated with generating excessive quantities of steam over and above the stoichiometric requirements for the WGS reaction. This paper presents data from over 2000 hrs of laboratory-scale reactor operation with a CeO2-La2O3based Cu catalyst for the water-gas shift conversion of CO in coal-derived syngas conditions over a range of steam to carbon ratios.
2011 ICCS&T Oviedo, Spain A CeO2-La2O3 based Cu catalyst for application in high temperature water gas shift reaction.
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2. EXPERIMENTAL The specifications for the water-gas-shift catalysts and the experimental conditions used in this study are listed in Table 1: The HT2 commercial catalyst is used as a reference throughout our studies as it is the state of the catalyst for the water-gas shift reaction. Table 1 - Materials and experimental conditions
WGS Catalyst
HT2 Commercial – Fe2O3 (80-95%), Cr2O3 (5-10%), CuO (1-5%) Copper Ceria Catalyst - (10at%Cu)/Ce(30at%La)Ox Weight: 0.20 ± 0.005 g; Fraction: 53 – 150 µm
Inert Packing
Alpha alumina, Weight - 1.00 ± 0.005g, Fraction 53 – 150 µm
Syngas Inlet composition (vol %)
65% CO; 30% H2 ; 2% CO2; 3% N2
Temperature / Pressure
450-600oC / 1atm
H2O:CO molar ratio (S/CO)
3.09, 2.71, 2.32, 1.93, 1.55, 1.25, 1.00 /1
Feed flow rate (dry basis)
797ml/min
2.1 Catalyst preparation The ceria-lanthana-based catalyst was synthesized using the urea method [4]. Pre-determined amounts of (NH4)2Ce(NO3)6, La(NO3)3.6H2O, Cu(NO3)2.2.5H2O
or
Cu(NO3)2.2.5H2O plus
Fe(NO3)3.9H2O and urea were dissolved in distilled water and then heated to 100°C while stirring vigorously. The mixture was kept at this temperature with periodic addition of water for 8 hours. The resulting solution was filtered and washed with hot distilled water twice, followed by drying at 100°C overnight and calcining in flowing air at 650°C for 5 hours.
2.2 Water Gas Shift Reactor System Experiments were undertaken in a laboratory scale fixed bed differential reactor where CO conversion was ~10%. This level of conversion was achieved by using low catalyst loadings and high gas flows of 797 ml/min (NTP, dry basis) to allow discrimination of both catalyst activity and stability. The simulated synthesis gas was fed into the reactor through individual gas lines for CO, H2, CO2 and H2. The inlet gas composition was adjusted through individually controlled mass flow controllers. The gas was electrically heated and maintained at ~180oC prior to entry to the furnace. A steady flow of de-ionised water was metered by a HPLC pump (Shimadzu LC 20-AT) and was vaporised inside an electrically-heated pre-heating tube and mixed with the gases prior to entering the 15mm ID stainless steel reactor. Inside the reactor 0.20 ± 0.005 g of catalyst was mixed with 1.00 ± 0.005 of inert α-Al2O3 and packed into the reactor between two ceramic sheets and a loose packing of quartz wool was placed at the inlet. Steam in the product gas was trapped using a water2011 ICCS&T Oviedo, Spain A CeO2-La2O3 based Cu catalyst for application in high temperature water gas shift reaction.
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cooled condenser and outlet gases were analysed by a micro gas chromatograph (Varian CP-4900). Four components (CO, H2, CO2 and N2) were measured every third minute during the experiment and accurate carbon balances were observed throughout.
3 RESULTS AND DISCUSSION 3.1 Preliminary Formulation Studies Figure 1 shows the conversion of the commercial reference HT2 catalyst at 450°C and 600°C. At the 450°C operational temperature limit of this catalyst, its performance is observed to stabilise to a conversion of 15.2 ± 0.1 % or a conversion rate of 2.85E-04 mol/g/s. Whilst there is some variation in the conversion stability, the level is in agreement with its previously observed performance given the duration of exposure and stabilisation time. The system temperature was then ramped beyond the specified catalyst operating temperature at a rate of 2°C/min to the desired elevated temperature limit of 600°C. The conversion is observed to increase dramatically to ~80% which approaches the equilibrium value for these operating conditions (black line). Following this and after three days testing, the performance is observed to degrade asymptotically to a value of ~3%. This conversion is slightly higher than the 1.72 ± 0.05% background level (due to the gas phase reaction) indicating a small amount of residual activity. The key result is that HT2 has been shown to be unstable at the high temperature, with its performance degrading to very low levels at 600°C. The inlet gas concentration is shown for reference (dashed line).
Following observation of the commercial catalysts operational limit, the performance of the highly active WGS ceria based catalyst was measured at 600oC. The conversion result, shown in Figure 2 for the S/CO ratio of 3.09/1, demonstrate its stability over a 4 day period at a level of 15.0 ± 0.2%. 3.2 Kinetic and Reaction Order Study A new sample of ceria catalyst was loaded in the reactor and a kinetic study performed over a three day period. Following activation in 150ml/min of H2 at 650°C at 1atm over a 2hr period, the conversion profile trended downwards from 21-22% to a stable 15% over a 2 day period. This trend was consistent with previous observations of catalysts conditioning and was re-assuring as it confirmed the reproducibility of both the catalyst preparation procedure and the experimental technique used in this study. Next, a program of 22 feed gas compositional changes with a constant wet gas velocity was conducted to quantify the effects of CO, H2O, CO2 and H2 concentration on the WGS reaction rate. Nitrogen was used to supplement the presence of one component in the stream whilst allowing the other components to remain at the same concentration. To obtain the 2011 ICCS&T Oviedo, Spain A CeO2-La2O3 based Cu catalyst for application in high temperature water gas shift reaction.
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apparent activation energy, an experiment was carried out where the temperature was ramped from 550 to 600°C with a heating rate of 0.4°C/min. The results are shown in Table 2 relative to a high temperature commercial catalyst evaluated in a separate study [5] . Over the temperature range of 550-600°C, the reaction order for CO was found to be around 1, the reaction rate retarded by increasing CO2 concentration and increasing H2 concentration according to the reactions orders shown in Table 2. The effect of varying H2O concentration on WGS reaction rates was found to be insignificant (reaction order close to zero) when the steam to CO ratio was between 3.09/1 and 1.55/1. This is similar behaviour to the commercial high temperature catalysts but at the higher operating temperature. The conversion profile (not shown) shows an increase from 4.0% to 15% over the temperature range 550°C to 600°C. The strong increase in reactivity with temperature begs the question of maximum operating temperature and maximum conversion level, however this was not pursued in this study as 600oC is likely the be the operating limit for any membrane separator in the foreseeable future. The stable conversion of 15.0 ± 0.2 % was achieved over a run that lasted for a further 2 days, indicating the excellent stability of ceria catalyst at the operating conditions. Table 2 - Reaction orders of gases components over the ceria based copper catalyst at 600°C and high t emperature commercial catalyst at 450°C Catalyst
Apparent Reaction Order
E (KJ mol-1)
Source
a [CO]
b [H2O]
c [CO2]
d [H2]
Ceria-Cu
0.95 ± 0.04
0.00 ± 0.18
-0.06 ± 0.04
-0.08 ± 0.03
92.3 ± 2.0
This study
HT
1.00 ± 0.03
0.00
-0.36 ± 0.04
-0.09 ± 0.007
111 ± 2.6
[5]
Figure 3 shows the plot of logarithmic values of reaction rates versus inverse temperature and the apparent activation energy was found to be 92.3±2.0 kJ mol–1. This level is consistent with our previous investigations on the high temperature commercial catalyst [5]. 3.3 Steam to CO Study A study of the effect of feed gas steam to CO ratio was undertaken using the stable ceria catalyst sample in the kinetic study. The gas flows shown Table 3 were maintained over the catalyst whilst the H2O:CO ratio was varied from 3.09/1 the ratio of 1.00/1. Each ratio was held for 12-18 hrs.
2011 ICCS&T Oviedo, Spain A CeO2-La2O3 based Cu catalyst for application in high temperature water gas shift reaction.
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Table 3 - Water %, S/CO Ratio and Gas and Water flows for the Steam to CO Study. The H2O flow is liquid. The italicised values were those changed during the experiment. H2O (%)
67
59
50
42
33
27
22
H2O:CO ratio (S/CO)
3.09
2.71
2.32
1.93
1.55
1.25
1.00
CO Flow (ml/min)
518
518
518
518
518
518
518
H2 Flow (ml/min)
239
239
239
239
239
239
239
CO2 Flow (ml/min)
16
16
16
16
16
16
16
N2 Flow (ml/min) H2O Flow (ml/min)
24
224
424
624
824
974
1104
1.288
1.17
0.966
0.805
0.643
0.523
0.418
Figure 4 shows the stable CO conversion levels at each S/CO ratio. The results show conversion at the starting S/C ratio of 3.09/1 to be the equivalent to the stabile level at the end of the kinetic study. Next an increase to ~19% was observed for the 2.32/1 ratio. The conversion for the 2.71/1 ratio was noted to be less than fully stabile, however our focus was the lower S/CO ratio gas conditions. At the lower S/CO ratios of 1.93, 1.55 and 1.25 to 1 the CO conversion is seen to be consistent and stable at around 16-17%. These results indicate good stability over appreciable periods of time on the laboratory scale at lower S/CO ratios. The ratio of 1.0/1 was then considered even though it is impractical in real world applications as it would be most probably to lead to the formation of carbon through the bed and downstream in the system. While stable conversion was observed at this ratio, a return to 3.09/1 revealed degradation in performance down to 10% conversion and below over the final two days of testing (not shown). Post run inspection revealed that carbon had formed in the bed indicating that the system had been pushed beyond its designed operating window in terms of S/CO ratio. However, the ceria based catalyst had demonstrated stable performance at 600oC with low steam to carbon ratios over an extended period of operation. 4. CONCLUSIONS
In this paper results from a novel ceria-based catalyst have been presented. The catalyst demonstrated good stability and strong conversion performance at 600oC, well beyond the 450oC operating temperature of the commercial catalyst HT2. The comparison was conducted at a S/CO ratio of 3.09/1 and over a period of 4-5days. In a subsequent kinetic study reaction orders for CO, H2O, CO2 and H2 were determined as was an activation energy value of 92.3 ± 2.0 kJ.mol-1. The values obtained were consistent with those observed in our previous high temperature catalyst work. A detailed steam to carbon ratio study with 12-18 hr exposure periods revealed stable conversion performance down to S/CO ratios as low as 1.00. The results are encouraging as elevation of the processing temperature and reduction in the amount of steam required could lead to 2011 ICCS&T Oviedo, Spain A CeO2-La2O3 based Cu catalyst for application in high temperature water gas shift reaction.
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improvements in both the efficiency and cost of an IGCC plant. Detailed modelling would be required to calculate the the optimum system steam to carbon ratio and the effect of reduced steam consumption on the overall process performance and economics. ACKNOWLEDGMENTS The authors wish to acknowledge the financial support of the Centre for Low Emission Technologies.
2011 ICCS&T Oviedo, Spain A CeO2-La2O3 based Cu catalyst for application in high temperature water gas shift reaction.
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Figure 1 - Conversion of 0.20g of HT2 catalyst over a 5day period; H2O/CO /CO ratio of 3.09/1 at the operating maximum of 450°C then 600oC. The inlet gas concentration of 65% CO, 30% H2, 2% CO2, 3% N2 (dashed) and empty reactor conversion levels (orange) are shown for reference. The inlet concentration was confirmed at the end of the experiment. experiment
CO Conversion (%)
30
20
Ceria Based Cu
15.0 ± 0.2%
Blank
1.72 ± 0.05%
10
0 0
1
2
3
4
Duration (days)
Figure 2 - Water-gas gas shift performance of synthesized catalysts with the H2O/CO CO ratio of 3.09/1. The ceria based copper catalyst produced a stable conversion of 15.0 ± 0.2% at the elevated temperature of 600°C. This compares with the background level of 1.72% and the commercial 450°C 450 limited catalysts performance of ~3%.
2011 ICCS&T Oviedo, Spain A CeO2-La2O3 La2O3 based Cu catalyst for application in high temperature water gas shift reaction.
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-8.0 Ln(CO rate)
Ln (reaction rate)
-8.2 -8.4
600oC
-8.6 -8.8
Activation Energy = 92.3 ± 2.0 kJ.mol-1 -9.0
550oC -9.2 1.14E-03
1.16E-03
1.18E-03
1.20E-03
1.22E-03
1/T (1/K)
Figure 3 - The Arrhenius plot for the WGS reaction over ceria based catalyst with a synthesis gas representative of coal derived synthesis gas (65% CO, 30% H2, 2% CO, 3% N2), temperature of 550 - 600oC and H2O:CO ratio of 3.09/1. 25
19.4
20
16.7
Conversion (%)
16.5
17.8
15.2
15.9
15.5
1.25/1
1.00/1
15
10
5
0 3.09/1
2.71/1
2.32/1
1.93/1
1.55/1
Steam to CO ratio (S/CO)
Figure 4 - Conversion performance of the ceria based catalyst as a function of H2O:CO ratio at 600oC. Error bars shown are the standard deviation in the final 2 hours of data from each S/CO ratio.
2011 ICCS&T Oviedo, Spain A CeO2-La2O3 based Cu catalyst for application in high temperature water gas shift reaction.
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REFERENCES [1]
[2]
[3]
[4]
[5]
Sun, Y., et al., A comparative study of CeO2-La2O3-based Cu catalysts for the production of hydrogen from simulated coal-derived syngas. Applied Catalysis A: General, 2010. 390(1-2): p. 201209. Sun, Y., et al., High temperature water-gas shift Cu catalysts supported on Ce-Al containing materials for the production of hydrogen using simulated coal-derived syngas. Catalysis Communications, 2010. 12(4): p. 304-309. Sun, Y., et al., Effect of Ce on the structural features and catalytic properties of La(0.9-x)CexFeO3 perovskite-like catalysts for the high temperature water-gas shift reaction. International Journal of Hydrogen Energy, 2011. 36(1): p. 79-86. Qi, X. and M. Flytzani-Stephanopoulos, Activity and Stability of Cu−CeO2 Catalysts in HighTemperature Water−Gas Shift for Fuel-Cell Applications. Industrial & Engineering Chemistry Research, 2004. 43(12): p. 3055-3062. Hla, S.S., et al., Kinetics of high-temperature water-gas shift reaction over two iron-based commercial catalysts using simulated coal-derived syngases. Chemical Engineering Journal, 2009. 146(1): p. 148-154.
2011 ICCS&T Oviedo, Spain A CeO2-La2O3 based Cu catalyst for application in high temperature water gas shift reaction.
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Coal plasma gasification for clean synthesis gas production V.E. Messerle1, A.B. Ustimenko1, N. Slavinskaya2, O.A. Lavrichshev3, E.F. Ossadchaya3 1 2 3
Research Department Plasmotechnics, Almaty, Kazakhstan
Institute of Combustion Technology, German Aerospace Centre, Stuttgart, Germany
Research Institute of Experimental and Theoretical Physics al-Farabi Kazakh National University, Almaty, Kazakhstan E-mail: [email protected]
Abstract This paper describes numerical and experimental investigation of coal gasification in combined arc-plasma reactor. The gasifier is an entrained flow reactor. The experimental installation is intended for work in the electric power range of 30-100 kWe, mass averaged temperature 1800-4000 K, coal dust consumption 3-10 kg/h and gas-oxidant flow 0.5-15 kg/h. The numerical experiments were conducted with the aid of PLASMA-COAL computer code. It was designed for computation of the processes in the plasma gasifier. This code is based on one-dimensional model, which describes two-phase chemically reacting flow with an internal plasma source. The thermo chemical conversion of coal-oxidant mixture is described through formation of primary volatile products, conversions of evolved volatile products in the gas phase and the coke residue gasification reactions. Kazakhstan Kuuchekinski bituminous coal of 40% ash content, Germany Saarland bituminous coal of 10.5% ash content and 14% ash content bituminous coal from the Middleburg opencast mines, South Africa, were used for the investigation. The investigation demonstrated that high quality synthesis gas can be produced using different power coals and plasma assisted technology. 1. Introduction Coal is the major source of energy. It provides 24% of thermal generation and 40% of electric energy in the world [1]. The share of coal in the world's proven reserves of fossil fuels is about 64% [2]. In particular, Kazakhstan is ranked eighth in the world in coal production, and its proven reserves reach 177 billion tons. In the near future increase of coal use is expected. According to forecasts [3] by 2020 the share of coal in the global fuel balance will be 56%.
Coal, being one of the most complex composition fossil fuels, is the richest source of valuable chemical products. In addition to power generation in world practice technologies are mastered by which from coals more than 500 products (synthesis gas, fuel oil, methanol, sorbents, etc.) are obtained. One of the new promising technologies for processing of coal is its plasma gasification [46], intensively developed in Kazakhstan. This paper considers numerical and experimental studies of plasma gasification of three kinds of coals: Ekibastuz coal (EC), ash content 40% (Kazakhstan), Saarland coal (SC) ash content of 10.5% (Germany) and Middleburg coal (MC), ash content 14% (South Africa). Experimental studies of plasma gasification were carried out by the example of EC (Table 1). Then we calculated its plasma gasification on the kinetic program PLASMA-COAL and compared the results of calculations with experiment. Then, using the verified program PLASMA-COAL comparative numerical study of plasma gasification of the three above coals was performed. Table 1. Chemical composition of Ekibastuz bituminous coal, mass. % Аd
С
О2
Н2
N2
S
40
48.86
6.56
3.05
0.8
0.73 23.09
SiO2 Al2O3 Fe2O3 CaO MgO K2O Na2O 13.8
2.15
0.34
0.31
0.16
0.15
Higher calorific value of coal on dry weigh Qd=16632 kJ/kg, moisture of coal Ww= 5.8%, Аd – ash on dry weight of coal
2. Experiment Experiments and further numerical calculations were performed applied to a plug flow plasma gasifier of the combined type of 100 kW nominal electrical power. The experimental setup is shown in Fig. 1 [4]. The electric arc is ignited between the rod and ring graphite electrodes in the combined plasma reactor 1, Fig. 1. The inner diameter of the reactor (i.e. of the graphite lining) is 0.15 m and its height is 0.3 m. The arc rotates under the influence of the magnetic field. The coal dust is fed to the reactor from dust feeder 7 through ejectors in the top cover of the reactor. Dust is sprayed in the reactor arc zone by plasma-forming gas (steam or air), also introducing into the reactor through injectors in the top cover. The oxidant-pulverised coal mixture entering the arc zone is heated to high temperatures by the rotating arc to produce a twophase plasma flow where the coal gasification process occurs. The gaseous products are derived through the slag and gas separator chamber 2, the chambers of synthesis gas cooling 4 and
hydration 6. The solid residue is removed through the diaphragm 2 into slag catcher 3. The distance from the top cover of the plasma reactor 1 to the exit of the gasifier (orifice 5) is 0.9 m. As a result of experiments on the basis of material and heat balancing of the plant, the main parameters of the plasma gasification of coal were measured. They are mass averaged temperature, coal gasification degree, specific power consumptions, synthesis gas yield. To conduct high temperature measurements in the reactor optical pyrometers were used. They allowed measuring the temperature to 4000 K. Sieve analysis of dust showed that the average particle size of coal was 75 microns. Fig. 1. Laboratory scaled plant for plasma gasification of coal: 1 – plasma reactor; 2 – orifice, synthesis gas and slag separator chamber; 3 – slag catcher; 4 – synthesis gas cooling chamber; 5 – orifice; 6 – hydration chamber; 7 – dust feeder; 8 – cooling water system; 9, 10 – power supply system; 11, 12 – rod electrode moving system; 13 – steam generator; 14 – safety valve; 15 – slag catcher elevator. The experimental results, obtained for EC, are summarized in Table 2. Reactor power (P) has been ranged from 25 to 52.8 kW. The measured efficiency of the reactor was 76%. The process specific power consumption (Qsp) is related to one kilogram of the reacting mass. As it is seen from Table 2, experiments No 1 and 2, coal gasification degree (Xc) at plasmaair gasification of coal increases from 89.6 to 95.8% with power consumption elevation from 2.1 to 3.3 kW·h/kg. The yield of synthesis gas varies from 43.3 to 56.3%. In plasma-steam gasification of coal, Table 2, experiments No 3 and 5, the specific power consumption is higher: 4.2 - 7.7 kW·h/kg. The degree of gasification remains at a high level, 92.0 – 94.2%. It is significant that the yield of synthesis gas is much higher, 90.0 – 97.3%, at plasmasteam gasification of coal. 3. Verification of computer-code PLASMA-COAL Computer-code PLASMA-COAL was designed for computation of the processes of moving, heating, and kinetics of thermochemical conversion of a coal–oxidant mixture in a plasma
gasifier alike, as shown in Fig. 1. This code is based on 1-D which describes a two phases (coal particles and gas oxidizer) chemically reacting flow in a reactor with an internal heat source (electric arc) [5, 6]. Coal particles and gas are admitted into the reactor with equal temperatures. There is a heat–mass exchange, particle-to-particle exchange, gas-to-particle exchange and gasto-electric arc exchange. In addition, heat and impulse exchange between the flow and the wall of the reactor is accounted for. Some chemical fuel transformations are also considered. They are the formation of primary volatile products, the conversation of evolved volatile products in the gas phase and the coke residue gasification reactions. Table 2. The main indices of Ekibastuz coal plasma gasification. Consumption, kg/h
N 1
EC Steam 8.0 -
Air 8.0
2
4.0
-
3
4.0
4 5
P, kW
Q, T, K kW·h/kg
CO
H2
N2 55.3
89.6
XC, %
33
2.1
2100
27.4
Vol. % 15.9
5.1
30
3.3
2850
38.1
18.2
43.7
95.8
1.9
-
25
4.2
3100
41.6
55.7
2.7
94.2
6.5
3.0
1.9
52.8
4.6
3150
38.6
51.4
9.8
92.0
4.0
2.44
0.43
52.3
7.7
3500
41.5
55.8
2.7
93.7
The plug flow assumption was applied to the entrained flow reactor. The resulting set of ordinary differential equations includes equations for species concentrations (chemical kinetics equations) in conjunction with the equations for gas and particle velocities and temperatures, respectively. The energy contribution from the plasmas had been found empirically and included into the energy equation as an internal heat source. Moreover, the model was distinguished by its detailed description of the chemical reactions previously mentioned [5, 6]. Kinetic scheme consists of 51 chemical reactions. Arrhenius equation describes the temperature dependence of
Ej n ⋅ T , where n is the temperature the rate constants of chemical reactions: k j = A j ⋅ exp − RT factor, A is the preexponential factor, Ej is the activation energy, and index j is the number of the reaction. The coal composition is presented in the model by its organic and mineral parts. The organic mass of coal is specified by the set of the functional groups (CO, CO2, CH4, H2O, and tar) and
Table 3. Kinetic parameters of the reactions of coal gasification. LgA b)
n
E
LgA b) n
E
1 [H2]S = H2
18.2
0
88.8
26 H+O2= O+OH
11.27 0
16.8
2 [H2O]S =H2O
13.9
0
51.4
27 H+H2O= H2+OH
10.98 0
20.3
3 [CO]S = CO
12.3
0
44.4
28 H2+O=H+OH
7.26 1.0
8.9
4 [CO2]S =CO2
11.3
0
32.6
29 H2O+M= H+OH+M
13.3
0
105.0
5 [CH4]S =CH4
14.2
0
51.6
30 H2O+O= OH+OH
10.53 0
18.3
6 [C6H6]S =C6H6
11.9
0
37.4
31 CO+OH= CO2+H
4.11 1.3 -0.77
7 C+H2O= CO+H2
11.32
0
60.8
32 CO+O2= CO2+O
11.5
8 C+CO2= CO+CO
13.2
0
83.6
33 CO2+H= CO+OH
6.15 1.3 21.6
9 C+O2=CO2
9.42
0
38.0
34 CO+O+M= CO2+M
12.77 0
10 C+C+O2= CO+CO
9.72
0
41.8
35 C2H2+M= C2H+H+M
11.0
0
114.0
11 CH4+H= CH3+H2
11.1
0
11.9
36 C2H2= C+C+H2
6.0
0
30.0
12 CH4+OH= CH3+H2O
0.54
3.1
2.0
37 C2H2+O2= HCO+HCO
9.6
0
28.0
13 CH4+M= CH3+H+M
14.15
0
88.4
38 C2H2+H= C2H+H2
11.3
0
19.0
14 CH4+O= CH3+OH
10.2
0
9.2
39 C2H2+OH= CH3+CO
9.1
0
0.5
15 CH3+H2O= CH4+OH
9.84
0
24.8
40 C2H2+O= CH2+CO
10.83 0
4.0
16 СH3+H2= CH4+H
9.68
0
11.4
41 CH2+H2O= CH2O+H2
11.0
0
3.7
17 CH3+M= CH2+H+M
13.29
0
91.6
42 CH2+O2= HCO+OH
11.0
0
3.7
18 CH3+O2= CH30+O
10.68
0
29.0
43 C2H+O2= HCO+CO
10.0
0
7.0
9.6
0
0
44 C2H+H2O= CH3+CO
9.08
0
0.5
20 СH3+O= CH2O+H
11.11
0
2.0
45 C6H6=C2H2+ C2H2+C2H2 12.0
0
85.0
21 CH3O+M=CH2O+H+M
10.7
0
21.0
46 OH+OH= H2O+O
0
1.1
22 CH2O+M= HCO+H+M
13.52
0
81.0
47 H+OH+M= H2O+M
10.56 0
0.0
23 HCO+M= H+CO+M
11.16
0
19.0
48 H+H+M= H2+M
9.56
0
0.0
24 O2+M= O+O+M
12.7
0
115.0
49 CH2O+OH= HCO+H2O
10.5
0
1.5
25 H2+M= H+H+M
11.34
0
96.0
50 H+OH= H2+O
9.84
0
7.04
51 H2+OH= H2O+H
11.4
0
10.0
j
Reaction a)
19 CH3+OH= CH2O+H2
j
Reaction a)
9.5
0
37.6
4.1
a)
Equations 1-6 are the devolatilization reactions. Dimensions of Aj are [s-1] for the first-order reactions and [10-3m3mol-1s-1] for the second-order reactions, and dimension of E is [kcal mol-1]. b)
carbon. According to the assumed scheme, the first chemical stage of the process is coal thermal destruction (reactions 1–6 in Table 3). Heating the coal particles generate volatile and tar components, which are presented by benzol C6H6 [6]. The interactions between the char carbon
and water vapor, oxygen, and carbon dioxide (reactions 7–10) are the rate-limiting stages of the process. These reactions present the complex heterogeneous processes, which include the different elementary stages: a reagent adsorption on the particle surface, dissociation, the reactions in the gaseous phase, desorption, etc. The detailed mechanism of these processes is not known at today, but such presented global steps can describe the main feature of the processes based on the empirical information about the rates of reaction paths. To verify the program PLASMA-COAL variant 5 from the Table 2 was calculated. Results of the comparison are presented in Table 4. Note that in the calculation the experimentally observed air leak at a rate of 0.43 kg / h was taken into account. It is seen that discrepancy in the concentrations of gas phase components is not more than 4 %, coal gasification degree – 1 % and mass average temperature at the exit of plasma reactor – 2 %. Such a relatively small discrepancy of the computed and experimental data confirms validity of the accepted physical and mathematical models and legitimacy of the kinetic code PLASMA-COAL application for numerical investigation of coal plasma gasification. Table 4. Comparison of the numerical results on PLASMA-COAL code with experimental one. Gas composition at the exit of gasifier, vol. % Method
X c, %
T, (K)
0.0
93.7
3500
0.0
93.9
3559
H2
CO
N2
О2
Experiment
55.8
41.5
2.7
Computation
53.5
42.1
2.71
4. Numerical simulation of the coals plasma gasification Three widely used in the energy sector of Kazakhstan and Germany coals were selected for the numerical study. They are Kazakhstan EC, German SC and the South African MC. Their characteristics are given in Table 5. For all the variants power of the plasma reactor was 52.3 kW, coal consumption – 10 kg/h, steam consumption was selected from an evaluation of the complete coal gasification. It was 7, 9.2 and 9.7 kg/h for EC, SC and MC respectively. Results are presented in Figs. 2 - 9. Figs. 2 - 4 show the composition of gases obtained by the coals plasma gasification. It is evident that high-quality synthesis gas is obtained from all three coals. Concentration of the synthesis gas at the outlet of the gasifier is 98.7, 96.4 and 97.15 vol. % for EC, SC and MC, respectively. In all variants of the calculation concentration of
hydrogen (H2) significantly exceeds that of carbon monoxide (CO). This excess is for EC 12.23%, for SC - 5.42% and for MC - 4.35%. Note that in the products of SC and MC gasification there is methane (CH4), 1.53 and 1.27 % respectively, whereas there is no methane at the outlet of the gasifier at EC gasification. Typically there is a methylene radical (CH2) , in the products of the coals plasma gasification with a concentration in the range of 1 to 2 %. The concentration of oxidant (H2O) to the plasma reactor exit (X = 0.3 m) tends to zero. The gasifying agent (water vapor) flow rate increasing can lead to an increasing in the yield of the synthesis gas due to the conversion of hydrocarbon impurities (CH4 and CH2). Table 5. Coals characteristics, mass. %. Coal
Ad
C
H2
H2 O
CO
CO2
СH4
C6H6
Vdaf
Qd, kJ/kg
EC
40.0
46.18
2.63
1.84
3.95
1.4
0.55
3.45
34.55
16632
SC
10.5
57.28
1.95
4.5
9.67
1.6
2.58
11.92
32.22
29277
MC
14.0
60.2
1.5
2.9
7.7
1.3
2.1
10.3
25.8
27321
100
100
H2 CO
10
10
Ci, vol.%
Ci, vol.%
H2O CH2 1
H2 CO
H2O
CH4
CH2
1
CH4
H
H
C6H6 C6H6 0.1 0.0
0.2
0.4
0.6
0.8
X, m
0.1 0.0
0.2
0.4
0.6
0.8
X, m
Fig. 2. Distribution of gas composition
Fig. 3. Distribution of gas p composition
along the gasifier at EC gasification.
along the gasifier at SC gasification.
The coals gasification degree (Fig. 5) increases with the length of the gasifier and at the reactor output it is 100%. This indicates the completion of the gasification process of all three coals. Complete conversion of EC is reached faster than for SC and MC, which is associated with higher temperatures (Fig. 6), and, accordingly, the specific power consumption for the gasification of EC (Fig. 7).
Fig. 6 shows that all curves have a maximum temperature in the range 2000 - 2600 K at the exit of the reactor (X = 0.3 m). Moreover, for all the coals gas temperature near the maximum exceeds the temperature of coal particles by 250 - 300 degrees. This is due to the dominance of the heterogeneous endothermic reactions (7) and (8) over exothermic reactions (9) and (10), Table 3, at the coals plasma-steam gasification. At the gasifier exit the difference between gas and particles temperature decreases to 40-60 degrees, the temperature of gasification products is reduced to 1270 - 1400 K.
100
100
H2 CO
H2O
1
80
2 3
CH4
60
Xc, %
Ci, vol.%
10
CH2
40
1
0.1 0.0
20
H
C6H6 0.2
0.4
0.6
0 0.0
0.8
0.1
X, m
0.2
0.3
X, m
Fig. 4. Distribution of gas composition along
Fig. 5. Coal gasification degree distribution
the gasifier at MC gasification.
along the gasifier: 1 – EC, 2 – SC, 3 – MC.
3.0
1
2500
1
2
2.5
2
T, K
6
3 5
4
1500
Qsp, kW ⋅h/kg
2000
1000
3
2.0 1.5 1.0 0.5
500 0.0
0.2
0.4
0.6
0.8
X, m
0.0 0.0
0.1
0.2
0.3
0.4
X, m
Fig. 6. Gas (1, 3, 5) and coal particles (2, 4, 6)
Fig. 7. Specific power consumption
temperature distribution along the gasifier: 1,
distribution along the gasifier: 1 – EC, 2 –
2 – EC; 3, 4 – SC; 5, 6 – MC.
SC, 3 – MC.
Specific power consumption (Fig. 7) increases in the length of the gasifier, reaching a maximum at the exit of the reactor (X = 0.3 m). Specific power consumption for EC gasification reaches 2.75 kW·h/kg, which is considerably higher than for SC and MC gasification (2.45 and 2.39 kW·h/kg, respectively). Specific yield of the gas determined as the ratio of flow of product gas to coal consumption, increases along the gasifier, peaking toward the exit of the reactor (Fig. 8). The yield of gas for low ash content coals (SC and MC) is 30 % higher than for the high-ash EC, although even in the latter case, 1.3 kg of gas is produced from 1 kg of coal. It follows from Fig. 9 that calorific value of the product gas for all three coals reaches a considerable value at the gasifier exit and varies in the range of 4,358 - 4,555 kcal/kg. Greater calorific value of gas produced at the high-ash EC plasma gasification is associated with a higher concentration of hydrogen in the obtained synthesis gas, in comparison to those obtained from SC and MC (Figs. 2 - 4).
5000
3 2
1.6
4500
1.4
Qg, kcal/kg
Gas Yield, kg/kg
1.8
1
1.2 1.0
1 2 3
4000
3500
0.8 0.0
0.2
0.4
0.6
0.8
3000 0.0
0.2
X, m
0.4
0.6
0.8
X, m
Fig. 8. Specific yield of gas distribution
Fig. 9. Gas calorific value distribution along
along the gasifier: 1 – EC, 2 – SC, 3 – MC.
the gasifier: 1 – EC, 2 – SC, 3 – MC.
5. Conclusions Numerical and experimental investigations of three different power coals plasma gasification showed the possibility to produce the high-quality synthesis gas, regardless of the quality of the gasified coal. The produced gas can be used as a high-energy gas, high potential gas – reducing agent and as a raw material for methanol and dimethyl ether synthesis.
Performed verification of the computer-code PLASMA-COAL confirmed its validity for simulations of solid fuels plasma gasification. Numerical simulations have shown that the synthesis gas concentration at the outlet of the gasifier reaches high values for all studied coals and varies in the range of 96.4 - 98.7%, whereas specific power consumption for the gasification process does not exceed 2.75 kW·h/kg. Regardless of the coal quality, its steam plasma gasification provides the high calorific value gas (4358 - 4555 kcal/kg) at the gas specific yield of 1.3-1.83 kg/kg. References [1] Key World Energy Statistics 2003 Edition, International Energy Agency, OECD/IEA, Paris, www.iea.org [2] British Petrol Statistical Review of World Energy, June 2002, British Petrol, London, www.bp.com [3] Cletcins K. World power policy. Using a technology of a three-stage combustion for NOx suppression on solid fuel boilers in Europe and CIS // Opening Rep. Europ. Commission for Power Engineering and Transport. – Moscow: Russian J.S.Co. “United Power System of Russia”. All-Russian Technical Institute. 2000. P. 4-17. [4] Messerle V.E. Ustimenko A.B. Solid Fuel Plasma Gasification. // Advanced Combustion and Aerothermal Technologies, N.Syred and A.Khalatov (eds.), Springer. 2007. P.141-156. [5] Gorokhovski M., Karpenko E.I., Lockwood F.C., Messerle V.E., Trusov B.G., Ustimenko A.B. Plasma Technologies for Solid Fuels: Experiment and Theory. // Journal of the Energy Institute. 78 (4), 2005. P. 157-171. [6] Kalinenko RA, Levitski A.A., Messerle V.E., Polak L.S., Sakipov Z.B., Ustimenko A.B. Pulverized Coal Plasma Gasification // Plasma Chemistry and Plasma Processing. V. 13. N 1. 1993. P. 141-167. New-York, London, Paris.
Oviedo ICCS&T 2011. Extended Abstract
Coprocessing of Low-Rank Coal and Biomass Utilizing Mild Solvent Treatment at around 350°C X. Li1, J. Wannapeera2, N. Worasuwannarak2, R. Ashida1 and K. Miura1* 1
Department of Chemical Engineering, Kyoto University, Kyoto 615-8510, Japan The Joint Graduate School of Energy and Environment, King Mongkut's University of Technology Thonburi, 126 Pracha-Uthit Road, Bangmod, Tungkru, Bangkok, 10140, Thailand *e-mail: [email protected] 2
Abstract The authors have been proposing methods to dewater, and fractionate coal by using the sequential thermal solvent extraction. They have recently showed that the degradative extraction at around 350°C using a non-polar solvent such as 1-methylnaphthalene under 10 MPa is effective to recover several fractions having similar chemical and physical properties from a wide range of low-rank coals. In this paper the applicability of the method to much lower grade carbonaceous resources is examined. Two biomasses, cellulose, lignite, a peat and four low rank coals, and mixed samples of biomass and two coals were treated by 1-methylnaphthalene at 350°C to upgrade and to fractionate the raw materials into several fractions having similar chemical and physical properties. The extracted fractions, Solubles and Deposits, were very close to each other in elemental composition, chemical structure, molecular weight distribution, pyrolysis behavior, and softening/melting behavior.
Thus, the proposed degradative solvent
extraction method was found to be effective to convert low grade carbonaceous resources into solid fuels having higher heating values and significantly upgraded compounds having similar chemical and physical properties.
1. Introduction It is beyond question that coal is a valuable resource expected as not only fuels but also chemical feedstocks in this century. On the other hand, the minable reserves of high grade coal, bituminous coal, have been depleting very rapidly due to rapid increase of worldwide coal consumption. This inevitably requests to utilize low-rank coals, brown coal/lignite and sub-bituminous coal, instead of high grade coal, because the minable reserve of the low-rank coals is as large as that of high grade coal. Peat being distributed worldwide and biomass wastes are also low-grade but valuable carbonaceous resources that must be utilized effectively
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1
Oviedo ICCS&T 2011. Extended Abstract
These low-grade carbonaceous resources contain a large amount of water (~ 60 %) in general, resulting in low calorific value. This means that dewatering or drying is essential when these resources are transported and stored to be utilized. Dewatered samples unfortunately tend to have high spontaneous combustion tendency as compared to non-dewatered raw materials.
It is required to reduce the amount of oxygen
functional groups to suppress the spontaneous combustibility. The process to reduce the oxygen functional groups is called upgrading.
Therefore, both dewatering and
upgrading are necessary to utilize these resources more effectively even as just fuels. If we utilize these resources as feedstocks of chemicals and materials, we might have to develop methods that enable us to effectively recover precursors of chemicals and/or materials from the raw low-grade resources. The authors have been proposing methods to dewater, and fractionate coal by using the sequential thermal solvent extraction (1-4). They have recently showed that the degradative extraction at around 350°C using a non-polar solvent such as 1methylnaphthalene (1-MN) under 10 MPa is effective to recover several fractions having similar chemical and physical properties from a wide range of low-rank coals (5, 6). In this paper the applicability of the method to much lower grade carbonaceous resources is examined.
2. Experimental section 2.1 Samples and solvents used Table 1 summarizes the ultimate and proximate analyses of the samples used. Reagent grade cellulose (Cell) and lignin (LN), two biomasses (leucaena (LC) and rice straw (RS) from Thailand), a peat (PE) from Belarus, a lignite from Thailand (Mae Moh; MM), and brown coals from Indonesia (WA and BB) and Australia (LY) were employed as low-grade carbonaceous resources. LC and RS were pre-dried, PE and coal samples contained water by more than 12.2 %. Four combinations of one to one mixed samples of biomass and coal, LC/MM. RS/MM, LC/LY, and RS/LY were also used as samples to examine their synergetic effect during the solvent treatment. 1-methylnaphthalene (1MN) was used as a non hydrogen donor solvent for the degradative extraction. 2.2 Experimental procedure Since the treatment at around 350°C was found to be effective to upgrade various low rank coals in the previous works, the degradative extraction of the sample was Submit before 31 May 2011 to [email protected]
2
Oviedo ICCS&T 2011. Extended Abstract
Table 1. Analyses of samples used Sample (abbreviation) Cellulose (CELL) Lignin (LN) Leucaena (LC) Rice straw (RS) Peat (PE) Mae Moh (MM) Loy Yang (LY) Wara (WA) Berau BInungan (BB)
Ultimate analysis (wt%, d.a.f.) C H N O+S(diff) 41.2 6.1 0.3 52.4 60.3 4.9 0.3 34.5 49.5 5.9 0.8 43.9 45.7 5.9 0.9 47.5 59.7 4.4 2.2 33.7 66.4 3.9 1.9 27.8 66.7 4.7 0.9 27.7 67.1 5.1 1.0 26.9 71.0 4.9 1.3 22.8
Proximate Analysis (wt%, d.b.) VM FC Ash 92.4 7.6 0.0 66.1 20.6 13.3 85.1 14.1 0.8 69.5 11.4 19.1 46.1 40.2 13.6 50.2 24.0 25.8 51.5 47.0 1.5 50.5 47.9 1.5 43.4 52.5 4.1
Moisture [wt%] 0.0 13.0 3.3 5.5 32.4 12.2 56.3 37.0 21.9
performed using a stainless steel autoclave (350 cm3, 55 mm I.D.) at 350°C. The autoclave was charged with samples (13 g on dry basis) and 310 cm3 of 1-MN. A stainless filter (65 mm O.D. and 0.5 µm opening) was equipped at the bottom of the autoclave. After sufficiently purging the autoclave with N2, the autoclave was heated up to 350°C, where it was kept for 60 min. The extract and the residue (Residue) were separated by opening the valve connected below the filter at the extraction temperature. The extract with the solvent was collected in a stainless steel vessel equipped under the valve which was cooled by water. The solvent containing the extract was filtrated using a PTFE membrane filter (0.5 μm opening) to separate the extract into the extract that precipitates at room temperature (Deposit) and the extract soluble in solvent even at room temperature. The latter fraction was treated by a rotary evaporator at around 140°C under reduced pressure to remove 1-MN and to recover the extract in this fraction as solid. The extract recovered as solid was called “Soluble” and that evaporated with 1MN was called “Liquid”. The yields of Residue, Deposit, and Soluble were determined by weight, and the Liquid yield was estimated by difference. The yield of H2O formed was determined by the oxygen balance.
3. Results and Discussion 3.1 Yields of products As stressed in the previous works, the water in the sample was completely removed while the sample was heated up to 350°C without phase change. The water removed is easily separated from solvent by decantation. Figure 1 shows the yields obtained for the nine samples through the degradative extraction on weight basis. The yields were all represented on the basis of free from ash and water (d.a.f.). The yields of Liquid look very large for Cell, LC, and RS. Elemental Submit before 31 May 2011 to [email protected]
3
Oviedo ICCS&T 2011. Extended Abstract
CO2
Yield [wt%, d.a.f.]
80
100
Carbon distribution [wt%, d.a.f.]
Other gas 100
H2O Liquid (diff.)
60
40
Soluble Deposit
20
CO2
60
Soluble 40
Deposit 20
Residue
Residue 0
0 Cell
LN
LC
RS
PE
MM
LY
WA
Figure 1 Yields of the products.
BB
Liquid
Other gas 80
Cell
LN
LC
RS
PE
MM
LY
WA
BB
Figure 2 Carbon distributions in the products.
balance shows, however, the Liquid consists mostly of oxygen. The Liquid from Cell, for example, consists 8.2 kg of C, 3.8 kg of H, and 42.5 kg of O on the basis of 100 kg of d.a.f Cell. This suggests that the yield of organic compound in Liquid is very small and that most of Liquid is H2O. Then from both practical and fundamental viewpoints the distribution of carbon in the product is more informative than the yield on weight basis. The yields shown in Figure 1 were then converted to the carbon distribution to the product, and they are shown in Figure 2. The formation of CO2 and hence the loss of carbon as CO2 is inevitable, because the treatment is intended to remove oxygen functional groups as CO2 and H2O. The losses of carbon as CO2 were around 5 % for Cell, LN, and biomasses, and less than 3 % for coals. About 20 % of carbon was recovered as Liquid for Cell and LC, but only 2.6 % as Liquid for LN. These results show that most of the products are recovered as solids. The largest extract faction is Soluble for every sample. The carbons in Soluble were respectively as large as 58 %, 49 %, and 42 % for Cell, RS, and LC and were 20 to 30 % for other samples. Figures 3 and 4 respectively show the yield of each product and the carbon distribution in the product for the four mixed samples. The experimental results are compared with the distribution calculated assuming no interaction. The experimental and calculated distributions are slightly different for all of the samples, showing the existence of some interaction during the solvent treatment.
The amounts of
experimentally obtained extract were larger for LC/MM and RS/MM and smaller for LC/LY and RS/LY than the calculated ones. More detailed examination is necessary to estimate the importance of the interaction.
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4
Oviedo ICCS&T 2011. Extended Abstract
Other gas
CO2
[wt%, d.a.f.] Yield
80
60
40
100
Carbon distribution [wt.%, d.a.f.]
100
Liquid (diff.)
Soluble Deposit
20 Residue 0 Exp. Cal. LC/MM
Exp. Cal. RS/MM
Exp. Cal. LC/LY
Other gas
Liquid
CO2
80
Soluble 60 Deposit 40 Residue
20 0
Exp. Cal. RS/LY
Exp. Cal. LC/MM
Exp. Cal. RS/MM
Exp. Cal. Exp. Cal. LC/LY RS/LY
Figure 4 Carbon distribution in the product for the mixed samples.
Figure 3 Yields of the products for the mixed samples. 3.2 Properties of products
The carbon contents of the Solubles were as large as 80.0-84.9 % and the hydrogen contents were as large as 6.4-7.8 % on d.a.f. basis. Figure 5 shows the elemental compositions of the raw materials, Solubles, Deposits, and Residues on the H/C vs. O/C diagram. For the Solubles, all of the data points converged. It was also the case with the Deposits. These results show that the proposed degradative extraction can convert the wide range of low grade carbonaceous resources into Solubles and Deposits that have respectively similar elemental compositions. Figure 6 compares the higher heating values (HHV) between raw material and the product (sum of HHV of Soluble, Deposit and Residue on raw material basis). The HHV of the product was slightly larger than that of raw material. This indicates that the 2.0
30
RS LC
1.5
H/C
Soluble 1.0
RS+MM
Lignin
Deposit
LC+MM Peat
Coal
Raw coal Residue Deposit Soluble
0.5 Biomass and BM + Coal 0.0 0.0
0.2
0.4
0.6
Raw Material Residue Deposit Soluble 0.8
HHV [MJ/kg-Raw material, d.a.f.]
Cell
1.0
Raw Material Liquid Soluble Deposit Residue
25 20 15 10 5 0
Cell
LN
LC
RS
PE
MM
LY
WA
BB
O/C
Figure 5 Elemental compositions of raw coals and extraction products on the H/C vs. O/C diagram.
Figure 6 Comparison of higher heating values (HHV) between raw coals and upgraded coals.
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5
Oviedo ICCS&T 2011. Extended Abstract
Cell LN LC
RS LC/LY RS/LY PE MM LY
Normalized displacement [-]
RS/MM
0.0
WA
-0.2
1000 1500 2000 Mass/Charge
Soluble 0.8
MM RS/MM
-0.4
RS -0.6
LY
-0.8
0.6
LC WA
0.4
LC/LY 0.2
RS/MM
Cell -1.0
0
100
200
300
400
500
0.0 600
o
BB
500
1.0
BB PE LN RS/LY
Relative weight [kg/kg-sample, d.a.f.]
Intensity [a.u.]
LC/MM
Temperature [ C]
Figure 8. TG curves (right) and TMA profiles (left) of the Solubles.
Figure 7 Molecular weight distributions of the Solubles. proposed degradative extraction is an endothermic process for these samples, which means the treatment does not lose heating value of raw materials. Figure 7 shows the molecular weight distributions (MWDs) measured by LDTOFMS of the Solubles. The MWDs are very close to each other, and the Soluble consisted of low-molecular-weight compounds of less than 500 in MW. The MWDs of Deposits are also rather close to each other, and Deposit consisted of compounds having less than 800 in MW. Figure 8 shows thermogravimetric (TG) curves (right axis) and thermomechanical analysis (TMA) profiles (left axis) of the Solubles during the heating to 900°C at a heating rate of 10 K/min. Both TG curves and TMA curves were respectively so close to each other.
Since the weight decrease below 400°C was judged to be due to
devolatilization, the Solubles are judged to consist of components having low-molecular weights. All of the Solubles completely melt at less than 100°C before devolatilaization starts. All of the Deposits have similar properties, but they melted at around 250°C.
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6
Oviedo ICCS&T 2011. Extended Abstract
4. Conclusions Two biomasses, cellulose, lignite, a peat and four low rank coals, and mixed samples of biomass and two coals were treated by 1-methylnaphthalene at 350°C to upgrade and to fractionate the raw materials into several fractions having similar chemical and physical properties. The Solubles and Deposits were very close to each other in elemental composition, chemical structure, molecular weight distribution, pyrolysis behavior, and softening/melting behavior. Thus, the proposed degradative solvent extraction method was found to be effective to convert low grade carbonaceous resources into solid fuels having higher heating values and significantly upgraded compounds having similar chemical and physical properties.
References [1] Miura K, Shimada M, Mae K. Extraction of coal at 300 to 350°C to produce
precursors for chemicals. Proceedings of the 15th Pittsburgh Coal Conference, Pittsburgh, 1998, Paper No. 30–1. [2] Miura K, Shimada M, Mae K, Huan YS. Extraction of coal below 350°C in flowing non-polar solvent. Fuel 2001;80:1573–82. [3] Miura K, Nakagawa H, Ashida R, Ihara T. Production of clean fuels by solvent skimming of coal at around 350°C. Fuel 2004;83:733–8. [4] Miura K, Mae K, Ashida R, Tamura T, Ihara T. Dewatering of coal through solvent extraction. Fuel 2002;81:1417–22. [5] Ashida R, Umemoto S, Hasegawa Y, Miura K, Kato K, Saito K, Nomura S. Upgrading of low rank coal through mild solvent treatment at temperatures below 350°C. Proceedings of the 26th Annual International Pittsburgh coal conference, Pittsburgh, 2009, 51/1–51/12. [6] Li X, Hasegawa Y, Morimoto M, Ashida R, Miura K. Conversion of low-rank coals into upgraded coals and extracts having similar chemical and physical properties using degradative solvent extraction. Preprints of Symposia - American Chemical Society, Division of Fuel Chemistry 2010;55:212–3.
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7
Upgrading and dewatering of low rank coals realizing the suppression of self-ignition tendency through solvent treatment at around 350°C H. FUJITSUKA, R. ASHIDA, and K MIURA Department of Chemical Engineering, Kyoto University, Kyoto 615-8510, Japan [email protected]
Abstract We have recently presented a novel method which can not only dewater but upgrade low rank coals under rather mild conditions. The method treats coal in non-polar solvents, such as 1-methylnaphthalene, at temperatures below 350°C. One of the remaining important questions to be considered for the method is if the self-ignition tendency can be suppressed by this solvent treatment method. In this study, self-ignition tendencies were examined for the samples prepared from an Australian brown coal by the proposed treatment. It was found that the treated coals obtained by the solvent treatment had little moisture content and had the heating values corresponding to subbituminous or bituminous coal. The self-ignition tendencies of the extracted fractions were significantly suppressed because of their small pore surface areas. The suppression of self-ignition tendency of the solvent treated coal was also explained by its small pore surface area which was probably caused by the coating of the extracted fractions. Thus, the validity of the proposed solvent treatment method as the dewatering and upgrading method of low rank coal was clarified.
1. Introduction Low rank coals such as brown coals and lignites, abundantly deposited and distributed worldwide, must be important resources for both energy and chemicals in this century. Low rank coals, however, have not been used on a large scale except for at mining sites. This is because low rank coals contain a large amount of water and show high self-ignition tendency when dewatered, which makes their transportation and storage extremely difficult. Therefore, not only dewatering but suppressing spontaneous combustibility of low rank coals is essential for their transportation and storage. Moreover, increasing their energy density, in other words increasing heating values on weight basis, is also needed for their transportation. The last process is generally called
upgrading. Thus, dewatering, upgrading, and suppression of self-ignition tendency are requested for effective use of low rank coals. We have recently presented a novel method which not only dewaters but upgrades low rank coals [Ref. 1, 2]. The method treats coal in non-polar solvents, such as 1-methylnaphthalene, at temperatures below 350°C, and separates the coal into extract, residue, and gaseous product consisting of CO2 and a negligible amount of hydrocarbon gases at the treatment temperature. Inherent water and water produced are almost completely removed from coal as liquid without phase change and separated from solvent by decantation at room temperature. The extract is further separated into solvent-soluble fraction, Soluble, and solvent-insoluble fraction, Deposit, at room temperature. The whole process of this treatment is therefore an efficient dewatering and removal of most of oxygen functional groups as either CO2 or H2O without affecting hydrocarbon moieties of coal. Then all of the three fractions obtained have the heating values corresponding to subbituminous or bituminous coal. Furthermore, the sum of the heating values of the three fractions multiplied by their yields slightly exceeds the d.a.f. basis heating value of original coal. Thus the proposed method was found to be an effective dewatering and upgrading method of low rank coal. One of the remaining important factors to be considered is if the self-ignition tendency of the three fractions obtained is suppressed. In this study self-ignition tendencies were examined for the three fractions and their mixture obtained from an Australian brown coal by the proposed method. Then factors affecting their self-ignition tendency were discussed.
2. Experimental section 2. 1. Solvent treatment procedure An Australian brown coal, Loy Yang (abbreviated to LY), was used in this study. Its analyses are given in Table 1. Its moisture content and O content are respectively as high as 50.6 % (a. r.) and 27.7 % (d. a. f.). The coal was ground into fine particles of less than 53 μm in diameter. Figure 1 shows a schematic diagram of the apparatus used for the treatment of coal in 1-methylnaphthalene (1-MN). A stainless steel autoclave Table 1 Ultimate and proximate analyses of LY coal Ultimate analysis wt%, d. a. f. C H N O
LY
66.7
4.7
0.9
27.7
Atomic ratio ‐ H/C O/C
0.85
0.31
Proximate analysis wt%, d. b. Moisture VM FC Ash wt%, a. r.
51.7
47.8
0.6
50.6
HHV MJ/kg
25.4
Autoclave 350 mL TC
Pressure gauge
Raw Coal Residue
Impeller Furnace
Valve
N2
Stainless Steel Filter 0.5 mm
Gas Soluble
Deposit
Reservoir 350 mL Figure 1 Schematic diagram of the apparatus for solvent treatment
(350 cm3, 55 mm I. D.) was charged with 30 g of as-received LY coal (14 g on d. a. f. basis) and 300 cm3 of 1-MN. A stainless filter (65 mm O. D. and 0.5 μm opening) was equipped at the bottom of the autoclave. After sufficiently purging the autoclave with nitrogen, the sample in the autoclave was heated up to 350°C at which it was kept for 1 h under sufficient agitation. Then the coal was separated into extract, residue, and gaseous product at the treatment temperature by opening the valve connecting the autoclave and the reservoir. The extract was further separated into solvent-soluble fraction, Soluble, and solvent-insoluble fraction, Deposit, at room temperature by filtration. Solvent treated coal (abbreviated to STC) was also prepared under the same conditions but without separating the coal into fractions. For comparison purpose, LY char was prepared by pyrolyzing LY coal at 350°C for 1 h in a helium stream. Carbon type distribution of each sample was measured by 13C-NMR and pore surface area was measured by CO2 adsorption at 25°C.
1.2
600
1
Relative weight
0.8
22 % O2/He
He
0.6
400 200 0 0
50
Oxidation 65°C, 2 h
0.4
Temperature
0.2
100 150 Time min
200
Relative weight kg/kg – d. a. f.
Temperature °C
800
0 250
Figure 2 Typical experimental procedure; temperature profile and weight change 2. 2. Procedure to estimate self-ignition tendency Self-ignition tendency of the solvent treated samples was examined using a thermogravimetric analyzer (Shimadzu TGA-50) connected to a micro gas chromatograph (Varian micro GC, CP-4900). The product gas composition was analyzed in every 80 s using the micro-GC. Experimental procedure is shown in Figure 2. Residue, Deposit and STC were first heated up to 250°C in a helium stream to remove solvent remaining, and then cooled to 65°C. Soluble and char were heated up to 65°C in a helium stream. Then the gas stream was changed to a 22 % oxygen-containing helium stream to absorb oxygen onto samples. The amount of oxygen adsorbed onto the sample after 2 h of adsorption, nO, was estimated from the weight change and the amount of gases formed. The nO value thus estimated was employed as the index of self-ignition tendency. After the adsorption step, the sample was heated up to 700°C at the rate of 20 K/min to estimate its gasification rate.
3. Results and Discussion 3. 1. Yields and properties of solvent treated samples Figure 3 shows the yields, the elemental compositions and the heating values of the prepared samples. All of the products were almost completely free from water as expected. The yields of Soluble and Deposit were 0.13 kg/kg-coal (d. a. f.) and 0.22 kg/kg-coal (d. a. f.) respectively. These two fractions were almost free from ash. The oxygen contents were as small as 10.2 kg/100 kg-coal (d. a. f.) and 16.9 kg/100 kg-coal (d. a. f.) for Soluble and Deposit, respectively. The sum of the yields of Soluble, Deposit
Yield
HN
Sample
C
Raw coal
66.7
STC
66.6 38.5
Residue
4.7 4.4
O
—
25.4
174.1
0.85
26.6 / 31.3
66.0
0.50
29.2 / 14.5
151.8
0.13
30.8 / 4.0
30.3
0.1 2.2
0.22
36.7 / 8.1
18.7
18.5
0.84
23.0 / 27.4
206.2
0.9 27.7 0.7
13.1
2.0 0.5 8.7
Deposit
10.1 0.7 0.1 2.2
Soluble
18.0
Char
HHV Surface area on the raw/ treated Sp kg/kg‐coal coal basis d. a. f. MJ/kg‐d. a. f. m2/g‐sample
61.0
1.7
3.7
0.7
0 20 40 60 80 100 Elemental composition kg/100 kg‐coal, d. a. f.
Figure 3 Elemental compositions, yields and heating values of the treated coals and Residue corresponds to the yield of char, but the elemental compositions are significantly different between the two. The oxygen content of STC was much smaller than that of char. Figure 4 shows the 13C-NMR spectra of the samples prepared. Soluble was richer in aliphatic carbons than the original coal, while other treated coals were less rich in aliphatic carbons than the raw coal. It is also clearly shown that the solvent treated products are free from –COOH. Figure 3 also gives the higher heating values (HHVs) of the samples prepared on
Intensity a.u.
200
150
100
50
Chemical Shift ppm
0
‐40
O‐CH3 ‐CH2 ‐CH3
COOH Ar‐O Ar‐C Bridgehead Ar‐H
Raw coal Residue Deposit Soluble
Raw coal STC Char
Intensity a.u.
O‐CH3 ‐CH2 ‐CH3
COOH Ar‐O Ar‐C Bridgehead Ar‐H
both the sample basis and the raw coal basis. The HHVs on the sample basis,
200
150
100
50
Chemical Shift ppm
0
‐40
Figure 4 Carbon type distributions of the prepared samples measured by 13C-NMR
representing energy density, are 28.7 to 37.4 MJ/kg-d. a. f. for the solvent treated samples. These values are comparable to those of subbituminous or bituminous coal. The sum of the HHVs of the three solvent treated fractions on the raw coal basis is 26.6 MJ/kg-d. a. f. and is slightly larger than the HHV of the raw coal. The above results clearly show that the proposed solvent treatment is very effective for dewatering and upgrading of LY coal. In addition it was found that the treatment does not lose the heating value.
3. 2. Examination of self-ignition tendency of the solvent treated samples Figure 5 shows how oxygen is adsorbed on the samples when they are exposed to the 22 % oxygen-containing helium stream at 65°C. Oxygen was rapidly adsorbed on both char and Residue, and the nO values were 2.3 and 2.2 mg-O/g-sample (d. a. f.) for char and Residue respectively. Oxygen was rapidly adsorbed on the raw coal at only the initial stage, and the nO value was 1.1 mg-O/g-sample (d. a. f.) for the raw coal. Oxygen was steadily adsorbed on Soluble at relatively small adsorption rate, and the nO value was 1.0 mg-O/g-sample (d. a. f.) for Soluble. Oxygen uptake was saturated at around 30 min for Deposit, and its nO value was only 0.3 mg-O/g-sample (d. a. f.). The oxygen uptake behavior of STC could be explained by the oxygen uptake behaviors of Soluble, Deposit, and Residue. The nO values in Figure 5 show that the self-ignition tendency of Deposit and Soluble are respectively much lower and slightly lower than that of the raw coal, suggesting that the proposed solvent treatment is effective also to suppress the self-ignition tendency of the extract. nO
mg‐O/g‐sample d. a. f.
Oxygen adsorption mg‐O/g‐sample d. a. f.
2.5 2.0
65°C, 22 % O2/He
Residue STC
1.5
1.6 1.1 1.0
Raw coal
1.0
Soluble
0.5 0 0
2.3 2.2
Char
Deposit 30
60 Time min
90
0.3
120
Figure 5 The oxygen uptake profiles and nO values of the treated coals
The self-ignition tendency of low rank coal is believed to be governed by many factors [Ref. 3]. One of the largest chemical factors is the abundance of aliphatic carbons, and one of influential physical factors is the pore surface area. Soluble was richest and Residue was poorest in the aliphatic carbons, but the nO values in Figure 5 do not reflect the abundance of the aliphatic carbons. Then the pore surface areas of the treated samples, Sp, were measured by the CO2 adsorption method at 25°C, and the Sp values are listed in Figure 3. The Sp value of Soluble was 18.7 m2/g-sample and was the smallest of all of the samples. The Sp value of Deposit was also as small as 30.3 m2/g-sample, whereas the Sp values of the raw coal, char, and Residue were much larger than the Sp value of either Soluble or Deposit. These results show that the small nO value of Soluble is due mainly to its small Sp value and the smallest nO value of Deposit comes from the low content of aliphatic carbons and the small nO value. Summing up the above results, it was clarified that the proposed solvent treatment method is not only very efficient for both dewatering and upgrading but promising for the suppression of self-ignition tendency of low rank coal.
4. Conclusion The posibility of non-polar solvent treatment at 350°C, which realizes dewatering and upgrading, as self-ignition suppression method of low rank coals was examined. It was found that the solvent treated samples contained little amount of water and had the heating values corresponding to subbituminous or bituminous coal. The extracted fractions, Deposit and Soluble, showed significantly low self-ignition tendency because of their small pore surface areas even though Soluble is richer in aliphatic carbons. Thus, the validity of the proposed solvent treatment method as the dewatering and upgrading method of low rank coal was clarified.
References [1] K. Miura, et al., ACS Div. Fuel Chem. 2009; 54-2; 212 [2] X. Li, et al., ACS Div. Fuel Chem. 2010; 55-2; 870 [3] Wang H, Dlugogorski BZ, Kennedy EM. Coal oxidation at low temperatures: oxygen consumption, oxidation products, reaction mechanism and kinetic modeling. Prog. Energy Combust. Sci. 2003; 29; 487–513
Upgrading of Low-quality Coals by Thermal Extraction
T. Takanohashi1, N. Sakimoto1, K. Koyano1, Y. Harada2 and H. Fujimoto3 1 Energy Technology Research Institute, National Institute of Advanced Industrial Science and Technology, 16-1 Onogawa, Tsukuba 3058569, Japan 2 Sakaide Plant, Mitsubishi Chemical Co., 1 Bannosu-cho, Sakaide-shi, Kagawa 762-8510, Japan 3 Steel Research Lab., JFE Steel Co., 1 Kokan-cho, Fukuyama, Hiroshima 721-8510, Japan E-mail: [email protected] Abstract “HyperCoal” (ash-free coal) is produced from low-quality coals, such as subbituminous coals and lignites, by thermal extraction using cost-effective industrial solvents below 400 °C in an inert atmosphere. It was found that some upgrading reactions took place during the thermal extraction at 380 – 420
o
C; degradation of oxygen functional groups and
aromatization of unit structure in coals. As a result, the chemical properties of HyperCoals became similar to those of bituminous coals.
1.
Introduction Coke making is becoming more expensive because of the sudden rise in the price of caking
coals due to the decreasing supply, and so a technology to manufacture good-quality coke from coal blends containing low-quality slightly caking or non-caking coals is strongly required. “HyperCoal” (ash-free coal, HPC) is produced by thermal extraction using cost-effective industrial solvents below 400 °C in an inert atmosphere. Thermal extraction using industrial solvents such as light cycle oil (LCO) and crude methyl naphthalene oil (CMNO) has been applied to produce ash-free coal called HyperCoal (HPC) [1-4]. It has originally a wider temperature range of thermoplasticity during heating [5]. When the HPC was mixed instead of a caking coal in coal blends, a significant improvement in the thermoplasticity of coal blends was observed [6]. HPC can be produced from various ranks of coals including lignite and subbituminous coal, however, the chemical property of HPC is different depending on the coal rank [7]. In addition, the chemical and physical properties of HPC are also dependent upon the manufacturing condition of HPCs. Especially, for low-quality coals such as low-rank and high-sulfur coals, some chemical reactions take place during the thermal
extraction. In the current work, upgrading of low-quality coals by changing the manufacturing conditions such as the extraction temperature and solvent type was investigated.
2. Experimental section Materials
One subbituminous coal (PA) and one lignite (MUL) were used. The coals
were ground and sieved less than 1.0 mm of particle size, and dried under vacuum at 80 °C for 12 h.
In dynamic shear rheometric test (DSR), the samples were ground less than 0.15
mm. The ultimate analysis of coals are shown in Table 1. Reagent grade 1-methyl-naphthalene (1-MN) was used as the thermal extraction solvent without further purification. In addition, 20% of reagent grade indole (IN) was added to 1-MN and the mixture was used as the polar extraction solvent [8].
Table 1 Coal
Ultimate analysis of coal samples used. C
H
N
PA
73.5
5.3
(wt%, daf) 1.9
MUL
65.5
5.0
0.9
Thermal extraction.
S
O(diff)
Ash
0.2
19.1
(wt%, db) 4.4
0.1
28.5
3.2
Thermal extraction of coal was carried out using a flowing solvent
extractor. Details of the extraction procedure are described elsewhere [6]. The coal sample was extracted for 60 min at 300 - 420 oC. After the extraction, the extract was precipitated by adding an excess of n-hexane (400 mL) to the extract solution, and collected as the solid extract (called HyperCoal, HPC). While, the extraction residue (RES) was washed with toluene and acetone. The extract and residue samples were dried at 80°C for 12 hr in a vacuum. The extraction yield of coal was defined as Extraction yield = (1 – Wr/Wc )/(1 – Ash/100) × 100
(1)
where Wc (g), Wr (g), and Ash (wt%, db) are the initial mass coal, the mass of the residue, and the ash content of the initial coal, respectively.
Nuclear magnetic resonance (NMR) measurement
Solid-state 13C nuclear magnetic
resonance (NMR) measurements were made using the CP (cross-polarisation)/MAS (magic angle spinning) method using a Chemagnetics CMX-300 NMR spectrometer. The peaks were assigned according to the chemical shifts of model compounds: the peak from 185 to 170 ppm was assigned to carbonyl carbon (COOH, C=O); the peak from 170 to 148 ppm to phenolic carbon; the peak from 148 to 129 ppm to nonprotonated aromatic carbon; the peak from 129 to 93 ppm to protonated aromatic carbon; the peak from 75 to 50 ppm to aliphatic carbon connected to oxygen; the peak from 50 to 25 ppm to methylene and methyne carbon; and the peak from 25 to 0 ppm to methyl carbon. The fraction of each carbon was obtained from the area intensity ratio of each carbon peak divided by the chemical shifts described above.
Dynamic shear rheometric test. Around 0.4 g of sample was pressed under 100 MPa to a pellet with a diameter of 13.1 mm and height of around 3 mm.
The dynamic shear
rheometer test was carried out usually at 40-550˚C at a heating rate of 3˚C/min under nitrogen flow of 80 L/min.
3.
The detail of procedures was described elsewhere [9].
Results and discussion Table 2 shows H/C and O/C atomic ratios, and the extraction yield of HPCs when the
extraction temperature was changed from 300 to 420 oC. The H/C ratio for both coals was gradually decreasing with an increase in the extraction temperature. The value of HPCs produced at lower temperatures; 300 – 340 oC for PA coal, and 300 oC for MUL coal, was higher than that of raw coals. While, at higher extraction temperatures; 380 – 420 oC for PA coal, and 320 – 420 oC for MUL coal, the value became lower. These results show that low-molecular-weight components were mainly extracted at low temperatures, and heavier components were extracted at higher temperatures. As the extraction temperature is increasing, demethanation and dehydrogenation (aromatization) reactions might have occurred. The O/C ratio was also decreasing with an increase in the temperature. Over around 350 oC, decarboxylation reaction would have occurred. While, the extraction yield gave the maximum at 400 oC for both coals, suggesting that the ratio of heavier components was increasing up to 400 oC. While, retrogressive reactions take place simultaneously over around 400 oC. As a result, at 420 oC the extraction yield might have decreased.
Table 2
H/C and O/C atomic ratios of HPCs and the extraction yield. Temperature [oC]
H/C [-]
O/C [-]
Extraction yiled [wt%, daf]
PA raw coal PAHPC
300 320 340 360 380 400 420
0.86 0.93 0.91 0.87 0.86 0.82 0.79 0.77
0.19 0.15 0.14 0.13 0.13 0.13 0.13 0.09
18.6 25.6 31.2 37.8 46.0 49.4 39.1
MUL raw coal MULHPC
300 320 340 360 380 400 420
0.91 0.93 0.9 0.85 0.82 0.8 0.76 0.77
0.33 0.18 0.17 0.17 0.16 0.15 0.14 0.14
18.0 22.5 27.3 35.8 40.3 46.7 39.8
Figure 1 shows the relationship between the H/C atomic ratio and O/C one of raw coals and their HPCs produced at different extraction temperatures. For reference, the plots for three bituminous coals and their HPCs produced at 360 oC are also shown. For both low-grade coals, the plot was moving obliquely left downward as the extraction temperature was increasing. Consequently, the plots at 380, 400 and 420 oC for the both low-grade coals are almost within those for bituminous coal samples. This result suggests that HPCs produced at 380 – 420 oC for low-grade coals may have similar chemical properties.
Figure 2 shows the carbon distribution for PA coal. At 360°C, the distribution of PAHPC was similar to that of the raw coal. As the extraction temperature was increased to 380 and 400°C, the ratios of methylene (CH2) and methyl (CH3) groups decreased and the ratio of protonated aromatic carbon (Ar-H) increased. These results suggest that with increasing extraction temperature, the aromaticity of the extract components obtained became higher, in agreement with the ultimate analysis.
0.95 bituminous raw PA raw MUL raw
bituminous HPC PA HPC MUL HPC
H/C atomic ratio [-]
0.9 360
0.85 380
0.8
400
420
0.75
0.7
0
0.05
360 380 400 420
0.1 0.15 0.2 0.25 O/C atomic ratio [-]
0.3
0.35
Figure 1 Relationship between H/C and O/C atomic ratio for various kinds of raw coals and their HPCs.
PA R aw C =O P henolic A r-C A r-H -C H 2-O CH2 CH3
P AR ES (360) P AH PC (360) P AH PC (380) P AH PC (400) 0
20
40
60
80
100
Ratio [%] Figure 2 Carbon distribution of PA raw coal, its extraction residue and HyperCoals obtained at different extraction temperatures.
Conclusions HyperCoal (HPC) was produced from two low-grade coals (PA subbituminous coal and MUL lignite) at different extraction temperatures. For both coals, the H/C and O/C atomic ratios of HPCs were gradually decreasing with an increase in the extraction temperature from 300 oC to 420 oC. Their H/C and O/C ratios of HPCs obtained at 380 – 420 oC are almost within those for bituminous coals, suggesting that some upgrading reactions such as aromatization and decarboxylation reactions, took place during the thermal extraction.
Acknowledgment This work has been done in Development of Cokemaking Technology from Low-grade coals and Nonconventional Carbon Resources, Division of High-Temperature Processes, Academic Society, the Iron and Steel Institute of Japan. The authors would like to acknowledge the research group members gratefully.
References [1] Yoshida T, Takanohashi T, Sakanishi K, Saito I, Fujita M, Mashimo K, Energy Fuels, 2002, 16: 1006 [2] Okuyama N, Komatsu N, Shigehisa T, Kaneko T, Tsuruya S, Fuel Proc. Tech., 2004, 85: 947 [3]Yoshida T, Li C, Takanohashi T, Matsumura A, Sato S, Saito I, Fuel Proc. Tech., 2004, 86: 61 [4] Kashimura N, Takanohashi T, Saito I, Energy Fuels, 2006, 20: 1605 [5] Takanohashi T, Shishido T, Kawashima H, Saito I, Fuel, 2008, 87: 592 [6] Takanohashi T, Shishido T, Saito I, Energy Fuels, 2008, 22: 1779 [7] Koyano K, Takanohashi T, Saito I, Energy Fuels, in press. [8] Kashimura N, Takanohashi T, Saito I, Energy Fuels, 2006, 20: 2063 [9] Yoshida T, Iino M, Takanohashi T, Kato K, Fuel, 2000, 79: 399.
Arsenic leachability and speciation in fly ashes from coal fired power plants S. Kambara*, M. Endo, S. Takata, K. Kumabe, H. Moritomi Gifu University, Environmental and Renewable Energy Systems Division, Graduate School of Engineering, 1-1 Yanagido, Gifu, 501-1193, Japan *Corresponding author: [email protected] Abstract To determine dominant factors on arsenic leaching from the coal fly ash, arsenic leaching test under a constant pH was performed. Twelve fly ash samples were collected from two coal fired power plants (600 MWe) having different boiler types. Arsenic in the raw coal was almost all associated with the fly ash in both power plants; however arsenic leaching fraction was strongly differed in boiler types. It found that the dominant factors of arsenic leaching were calcium contents in fly ashes and ash contents in raw coals. 1. Introduction Major fractions of coal fly ash generated from pulverized coal combustion processes have been used as fill materials to reclaim land from the sea. Coal fly ash contains hazardous toxic elements such as arsenic, and often contains elevated concentrations of the toxic elements. An ash storage area is usually holding seawater and rainwater (called excess water), therefore some elements including arsenic in the fly ash are leach out the excess water. If arsenic concentration in the excess water exceeds the environmental limit (0.1 mg/L in Japan), the excess water can not be drained to the sea. This is a serious situation, because ash storage has to discontinue. In this concern, it is important to find leachability of arsenic from the fly ashes for various coal types. Another interest is effect of the boiler types on the leachability, because the ash properties and arsenic partitioning may be changed by types of coal fired boilers. Some researchers investigated As partitioning and its mechanisms during coal combustion [1, 2]. It concluded that As in raw coal was released as vapor at high temperature during combustion, and generated gaseous arsenic oxide reacted with calcium oxide on fly ash. Consequently, Ca 3 (AsO 4 ) 2 is formed on fly ash surface, which is the most thermodynamically stable calcium–arsenic compound under conditions of coal fired boilers [3]. Oxidation state of As (+3 and +5) is an important factor in controlling As leachability [4]: hence, As leachability may depend on combustion conditions. In this paper, As leachability was investigated for various coal fly ashes collected from two different power plants. Effects of Ca and boiler types on As leachability were discussed.
2. Experimental section 2.1. Fly ash samples Six fly ash samples were carefully collected from each coal fired power plants (Unit A and Unit B: 600 MWe). Fig. 1 depicts the process flow of the plants, ash collection locations, and typical gas temperatures between the boiler exit and the low temeprature electrostatic presipipator (ESP). The unit B has a DeNOx (SCR) system. To prevent contamination of samples, after enough time from coal switching, the ash sampling was began at each chamber (#1, #2, and #3). Table 1 lists coal properties and ash composition. Coal F and G, and coal H and I are same coal between unit A and B. 370℃
350℃
Boiler
145℃
A/H
GGH
ESP
FGD
DeNOx (Unit B only) #1
Eco-hopper Clinker
#2
#3
85% 10% 5% Multi Cyclone
(Ash partitioning)
Flue gas Ash collection
Fig. 1. Process flow of the coal fired power plants and ash collection points. Table 1. Properties of raw coals and fly ashes collected from #1 chamber of ESPs. Power station
Unit A
Unit B
Key E F H O P R G I K L M Q
Raw coal (on dry basis) C Ash As wt% wt% mg/kg 67.9 14.3 2.14 71.5 13.3 0.84 68.3 10.4 3.69 69.6 9.7 1.45 70.9 13.0 0.78 76.5 9.5 0.88 71.5 13.3 0.84 68.3 10.4 3.69 67.9 13.9 1.35 73.1 10.3 0.87 73.0 9.7 1.53 74.0 9.5 1.02
As mg/kg 12.16 3.16 26.46 15.65 4.96 8.23 4.53 39.22 8.85 9.46 10.41 7.48
SiO2 wt% 55.5 67.0 59.3 75.7 62.1 62.6 65.4 59.0 56.1 58.1 64.5 62.3
Fly ashes (on dry basis) Al2O3 Fe2O3 CaO Na2O wt% wt% wt% wt% 31.2 5.35 2.18 1.17 26.2 2.26 0.68 0.26 25.6 7.49 2.05 0.60 17.2 2.79 0.97 0.47 26.5 4.77 1.68 0.95 28.7 3.86 0.93 0.45 26.5 3.18 0.93 0.28 26.0 7.25 2.09 0.65 20.6 7.80 9.46 0.71 21.4 6.40 8.24 0.83 22.9 6.31 1.46 0.51 27.8 4.04 1.39 0.73
K2O wt% 1.18 0.60 1.56 0.94 0.98 0.69 0.56 1.50 2.04 1.86 1.74 0.89
SO3 wt% 0.29 0.24 0.42 0.00 0.15 0.00 0.64 0.51 0.80 0.84 0.34 0.04
2.2. Leaching tests To simulate pH of the excess water, a buffer solution adjusted pH = 10 was prepared as a leaching solvent. The ash sample (1.0 g) was added to the leaching solvent (10 mL), and it was shaken for 30 minute at 200 rpm. After shaking, the solid and the solvent were separated by filtration, and both arsenic concentration of the solid and the solvent was measured by ICP-AES.
3. Results and Discussion 3.1. Arsenic partitioning In pulverized coal combustion processes, arsenic has classified as Group II elements which are not incorporated into the bottom ash. It is believed that volatilized arsenic during combustion is chemically condensed on fly ash at low temperature processes [2]. For twelve fly ash samples, percentage of As partitioning was ranged from 95−157% as calculated from data listed in Table 1, which represented the behavior of Group II elements. To compare arsenic partitioning in the unit A and B, relation between modified arsenic concentration in the raw coals, [As 0 /Ash 0.65 ], and arsenic concentration in the fly ashes, As FA , for the unit A and B is shown in Fig. 2. As 0 and Ash are As concentration and ash content in the raw coals, respectively. It found that As FA can be accurately estimated by the parameter [As 0 /Ash 0.65 ], and arsenic partitioning was same behavior between the unit A and B.
Arsenic leaching fraction [%]
Arsenic conc. in FA [mg/kg]
40 Unit A Unit B
35 30 25 20 15 10 5 0 0
20
40 60 80 As0 /Ash0.65 ×100 [ - ]
100
Fig. 2. Relation between modified As concentration in the raw coals and As concentration in the fly ashes for the unit A and B.
2.5 2.0
Unit A
Unit B
1.5 1.0 0.5 0.0 E F H O P R G I K L M Q Fly ashes
Fig. 3. Variation in As leaching fraction for various fly ash samples and for the unit A and B. (pH of the leaching solvent was fixed on 10.)
3.2. Arsenic leaching Fig. 3 shows the arsenic leaching fraction, L As , for the fly ash samples of the unit A and B. With the unit A, L As was observed in the range of 0.3−3.0%, which was a wide range comparing that of the unit B. Particularly, L As of fly ash F and G, although the both raw coals were same, significantly differed. It clarify that arsenic leaching is affected by boiler types. 3.3. Dominant factors on arsenic leaching Ca 3 (AsO 4 ) 2 is a stable compound formed during combustion, which is an insoluble material: it seems that fly ash types having high CaO/Ash ratios generate much Ca 3 (AsO 4 ) 2 , and have low As leaching fraction. Fig. 4 shows variation in L As for all
fly ash samples as a function of CaO/Ash×100, where CaO is ash composition of the raw coal. L As increased suddenly below CaO/Ash = 50 in both units. CaO/Ash×100 of coal F and G is 6.3: it is reasonable that both fly ashes have high L As among the same unit. However, L As of coal F (unit A) is much higher than that of coal G (the unit B). The reason for the difference may be the difference in CaO content of the fly ash between unit A and B. Fig. 5 shows CaO% in the fly ashes of coal F and G. It found that the fly ash F from the unit A indicated low CaO% compared to the fly ash G. Therefore actual CaO/Ash ratio of the fly ash F was much lower than the appearance CaO/Ash ratio. It is supposed that high L As of the fly ash in unit A is owing to the loss of calcium during combustion.
1.4
Unit A Unit B
2.0
1.2
CaO % in FA
Arsenic leaching fraction [%]
2.5
1.5 1.0 0.5
1.0
Unit A, Coal F Unit B, Coal G
0.8 0.6 0.4 0.2
0.0
0.0
0
50
100 150 CaO/Ash×100 [-]
200
Fig. 4. Variation in LAs% as a function of CaO/Ash ratios for the unit A and B.
#1
#2 #3 Sampling location
Average
Fig. 5. Different in CaO% in the fly ashes between the unit A and B.
4. Conclusions Arsenic petitioning in the unit A and unit B represented the same behavior. Most arsenic in the raw coal associated with the fly ash for various coal types. However, arsenic leaching fraction of the fly ashes in the buffer solution (pH = 10) was strongly affected by coal types and boiler types. It was found that arsenic content, calcium content, and ash content were the dominant factors controlling As leachability. References [1] Seames WS, Wendt JOL, Regimes of association of arsenic and selenium during pulverized coal combustion, Proc. Combust. Inst., 2007;31:2839–2846. [2] Senior CL., Lignell DO., Sarofim AF., Mehta A., Modeling arsenic partitioning in coal-fired power plants, Combust. Flame, 2006;147:209–221. [3] Frandsen F., Dam-Johansen K., Rasmussen P., Trace elements from combustion and gasification of coal—An equilibrium approach, Prog. Energy Combust. Sci. 20 (1994) 115–138. [4] Jing C., Liu S., Meng X., Arsenic leachability and speciation in cement immobilized water treatment sludge, Chemosphere, 2005;59:1241–1247.
Low temperature SNCR by photochemical activation of ammonia
S. Kambara* 1 , M. Kondo 1 , N. Hishinuma 2 , M. Masui 3 , K. Kumabe 1 , H. Moritomi 1 1 Gifu University, Environmental and Renewable Energy Systems Division, Graduate School of Engineering, 1-1 Yanagido, Gifu, 501-1193, Japan 2 Ushio Inc., 1194 Sazuchi, Bessho-cho, Himeji, Hyogo 671-0224, Japan 3 Actree Co. Ltd., 375 Hakusan, Ishikawa, 924-0053, Japan *Corresponding author: [email protected] Abstract To broaden and lower the temperature window of the selective non catalytic reduction (SNCR) of nitric oxide (NO), the use of activated ammonia was examined. A wavelength of 172 nm was employed as the excitation source for molecular ammonia. Activated ammonia was injected into a model flue gas (NO/O 2 /N 2 ) at room temperature. The effects of reaction temperatures, oxygen concentrations, and NH 3 /NO molar ratios on NO removal were investigated in a lab-scale plug flow reactor. Reaction temperatures ranged from 500 °C to 850 °C. A temperature window enlargement of 150 °C was achieved at the lower boundary of the temperature window. Above 600 °C, NO removal was effected by injection of activated ammonia, while around 750 °C, conventional SNCR by injection of molecular ammonia was effective. An approximate 80% NO removal was attained at 700 °C with an MR = 2.0 and 8.3% O 2 . The formation of nitrous oxide (N 2 O) using activated ammonia SNCR technology was also investigated and was found to be strongly affected by O 2 concentrations, while the concentration of N 2 O increased with an increase in NO removal. 1. Introduction Selective non catalytic reduction (SNCR) techniques are a conceptually simple process that involves injecting molecular ammonia into the furnace without using a catalyst [1]. SNCR systems seem to be a promising technology because of their cost-effectiveness, although critical issues regarding their application still exist. In SNCR systems, NO x reduction occur at temperatures between 850 °C and 1175 °C (temperature window), however, high enough NOx reduction was not obtain in large-scale combustors [2]. To improve NOx reduction efficiency, the expansion of the temperature window is desired. It has been reported that chemical additives together with molecular ammonia can lower and widen the temperature window. Various additives have been studied, including hydrogen, hydrogen peroxide, hydrocarbons, and carbon monoxide [3], all of which are effective. However, utilization of chemical additives increases the cost
of NO reduction. A recognized research goal is to expand the temperature window without the need for additives. The aim of the research presented in this study was to find an alternative method of producing effective chemical species for NO removal without the use of argon gas. 2. Experimental section The experimental setup is shown in Fig. 1. The apparatus consists of two gold furnaces with quartz tubes, the gas mixing and flow control systems, the photochemical reactor, and the gas analyzers. The quartz tubes were connected via the mixing chamber. An NO/O 2 /N 2 gas mixture was prepared as the model flue gas, and fed into the pre-heater quartz tube. Ammonia gas diluted with nitrogen was used as the NO removal agent, which was fed into the photochemical reactor at room temperature. Molecular ammonia is excited by photons emitted from the excimer lamp in the photochemical reactor. In this paper, the chemical species generated by VUV radiation will be termed “activated ammonia”. Activated ammonia was introduced into the mixing chamber. The reaction temperature was varied from 500 °C to 850 °C. The total gas flow rate of the NO/O 2 /NH 3 /N 2 gas mixture was fixed at 3.0 L/min for all experimental conditions. The gas composition of the output stream was continuously measured by gas analyzers for NO x , O 2 , and N 2 O. Fig. 2 depicts the configuration of the photochemical reactor. The excimer lamp (USHIO Inc.) was placed on top of the center of the power unit. An aluminum cylindrical chamber coaxial in configuration to the excimer lamp was fitted around the excimer lamp. An NH 3 /N 2 gas mixture was fed into the gap between the excimer lamp and the inside wall of the cylindrical chamber. The radiation power of the VUV ray was 26 mW/cm 2 on the quartz glass surface of the excimer lamp.
Fig. 1. Experimental setup.
Fig. 2. Configuration of the photochemical reactor.
3. Results and Discussion 3.1. Effect of reaction temperatures The relationship between reaction temperatures and NO removal in both SNCR systems is presented in Fig. 3. In conventional SNCR system, it can be observed that an NO removal of 3–4% was effected at 750 °C and increased significantly above 800 °C for both molar ratios of 1.0 and 1.5. In the activated ammonia SNCR, slight NO removal began at the reaction temperature of 600 °C and increased almost proportionally with a further increase in the
reaction
temperature.
It
clearly
indicates that the injection of activated ammonia
broadened
window
and
the
lowered
temperature its
starting
temperature. The temperature shift was 150 °C at an NO removal of 20%. This result suggests that effective chemical species for NO removal were formed from activated ammonia generated by VUV
radiation
at
the
reaction
temperature above 600 °C.
Fig. 3. NO removal performances of conventional SNCR (Thermal) and activated ammonia SNCR.
3.2. Effect of molar ratios and oxygen concentrations Fig. 4 shows the variation in NO removal with the variation of NH 3 /NO molar ratios in the range of 1.0–3.5 at 700 °C. With the activated ammonia SNCR, NO removal proportionally increased with increasing the molar ratios up to MR = 2.0, beyond which a gradual increase in NO removal was observed. With conventional SNCR, because the reaction temperature of 700 °C was outside the temperature window, slight NO removal was observed even at high molar ratios. Approximately 80% NO removal was obtained at MR = 2.0 at 700 °C using the activated ammonia SNCR technology. The effect of the oxygen concentration on NO removal in the MR range of 1.0 to 2.0 is also shown in Fig. 4. A slight increase of about 5% NO removal was observed at 8.3% O 2 compared to that at 2.1% O 2 for the activated ammonia technology. With the conventional SNCR, the O 2 concentration had a weak effect on NO removal, although NO removal increased monotonically with O 2 concentration. This suggested that the reaction mechanisms for NO removal in the activated ammonia SNCR was caused by similar elemental reaction pathways to the conventional SNCR, although the reaction temperatures differed for both SNCR systems. 3.3. N 2 O formation In conventional SNCR systems, N 2 O is usually generated as a byproduct during the
reactions that effect NO removal. Fig. 5 shows the N 2 O formation with the activated ammonia SNCR as a function of NO removal with the different oxygen concentrations and NH 3 /NO molar ratios. It can be seen that N 2 O concentrations increased with increasing NO removal for all experimental conditions. The increase in the O 2 concentration significantly promoted N 2 O formation, while the molar ratio did not have much effect on N 2 O formation. The maximum N 2 O concentration of 17 ppm was detected at 80% NO removal with 8.3% O 2 and MR = 1.5.
Fig. 4. Variation in NO removal with NH 3 /NO molar ratios in activated ammonia SNCR and conventional SNCR at temperature of 700 °C.
Fig. 5. N 2 O formation with activated ammonia SNCR as a function of NO removal with different O 2 levels and NH 3 /NO molar ratios.
4. Conclusions With the activated ammonia SNCR, NO removal began at a reaction temperature of 600 °C, and NO removal almost proportionally increased with a further increase in reaction temperature. Approximately 80% NO removal was obtained at a molar ratio of 2.0 at 700 °C. Oxygen concentrations had little influence on NO removal: an increase of about 5% NO removal was observed with an increase from 2.1% to 8.3% O 2 . However, an increase in the oxygen concentration caused the formation of N 2 O. The molar ratios had a large impact on NO removal. Acknowledgment The authors would like to acknowledge that funding for this study was provided by the Japan Science and Technology Agency through the Adaptable and Seamless Technology Transfer Program (A-STEP). References [1] Lyon RK. Method for the reduction of the concentration of NO in combustion effluents using NH 3 . US. Patent 3900554. 1975. [2] Jodal, M., Nielsen, C., Hulgaard, T., Dam-Johansen, K., Pilot-scale experiments with NH 3 and urea as reductants in selective non-catalytic reduction of nitric oxide. 23rd Symp. (Int.) on Combus. 1990; 237–243. [3] Javeda MT, Irfana N, Gibbs BM. Control of combustion-generated nitrogen oxides by selective non-catalytic reduction. J. Env. Manage. 2007;83:251–289.
Oviedo ICCS&T 2011. Extended Abstract
Optimum temperature for sulphur retention in fluidised beds working under oxy-fuel combustion conditions A. Rufas, M. de las Obras-Loscertales, L.F. de Diego, F. García-Labiano, A. Abad, P. Gayán, J. Adánez. Dept. Energy and Environment, Instituto de Carboquímica (ICB-CSIC) Miguel Luesma Castán 4, 50018 Zaragoza. Spain Corresponding author: [email protected]
Keywords: Oxy-fuel combustion, SO2 retention, limestone, fluidised bed. Abstract In the oxy-fuel combustion process the fuel is burned in pure oxygen gas diluted with recirculated flue gas, mainly composed of CO2, producing a gas stream leaving the combustor highly concentrated on CO2, up to 95%, which greatly facilitates the capture of CO2. The fluidised bed combustors are very appropriate for this process, allowing additionally the in situ desulphurisation by feeding calcium-based sorbents into the combustor. In this work, the effect of the temperature of the combustor on the retention of the SO2 generated in the combustion of two coals with very different sulphur content (a lignite and an anthracite) has been studied. The experimental facility used has been a bubbling fluidised bed (BFB) combustor operating under oxy-fuel combustion conditions. Several tests were also carried out under enriched air combustion conditions for comparison reasons. A Spanish limestone “Granicarb” was used as Ca-based sorbent for sulphur retention. The temperatures tested were between 800 and 950 ºC using Ca/S molar ratios of 2 and 3. It was found that under oxy-fuel combustion conditions in BFB the optimum temperature to achieve the highest sulphur retention was 900-925 ºC, whereas in enriched air combustion the optimum was 850 ºC. Working at the optimum temperature for each case, the SO2 retentions were higher in enriched air combustion than in oxyfuel combustion conditions.
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Oviedo ICCS&T 2011. Extended Abstract
1. Introduction Carbon dioxide (CO2) and sulphur dioxide (SO2) emissions are a major concern in combustion processes using fossil fuels, as for example coal; the former is implicated in global climate change and the latter produces acid rain. CO2 is one of the major contributors to the build-up of greenhouse gases in the atmosphere. At the same time, fuels containing sulphur produce SO2 emission during combustion. The capture and storage of CO2, emitted in large quantities from power plants, is considered an option to be explored in the medium term for reducing CO2 levels released to the atmosphere. Oxy-fuel combustion is one of the possibilities under investigation within the different options for CO2 capture [1-3]. This technology uses for combustion pure O2 mixed with recirculated flue gases, instead of air used in conventional combustion, and so, the flue gas stream finally produced is highly concentrated on CO2. It is believed that an oxy-fuel circulating fluidised bed (CFB) combustor will be an important candidate for new coal fired power plants [4-5] mainly because the circulation of solids in the combustor can help to an effective control of the temperature. Other of the best known advantage of fluidised bed (FB) combustion is the in-situ desulphurisation of the flue gas with the addition of Ca-based sorbents such as limestone or dolomite. Sulphur capture with these sorbents is a process highly dependent on the temperature and CO2 concentration according to the thermodynamic equilibrium curve of CaCO3 calcination, plotted in Figure 1. 100
P CO2 (kPa)
80 OXY-FUEL COMBUSTION
60
CaCO3
40
CaO
20 AIR COMBUSTION
0 700
750
800
850
900
950
1000
Temperature (ºC)
Figure 1. Thermodynamic equilibrium curve of CaCO3 calcination.
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Oviedo ICCS&T 2011. Extended Abstract
In conventional air-combustion in FB boilers, characterised by low CO2 concentrations (~15%) and temperatures about 850 ºC, the operating conditions lead to a previous sorbent calcination (R1) and to the sulphation of calcines (R2), i.e. indirect sulphation: CaCO3
CaO + CO2
CaO + SO2 + ½ O2
(R1) CaSO4
(R2)
However, in oxy-fuel combustion, CO2 concentration in the flue gas may be enriched between 60 and 90%. Under so high CO2 concentration, the sorbent can behave in two ways depending on the temperature. In oxy-fuel combustion conditions at 850 ºC, the sulphur retention will be produced under direct sulphation (R3), being necessary higher temperatures to operate under indirect sulphation. CaCO3 + SO2 + ½ O2
CaSO4 + CO2
(R3)
The objective of this work was to analyse the behaviour of a limestone in a continuous FB combustor working in typical oxy-fuel operating conditions to determine the optimum temperature for SO2 retention using this technology.
2. Experimental section Coal combustion experiments with limestone addition were carried out in a bubbling fluidised bed (BFB) combustor working in oxy-fuel combustion conditions and also with enriched aire. Figure 2 shows a photograph and a schematic diagram of the installation used. The experimental installation consisted mainly of a gas supply (air, CO2, O2) system using mass flow controllers; a fluidised bed reactor, 10 cm internal diameter, that could operate with two bed heights (25 cm and 40 cm); and a freeboard with an internal diameter of 15 cm and a height of 45 cm. The solids were fed to the reactor by means of two screw feeders located just above the distributor plate: One to feed coal and sorbent and the other to feed sand as inert solid. The flow rate of solids entering the reactor was controlled by regulating the velocity of the screw feeders. The facility was equipped with pressure sensors and thermocouples to measure solid inventory, temperature and to control the system. To ensure good temperature control, a heat exchanger was installed. An air pre-heater was used to heat the bed up to the ignition temperature of coal during the start-up of the plant. The flow of combustion gases passed through a high efficiency
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Oviedo ICCS&T 2011. Extended Abstract
cyclone for the retention of elutriated solids and then was sent to the stack. The composition of gases leaving the reactor was analysed continuously by on-line gas analysers.
Heat-exchanger
On-line analysers
Condenser
SO2, O2, CO2, CO, NOx To stack
T5 P3
Cyclone
Cyclone deposit T4 Bubling fluidised bed T3 Coal + Limestone
Drain deposits
T2 Sand P2
Distributor plate
Refrigerated Screw O2
T1 P1
CO2 Air
Preheater
Figure 2. Bubbling Oxy-fuel Fluidized Bed Combustor at ICB-CSIC.
The coals used were a Spanish lignite from Teruel basin with high sulphur content and a spanish anthracite from Bierzo with lower content of sulphur. The calcium sorbent used was the Granicarb limestone. Sand was used as inert solid in the fluidised bed. Tables 1 and 2 show both the composition of the coals and the particle size of the solid materials used.
Table 1. Composition of coals. Proximate analysis (wt%)
Ultimate analysis (wt%)
Lignite
Anthracite
Lignite
Moisture
12.55
1.00
C
45.43
60.66
Ash
25.17
31.55
H
3.90
2.05
Volatiles
28.65
7.55
N
0.65
0.87
Fixed C
33.63
59.90
S
5.17
1.33
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Anthracite
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Oviedo ICCS&T 2011. Extended Abstract
Table 2. Particle size of the solids used in the experimental tests. Material
dp (mm)
Sand
0.2 - 0.6
Coal
0.2 - 1.2
Limestone
0.3 - 0.5
Experiments were carried out at different temperatures between 800 and 950 ºC under oxy-fuel combustion conditions, keeping constant the O2 concentrations (27% O2 -73% CO2 and 35% O2 - 65% CO2), and under enriched air combustion conditions (27% O2 73% N2). Limestone was fed to the reactor, for SO2 retention, using Ca/S molar ratios of 2 and 3. In all of the experiments, the operation conditions were held in a steady state for 2 hours while CO, CO2, SO2 and O2 were measured. Experimental SO2 retention (SR) was calculated, taking into account the feed rate of coal (F0,coal), its ultimate analysis (XSO2,coal), the gas flow rate (Q), and the SO2 concentration (CSO2) measured in the flue gas at the exit of the reactor, by the following equation: SR =
(F0 ,coal X SO 2 ,coal / M SO 2 ) − Q ⋅ CSO 2 F0 ,coal X SO 2 ,coal / M SO 2
being MSO2 the molecular weight of SO2.
3. Results The experiments with the two coals were performed in the BFB combustor at different temperatures between 800 and 950 ºC, with 27 and 35% O2 in the feed gas, and using Ca/S molar ratios of 2 and 3. Coal and limestone were added together to the combustor and the sand was fed continuously to control the residence time of sorbent in the combustor. Influence of temperature. When a Ca-based sorbent is added to a FB combustor, one of the most important parameters affecting the SO2 retention process is the temperature. Figure 3 shows the concentration of the SO2 measured at the exit of the combustor during typical tests
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5
Oviedo ICCS&T 2011. Extended Abstract
working with the lignite in oxy-fuel combustion conditions at the temperatures of 850, 875 and 900 ºC and with a Ca/S molar ratio of 3. An increase in the SO2 emissions was observed when the temperature decreased from 900 to 850 ºC.
SO2 Emissions (vppm)
7000
850 ºC
6000 5000
875 ºC 4000
900 ºC
3000 2000 1000 7
8
9
10
11
12
13
14
Time (h)
Figure 3. SO2 concentration in the outlet gas stream during oxy-fuel combustion of lignite. O2/CO2 : 35/65, Ca/S = 3. 100 90
No-calcining conditions
Calcining conditions
SO2 Retention (%)
80 70 60 50 40 30 20
Lignite Anthracite
10 0 825
850
875
900
925
950
975
Temperature (ºC)
Figure 4. Effect of temperature on the sulphur retention with Granicarb limestone. Ca/S=3; O2/CO2 : 35/65. Calcining and non-calcining conditions calculated with the gas inlet composition.
Figure 4 shows the effect of the temperature on the SO2 retention working with the two coals in oxy-fuel combustion conditions and using a Ca/S molar ratio of 3. For both coals, the sulphur retention initially increased with increasing temperature up to a maximum of 900-925 ºC and then, further increases in temperature caused a decrease in sulphur retention. These results are in good agreement with the results obtained in a TGA and in a batch FB reactor in previous works where the sulphur retention capacity of limestones was analysed [6-7]. In these works, it was found that the effect of the
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Oviedo ICCS&T 2011. Extended Abstract
temperature was different for the direct and indirect sulphation reactions (see Figure 5). In direct sulphation or non-calcining conditions, the sulphation reaction rate rose with increasing the temperature. On the contrary, in conditions of indirect sulphation or calcining conditions, the sulphation reaction rate increased with increasing the temperature up to 900 ºC and then decreased due to the sinterization of the sorbent. As a consequence, the maximum sorbent conversion was reached at temperatures about 900 ºC.
Sulphation conversion
0.6 925
0.5
900 ºC 950 975
0.4
850 0.3
800
0.2 0.1 0.0 0
2
4
6
8
10
Time (h)
Figure 5. Effect of the temperature on the Granicarb sulphation conversion in TGA. dp= 0.1-0.2 mm; 60 vol.% CO2; 3000 vppm SO2. Indirect sulphation (____) and direct sulphation ( - - -).
It can be also observed in Figure 4 that the behaviour of the limestone with both coals was qualitatively similar but higher SO2 retentions, and so, higher limestone sulphation conversions, were achieved working with the lignite. Previous studies in a TGA [6] and and in a batch FB reactor [7] demonstrated that the sulphation conversion of the sorbent increased as the SO2 concentration increased. As the lignite has a higher sulphur content than the anthracite, the SO2 concentration generated in the combustor was higher with the lignite and therefore the limestone was larger sulphated.
Influence of the Ca/S molar ratio As it is well know, the utilisation of Ca-based sorbents for SO2 retention in FB boilers is not complete due to the relatively large particle sizes used and the pore blockage there produced by CaSO4 formation. In the typical operating conditions used in FB combustors, the sulphation reaction usually takes place at the external surface and
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7
Oviedo ICCS&T 2011. Extended Abstract
around the pores of the sorbent particles. Since the molar volume of CaSO4 is higher than the molar volume of CaCO3 or CaO, the pores are blocked and the centre of the particles remains essentially unsulphated. Therefore, an important parameter in the SO2 retention process is the Ca/S molar ratio used in the plant. This Ca/S molar ratio has to be necessarily higher than 1 in order to achieve a significant retention. Figure 6 compares the sulphur retention achieved in oxy-fuel combustion working with the lignite with Ca/S molar ratios 2 and 3. Each data corresponds to 2 hours of steady-state operation in the oxy-fuel plant. Qualitatively the effect of the temperature was very similar for the two Ca/S molar ratios used. However, as expected, an increase in the Ca/S molar ratio produced higher sulphur retentions in all range of temperature studied.
100 90
SO2 Retention (%)
80 70 60 50 40 30 20
Ca/S=3 Ca/S=2
10 0 775
800
825
850
875
900
925
950
975
Temperature (ºC)
Figure 6. Effect of the Ca/S molar ratio on the SO2 retention working with the lignite in oxy-fuel combustion conditions at different temperatures. O2/CO2 : 35/65.
Oxy fuel combustion vs. air combustion In order to compare the oxy-fuel combustion and the conventional combustion with air technologies in fluidised beds, Figure 7 shows the results of the SO2 retention obtained with the Granicarb limestone in conditions of enriched air (O2/N2 : 27/73) and in oxyfuel combustion (O2/CO2 : 27/73) working with the lignite as fuel.
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8
Oviedo ICCS&T 2011. Extended Abstract
100 90
SO2 Retention (%)
80 70
AIR
60 50 40 30
OXY-FUEL
20 No-calcining conditions
10 0 775
800
825
850
Calcining conditions
875
900
925
950
975
Temperature (ºC)
Figure 7. Sulphur retention with Granicarb limestone in the BFB combustor. Coal: lignite; Ca/S=2. Calcining and non-calcining conditions calculated with the gas inlet composition.
As it can be observed in the Figure, the maximum SO2 retention under enriched air combustion was obtained at 850 °C, while under oxy-fuel combustion the maximum retention was obtained at 900-925 °C. In addition, considering the optimal temperature for each case, the SO2 retention obtained in oxy-fuel combustion are lower than that obtained in enriched air combustion conditions.
4. Conclusions - In oxy-fuel combustion with limestone addition in FB, the retention of SO2 were higher working at calcining (indirect sulphation) than at non-calcining (direct sulphation) operating conditions. - The optimum FB combustor temperature from of point of view of sulphur retention shifted from 850 ºC in combustion with air to 900-925 °C in oxy-fuel combustion conditions. - For Granicarb limestone, working at the optimum temperature of combustion for each technology, the SO2 retentions obtained in oxy-fuel combustion conditions were lower than those obtained under air combustion conditions.
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9
Oviedo ICCS&T 2011. Extended Abstract
Acknowledgements This work has been supported by Spanish Ministry of Science and Innovation (MICINN, Project: CTQ2008-05399/PPQ) by FEDER and by Fundación CIUDEN. A. Rufas thanks CSIC for the JAE fellowship and M. de las Obras thanks MICINN for the F.P.I. fellowship.
References [1] Wall T, Liu Y, Spero C, Elliott L. An overview on oxyfuel coal combustion – State of the art research and technology development. Chem Eng Res Des 2009;87:10031016. [2] Kanniche M, Gros-Bonnivard R, Jaud P, Valle-Marcos J, Amann JM, Bouallou C. Pre-combustion, post-combustion and oxy-combustion in thermal power plant for CO2 capture. Appl Therm Eng 2010;30:53-62. [3] Toftegaard MB, Brix J, Jensen PA, Glarborg P, Jensen AD. Oxy-fuel combustion of solid fuels. Prog Energ Combust 2010;36:581-625. [4] Myöhänen K, Hyppänen T, Pikkarainen T, Eriksson T, Hotta A. Near Zero CO2 Emissions in Coal Firing with Oxy-Fuel Circulating Fluidized Bed Boiler. Chem Eng Technol 2009;32:355-363. [5] Czakiert T, Sztekler K, Karski S, Markiewicz D, Nowak W. Oxy-fuel circulating fluidized bed combustion in a small pilot-scale test rig. Fuel Process Technol 2010;91:1617-1623. [6] García-Labiano F, Rufas A, de Diego LF, de las Obras-Loscertales M, Gayán P, Abad A, Adánez J. Calcium-based sorbents behaviour during sulphation at oxy-fuel fluidised bed combustion conditions. Fuel (in press,doi:10.1016/j.fuel.2011.05.001). [7] de Diego LF, de las Obras-Loscertales M, García-Labiano F, Rufas A, Abad A, Gayán P, Adánez J. Characterization of a limestone in a batch fluidized bed reactor for sulfur retention under oxy-fuel operating conditions. Int J Greenh Gas Con (in press).
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10
Oviedo ICCS&T 2011. Extended Abstract
Carbon based catalytic briquettes for NOx removal in flue gases M.J. Lázaro, M.E. Gálvez, S. Ascaso, I. Suelves, R. Moliner Instituto de Carboquímica, CSIC, Miguel Luesma Castán, 4, 50018 Zaragoza (Spain) Corresponding author: [email protected]
Abstract Catalytic briquettes were prepared using a low-rank coal as carbon precursor and vanadium compounds, i.e. V2O5 and petroleum coke ash (PCA), as the source of the active phase. The catalytic briquettes presented herein offer a feasible low-cost possibility for flue gas cleaning in medium and small industrial facilities. For their preparation, coal was first pyrolysed, then blended with tar pitch and cold pressed to produce the cylindrical briquettes. Pressed briquettes were submitted then to pyrolysis and activation in the presence of steam or CO2, at different temperatures and residence times. Activated briquettes were functionalized by means of HNO3 and H2SO4-wet oxidation. Physical and chemical properties of the produced briquettes were evaluated using N2 adsorption, temperature programmed desorption (TPD) and NH3 chemisorption. Important mechanical properties such as Impact Resistance Index (IRI) and Water Resistance Index (WRI) were also determined for this series of briquettes. Briquettes showed considerable activity in the SCR of NO in a wide temperature range (75-350ºC) and high selectivity towards N2. The several steps involved in the preparation of the catalytic briquettes influence both their structural-chemical and mechanical properties, as well as their catalytic activity in the SCR of NO. 1. Introduction Increasingly stricter environmental regulations concerning the emission of nitrogen oxides (NOx), have forced the development of each time more efficient technologies to reduce the discharge of this pollutants from small and medium industrial facilities. In response, researchers and manufacturers are considering Selective Catalytic Reduction (SCR) as one of the most promising NOx reduction options. While SCR performance has been well established through power plants and evaluation under laboratory conditions, there exists little data characterizing SCR performance under small and medium industrial facilities operating conditions over time. The promulgation of more stringent European regulation of NOx coupled with considerable interest by government, industry,
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Oviedo ICCS&T 2011. Extended Abstract
and other stakeholders to reduce NOx from the existing small and medium facilities, has prompted renewed interest in SCR [1,2]. The use of carbon materials presents unique advantages over other catalyst supports, such as the simplicity of their preparation process, their potentially low-cost, and the easy availability of carbon precursors [3]. Moreover, the carbon support can be shaped as low-pressure drop aggregates, like carbon briquettes and pellets, with good mechanical resistance to withstand abrasion phenomena. Carbon-based catalysts doped with various transition metal oxides. i.e. Cu [4,5], Fe [5], Mn [6-8] and V [9], have shown large potential for their application as SCR catalysts at low temperatures (100250ºC). In previous works we reported the preparation of powder carbon-based catalysts, using a low-rank coal that was subsequently pyrolysed, activated and loaded with different vanadium compounds [10,11]. These catalysts showed high activity and selectivity in the SCR of NO in the presence of NH3, at temperatures between 150 and 200°C [12,13]. Both the characteristics of the carbon support and the preparation method were found to be decisive for optimizing the catalytic activity of the materials prepared [12,14,15]. The present work considers the preparation of catalytic briquettes for their application in small or medium scale deNOx facilities. The several steps involved in their preparation and affecting the final properties of the briquettes, such as pyrolysis and activation, functionalization and calcination after active phase loading, are being considered. Carbon-based catalytic briquettes were prepared using either vanadium pentoxide or the ashes of a petroleum coke (PCA) as active phase precursors. The use of this petroleum residue for producing catalytic briquettes is supposed to be a totally innovative way of upgrading the value of this waste material. In addition, both the use of a low-rank coal and PCA could make the production cost of these catalytic briquettes considerable lower. This fact may involve a considerable reduction of the economic effort that medium and small industrial facilities will be forced to face in order to meet to the upcoming legislation in the near future. In addition, some other pollutants present in the flue gas stream, such as volatile metals or organic compounds, could be at the same time retained within the activated carbon matrix of these catalysts, due to its adsorptive properties. 2. Experimental section SAMCA coal from SAMCA mines in Teruel, Spain [16], was used as carbon support precursor. This coal was pyrolysed at 800°C in nitrogen in a fixed bed. The obtained Submit before January 15th to [email protected]
2
Oviedo ICCS&T 2011. Extended Abstract
char was blended with ground commercial tar pitch, SP-110, used as binder agent. Coal and binder were blended and cold pressed at 125 MPa in a plug and mould press to produce cylindrical briquettes of approximately 10.5 mm diameter, 13.5 mm height and 1.2 g weight. Produced briquettes were subsequently cured in air for 2 hours in an electric oven, rising the temperature up to 200ºC at a heating rate of 2ºC/min. Pyrolysis was performed at 800°C and pressure of 0.1MPa under N2, in a bench-scale reactor described elsewhere [16]. Finally, the briquettes were activated either using 20% of water vapor in N2 at temperatures between 600 and 750°C, or in pure CO2 between 700 and 800°C. Functionalization of the carbon briquettes was performed by means of wet oxidation using sulphuric and nitric acid. Briquettes were immersed in a 2 N HNO3 or H2SO4 aqueous solutions and stirred for 3 hours, then rinsed with distilled water and finally dried at 108ºC for 24 hours. Vanadium pentoxide (V2O5, Aldrich, 99.99% pure) and a refinery waste, i.e. the ashes of a petroleum coke (PCA), were used as the active phase precursors. PCA were obtained from a petroleum coke from the Delayed Coke Unit in the REPSOL in Puertollano (Spain), by combustion under air at 650°C. The PCA contain 23% (w/w) of V, 3.5% (w/w) of Fe and 3% (w/w) of Ni, determined by atomic absorption spectroscopy. A more complex description of PCA can be found in [10]. The impregnation was carried out by equilibrium adsorption of vanadium. Briquettes were immersed into suspensions containing 3% (w/w) V, either of V2O5 or the PCAs, stirred for 3h and carefully washed up afterwards in distilled water. Finally, the catalysts were dried overnight in an oven at 108°C and subjected to a final thermal treatment in Ar for 4-6 hours and at temperatures ranging from 200 to 600 ºC. Carbon briquettes were physically and chemically characterized by means of N2 adsorption at -196ºC (Micromeritics ASAP-2000), temperature programmed desorption (TPD), scanning electron microscopy (SEM-EDX, Hitachi S-3400 coupled to Röntec XFlash analyzer) and NH3 chemisorption (Micromeritics Pulse Chemisorb 2700). Mechanical strength was tested by means of Impact Resistance Index (IRI) and Water Resistance Index (WRI). Catalytic briquettes were tested for NO reduction in a laboratory scale installation consisting of a tubular quartz reactor of 7 mm internal diameter, heated up with the help of an electric oven and connected to a mass spectrometer (Quadrapole Balzers 422) as online gas analysis system. Experiments were performed feeding 22,2 ml/min of a reactant mixture containing 1000 ppm of NO, 1500 ppm of NH3 and 3,5% O2 in Ar, at temperatures between 75 and 350°C and 1,4 s of residence time.
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3
Oviedo ICCS&T 2011. Extended Abstract
3. Results and discussion Activation provided the adequate textural properties and modified the support’s surface chemistry leading to a proper dispersion of the active phase. Activation in the presence of steam led to the production of carbon briquettes with a more developed porous structure in comparison to CO2-activated ones. Maximal development of surface area, micropore and mesopore volumes was gained at 750ºC and 2 hours residence time in both cases. Surface chemistry was also modified upon activation. In general, activation in the presence of CO2 led to the formation of a higher amount of surface groups, particularly of thermally more stable functionalities, i.e. carbonyls and quinones, as determined by TPD analyses. Mechanical properties of the briquettes such as, Impact Resistance Index (IRI) and Water Resistance Index (WRI), were notably influenced as a consequence of activation. IRI decreased with increasing severity of the activation process, Figure 1 shows the surface area and IRI determined for both steam and CO2activated briquettes, as a function of burn-off. IRI decreases almost linearly with burnoff whereas for surface area, an optimal burn-off around 10% can be observed, resulting in an adequate development of porosity, SBET ≈ 250 m2/g, and IRI around 400. On the other hand, the presence of new surface functionalities introduced upon activation slightly improved briquette’s resistance. WRI decreased also with increasing activation temperature and residence time, due to the loss of hydrophobic structures which are thought to avoid water adsorption on surface and briquette disintegration. 400
2
SBET (m /g)
800
IRI
600
200
IRI (-)
400
2
SBET (m /g)
300
200
100
0
4
8
12
16
0 20
Burn-off (%)
Figure 1. Surface area (SBET) and Impact Resistance Index (IRI), as a function of burn-off Wet oxidation of the carbon briquettes, either with HNO3 or H2SO4, led to a partial
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Oviedo ICCS&T 2011. Extended Abstract
collapse of the porous structure, with H2SO4-treatment causing even a more dramatic decrease of the textural parameters vis-à-vis HNO3-oxidation. Surface chemistry was substantially modified after acid treatment. New surface functionalities were introduced, especially carboxylic and lactones in the case of HNO3-treated briquettes, and carboxylic and phenolic, for the H2SO4-oxidized ones. Impact Resistance Index (IRI) was found to decrease considerably for the oxidized briquettes in comparison to the nonoxidized ones, most probably due to the enlargement of pore sizes as a consequence of the functionalization treatment. Lower values of Water Resistance Index (WRI) were found also for the acid-treated briquettes, suggesting the presence of carboxylates that might have enhanced briquette wetting, or just due to the generalized loss of textural properties, which resulted in an easier swelling of the compacted carbon particles. Briquettes either loaded with V2O5 or PCA showed similar activity, demonstrating the feasibility of using this petroleum residue as a source of the active phase. In fact, most of the catalytic briquettes prepared showed significant activity in the SCR of NO (40-90% conversion) at temperatures from 75 to 350ºC. NO reduction was found to be completely selective towards N2. As shown in Figure 2 a, activity increases with surface area. This is due to the fact that a considerable initial surface area of the support seems to be necessary, in order to avoid pore blockage and allow an adequate distribution of the active phase. Surface chemistry does play also a definitive role in activity, contributing to metal anchoring on surface, concretely on phenolic (-OH) groups, as can be observed in Figure 2 b. Thus, activity increased when using the HNO3 and H2SO4-treated briquettes as support. Moreover, ammonia adsorption was favored by the presence of the newly created oxygen functionalities. As shown in Figure 3, thermal treatments of the catalytic briquettes, in Ar at temperatures from 200 to 600ºC, substantially enhanced their catalytic activity. Vcompounds are oxidized by the radicals generated as a consequence of oxygen surface group evolution upon heating. Increasing calcination temperature results in the formation of acidic sites of different strength, which might explain the differences in activity observed for the catalytic briquettes treated at different temperatures and residence times.
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5
Oviedo ICCS&T 2011. Extended Abstract
1.4
1.4 a)
PCA-loaded briquettes
1.2
V2O5-loaded briquettes PCA-loaded briquettes
1.2 1.0
3
k (cm /s·g)
1.0
3
k (cm /s·g)
b)
V2O5-loaded briquettes
0.8 0.6
0.8 0.6
0.4 100
150
200
250
300
0.4 0.04
350
2
0.08
0.12
0.16
0.20 3
SBET (m /g)
0.24
2
Basic surface functionalities (cm /m )
Figure 2. Apparent kinetic constant as a function of a) surface area (SBET) and b) surface area normalized basic surface group concentration, for both steam and CO2-activated briquettes loaded either with V2O5 or PCA 100
100
b)
a) 80
NO conversion (%)
NO conversion (%)
80 60 40 20 0
60 40 20 0
not calcined 200ºC,4 h 200ºC,6 h 400ºC,4 h 500ºC,4 h 600ºC,4 h
not calcined 200ºC,4 h 200ºC,6 h 400ºC,4 h 500ºC,4 h 600ºC,4 h
Figure 3. NO conversion at 150ºC after 1 hour time-on-stream, for a) PCA-loaded steam activated (700ºC, 2 h) briquettes and b) PCA-loaded steam activated (700ºC, 2 h), HNO3-treated briquettes, after calcination at different temperatures and residence times 4. Conclusions The catalytic briquettes presented in this work represent a low-cost option for flue gas treatment in medium and small industrial facilities. They have shown significant activity in the SCR of NO (40-90% conversion) in a wide range of temperatures (75-350ºC). NO reduction was found to be completely selective towards N2. Both catalytic activity and mechanical properties (IRI and WRI) of the briquettes were strongly influenced by the several steps in their preparation procedure. Both an adequate surface area development and the presence of oxygen surface groups were found to be fundamental for achieving
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6
Oviedo ICCS&T 2011. Extended Abstract
an optimal active phase distribution, and minimizing pore blockage towards an adequate diffusion of both reactants and products. Functionalization by means of wet oxidation treatments led to the introduction of phenolic structures which are thought to contribute to active phase fixation, as well as more acidic surface functionalities which enhance NH3 chemisorption ability of the carbon surface, both facts resulting in increased catalytic activity. Thermal treatments upon the introduction of the active phase also resulted in a higher deNOx activity. Acknowledgement. Authors acknowledge AECI (A/18200/08, A/9858/07), Gobierno de Aragón-La Caixa (GA-LC-030/2008), M.E. Gálvez is indebted to the Spanish Ministry of Science and Innovation for her Juan de la Cierva post-doctoral fellowship. S. Ascaso acknowledges CSIC for her JAE pre-doctoral grant. References [1] Hjalmarsson AK. Report 24: NOx control technologies for coal combustion. London IEA Coal Research Publication, 1990, p. 40. [2] Wood SC. Chem Eng Prog 1994;90:32-38. [3] Rodriguez-Reinoso F. Carbon 1998;36:159-175. [4] Teng S, Tu YT, Lai YC, Li CC. Carbon 2001;39:575-582. [5] Zhu Z, Liu Z, Liu S, Niu H, Hu T, Liu T, Xie Y, Appl Catal B: Environ 2000;26:25-35. [6] Grzybek T, Pasel J, Papp H. Phys Chem Chem Phys 1999;1:341-348. [7] Marbán G, Fuertes AB. Appl Catal B: Environ 2001;34:43-53. [8] Marbán G, Fuertes AB. Appl Catal B: Environ 2001;34:55-71. [9] Zhu Z, Liu Z, Liu S, Niu H. Fuel 2000;79:651-658. [10] Vassilev SV, Braekman-Danheux C, Moliner R, Suelves I, Lázaro MJ, Thieman T. Fuel 2002;81:1281-1296. [11] Lázaro MJ, Suelves I, Moliner R, Vassilev SV, Braekman-Danheux C. Fuel 2003;82:771782. [12] Lázaro MJ, Gálvez ME, Suelves I, Moliner R, Vassilev SV, Braekman-Danheux C. Fuel 2004;83:875-884. [13] Gálvez ME, Lázaro MJ, Moliner R. Cat Today 2005;102-103:142-147. [14] Lázaro MJ, Gálvez ME, Artal S, Palacios JM, Moliner R. J Anal Appl Pyrolysis 2007;78:301-315. [15] Lázaro MJ, Gálvez ME, Ruiz C, Juan R, Moliner R. Appl Catal B: Environ 2006;68:130138. [16] Moliner R, Ibarra J, Lázaro MJ. Fuel 1994;73:1214-1220.
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7
Oviedo ICCS&T 2011. Extended Abstract
Identification of Operational Regions in the Chemical-Looping with Oxygen Uncoupling (CLOU) Process with a Cu-based Oxygen-Carrier I. Adánez-Rubio, A. Abad, P. Gayán, L. F. de Diego, F. García-Labiano, J. Adánez Instituto de Carboquímica (ICB-CSIC), Dept. of Energy & Environment, Miguel Luesma Castán, 4, Zaragoza, 50018, Spain Tel.: +34 976 733977; fax: +34 976 733318 Email address: [email protected] Abstract Chemical-Looping with Oxygen Uncoupling (CLOU) is an alternative chemical-looping process for the combustion of solid fuels with inherent CO2 capture. The CLOU process demands a material as oxygen-carrier with the ability to decompose through O2 release at suitable temperatures for solid fuel combustion, e.g. copper oxide. In this work, the combustion of coal by using a promising Cu-based oxygen-carrier prepared by the spray drying method was tested. The oxygen-carrier (Cu60MgAl) was composed of 60 wt% CuO and MgAl2O4 was used as supporting material. The capability for oxygen generation in a CLOU process was evaluated in a batch fluidized-bed reactor at temperatures ranging from 900 to 980 ºC, as well as the oxygen generation rate was determined. Three different regions were identified depending on the oxygen-carrier to coal mass ratio. For oxygen-carrier to coal ratios higher than 50 (Region I), coal was fully converted to CO2 and H2O. In addition, an excess of oxygen was present in the flue gases, which was close to the equilibrium concentration. When this ratio was in the range 50-25 (Region II), the concentration of oxygen was decreasing whereas some CO was observed as the only unconverted gas. Further decrease in the oxygen-carrier to coal ratio below 25 (Region III) caused the depletion of oxygen in the exhaust gases but CO remained as the only unconverted gas. The rate of oxygen generation calculated in every case was related to the solid inventory in the fuel-reactor to get complete combustion of the fuel. The estimated solids inventory in the fuel-reactor was 39, 32 and 29 kg/MWth at Tmax of 930, 955 and 980 ºC, respectively. The results obtained in this work showed that the use of the Cu60MgAl oxygen-carrier is suitable for the coal combustion by CLOU process. 1. Introduction Chemical-Looping with Oxygen Uncoupling (CLOU) is one of the most promising technologies to carry out CO2 capture at low cost when solid fossil fuels are used in
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1
Oviedo ICCS&T 2011. Extended Abstract
energy generation. In the CLOU process the oxygen required for combustion is supplied by a solid oxygen-carrier (OC) which is able to evolve gaseous oxygen. The OC is continuously circulated between two interconnected fluidized-bed reactors, namely the fuel-reactor and the air-reactor. Figure 1 shows a CLOU system schematic design. In the CLOU process several reactions take place in the fuel-reactor, as it is showed for coal: 2 MexOy ↔ 2 MexOy-1 + O2
(1)
Coal → Volatiles + Char + H2O
(2)
Char + O2 → CO2
(3)
Volatiles + O2 → CO2 + H2O
(4)
First the oxygen-carrier releases oxygen according to reaction (1) and the solid fuel begins devolatilization producing a porous solid (char) and a gas product (volatiles), reaction (2). Then, the char and volatiles are burnt as in usual combustion according to reactions (3) and (4). After that the oxygen-carrier is re-oxidized in the air-reactor: O2 + 2 MexOy-1 ↔ 2 MexOy
(5)
The overall heat release over the fuel- and air-reactors is the same as for conventional combustion. The advantage is that the CO2 and H2O are inherently separated from the nitrogen in the combustion air, and thus no energy is needed to obtain pure CO2. N2 (+O2)
CO2 + H2O
CO2
Condenser
MexOy
Fuel Reactor
Air Reactor
CO2
H2O(l)
Coal
MexOy-1
CO2 Air
Ash
Figure 1: CLOU system schematic design. Three metal oxide systems have been identified with CLOU properties: CuO/Cu2O, Mn2O3/Mn3O4, and Co3O4/CoO [1]. Focused in Cu-based materials, cyclic testing with solid fuels verified a high reactivity during oxidation and reduction, where very rapid release of oxygen was observed [2]. In addition, Cu-based materials have other advantages as its high oxygen transport capacity (RO = 10 wt%) and the global process in the fuel-reactor is exothermic.
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2
Oviedo ICCS&T 2011. Extended Abstract
4 CuO + C → 2 Cu2O + CO2
D H r950ºC = - 134 kJ / mol O2
(6)
Previously to this work, a screening of different Cu-based materials was carried out [2, 3]. From this work, it was shown that particles prepared by mechanical mixing followed by pelletizing by pressure containing 60 wt% CuO and using MgAl2O4 as supporting material were adequate for its use as oxygen-carrier for the CLOU process. The aim of this work was to investigate the performance of a CuO/MgAl2O4 material as oxygen-carrier for the CLOU process. The rate of oxygen release was determined when batches of “El Cerrejon” coal particles were added to a fluidized-bed reactor containing the oxygen-carrier material. Different oxygen-carrier to fuel ratios, as well as several temperatures between 900 and 950 ºC were tested. The needed solids inventory in the fuel-reactor for complete fuel combustion was inferred from the results obtained. 2. Experimental section 2.1 Materials
The material used was a Cu-based oxygen-carrier prepared by spray drying manufactured by VITO. The CuO content was 60 wt% and MgAl2O4 was the supporting material. The oxygen transport capacity of the oxygen-carrier was ROC=6 wt% and the particle size was 100-200 μm. From now on the oxygen-carrier was named as Cu60MgAl. The fuel used was a bituminous Colombian coal “El Cerrejón”. Ultimate and proximate analysis are shown in Table 1. The coal particle size used for this study was +200-300 μm. The Low Heating Value was 21899 kJ/kg. Table 1. Properties of pre-treated “El Cerrejón” coal.
C
65.8 %
Moisture
2.3 %
H
3.3 %
Volatile matter
33.0 %
N
1.6 %
Fixed carbon
55.9 %
S
0.6 %
Ash
8.8 %
2.2 Experimental setup: batch fluidized-bed reactor
Reduction-oxidation multi-cycles were carried out in a fluidized-bed reactor to know the oxygen release behaviour of the oxygen-carrier in similar operating conditions to that existing in the CLOU process. A schematic layout of the laboratory set-up is presented in Figure 2. Solids are placed in the reactor (55 mm inner diameter and 700 mm height )
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3
Oviedo ICCS&T 2011. Extended Abstract
above the distributor plate, assuring a bed height about 50 mm at static conditions. The total fluidizing flow was 200 LN/h, which corresponds to a gas velocity of 0.1 m/s at 900 ºC. The oxygen-carrier particles were exposed sequentially to reducing and oxidizing conditions. The reducing period consisted in loads of a certain amount of coal particles in N2, whereas oxidation was done by air.
N2
v1
v2
Solids feeding system
P
Filter v3 Gas analysis
Furnace H2O (l) Thermocouple
ΔP
Distributor plate N2
P Air
Figure 2. Schematic layout of the laboratory setup.
Three series of experiments were performed to increase the oxygen-carrier to coal ratio covered. The first tests were done with an amount of 240 g of oxygen-carrier material in the bed. Loads of coal between 0.2 and 2 g were added, corresponding to OC/coal = 1200-120. The second and third series were carried out diluting the oxygen-carrier in alumina particles. Thus, the mass fraction of the OC in the bed was 10 wt% and 2.5 wt%, respectively. Loads of coal between 0.4 and 1.2 g (OC/coal = 60-20) in the first case, and 0.03 and 0.5 g (OC/coal = 200-12) in the second case were used. The experimental work has been carried out at temperatures between 900 and 950 ºC. The temperature of the reactor was fixed before starting the reducing or oxidizing period, but during reaction the temperature could increase up to 50 ºC because the exothermic reaction when coal is burnt. Different gas analyzers measured continuously the gas composition (CO, CO2, CH4, H2, O2) at the reactor exit after water condensation. 2.3 Data Evaluation
During the reduction period, the rate of oxygen generation per amount of oxygen-carrier,
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Oviedo ICCS&T 2011. Extended Abstract
rO2 , was calculated from a mass balance to the oxygen in the reactor. rO2 (t ) =
(
M O2
)
⎡ FO + FCO + 0.5 FCO + FH O − 0.5 FO ,coal ⎤ 2 2 ⎦ mOC ⎣ 2
(7)
being Fi the gas flow of each component exiting the fuel-reactor. Possible gas i includes O2, CO2, CO and H2O. The water concentration was not measured, but the H2O flow was calculated as:
(
FH 2O = 0.5 f H / C FCO2 + FCO
)
(8)
fH/C being the hydrogen to carbon molar ratio in the coal (fH/C = 0.61). Equally, the flow of oxygen coming from coal, FO,coal, was calculated as:
(
FO ,coal = fO / C FCO2 + FCO
)
(9)
fO/C being the oxygen to carbon molar ratio in the coal (fO/C = 0.20). The oxidation conversion was calculated from the integration of rO2 (t ) with time: X o (t ) = 1 −
1 N O2
t
∫r 0
O2
(t )dt
(10)
being N O2 the molar amount of oxygen in the oxygen-carrier active for CLOU process, i.e. from reduction of CuO to Cu2O expressed as mol of molecular O2. 3. Results and Discussion
As example, Figure 3 shows the concentration of O2, CO2 and CO measured at the outlet of the reactor and the bed temperature during a typical reduction and oxidation cycle at initial temperature T0 = 925 ºC with an OC inventory in the reactor of 240 g. The time t = 0 corresponds to the initial time of the reduction period. At the beginning a rapid oxygen release occurred close to the oxygen concentration equilibrium for the measured bed temperature. After a short period, a batch of 2 g of coal particles were fed to the reactor, and only CO2 and O2 were observed in the outgoing gases, indicating full combustion of the volatiles and char. The CO2 concentration in this case was as high as 76 vol% which was maintained constant during ~8 s, and eventually decreased to zero when the complete coal combustion was reached. During coal combustion the bed temperature increased being the maximum temperature (Tmax) reached about 30 ºC higher than T0 due to the exothermic reaction of CuO with coal, see equation (6). The oxygen concentration was correspondingly increased, remaining close to the equilibrium condition when temperature varied. This result suggests a fast combustion of coal. In
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Oviedo ICCS&T 2011. Extended Abstract
addition, the oxygen release rate was high enough to supply an excess of gaseous oxygen in the combustion gases. In Figure 3, the variation of solids conversion, X0, with reacting time is also shown. It can be seen that the oxygen-carrier was slowly converted during the initial period before coal addition. A sharp decrease in the OC conversion occurred when coal was fed to the reactor because the fast oxygen release. The variation in the solids conversion during the reducing period was 41%. Then, oxidation period starts at t = 360 s and a quick increase in the temperature occurred due to the highly exothermic reaction. Equally to reduction period, oxygen concentration was close to the equilibrium condition until the OC was fully oxidized. reduction
1.0
100
0.8
80
oxidation
1000
0.4
0.2
0.0
980
T
60
960
40
20
940
Coal addition
CO2
O2
0 0
300
600
Temperature (ºC)
0.6
Concentration (vol%)
Solids conversion (-)
Xo
920
900 900
time (s)
Figure 3. Variation of solids conversion, bed temperature and concentration of O2 and
CO2 during redox cycle with Cu60MgAl. T0 = 925 ºC; mOC: 240 g; mcoal: 2 g. Similar behaviour was observed in redox cycles when different amount of coal was added to the bed in the range 0.2-2.0 g. From the CO2 and O2 evolution in the gas phase, it was possible to calculate the oxygen generation rate, rO2 , from equation (7). In every experiment the highest value of rO2 was reached when CO2 is maximum. Figure 4 shows the highest value of rO2 as a function of the OC/coal ratio. It can be observed a similar trend of the data obtained with different oxygen-carrier dilutions. When the OC/coal ratio decreased until a value of 25 the oxygen generation rate increased. However, at lower OC/coal values it seemed that the oxygen generation rate reached a maximum. Thus, for OC/coal < 25 the coal conversion was limited by the oxygen generation rate from the oxygen-carrier. The calculated average value for the oxygen generation rate
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Oviedo ICCS&T 2011. Extended Abstract
was 2.6·10-3 kg O2/s per kg of OC when the temperature increased up to Tmax = 955 ºC. Analogous experiments were carried out at initial temperatures, T0, of 900 ºC and 950 ºC. In these cases, temperature increased until Tmax of 930 ºC and 980 ºC, respectively, when coal was added to the bed the reactor. Similar trend was found for the oxygen generation rate with the OC/coal ratio. Nevertheless, the maximum oxygen generation rate increased with the reacting temperature. The oxygen generation rate was 2.1·10-3 and 2.8·10-3 kg O2/s per kg of oxygen-carrier at Tmax of 930 and 980 ºC, respectively.
rO x103 (kg O2/s per kg OC) 2
10
1
0.1
0.01 1
10
100
1000
10000
OC/coal (kg/kg)
Figure 4. Maximum oxygen generation rate, rO2 , for the OC as a function of the
OC/coal ratio. T0 = 925 ºC. OC in the reactor: (□) 100 wt%; (▲) 10 wt%; (○) 2.5 wt%. The maximum rate of oxygen generation calculated by the procedure above described can be related to the solids inventory needed to fully convert the fuel. Assuming that there was an excess of oxygen in the circulating solids in a CLOU system the coal combustion was not limited by the availability of oxygen transported from the airreactor. Thus, the solids inventory, mFR, depends on the flow of coal which is able to process the oxygen-carrier at its maximum rate of oxygen generation, rO2 ,max , and taking as reference 1 MWth for the fuel feeding rate, it can be calculated as:
mFR = 103
mO rO2 ,max LHV
(11)
being mO the mass of oxygen required per kg of coal to fully convert it to CO2 and H2O, as for the case of the conventional combustion with air. LHV is the lower heating value of the solid fuel. From coal analysis showed in Table 1, a value for mO = 2.1 kg O2 per
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7
Oviedo ICCS&T 2011. Extended Abstract
kg of coal was calculated. The solids inventory in the fuel-reactor was dependent on the reactor temperature because the increase of the reactivity with the temperature. Thus, 39 kg/MWth would be necessary in the fuel-reactor at a Tmax = 930 ºC, being decreased to 29 kg/MWth if Tmax was 980 ºC. Here, the temperature is the maximum temperature reached during the reduction period. Figure 5(a) shows the CO2 yield, yCO , defined as carbon fraction in the form of CO2 in 2
the outgoing gases, whereas the ratio between the oxygen concentration at the reactor exit and that present at equilibrium conditions, O2/O2,eq, was plotted in Figure 5(b). It can be seen as both the CO2 yield and the oxygen concentration decreases when the OC/coal mass ratio was lower than 50. The observed decrease in the CO2 yield was due to the presence of CO in the gases, but CH4 or H2 were not observed in any case, indicating that volatiles had enough contact time to be burnt in the bed. Moreover, the oxygen concentration becomes zero when the maximum rate of oxygen generation was reached. This fact determines a change in the limiting process during conversion of coal towards the rate of oxygen production by the oxygen-carrier. Thus, different region can be differentiated depending of the OC/coal ratio.
1.0
100
(b)
(a) 0.8
O2/O2,eq (-)
CO2 yield (%)
95
90
85
0.6
0.4
0.2
0.0
80 1
10
100
1000
10000
OC/coal (kg/kg)
1
10
100
1000
10000
OC/coal (kg/kg)
Figure 5. (a) CO2 yield and (b) ratio between the oxygen concentration at the reactor
exit and at equilibrium conditions, O2/O2,eq, as a function of the OC/coal mass ratio. T0 = 925 ºC. OC in the reactor: (□) 100 wt%; (▲) 10 wt%; (○) 2.5 wt%. Figure 6 shows the limiting conditions for these regions and the calculated solids inventory with different OC/coal ratios used in this work. The Region I was defined as
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Oviedo ICCS&T 2011. Extended Abstract
the region where CO was not present in the exhaust gases, corresponding to OC/coal ratios higher than 50. The minimum solid inventory in this region was 58 kg/MWth. In the Region II, O2 and CO simultaneously appeared in the outgoing gases. At this condition, the minimum solid inventory was about 43 kg/MWth at Tmax = 955 ºC. Finally, in the Region III no oxygen was present in the exhaust gases. The maximum rate for oxygen generation of the OC is reached, and lowering the solids inventory does not permitted to fully convert the coal in the fuel-reactor. This region was observed for OC/coal ratios lower than 25.
1000
mFR (kg/MW th)
Regime III Regime II
Regime I
100
10 10
100
1000
OC/coal (kg/kg)
Figure 6. Calculated solids inventory in the fuel-reactor for the OC as a function of the
OC/Coal ratio. T0 = 925 ºC. OC in the reactor: (□) 100 wt%; (▲) 10 wt%; (○) 2.5 wt%. From the results showed in this work, to avoid the presence of CO and thus the necessity of an oxygen polishing step, a small oxygen amount is desirable in the flue gases coming from the fuel-reactor. Thus, it can be suggested that a CLOU system should be operated in the Region I, where complete combustion of coal can be obtained. However, the excess of oxygen could be problematic for the transport and storage of the CO2, and it must be removed from the exhaust gas stream. To address the excess of oxygen one option would be to separate the oxygen in the purification and compression process of the CO2 before being transported. This situation is similar to that present in coal oxycombustion units. 4. Conclusions
A Cu-based oxygen-carrier prepared by spray drying was tested for the CLOU process
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Oviedo ICCS&T 2011. Extended Abstract
in a batch fluidized-bed reactor. The oxygen-carrier contained 60 wt% CuO and MgAl2O4 was used as inert material. A bituminous Colombian coal “El Cerrejón” was used as fuel. The capability of particles to evolve gaseous oxygen in the fuel-reactor was evaluated as a function of the oxygen-carrier to coal mass ratio and reactor temperature. A decrease in the oxygen-carrier to coal mass ratio caused a continuous increase in the oxygen generation rate of the Cu60MgAl material until the maximum rate of oxygen generation was reached. Three operating regions were identified depending on the solids inventory. Unburnt compounds were not present in the outgoing gases in Region I when the calculated solids inventory was higher than 58 kg/MWth at Tmax = 955 ºC. CO2 and H2O were the only products of coal combustion and an excess of O2 was observed, which was found to be close to the equilibrium concentration. At these conditions, the oxygen concentration from the reactor increased with the temperature. The Region II was confined between 32 and 58 kg/MWth, at Tmax = 955 ºC, where both CO and O2 were present in the exhaust gases together CO2 as majority compound. The maximum rate of oxygen generation was found in the so-called Region III. Oxygen was not present in the flue gases, and a certain concentration of CO was present as unburnt compound. To operate in Region III, the estimated solids inventory in the fuel-reactor was 39, 32 and 29 kg/MWth at Tmax of 930, 955 and 980 ºC, respectively. The results obtained in this work showed that the use of the Cu60MgAl oxygen-carrier is suitable for coal combustion by CLOU process. Acknowledgement
This work was partially supported by the European Commission, under the RFCS program (ECLAIR Project, Contract RFCP-CT-2008-0008), ALSTOM Power Boilers (France) and by the Spanish Ministry of Science and Innovation (ENE2010-19550). I. Adánez-Rubio thanks CSIC for the JAE fellowship. References
[1] Mattisson, T.; Lyngfelt, A.; Leion, H. Chemical-looping with oxygen uncoupling for combustion of solid fuels. Int. J. Greenhouse Gas Control, 2009, 3, 11-19. [2] Adánez-Rubio, I.; Gayán, P.; García-Labiano, F.; de Diego, L.F.; Adánez, J.; Abad, A. Development of CuO-based oxygen-carrier materials suitable for Chemical-Looping with Oxygen Uncoupling (CLOU) process. Proc 10th Int Conf Greenhouse Gas Technology (GHGT-10). Amsterdam, The Netherlands; 2010. [3] Adánez-Rubio, I.; Gayán, P. Abad, A; García-Labiano, F.; de Diego, L.F.; Adánez, J.CO2 Capture in Coal Combustion by Chemical-Looping with Oxygen Uncoupling (CLOU) with a Cu-based Oxygen-Carrier. Proc 5th Int Conf on Clean Coal Technology (CCT-11). Zaragoza, Spain; 2011.
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Oviedo ICCS&T 2011. Extended Abstract
Influence of hydrogenation on the mercury capture by activated carbons Jorge Rodríguez-Pérez, M. Antonia López-Antón, Roberto García, Mercedes DíazSomoano and M. Rosa Martínez-Tarazona Instituto Nacional del Carbón (CSIC), Francisco Pintado Fe, 26, 33011, Oviedo, Spain. Abstract In this study, activated carbon (AC) sorbents have been used for elemental mercury retention experiments at a laboratory scale. A previous hydrogenation treatment of the parent AC resulted in improved performances, in terms of mercury retention capacity (RC) and efficiency. The surface functional groups of the untreated and hydrogenated AC have been investigated by temperature programmed desorption (TPD) experiments and by X-ray photoelectron spectroscopy (XPS) analysis. The presence of an increased concentration of protons in the carbon surface seems to have a greater influence on the enhanced Hg retention capacity than the variation in the proportion and types of oxygen functional groups.
1. Introduction Power generation by coal combustion is the largest anthropogenic source of mercury emissions in the world. Then, removing mercury from power plant flue gas streams is a priority environmental issue. Mercury can be emitted in oxidized or elemental forms, and the capture of elemental mercury is especially problematic, due to its volatility, chemical inertness, and insolubility in water. The typical air-pollution control systems are ineffective and it has to be addressed with retention devices ad hoc. A possible approach is the use of carbonaceous sorbents with and without supported catalysts. Activated carbons have been considered as sorbents for Hg0 in many studies [1-7]. However, the mechanism of the adsorption process is not well understood yet. The influence of the textural properties and of the presence of several types of surface functional groups has been studied and both physisorption and chemisorption effects have been proposed. It seems clear that, in addition to a good surface area, the surface chemistry turns out to be very relevant when choosing the appropriate carbon material, to the extent that the presence of a metallic catalyst could not be strictly necessary.
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Oviedo ICCS&T 2011. Extended Abstract
In this contribution, an activated carbon has been tested for the capture of Hg0, both in its fresh state and after hydrogenation. The influence of hydrogenation in the surface distribution of oxygenated functional groups and the effect of such groups in Hg0 capture have been studied.
2. Experimental section A commercial activated carbon (RB3, from Norit), ground to a size of 0.2-0.5 mm, was used as Hg0 sorbent. The surface chemistry of the sorbent samples was analysed by temperature programmed desorption (TPD) and X-ray photoelectron spectroscopy (XPS) to study the oxygenated surface functional groups. BET surface area was determined by means of volumetric adsorption of nitrogen at 77 K. Hydrogenation of RB3·was carried out in a stainless steel reactor, at 350 °C in 3 hourexperiments, using 0.6 g of activated carbon, ground to 0.2-0.5 mm, under 35 bar of H2. The experimental device used for the Hg retention experiments consists of a glass reactor heated by a furnace (Figure 1). Gas phase Hg0 was obtained from a permeation tube and passed through the sorbent bed in an air flow of 0.5 L min-1. The concentration of Hg in the air stream was 100 μg m-3 and the sorbent temperature was kept at 120 ºC. A continuous Hg emission monitor (VM 3000) was used to obtain the Hg adsorption curves. The concentration of Hg retained (retention capacity) was determined by a postretention analysis of the sorbents using an automatic mercury analyzer (AMA). VM 3000
R1 N2
Sorbent 120ºC
Permeation tube activated carbon
R2
R3 Air
Figure 1. Schematic diagram of the experimental device for mercury retention.
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Oviedo ICCS&T 2011. Extended Abstract
3. Results and Discussion The continuous mercury analyzer employed in this study is only able to detect elemental mercury. Then, the fact that the sample curves do not reach the background line (Cout/Cin=1) could indicate that the samples have not reached their maximum retention capacity, but the same effect would be observed if the mercury in the gas phase leaving the sorbent bed was oxidised. To ensure that all the Hg leaving the sorbent was Hg0, a study of Hg speciation in the RB3 samples was carried out. For this purpose, the gas at the sorbent bed outlet was collected in an Ontario Hydro impinger train device, instead of the VM 3000 analyzer. This method is able to distinguish between Hg0 and Hg2+. The results obtained clearly show that no oxidation of Hg0 took place during the retention experiments.
1 0,9 0,8
Hg C out/C in
0,7 0,6 0,5 0,4 0,3
RB3
0,2
RB3 (red. H2)
0,1 0 0
500
1000
1500
2000
2500
t (min)
Figure 2. Hg0 adsorption curves of the activated carbon before (RB3) and after hydrogenation (RB3 (red. H2)). Figure 2 displays the Hg0 adsorption curves obtained with RB3 activated carbon before and after hydrogenation. The curves represent the outlet/inlet Hg concentration ratio (Cout/Cin) versus time. The Hg0 retention capacities are listed in Table 1. The results indicate that hydrogenation of the activated carbon promotes a significant improvement of its Hg0 retention capacity.
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Oviedo ICCS&T 2011. Extended Abstract
Table 1. Retention capacity of the activated carbon-based sorbents. Sample RB3 RB3 (red. H2)
texp (min) 4320 2796
RCAMA (μg g-1) 10 934
The values of the BET surface area of the activated carbon and its hydrogenated counterpart are listed in Table 2. No significant differences can be noticed, indicating that under the conditions used for the Hg0 retention experiments physisorption does not play a significant role in the mercury uptake. At least, it cannot account for the differences observed between the two kinds of sorbents. Then, the retention of elemental mercury must be ruled by its interactions with the surface functional groups present in the activated carbons. Table 2. BET surface area of the sorbents. Sample RB3 RB3 (red. H2)
SBET (m2 g-1) 1183 1137
The TPD curves of the activated carbon undergo significant changes after hydrogenation (Figure 3). RB3 shows the presence of several kinds of oxygenated groups. The CO evolution curve indicates the presence of C=O groups with a wide band in the 700-900 °C zone. The shoulder at ~600 °C can arise from OH groups in phenols and anhydrides. The evolution of CO2 indicates the presence of anhydrides and lactones (two peaks at ~600 °C) and also carboxylic groups (wide band in the 100-400 °C area). In the case of RB3 (red H2), the evolution curves displays less features, and only C=O groups seem to be present.
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Oviedo ICCS&T 2011. Extended Abstract
TPD - CO curves
TPD - CO2 curves
0,05
RB3 (red. H2) RB3
0,08
Concentration (%)
Concentration (%)
0,1
0,06
0,04
RB3 (red. H2) RB3
0,04
0,03
0,02
0,01
0,02
0
0 0
200
400
600
800
1000
0
200
Temperature (º C)
400
600
800
1000
Temperature (º C)
Figure 3. CO and CO2 TPD curves. XPS results are shown in Figure 4. C 1s peaks are very similar for both samples, but the as received RB3 displays a slightly higher intensity at high binding energy (~288 eV), next to the main peak centred at 284 eV, that arises from oxygenated carbon [6]. Accordingly, in the O 1s spectrum, the band of RB3 is wider than that of the hydrogenated sample, indicating the presence of more types of oxygen functional groups.
O1s
C1s RB3 RB3 (red. H2)
298
296
294
292
RB3 RB3 (red. H2)
290
288
286
284
282
280 540
538
Binding Energy (eV)
536
534
532
530
528
Binding Energy (eV)
Figure 4. C 1s and O 1s XPS spectra. It has been previously reported that carbonyl groups [5,7] can be activated sites for Hg0 capture. In our case, however, although such groups are the most abundant oxygenated groups in the reduced activated carbon, their concentrations are much lower than in the as received RB3 sample, and therefore they alone cannot account for the increased Hg0 adsorption. However, the same study [5,7] identifies OH and phenolic groups as inhibitors of Hg0 adsorption, indicating that the best adsorption capacities are observed
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Oviedo ICCS&T 2011. Extended Abstract
in activated carbons with the lowest CO/CO2 ratios released in TPD. In our case, this ratio is very similar for both samples (RB3, 2.51; RB3 (red H2), 2.66), and even higher in the hydrogenated activated carbon. Apparently, the beneficial effect of the loss of phenolic groups and the reduction of their inhibitory effects for mercury capture is more significant that the decrease in the concentration of C=O groups. It has been suggested that the mechanism of Hg0 adsorption involves an electron transfer process with the carbon acting as an electrode, accepting electrons from mercury, which oxidises to Hg2+ [5,8-9]. The equilibrium reaction proposed is as follows, involving quininoid complexes: C6H4O2 + 2 H+ + 2 e- ↔ C6H4(OH)2 In such a scenario, the presence of additional protons, such as those supplied by the previous hydrogenation of the activated carbon, would be beneficial. The potential (Eh) of the electrode can be calculated as [5,8-9]: Eh = E0 – (R T/2 F) ln(a[C6H4(OH)2]/(a[C6H4O2] a[H+]2) where E0 is the characteristic constant potential, R is the molar gas constant, T is the temperature, F is the Faraday constant, and a is the activity (concentration × activity coefficient). Higher H+ concentration as in the hydrogenated activated carbon could increase the potential of the electrode and also the potential difference (ΔE) required for the oxidation-reduction reaction.
4. Conclusions The low Hg0 retention observed in activated carbon RB3 increases significantly after hydrogenation. The modification of the concentration and types of oxygen functional groups in the surface of the carbon does not seem to explain this enhancement. However, the increased presence of protons in the surface can improve the potential difference required for the oxidation-reduction reaction that may be responsible for Hg0 capture.
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Oviedo ICCS&T 2011. Extended Abstract
Acknowledgement. The authors thank the CSIC (PIF-06-050) and the Spanish Ministerio de Ciencia e Innovación (CTQ2008-06860-C02-01) for financial support.
References [1] [2] [3] [4] [5] [6] [7] [8] [9]
Granite EJ, Pennline HW, Hargis RA. Novel sorbents for mercury removal from flue gas. Industrial & Engineering Chemistry Research 2000;39:1020-9. Li YH, Lee CW, Gullett BK. The effect of activated carbon surface moisture on low temperature mercury adsorption. Carbon 2002;40:65-72. Kwon S, Borguet E, Vidic RD. Impact of surface heterogeneity on mercury uptake by carbonaceous sorbents under uhv and atmospheric pressure. Environmental Science & Technology 2002;36:4162-9. Ghorishi SB, Keeney RM, Serre SD, Gullett BK, Jozewicz WS. Development of a cl-impregnated activated carbon for entrained-flow capture of elemental mercury. Environmental Science & Technology 2002;36:4454-9. Li YH, Lee CW, Gullett BK. Importance of activated carbon's oxygen surface functional groups on elemental mercury adsorption. Fuel 2003;82:451-7. Maroto-Valer MM, Zhang Y, Granite EJ, Tang Z, Pennline HW. Effect of porous structure and surface functionality on the mercury capacity of a fly ash carbon and its activated sample. Fuel 2005;84:105-8. Liu J, Cheney MA, Wu F, Li M. Effects of chemical functional groups on elemental mercury adsorption on carbonaceous surfaces. Journal of Hazardous Materials 2011;186:108-13. Leon Y, Leon CA, Radovic LR. Interfacial chemistry and electrochemistry of carbon surfaces. In: Thrower PA (editor). Chemistry and physics of carbons, New York: Marcel Dekker; 1994, p. 213-310. Puri BR. Surfaces complexes on carbons. In: P.L. Walker J (editor). Chemistry and physics of carbon, New York: Marcel Dekker; 1970, p. 191-282.
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Oviedo ICCS&T 2011. Extended Abstract
SUB-PRODUCTS OF GASIFICATION AS SORBENTS FOR MERCURY RETENTION A. Fuente-Cuesta, M. Diaz-Somoano, M.A. Lopez-Anton, M.R. Martinez-Tarazona Instituto Nacional del Carbón (CSIC). C/ Francisco Pintado Fe Nº26, 33011, Oviedo, Spain. Phone: +34 985119090. Fax: +34 985297662 E-mail: [email protected]
Abstract Different technologies can be used to limit the release of mercury during coal combustion, such as particulate control devices and flue gas desulphurization systems. Even so, it is very difficult to remove vapour elemental mercury because a large proportion of it eludes the typical air-pollution control devices. Activated carbon injection is currently considered as the most promising specific technology in terms of efficiency and reliability for mercury removal. However, it is still necessary to overcome the high operational costs of this technology. The aim of this work is to evaluate the mercury retention capacities of low cost biomass gasification chars. The tested solids were sub-products from the gasification of sunflower husks, chicken manure, wood and plastic-paper wastes. The capture of mercury was evaluated in a labscale device under inert and oxygen atmospheres. The mercury retention capacities of these chars were compared with those of a commercial activated carbon (impregnated with sulphur) considered as a reference material. The results showed that the highest mercury capture was attained by the chars obtained from plastic-paper wastes. The chars with a high chlorine content and low pH present high mercury retention capacities. For all the chars studied, retention was higher under oxygen than under nitrogen, which indicates that mercury adsorption capacity depends not only on the properties of the chars, but also on the composition of the flue gas.
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Oviedo ICCS&T 2011. Extended Abstract
1. Introduction Coal combustion in power stations releases different gaseous trace elements, such as mercury, which are harmful to the environment and to human health. Several technologies and policies have been developed over the last two decades for the control of mercury emissions from coal-fired power plants [1], the major source of anthropogenic mercury emissions. The leading technology is the injection of powdered activated carbon into the flue gas prior to the particulate control devices (PCD). Nevertheless cheaper sorbents need to be developed in order to reduce costs and make this technology economically competitive.
Nowadays biomass gasification is a viable alternative for energy production from renewable sources. Chars are the solid wastes left behind after the gasification process. They contain part of the carbonaceous material, nitrogen and sulphur of the original biomass, as well as almost all its mineral matter. These characteristics make gasification chars potential sorbents for mercury capture. Although some studies have used biomass chars after the activation process for mercury removal [2], there is a lack of knowledge about the behaviour of untreated biomass chars as mercury sorbents.
The final objective of this research is to develop low-cost sorbents for direct injection in power plants. For this purpose a study at lab-scale was carried out in nitrogen atmosphere using five different gasification chars to evaluate the influence of the physico-chemical properties of the char on mercury adsorption capacity. After that, an air atmosphere was used to determine whether the gaseous components have any influence on mercury capture. The mercury adsorption capacities of the chars were compared with those of a commercial sulphur-activated carbon (Filtracarb).
2. Experimental section The five biomass gasification chars used in this study were produced at the Energy Research Centre of the Netherlands (ECN) in a pilot plant of 500 kW of power with a circulated fluidized bed gasifier (BIVKIN). The samples were labelled as follows: SH is the gasification char obtained from sunflower husks, PL is the one from chicken manure, WW the one from wood wastes and PW1 and PW2 the ones from plastic-paper wastes. The only difference in the production method of samples PW1 and PW2 was the gasification temperature. The commercial activated carbon used as reference was Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
Filtracarb D47/7+S.
Before the retention experiments, the solids were characterised using different techniques. The sulphur content was determined by a LECO automatic analyzer, the chlorine content by ionic chromatography and the total organic carbon content (TOC) using a TOC V-CPH E200V equipment. The Brunauer-Emmett-Teller (BET) surface area was measured by volumetric adsorption of nitrogen at 77 K. The pH of the chars and activated carbon sample suspensions was determined after 48h of stirring in water, once the samples had reached equilibrium. Finally, the mercury content before and after the retention experiments was determined by means of an automatic mercury analyzer (AMA254).
The lab-scale device employed for the mercury retention experiments is shown in Figure 1. Four main parts can be distinguished in this device: i) a system to generate gaseous elemental mercury from a permeation tube, ii) a fix bed reactor, iii) a system to produce the atmospheres in which the materials were tested (N2 or O2 (12.6 % v/v) and iv) a continuous mercury emission analyzer to monitor Hg concentration at the reactor outlet (VM3000). The experiments were carried out at 150ºC and the flow rate through the reactor was 0.5 L min-1. The sorbent bed consisted of a mixture of the sample and fine sand at a weight ratio of 1:3. The mercury concentration in the gas phase was approximately 100 µg m-3.
Rotameter VM 3000 N2
Mercury analyzer
Permeation tube Clean flue gas Reactor
N2 / Air Dilution Air
Activated carbon
Figure 1. Schematic diagram of the lab-scale experimental device for mercury retention.
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Oviedo ICCS&T 2011. Extended Abstract
3. Results and discussion The total organic carbon (TOC), sulphur and chlorine contents, pH, surface area (SBET) and mercury content of the raw biomass gasification chars and the activated carbon are shown in Table 1, whereas the mercury retention capacities in the two atmospheres used in this study are presented in Table 2.
Table 1. Total organic carbon (TOC), sulphur and chlorine contents, pH, surface area (SBET) and mercury content in the biomass gasification chars and the activated carbon. SAMPLES
TOC (%)
S (%)
Cl (%)
pH
SBET (m2·g-1)
Hg (µg·g-1)
SH
55
0.5
0.8
9.9
5
0.01
PL
11.7
1.7
1.8
11.4
12
0.02
WW
37.8
0.7
2.6
10
2
0.02
PW1
32.7
0.09
4.7
7.9
65
0.01
PW2
24.6
0.08
5.2
8
42
0.01
FILTRACARB
68.3
2.5
0.07
5.8
560
0.34
The mercury content in the raw biomass chars was very low (<20 ng g-1). As can be seen, the gasification chars employed in this study have different characteristics which allows us to evaluate the main parameters that influence mercury retention. The mercury retention capacity of the chars follows the order PW1>PW2>WW>SH>PL in both atmospheres (Table 2). Although the highest mercury retention capacity was achieved by Filtracarb activated carbon, the PW chars, which have a much smaller surface area than that of the activated carbon (Table 1) and sulphur contents, can reach similar retention capacities.
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Oviedo ICCS&T 2011. Extended Abstract
Table 2. Mercury retention capacities of the materials studied in the nitrogen and oxygen atmospheres. SAMPLES
Hg retained N2 (µ µg·g-1)
Hg retained O2 (µ µg·g-1)
SH
1.1
4.1
PL
<1
<1
WW
2.7
6
PW1
135
570
PW2
110
530
FILTRACARB
145
840
Several studies have been carried out on the influence of carbon content on mercury retention [3] and it is usually accepted that the retention of mercury increases with the increase in carbon content. However, in this study we can observe that the SH char with the highest carbon content (TOC: 55%) shows a low mercury capture capacity. Therefore other characteristics need to be considered.
It is also well known that mercury species may react with sulphur and chlorine species [4]. For this reason, special attention was paid to the content of these two elements in the samples studied. In general, no relation was found between mercury adsorption capacity and the sulphur and chloride content of the chars. However, it must be emphasized that the highest mercury retentions were obtained with the chars that had the highest chlorine contents (PW chars).
The pH of the sorbents was measured in order to evaluate its possible influence on mercury capture. Whereas the activated carbon presents a slightly acid character (5.8), the chars show a neutral or basic character (pH ranging from ~8 to 12). The PW chars, which have the lowest pH value, show the highest mercury retention value. The results suggest then that the neutral or acid character of the samples favours mercury capture.
Although the char samples are solids with a low surface area (Table 1) and their values are obviously not as high as those of the activated carbon (65 vs 560 m2·g-1), in general
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Oviedo ICCS&T 2011. Extended Abstract
the increasing surface area of the char samples would enhance mercury capture.
When the results in both the inert and oxygen atmospheres are compared it can be seen that higher retention capacities are found in presence of an oxygen than in a nitrogen atmosphere (Table 2). The only exception is the char from chicken mature (PL) which exhibits a mercury retention capacity which is so low that no difference can be detected between the two atmospheres. This suggests that mercury-sorbent interactions may take place in presence of certain reactive gases, like oxygen, that favour mercury oxidation. This matter will be discussed in some depth in future studies in which special emphasis will be placed on a range of gases that may be present during coal combustion processes.
4. Conclusions The gasification chars obtained from plastic-paper wastes used in this work are promising candidates for mercury capture in the experimental conditions evaluated. Mercury retention capacities comparable to those obtained with a commercial activated carbon specifically designed to capture elemental mercury were achieved. The results show that a high chlorine content and lower pH benefit mercury capture. This study also reveals that reactive components of the atmosphere, such as oxygen, may favour the adsorption of mercury.
Acknowledgments The financial support for this work was provided by the project MERCURYCAP (RFCR-CT-2007-00007). The authors thank the Energy Research Centre of the Netherlands for supplying the chars employed in this study.
References [1] Milford JB, Pienciak A. After the Clean Air Mercury Rule: Prospects for reducing mercury emissions from coal-fired power plants. Environ. Sci. Technol. 2009; 43(8): 2669-73. [2] Klasson KT, Lima IM, Boihem Jr LL, Wartelle LH. Feasibility of mercury removal from simulated flue gas by activated chars made from poultry manures. J. Environ. Mang. 2010, 91: 2466-70. [3] Lei CH, Yufeng D, Yuqun Z, Liguo Y, Liang Z, Xianghua Y et al. Mercury Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
transformation across particulate control devices in six power plants of China: The coeffect of chlorine and ash composition. Fuel 2007, 86: 603-10. [4] Yang H, Xu Z, Fan M, Bland AE, Judkins RR. Adsorbents for capturing mercury in coal-fired boiler flue gas. J. Hazard. Mater. 2007, 146: 1-11.
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Oviedo ICCS&T 2011. Extended Abstract
Biomimetic Sequestration of CO2 and Conversion to CaCO3 Using Enzyme Extracted Oyster S.K. Jeong, Y.I. Yoon, S.C. Nam Korea Institute of Energy Research, 71-2 Jang-dong, Yusung-ku, Daejeon 305-343, Korea ([email protected]) Abstract In this study, carbon dioxide sequestration activity was compared and evaluated using bovine carbonic anhydrase (BCA) and a water soluble protein extract derived from hemocytes from diseased shell (HDS). Para-nitrophenyl acetate (p-NPA) was used to measure the reaction rate. The kcat/Km values obtained from the Lineweaver-Burk and Michaelis-Menten equations were 230.7 M-1s-1 for BCA and 194.1 M-1s-1 for HDS. Without a biocatalyst, CaCO3 production took 15 seconds on average, while it took 5 seconds on average when BCA or HDS were present, indicating an approximately 3-fold enhancement of CaCO3 production rate by the biocatalysts. The biocatalytic hydration of CO2 and its precipitation as CaCO3 in the presence of biocatalysts were investigated.
1. Introduction Since the industrial revolution, fossil fuel consumption has drastically increased. Carbon dioxide discharged from combustion processes is now believed responsible for global warming and an increase in the earth's average temperature [1]. The carbon dioxide concentration in the atmosphere has increased by about 35% over the past 100 years, and it continues to increase by more than 0.4% each year, resulting in elevation of the sea level, abnormal air temperatures, and worldwide destruction of the ecosystem. In order to reduce the emission of carbon dioxide, the greatest contributor to greenhouse effects among the 6 major greenhouse gases listed by the IPCC (Intergovernmental Panel on Climate Change), substantial efforts need to be aimed at increasing energy efficiency, use of nuclear power, the introduction of new and renewable energy sources, and the installation of CCS (carbon dioxide capture and storage) process. In the case of CCS, the ETP (Energy Technology Perspectives) report of the IEA (International Energy Agency) predicted that CCS processes will account for 19~22% of the carbon
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Oviedo ICCS&T 2011. Extended Abstract
dioxide emission reduction by 2050 [2], indicating a significant role for CCS in reducing global warming. Among the technologies of CO2 capture, the absorption process is the most actively studied process, since it can conveniently capture large masses of carbon dioxide. However, absorption process has some drawbacks such as erosion of the reaction apparatus by the applied solvent, loss of solvent due to volatility, and the need for high heat duty for regeneration of the solvent in stripper [3,4]. CCS also requires additional space for storage of the captured carbon dioxide. Currently, CCS is very expensive to install and to operate, resulting in a potential doubling of electricity costs. For this reason, many research programs are aimed at developing more economical CCS processes. One of new technologies that can replace conventional CCS processes is carbon dioxide sequestration using a biocatalyst. Carbon dioxide capture by biocatalysts has a number of advantages, such as rapid carbon dioxide absorption, low energy requirements, no need for additional storage, and economic payback in the form of useful minerals following addition of an appropriate cation. In this study, we studied carbon dioxide sequestration using hemocytes from diseased shell (HDS) extracted from oyster (Crossostrea Gigas) that are widely distributed in sea. Carbon dioxide sequestration was studied by varying the conditions of HDS and comparing their effectiveness to that of BCA. The study also investigated the kinetics of the biocatalyst reaction using para-nitrophenyl acetate (p-NPA) and followed biomineralization using calcium ions.
2. Experimental section The biocatalysts employed for the hydration of CO2 and mineralization in this study was BCA and HDS. The molecular weight of the individual biocatalysts measured by MALDI-TOF: BCA 29 kDa and HCA 25.5 kDa. 0.1M Tris-HCl buffer (SigmaAldrich Co.) solution was used to maintain pH 7.0 for the biocatalytic activity test. Biocatalyst catalyzes the hydration of CO2, and consequently, hydrogen ions are transferred between the active site of the biocatalyst and the surrounding buffer solution. This results in a change in pH. The pH variation over time was monitored in real-time using a pH meter (Thermo Scientific Co. Orion Star Series). The experiment was carried out as a batch reaction. The desired reaction temperature was maintained by water
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Oviedo ICCS&T 2011. Extended Abstract
circulating pump. Tipically, 500 ml of 0.1M Tris-HCl buffer was stirred at 150 rpm to enhance diffusion. Nitrogen was injected uniformly to the bottom of the reactor through a porous sintered metal plate to eliminate all residual gas-phase substances in the TrisHCl buffer. The selected biocatalyst was then added and dissolved by stirring for 15 minutes. A mixture of carbon dioxide, oxygen, and nitrogen was then injected at a rate of 500cc/min. Progress of the reaction was monitored by measuring pH of the mixture. The experiment was stopped when there was no further change in pH. Carbon dioxide concentration was mainly set at 12%, which is the average concentration in flue gas from a coal-fired power plant. The biocatalytic activities of the biocatalysts were estimated spectrophotometrically using p-NPA as a substrate, as reported previous paper [5]. CaCO3 precipitation was carried out after hydration of CO2 in a batch reactor. To obtain the final product, CaCO3, a saturated carbon dioxide solution was prepared. 500 ml of distilled water was stirred in the reactor at 150 rpm and 25°C, while carbon dioxide was injected into it for 3 hours. CaCl2 (Sigma-Aldrich Co) was prepared as 0.1 M solution and the selected biocatalyst was dissolved into these solutions by mixing for 30 minutes. Then, a 20 ml sample of the previously prepared carbon dioxide solution was added to the mixture and stirred for formation of CaCO3. The amount of CaCO3 was determined after filtration and drying. After weighing of CaCO3, the particle size and morphology of particles was measured by SEM, and the structure of the CaCO3 was analyzed by XRD.
3. Results and Discussion The biocatalysts produce p-NP and acetic acid through an acyl-enzyme intermediate in the hydration reaction with p-NPA. The reaction path way of p-NPA and CO2 hydration for metalloenzymes is almost same. Therefore, we can estimate the CO2 hydration efficiency of metalloenzymes from the results of p-NPA hydrolysis. The Km and kcat values were obtained by the Michaelis-Menten (equation 1) and Lineweaver– Burk (equation 2) equations, and were used to calculate the rate constant, kcat/Km.
k cat [ S ][ Eo] K m + [S ]
(1)
K 1 1 1 = + m V Vmax Vmax [ S ]
(2)
V=
where V is the rate of p-NP formation, Vmax is the maximum rate, kcat is the catalytic rate
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3
Oviedo ICCS&T 2011. Extended Abstract
constant, [E0] the enzyme concentration, [S] is the p-NPA concentration, and Km is the rate constant when the rate is equal to Vmax/2, and kcat/Km is the kinetic constant. The carbon dioxide reaction rate of BCA and HDS, measured through the p-NPA reaction, was determined by the absorbance, which is proportional to the reaction rate of BCA and HDS. The reaction rate of BCA was higher than that of HDS. Here, HDS is not a pure enzyme, it consisted of several proteins. The reaction rates of HDS and BCA were similar at biocatalyst concentrations above 1 μM. Table 1 summarizes all of the kinetic parameters. The kcat/Km rate constant was 230.7 M-1s-1 for BCA and 194.1 M-1s-1 for HDS. Considering that HDS is a water-soluble, conjugated protein, the kcat/Km value of HDS may be regarded as similar to that of BCA. This means that BCA, which is expensive and difficult to extract, can be replaced by the more economical HDS biocatalyst extracted from oysters.
Table 1. Summary of the kinetic parameters of p-NPA hydrolysis by BCA and HDS Enzyme
BCA
HDS
Eo (μmol/l)
0.18276
0.05
1/Vmax
0.0213
0.61297
Slope
0.3953
1.71735
Vmax(μmol/min)
46.9484
1.63151
Km(mM)
18.5587
2.80187
kcat(S-1)
4.2814
0.543836
kcat/Km (M-1S-1)
230.7
194.1
CaCO3 was formed by adding Ca2+ ion to the CO2 hydrated solution containing CO32-. The amount of CaCO3 was also measured at different times during the reaction. The initial formation of CaCO3 required about 15 seconds in the CaCO3 precipitation experiment without the biocatalysts. In contrast, when BCA and HDS were added, CaCO3 formation occurred within 5 seconds, on average, for a 3 times faster formation
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4
Oviedo ICCS&T 2011. Extended Abstract
rate. The final quantity of the produced CaCO3 was the same since it was strongly dependent on the amount of CO32- originally dissolved in the water. Figure 1 shows the size of the produced CaCO3 analyzed by SEM. CaCO3 particles produced without the biocatalysts were larger than 10 µm in diameter, but were less than 4 µm when BCA and HDS were used. There are two models that can describe size enlargement in precipitation system. One is the deposition of ionic or molecular species on crystal surfaces and the other is aggregation mode. The later process is found to be most important parameter to increase particle size. The aggregation mode depends on hydrodynamic conditions like agitation. In this study there is same experimental condition. Therefore, we can conclude that the particle size partially depends on the existence of enzyme. Berman et. al. reported that a soluble protein, extracted from the shell of a mollusk, inhibited the growth of particle [6]. However, in this study, the particle size with enzyme is smaller than that without enzyme. It might be that additional study may be required in this regard. In addition, the reported that the CaCO3 was a calcite structure, as shown by the calcite peak with 2θ=29o in the XRD analysis.
Figure 1. SEM images of CaCO3 by (A) BCA (B) without BCA, (C) HDS, (D) without HDS.
4. Conclusions In this study, we used biocatalysts to capture carbon dioxide in cost-effective and ecofriendly manner, in an effort to eliminate some of the drawbacks of conventional CCS
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5
Oviedo ICCS&T 2011. Extended Abstract
technologies for carbon dioxide sequestration. Using a p-NPA, the reaction rates for BCA and HDS were determined as 230.7 M-1s-1 and 194.1 M-1s-1, respectively. Formation of CaCO3, the final product of the carbon dioxide sequestration process, took about 5 seconds when the biocatalysts were used and 15 seconds when they were not used, indicating that the biocatalysts increased the CaCO3 formation rate approximately 3-fold. At present, the study of carbon dioxide sequestration using biocatalysts is essentially only fundamental research. However, carbon dioxide sequestration using biocatalysts is likely to be the most stable method for carbon dioxide capture and storage. In addition, CaCO3 can be utilized as road pavement or paper coating materials.
References [1] B. Metz, O. Davidson, H. de Coninck, M. Loos and L. Meyer, IPCC Special Report on Carbon Capture and Storage, 2005. [2] OECD/IEA, Energy Technology Perspectives, 2008 [3] S. G. Bishnoi and T. Rochelle, AlChE J, 2002;48. [4] M. Aineto, A. Acosta, J. M. Rincon and M. Romero, Fuel, ; 85; 2352. [5] M. Vinoba, D. H. Kim, K. S. Lim, S. K. Jeong, S. W. Lee and M. Alagar, Energy Fuels 2001; 25; 438 [6] A. Berman, L. Addadi, and S. Weiner, Nature, 1988; 331; 546
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6
Carbon dioxide sequestration by aqueous mineral carbonation of serpentine and explanation of experimental results Kimia Alizadehhesari1, Karen Steel
School of Chemical Engineering, The University of Queensland, St Lucia, QLD 4072, Australia
Abstract As the concentration of CO2 in the atmosphere increases there have been numerous studies to find economically feasible processes for the long-term sequestration of CO2. This research investigates an understanding of, and control of, the ionic equilibria in solution to enable a new process to be developed for the conversion of Mg-silicate ores to magnesium carbonate. This paper presents results for the dissolution of Mg from natural serpentine. The next phase of the research involves using novel methods to adjust pH and enable carbonate precipitation. 1. Introduction The potential threat of climate change due to an increase in greenhouse gas emissions requires a feasible, leakage-free, and economically viable method of CO2 storage [1, 2]. The reaction of magnesium-rich minerals such as serpentine with CO2 to form stable mineral carbonates is a promising approach for long-term storage of CO2 as a solid benign deposit [3-5] and the raw materials are available world-wide in vast quantities. Serpentine draws attention for the sequestration of CO2 as a suitable raw material due to its worldwide existence [6-8]. Nevertheless mineral sequestration has a major drawback owing to the difference in pH needed for Mg dissolution and MgCO3 precipitation. The most effective method of carbonation involves firstly dissolving magnesium ions by a solvent and then magnesium carbonate is precipitated by a reaction between magnesium ions and aqueous CO2. The CO2 should undergo a number of transformations such as the dissolution in an aqueous phase, hydration by water, ionization, carbonate formation, and finally the precipitation by the available Mg2+ ions. The reactions are: 1
Corresponding author email: [email protected]
1 2 3 4 5 An acidic medium is needed to increase the efficiency of mineral leaching however for reaction 4 forming the
the solution must be basic [9]. Reaction 1, dissolving CO2 in water, is quite
fast. When pH is high enough between 8 and 10 reaction 4 occurs to a substantial extent. The key issue in forming the carbonates is to control the pH. O’Connor et al. [10] at Albany Research Center proposed a high pressure and temperature reactor (185°C and 115 atm) utilizing distilled water as the liquid medium and bicarbonate/salt mixture to buffer the pH of solution in the range of 7.7-8. Although they could achieve up to 78% stoichiometric conversion of silicate minerals in 30 minutes, it is too expensive to be applied in industrial scale in terms of high energy usage and large amount of additives. Another approach is the pH swing process which was first developed in Japan presented in a patent by Yogo [11]. In this process at the first step the pH of the solution is lowered which results in extraction of metal ions followed by the second step where the pH is raised in order to enhance the precipitation of carbonate. Park et al.[12] suggested a pH swing process in a fluidized-bed reactor using weak acids to increase the dissolution of serpentine and then by raising the pH to 9.5 by means of NH4OH forming Nesquehonite (MgCO3). In other work by Teir et al. [13] HCl or HNO3 were used for mineral leaching and then by addition of NaOH the pH was raised to 9 to produce carbonates. The chemicals needed in these processes make it completely unviable. Kodama et al. [9], used steel slags as a source of metal ions to react with (NH4)2SO4 for the first step and then purged CO2 to precipitate carbonate. However the capacity of steel slags is quite limited to capture giga tones of CO2 in large scale. There have been several studies that propose a biological catalyst, carbonic anhydrase (CA), be used to increase the kinetics of reactions 2-4 [14]. Bond et al. [14] showed that by employing the
enzyme CA the rate of CO2 hydration for the mineral sequestration was accelerated. Liu et al. [15] investigated the precipitation of calcium carbonates from produced water in the presence of the CA enzyme. The same problem of needing to raise the pH exists with this approach. Besides, some barriers exist for applying enzyme like: enzymes are pricey due to their purification costs and they need to be immobilized for their repeatable usage [16]. Although there have been various studies on the dissolution kinetics and mechanisms of magnesium from serpentine in different solvents to improve the efficiency [12, 17], most of provided experimental data is not consistent due to the varying properties of natural minerals as well as low aqueous concentrations [18]. In the present work, first the dissolution of serpentine in distilled water and hydrochloric acid is studied. The effects of the HCl concentration, temperature, and residence time on the mineral leaching were investigated. The study offered useful data for the optimization of Mg dissolution.
2. Experimental 2.1 Characterization of Serpentine In order to make a study of the potential of the minerals for carbon storage it is important to use natural minerals. The serpentine [Mg3Si2O5(OH)4] used for this study was obtained from the naturally occurring materials in northern Queensland, Australia. The serpentine was initially ground by hand in a pestle and mortar and then in a laboratory attrition mill. Ground serpentine was sieved with ASTM standard sieves to get d<57 μm. The mineral phases present in the sample was determined by XRD analyses. A diffractometer (PANanalytical XPERT-PRO) with Cu-Kα target (λ=0.15406 nm) was used at room temperature. Measurements were made in a step scan mode (0.1°/step) over the 2θ range of 1090°. The XRD pattern of serpentine is shown in Figure 1. As indicated by the XRD pattern, antigorite is the primary phase present in original serpentine and no excess phase was detected.
60000
50000
In ten sity
40000
30000
20000
10000
0
0
20
40
60
80
100
Angle, 2-Theta
Figure 1. XRD pattern of serpentine sample.
EDS spectrums were collected from the prepared samples. EDS profile is given in figure 2. The peak at 8 keV on the spectrum is from the Ag coating and can be neglected. The key elements of serpentine are identified as magnesium (Mg), silicon (Si) and oxygen (O) and iron (Fe). The amount of calcium (Ca) was found to be insignificant.
Figure 2. EDS spectrum of the natural serpentine.
2.2 Magnesium extraction from serpentine
Extraction experiments on dissolution of serpentine were carried out using a 250 ml round bottom flask reactor in 100 ml solution using concentrated hydrochloric acid with 0.5 g of <57 μm fraction of serpentine, which was heated 3 hours by a temperature-controlled stirrer at 625 rpm and equipped with a water-cooled condenser to minimize solution losses due to evaporation. Solution was heated up to the temperature required for boiling. For the kinetics studies, the t0 started when boiling started. The effects of temperature, concentration of HCl, residence time and volume of the solution were investigated to find the best conditions for the leaching of Mg. All the solutions were prepared individually for each reaction. Once reacted, the precipitation was vacuum filtered and dried over night and then weighed. Solutions were kept for measuring the pH. Since magnesium exists as serpentine [Mg3Si2O5(OH)4], extraction of magnesium can be described in the form of reaction 6: 6
3
6
2
5
6
3. Results and discussion The dissolution results for the effect of HCl concentration is shown in Fig 3(a). An interesting result was that a significant amount of Mg (20%) was dissolved in distilled boiling water when no HCl was used. From eq.(6) the stochiometric amount of acid needed is 0.12M which gives 46% magnesium extraction. The dash line in Fig 3 shows the stochiometric level needed for complete dissolution. A considerable amount of magnesium (77%) from serpentine was extracted after 3 hours boiling in 0.6 M HCl solution. The curve reaches a plateau at 0.25M and after this point increasing the concentration of HCl did not affect the magnesium extraction. Because it is desirable to not have excess acid in solution for the carbonation a compromise between Mg extraction and solution pH must be struck. Fig 3(b) shows the pH change when using different concentration of acids. The dotted line in Fig 3 illustrates the operating point for temperature and residence time experiments.
100
Mg Extraction (%)
80
60
40
Experimental results Stochiometric level Operating level
20
0 0.0
0.1
0.2
0.3
0.4
0.5
0.6
HCl concentration (M)
(a)
8 Experimental results Stochiometric level Operating level
pH
6
4
2
0 0.0
0.1
0.2
0.3
0.4
0.5
0.6
HCl concentration (M)
(b) Figure 3. (a)Effect of HCl concentration on the dissolution of Mg, (b)pH measurement at different HCl concentration (100°C, 3 hrs residence time).
The effect of residence time on dissolution of Mg in 0.25 M is shown in figure 4. Increasing the residence time from 15 min to 3 hours raises the Mg dissolution but with more than three hours the Mg dissolution levels off.
100
Mg Extraction (%)
80
60
40
20
0 0
1
2
3
4
5
6
7
Residence time (hr)
Figure 4. Residence time experiments at 0.25 M solution and 100°C.
100
Mg Extraction (%)
80
60
40
20
0 20
40
60
80
100
Temperature (C)
Figure 5. Effect of temperature on dissolution of Mg in 0.25 M solution and 3 hrs residence time.
The effect of temperature upon the dissolution of serpentine was tested at temperatures of 25 100°C in 0.25 M HCl solution (Figs 5) for 3 hours. As expected, temperature has a significant effect upon the solubility of magnesium from serpentine. For temperatures less than 50°C only a tiny amount of Mg is dissolved. At temperatures higher than 50°C a linear relationship is obtained suggesting that at higher temperatures above 100°C higher extraction efficiencies might be achieved. Extrapolation of the Mg extraction line to 100% occurs at approximately140°C.
4. Conclusion In this paper the dissolution of natural serpentine ore using distilled water and hydrochloric acid was investigated. It has been shown that distilled water at boiling temperature can extract considerable amount of magnesium (19.3%). Approximately 46% of the Mg is extracted with the stoichiometric level of HCl and at high HCl concentrations approximately 77% of the Mg can be extracted. Lower HCl concentrations are preferred as it is easier to raise the pH for the subsequent step of carbonation. Higher temperatures are also preferred and could further enable the acid requirements to be reduced. The next phase of this research will involve the use of a novel set of additives for controlling the pH and enabling magnesium carbonate to precipitate.
References [1] S. Solomon, Cambridge Univ. Press (2007) 21-91. [2] B. Metz, O. Davidson, H. C. de Coninck, M. Loos, and L. A. Meyer (eds.), Cambridge University Press (2005). [3] IPCC, Forth Assessment Report; Mitigation of climate change, Cambridge University Press, 2007. [4] P.C. Goldberg, Z.Y.;O'Connor, W.K.;Walters, R.;Ziock, H., Journal of Energy and environmental research 1 (2001) 117. [5] W. Seifritz, Nature 345 (1990). [6] R.K. Schulze, M.A. Hill, R.D. Field, P.A. Papin, R.J. Hanrahan, D.D. Byler, Energy Conversion and Management 45 (2004) 3169-3179. [7] J. Fagerlund, S. Teir, E. Nduagu, R. Zevenhoven, Energy Procedia 1 (2009) 4907-4914. [8] W.K.D. O'Connor, D.C.; Rush, G.E.; Gerdemann, S.J.; Penner, L.R.; Nilsen, R.P., (2005). [9] S. Kodama, T. Nishimoto, Energy 33 (2008) 776-784. [10] W.K. O'Connor, D.C. Dahlin, D.N. Nilsen, G.E. Rush, R.P. Walters, P.C. Turner, 5th International Conference on Greenhouse Gas Technologies, Cairns, Australia (2000). [11] K. Yogo, Eikou, T., Tateaki, Y., Method for fixing carbon dioxide, JP2005097072, 2005. [12] A. Park, L. Fan, Chemical Engineering Journal 59 (2004) 5241-5247. [13] S. Teir, Kuusik, R., Fogelholm, C.-J., Zevenhoven, R., Int. J. of Miner. Proc. 85 (2007) 1-15. [14] M.G.S. Bond, J.;Brandvold, D. K.; Simsek, F. A.; Medina, M.; Egelang, G., Energy & Fuels 15 (2001) 309-316. [15] N.B. Liu, G.M.; Abel, A.; McPherson, B.J.; Stringer, J., Fuel Process. Technol. 86 (2005) 16151625. [16] E. Ozdemir, Energy & Fuels 23 (2009) 5725-5730. [17] S. Teir, Revitzer, H., Eloneva, S., Fogelholm, C.-J., Zevenhoven, R., International Journal of Mineral Processing 83 (2007) 36-46. [18] J.T. Sipila, S.;Zevenhoven,R., Carbon dioxide sequestration by mineral carbonation literature review, Abo Akademi University Faculty of Technology Heat Engineering Laboratory, 2008.
CHARACTERIZATION OF A NOVEL FLAT-PANEL AIRLIFT PHOTOBIOREACTOR WITH AN INTERNAL HEAT EXCHANGER Luce Helena Kochem1, Nicéia Chies Da Fré1, Cristiane Redaeli1, Nilson Romeu Marcílio2, Rosane Rech1. 1
Bioengineering Laboratory, Food Technology Institute, Federal University of Rio Grande do Sul - UFRGS Porto Alegre, RS – Brazil, e-mail: [email protected], [email protected], [email protected] 2 Chemical Engineering Department, School of Engineering, Federal University of Rio Grande do Sul – UFRGS Porto Alegre, RS – Brazil, e-mail: [email protected]
ABSTRACT This work characterizes a novel type of flat-panel airlift photobioreactor with an internal heatexchanger (FPA-IHE) separating the riser and the downcomer zones. The overall heat and mass transfer coefficients, along the mixture and circulation times, were estimated. The FPAIHE has 2.2 L of working volume, high of 450 mm and light path of 50 mm. The overall heat transfer coefficients were estimated for the internal heat-exchanger (Ui) and for the external surface (Ue), resulting in Ui = 47.0 W m-2 K-1 and Ue = 7.1 W m-2 K-1. Further analysis showed that the main resistances to internal and external heat transfer were, respectively, the acrylic wall and the resistance of the air surrounding the reactor. The mass transfer coefficient (kLa) increased with the volumetric power input until around 0.01 s-1. The mixture time decreased as the superficial gas velocity in the riser increased. The circulation time ranged between 5 and 6 sec. The growth of the microalgae Dunaliella tertiolecta and Chlorella minutissima showed a high specific growth rate compared with earlier published values, making the FPA-IHE photobioreactor a promising environment to develop high productivity processes using microalgae. 1. INTRODUCTION In the last century, the combustion of fossil fuels has increased due to industrialization and population expansion. As a result, CO2 levels in the atmosphere have increased from 260 to 380 ppm. Climatic changes generated by greenhouse gases (GHG) such as CO2 have severe impacts, including reduced agricultural production, changes in fresh water supplies, species extinction, storms and floods (Morais and Costa, 2008). Therefore, technologies for the absorption or fixation of excess CO2 must be developed. In the biological carbon cycle, organic carbon can be oxidized to CO2 in several ways; however, photosynthesis, which is performed by plants and algae, is the only metabolic process that captures and converts CO2 to organic carbon (glicides, lipids, proteins and cellulose). Compared to plants, microalgae grow faster, have shorter life cycles and can be easily genetically modified. Recently, the development of photobioreactors for microalgae cultures has gained interest in the scientific community due to the ability of these microorganisms to use CO2 as the sole carbon source and produce high-value products such as biofuel, food, feed and bioactive compounds (Chisti, 2007). In several countries, microalgae are cultivated in open ponds to produce algal biomass, pigments and feed. However, open pond cultures have low biomass concentrations, low productivities, low efficiencies in light utilization, and high risk of contamination. In addition, the efficiency of CO2 transfer between the gas and liquid phase is low, and the process is
difficult to control. These limitations can be overcome in closed photobioreactors (Mata et al., 2010). In the present study, a novel flat-plate airlift photobioreactor with an internal heat exchanger separating the riser and downcomer zones was developed. The heat and mass transfer coefficients of the photobioreactor were determined, and the mixing and circulation times were evaluated. In addition, Dunaliella tertiolecta and Chlorella minutissima were grown to evaluate the performance of the proposed photobioreactor. 2. MATERIALS AND METHODS 2.1. Photobioreator The photobioreactor evaluated in the present study is a bench-scale flat-panel airlift reactor with an internal heat exchanger (FPA-IHE) separating the riser and downcomer zones (Figure 1). The photobioreactor was constructed from 4-mm acrylic sheets, and one side of the reactor was illuminated. The proposed FPA-IHE possessed the following specifications: height = 450 mm; width = 108 mm; length = 80 mm; working volume = 2.2 L, external superficial area = 1306.3 cm2; and internal heat exchanger area = 635 cm2.
Figure 1. Flat-panel airlift photobioreactor with an internal heat exchanger. 2.2. Heat transfer, mass transfer and hydrodynamic characterization The overall heat transfer coefficients were determined for the internal heat exchanger (Ui; heat transfer between the broth and the fluid circulating through the heat exchanger) and the outer surface of the reactor, which was exposed to air (Ue; heat transfer between the broth and the surroundings). The overall heat transfer coefficients were estimated from experimental data using EMSO software (Environment for Modeling, Simulation and Optimization) (Soares, 2003) according to the energy balance for the system. The mass transfer coefficient of the liquid phase, kLa, was determined according to the dynamic gassing-out method (Chisti, 1989). The superficial gas velocity in the riser, uGr, is the ratio between the air flow-rate, Q, and the cross-sectional area of the riser. The mixing time, tm, and the circulation time, tc, were calculated according to the acidtrace method (Mirón et al., 2004). The mixing time was defined as the time required to achieve
95% complete mixing, and the circulation time was defined as the time required for the fluid to completely loop around the internal heat exchanger (Chisti, 1989). 2.3. Microorganisms and growth conditions The microalgae Dunaliella tertiolecta and Chlorella minutissima were cultured in Guillard f/2 medium (Lourenço, 2006). Inocula were grown in an orbital shaker at 30°C and were illuminated with an electronic circular lamp (30 W day light, Avant). A 10-mL aliquot of algal stock culture was inoculated into 100 mL of culture medium. After 7 days, 100 mL of culture medium was added to the flasks, and the cultures were grown for 5 days. The FPA-IHE photobioreactors were sterilized with sodium hypochlorite (Andersen, 2005) and were subsequently inoculated with 200 mL of inoculum in 2.000 mL of culture medium. The temperature of the reactor was maintained at 30°C by circulating water from a thermostatic bath through the internal heat exchanger, and the air flow-rate was set to 0.5 L min-1. The riser side of the FPA-IHE photobioreactor was illuminated with a panel of electronic lamps (24 × 13 W cool light, Tashibra) at 18.000 lux. The concentration of biomass was measured by determining the optical density at 570 nm (Lourenço, 2006), and the data were correlated to the dry weight of the cells. 3. RESULTS AND DISCUSSION Culture temperature is a key parameter of photobioreactors, although many lack temperature control systems. The FPA-IHE photobioreactor use an internal heat exchanger to control the temperature. Table 1 shows the heat transfer coefficients estimated from the experimental data. The results revealed that Ui was greater than Ue, which suggests that heat transfer between the heat exchanger and the fluid was more efficient than heat transfer to the surroundings. However, compared to the results obtained in other studies, the heat transfer coefficient of the heat exchanger was low (Sierra et al., 2008). The acrylic wall of the heat exchanger may be highly resistant to heat transfer due to the poor thermal conductivity of the material. Table 1: Heat transfer coefficients of the FPA-IHE photobioreactor. Heat transfer coefficient between the broth and the fluid circulating through the heat exchanger, Ui between the broth and the surrounding, Ue
47.0 W m-2 K-1 7.1 W m-2 K-1
One of the most significant parameters governing the performance and design of bioreactors is the mass transfer coefficient of the liquid phase, kLa. Figure 2 shows the kLa data obtained in the present study. The mass transfer coefficient increases with the air velocity in the riser. The estimated values of kLa, between 0.002 and 0.010 s-1 are similar to those reported for rectangular airlift reactors and higher to those reported for rectangular bubble column reactors (Chisti, 1989; Reyna-Velarde, 2010). The mixing (tm) and circulation (tc) times were analyzed according to the acid trace method, and the results are shown in Figure 3 as a function of the gas velocity in the riser (uGr). The mixing times were comparable to those obtained for a flat-panel airlift reactor with a similar height (Reyna-Velarde, 2010). As the gas velocity increased, tm decreased until uGr was equal to 0.10 m s-1. At this point, an inflection was observed. Thus, when uGr is greater than the aforementioned value, the operating conditions are not optimal. Namely, the value of tm increases, and only a slight increase in kLa is observed. Similar behavior was also detected in a
flat-plate photobioreactor (Sierra et al., 2008). The circulation time oscillated between 4 and 7 seconds and was independent of the gas velocity.
Figure 2. Relationship between mass transfer coefficient (kLa) and gas velocity (uGr) of the proposed FPA-IHE reactor..
(a)
(b)
Figure 3: Relationship between the (a) mixing time (tm) and (b) circulation time (tc) with gas velocity (uGr) of the proposed FPA-IHE reactor. Table 2 shows the growth parameters for D. tertiolecta and C. minutissima in the FPAIHE photobioreactor. The microalgae reached the stationary phase at 80 and 110 hours of growth, and 0.35 and 0.38 g L-1 of cells were obtained, respectively. In previous studies conducted in photobioreactors without CO2 injections, 0.68 g L-1 of Chlorella sp. (Chiu et al., 2008) and 0.2 g L-1 of Nannochloropsis oculata (Chiu et al., 2009) were obtained. Although the biomass yield obtained in the present study was low, both strains presented high specific growth rates in the FPA-IHE reactor (0.77 d-1 for D. tertiolecta and 0.67 d-1 for C. minutissima). The reported specific growth rates for Chlorella range from 0.14 to 0.25 d-1 (Chiu et al., 2008; Converti et al., 2009), which suggests that the proposed FPA-IFE is an adequate environment for the growth of microalgae. To increase the biomass yield, air enriched with CO2 could be injected into the photobioreactor. This technique has shown promising results for various species of microalgae (Degen et al., 2001; Hsieh and Wu, 2009; Watanabe and Saiki, 1997; Jacob-Lopes et al., 2009).
Table 2: Growth parameters of D. tertiolecta and C. minutissima in the proposed FPA-IHE photobioreactor. Xm: maximum biomass; μm: maximum specific growth rate. Xm (g L-1 ) μm (d-1) D. tertiolecta C. minutissima
0.35 0.38
0.77 0.67
4. CONCLUSIONS The aim of the present study was to characterize a novel configuration of a flat-panel airlift photobioreactor with an internal heat exchanger separating the riser and downcomer zones. As the gas velocity of the riser, an increase in the volumetric mass transfer coefficient was observedThe mixing time decreased with an increase in the gas. The circulation time ranged from 5 to 6 sec, and a correlation between the circulation time and the gas velocity was not observed. The biomass concentrations of D. tertiolecta and C. minutissima in the FPA-IHE photobioreactor were similar to the published results for various species of microalgae; however, the tested strains presented high specific growth rates. Thus, the FPA-IHE photobioreactor is a promising environment for the development of microalgae processes with high productivities.
ACKNOWLEDGMENTS Authors are very grateful to CNPq (Brazilian National Council of Research and Development) and to RNC (Brazil Coal National Network, http://www.ufrgs.br/rede_carvao) by the financial support for this research.
BIBLIOGRAPHY ANDERSEN A.R., Algal Culturing Techniques, Elsevier Applied Science, Oxford, 2005. CHISTI M.Y., Airlift Bioreactors, Elsevier Applied Science, London and New York, 1989. CHISTI Y., Biodiesel from microalgae, Biotechnol. Advan. 25 (2007) 294-306. CHIU S.Y., KAO C.Y., CHEN C.H., KUAN T.C., ONG S.C., LIN C.S., Reduction of CO2 by a high-ensity culture of Chlorella sp. in a semicontinous photobioreactor, Bioresour. Technol. 99 (2008) 3389-3396. CHIU S.Y., KAO C.Y., TSAI M.T., ONG S.C., CHEN C.H., LIN C.S., Lipid accumulation and CO2 utilization of Nannochloropsis oculata in response to CO2 aeration, Bioresour. Technol. 100 (2009) 833–838. CONVERTI A., CASAZZA A.A., ORTIZ E.Y., PEREGO P., DEL BORGHI M., Effect of temperature and nitrogen concentration on the growth and lipid content of Nannochloropsis oculata and Chlorella vulgaris for biodiesel production, Chem. Eng. Process. 48 (2009) 1146–1151.
DEGEN J., UEBELE A., RETZE A., SCHMID-STAIGER U., TRÖSCH W., A novel airlift photobioreactor with baffles for improved light utilization through the flashing light effect, .f Biotechnol.y 92 (2001) 89–94. HSIEH C.H., WU W.T., A novel photobioreactor with transparent rectangular chambers for cultivation of microalgae, Biochem. Eng. J. 46 (2009) 300–305. JACOB-LOPES E., SCOPARO C.H.G., LACERDA L.M.C.F., FRANCO T.T., Effect of light cycles (night/day) on CO2 fixation and biomass production by microalgae in photobioreactors, Chem. Eng. Process. 48 (2009) 306–310. LOURENÇO S.O., Cultivo de Microalgas Marinhas: Princípios e Aplicações, first ed., Rima (Ed.), São Paulo, 2006 MATA M.T., MARTINS A.A., CAETANO S.N., Microalgae for biodiesel production and other applications: A review, Renew. Sustain. Energy Rev. 14 (2010) 217-232. MIRÓN A.S., GARCÍA M.C.C., CAMACHO F.G., GRIMA E.M., CHISTI Y., Mixing in Bubble Column and Airlift Reactors, Chem. Eng. Res. Des. 82 (2004) 1367-1374. MORAIS G.M., COSTA J.A.V., Bioprocessos para remoção de dióxido de carbono e óxido de nitrogênio por microalgas visando a utilização de gases gerados durante a combustão do carvão, Quím. Nova 31 (2008) 1038-1042. REYNA-VELARDE R., CRISTIANI-URBINA E., HERNÁNDEZ-MELCHOR D.J., THALASSO F., CAÑIZARES-VILLANUEVA R.O., Hydrodynamic and mass transfer characterization of flat-panel airlift photobioreactor with high light path, Chem. Eng. Process. 49 (2010) 97-103. SIERRA E., ACÍEN G.F., FERNÁNDEZ M.J., GARCÍA L.J., GONZÁLEZ C., MOLINA E., Charaterization of a flat plate photobioreactor for the production of microalgae, Chem. Eng. J. 138 (2008) 136-147. SOARES R.P., EMSO : an integrated tool for process modeling, dynamic simulation and optimization, American Institute of Chemical Engineeers. Annual Meeting (2003 : San Francisco, Calif.) Technical programs. New York : AIChE, 2003. WATANABE Y., SAIKI H., Development of a photobioreactor incorporating Chlorella sp. for removal of CO2 in stack gas, Energy Convers. Manag. 38 (1997) 499-503.
Oviedo ICCS&T 2011. Extended Abstract
CLEAN COAL TECHNOLOGIES SCENARIO AND EVALUATION OF PRESENT CO2 DWINDLING INITIATIVES TO APPROACH ZERO EMISSION POWER STATIONS BY COAL COMBUSTION. DEPLOYMENT SITUATION AND EVALUATION STUDY Francisco Guerrero, Carmen Clemente-Jul Department of Chemical Engineering and Fuels. ETS Ingenieros de Minas. Universidad Politécnica de Madrid (UPM), Alenza 4. 28003 Madrid (Spain).
Abstract In the present uncertain global context of reaching an equal social stability and steady thriving economy, power demand expected to grow and global electricity generation could nearly double from 2005 to 2030. Fossil fuels will remain a significant contribution on this energy mix up to 2050, with an expected part of around 70% of global and ca. 60% of European electricity generation. Coal will remain a key player. Hence, a direct effect on the considered CO2 emissions business-as-usual scenario is expected, forecasting three times the present CO2 concentration values up to 1,200ppm by the end of this century. Kyoto protocol was the first approach to take global responsibility onto CO2 emissions monitoring and cap targets by 2012 with reference to 1990. Some of principal CO2 emitters did not ratify the reduction targets. Although USA and China spur are taking its own actions and parallel reduction measures. More efficient combustion processes comprising less fuel consuming, a significant contribution from the electricity generation sector to a CO2 dwindling concentration levels, might not be sufficient. Carbon Capture and Storage (CCS) technologies have started to gain more importance from the beginning of the decade, with research and funds coming out to drive its come in useful. After first researching projects and initial scale testing, three principal capture processes came out available today with first figures showing up to 90% CO2 removal by its standard applications in coal fired power stations. Regarding last part of CO2 reduction chain, two options could be considered worthy, reusing (EOR & EGR) and storage. The study evaluates the state of the CO2 capture technology development, availability and investment cost of the different technologies, with few operation cost analysis possible at the time. Main findings and the abatement potential for coal applications are presented. DOE, NETL, MIT, European universities and research institutions, key technology enterprises and utilities, and key technology suppliers are the main sources of this study. A vision of the technology deployment is presented.
Oviedo ICCS&T 2011. Extended Abstract
A review and evaluation of the global initiatives carry out trough demonstration projects to aim CCS commercially available by 2020 – 2030. State and improvements of the different programs on going, the UK program, the European program EEPR and the like are part of the scenario analysis. Keywords: Carbon Capture & Storage (CCS), CO2, Clean Coal Technologies, demonstration projects. 1. Introduction A few main certainties currently well known have to be mentioned as starting point for further discussion: CO2 minimum concentration in the atmosphere is necessary for life existence on Earth; [1] with reference before industrial revolution of the 18th – 19th centuries, CO2 global emissions scaled up from around 280ppm to 450ppm values presented by 2005, with a forecast to 750ppm by 2050 and expecting to reach 1,200ppm by the end of this century; [2] Climate Change is taking place and human activity is directly related to it, with anthropogenic component of global warming and climate changes directly affecting species and natural biological systems. Current world population counts around 6.93 billion people [3] and is forecast to 50% increase by 2050. The world´s real income [4], valid approach concerning world GDP (Gross Domestic Product) and standard of living trend, apparently it has grown by 87% over the past 20 years, even the financial crisis faced in 2008. At a global level more people with more income means that the consumption and production of energy will raise. The rise in standard of living indicators and world real income observed is driven by the Non-OECD countries, mainly China, [5] with developing economies pushing up production, demanding infrastructures, hospitals, more comfortable homes, commercial services and the like. World energy consumption [5] is expected to move up more than 45% from 2007 to 2035, with electricity generation use accounting for 57% of the growth to 2030. According to the IEA global electricity will nearly double from 2005 to 2030, stating that fossil fuels will comprise about 70% of global and 60% of European electricity generation. Coal will remain a key player accounting today for about 27% of the global energy consumption, and provides the largest share of world electricity generation accounting for 42% in 2007 [5], it is about 7 trillion kWh. Currently in Europe some 30% of European electricity is generated from coal, it is 3,358 TWh for the EU27 [6]. According to the IEA forecast, coal consumption increases by 1.6% per annum on average from 2007 to 2035, but most of the growth in demand will occur after 2020. Only between 2002 and 2007 worldwide coal consumption increased by 35%, largely
Oviedo ICCS&T 2011. Extended Abstract
because of the growth in China´s coal use. China alone accounts for 78% of the net increase in world coal consumption, whereas India and the rest of non-OECD Asia account for 17% of the world increase. Strong growth on the use of coal, specially by the non-OECD countries, translates into continuous increase of CO2 global emissions. Values between 0.9 and 0.6 tCO2/MWh are released depending on the coal type and technology. Based on IEA tables, currently global emission figure is about 27 Gt CO2, and WEO 2009 [7] forecasts in its reference scenario that world emissions will reach 40.2 Gt CO2 by 2030.
2. Scenario and evaluation study of CCS deployment Application of Carbon Capture and Storage Technologies stands out from three promising measures set to contain global CO2 emissions under necessary and objective limits. To increase the use of renewable energy and to improve energy efficiency are the two other options for reducing greenhouse gas emissions. Carbon Capture and Storage (CCS) concept comprises different technologies for CO2 capture, transport and storage, to attend the three stages of the whole chain of CO2 removal from fossil fuel combustion, particularly applicable to coal fueled power stations. Potential impact of CCS is estimated in 2030 between 2 and 4 Gt/y of CO2 abatement globally [8]. For Europe it is forecast at 0.4 Gt/y, which is around 20% of the total European abatement potential. Post-combustion, Pre-combustion and Oxy-combustion are three ways to remove CO2 from coal fired power stations concerning capture stage, and they depend on where and when the capture process is located referred to CO2 formation. Post-combustion refers to the installation of capture units that process flue gas after coal is combusted. Pre-combustion capture is applicable to coal gasification power stations (normally termed IGCC or Integrated Gasification in Combined Cycle), in which coal is partially oxidised. The capture unit to process syngas can be installed after a shift reaction occurs to raise H2 and CO2 concentration, then CO2 is removed and syngas comprising mostly H2 will be burnt. Oxy-combustion implies the combustion of coal in oxygen as high-purity oxidant stream, hence in a nitrogen depleted atmosphere; flue gas is highly concentrated in CO2 which is then removed by dehydration and compression cycles. Considering post-combustion processes, chemical absorption stands like the most promising technology to be implemented, with large experience through many years deploying in refineries and fertilizer processing facilities. Its application to a 400 MWe coal fired power station tackles with relatively low CO2 concentration levels of about 12 – 14%, and with very high flue gas stream of about 1.3 million Nm3/h at atmospheric
Oviedo ICCS&T 2011. Extended Abstract
pressure. In the short term chilled ammonia and amine base solvents will result the most promising solvents to be implemented in a regeneration process comprising absorption and regeneration phases. Capture efficiency rates over 90% would be achieved. An influence on the overall efficiency (efficiency loss) of the coal fired power station is expected, mainly due to regeneration phase of the capture process and chemical composition of the solvent. Alternative adsorption processes are under research but CO2 capture efficiencies achieved at the time, by testing different adsorbents based on very porous materials under particular conditions, are still low compared to chemical solvents. Regarding pre-combustion, there is extensive experience in natural gas and synthesis gas treating industry, with available technology to process similar flows to coal gasification power stations or IGCC, it is about 200,000 Nm3/h. Main features of the application of this technology to IGCC are syngas entering unit capture process at an inlet pressure between 12 – 30 bar, with CO2 concentration levels between 30 – 40 % after shift reaction and variable presence of sulphur compounds. Capture efficiency rates over 90% would be achieved, by using available licensed processes based on chemical and physical absorption. Moreover, implementation of non-selective processes removing CO2 and sulphur compounds (H2S and COS) simultaneously would be possible. Regeneration phase of the chemical or physical solvent is also energy consuming and influences on the overall efficiency of IGCC coal power station. Alternative capture processes comprising physical adsorption and membranes are being researched, but still are not ready to compete with chemical processes in the short term. Technology associated with Oxy-combustion processes has grown and improved rapidly during the last five years. Oxy-coal combustion power station will require, but it will not be limited to, the following units: Coal Storage, ASU (Air Separation Unit), O2 Storage, Oxy-combustion boiler, Flue gas recirculation system, Flue gas cleaning unit (SO2 removal and ESP principally) and CPU (Compression and Purification Unit). These required units will have impact on the overall efficiency of the power station, depending on scale of integration of the different units that will be achieved. For instance oxygen produced [9] by ASU implies a specific energy consumption of 180 kWh/tO2 and about 160 kWh/tO2 would be achievable with heat integration of the unit. Separation energy values [10] around 140 kWh/tO2 are expected with heat integration by 2015. Additional efficiency penalty of the oxy-coal power station will come from energy consumed by the CO2 compression process; its integration would be study. Oxycombustion processes will draw oxygen concentration levels in stream varying from 27 to 75%. The application of oxy-combustion in coal fired power stations implies recirculation of flue gas to oxy-boiler to reach CO2 concentration levels of ca.90%.
Oviedo ICCS&T 2011. Extended Abstract
Then the flue gas is compressed into the CPU, where basically water condenses and CO2 is compressed up to 90 bar at 25°C (variable, depending on integration strategy), making it ready to put on to transport step. The next Table 1 shows possible application of CO2 capture processes and technologies to different coal fired power station configurations according to the systems of CO2 capture presented before. Table 1. Possible CO2 Capture Technologies Applicable PRE-COMBUSTION
POST-COMBUSTION
OXY-COMBUSTION
ABQ – ADF
ASU + CPU
WGS + (Selexol, Rectisol,
Amines / Chilled Amonia /
Oxy-mode /
Purisol) / WGS + Sulfinol /
Carbonation+Calcination
Carbonation+Calcination
Amines
Cycle
Cycle
ABQ – ABF – ABFQ – ADF
IGCC
PC / SC PC / USC PC / CFB
Abbreviations: i) Processes: ABQ: Spanish term for Chemical Absorption; ABF: Spanish term for Physical Absorption; ABFQ: Spanish term for Chemical-Physical Absorption; ADF: Spanish term for Physical Adsorption ii) Technologies: WGS: Water Gas Sift Reaction iii)Coal fired power station configuration: PC: Pulverised Coal; SC: Supercritical (250-300 bar & 600°C); USC: Ultra Supercritical (350-375 & >700°C); CFB: Circulating Fluidised Bed
It is important to evaluate the impact of implementing CO2 capture processes on the overall energy efficiency of coal fired power stations. Next Table 2 presents a comparison of expected energy efficiency % for different power station configuration with (w) and without (w/o) CO2 capture implementation, made by data evaluation of several reports [9, 10, 11, 12, 13].
Oxy SC PC Oxy USC PC
Oxy CFB
Oxy PC
SC PC (w) USC PC (w/o) USC PC (w)
SC PC (w/o)
IGCC (w/o) IGCC (w)
CFB (w)
CFB (w/o)
PC (w)
PC (w/o)
Table 2. Evaluation of %Efficiency for Coal Fired Power Stations
36– 26– 36– 25– 42 – 32– 38 – 28 – 43 - 33 - 26- 25 - 30- 3238 28 40 28 44 34 40 29 45 34 28 29 34 38 Notes: Number range in %; for abbreviations meaning please referred to Table 1
Oviedo ICCS&T 2011. Extended Abstract
CO2 Transport might not represent a big issue within the CCS deployment, apart from infrastructure investment costs, adequate monitoring during transport and availability of clear and efficient regulation. Transportation of carbon dioxide can be by tanks, onto ships or trains, and by pipe lines depending on the CCS strategy, integrated project definition and cost analysis. [11] Pipelines for transporting nearly 30 million tCO2/y for EOR are available in some regions of the United States. Storage of carbon dioxide is a crucial link of the chain to make CCS feasible. Many official studies have been carried out in order to evaluate necessary removal capacity of the total CO2 amount, which results from the potential application of CO2 capture technologies [14]. Due to the large amount of CO2 expected to be necessarily removed, only little portion would be used by industrial sector (food / beverage industry and fertiliser production). Storage appears the only option short term to reduce the huge amount of CO2 emitted. Three ways of CO2 storage are considered: geological storage, ocean storage and mineral carbonation [11]. Only geological storage would be applicable in the short term. The options for geological storage are sequestering into deep saline aquifers and EOR (Enhance Oil Recovery) or EGR (Enhance Gas Recovery), which consists in pumping CO2 at particular conditions into depleted oil or gas fields to improve fossil fuel extraction. World storage required in 2050 will be about 144.7 GtCO2 [14]. Potential viable capacity of 1,680 GtCO2. Only for OECD Europe this potential is 94 GtCO2.
3. Results and Discussion According to the vision provided by the IEA CCS Technology Roadmap 2009, some 100 commercial-scale CCS projects must be operational worldwide by 2020 and 3,400 by 2050 if global warming is to stay below 2°C [15]. The principal metric used to define commercial scale integrated projects are those with a storage capacity rate of 1 million tCO2/y or greater [16]. Achieving 0.4 GtCO2 abatement per year from CCS in Europe by 2030, would require the installation of between 80 and 120 commercial scale CCS projects [8]. It is likely to develop as a series of capture clusters, all connected into a common transport and storage network. To reach this range of power stations in operation would be necessary to evaluate and prove the technical and economical feasibility of integration projects comprising capture, transport and storage of carbon dioxide, hence in a previous demonstrations scale.
Oviedo ICCS&T 2011. Extended Abstract
For European market, studies shows that first demonstration projects will have a CO2 reduction cost between 60 – 90 €/tCO2, considering capture, transport and storage and assuming similar costs for capture process selected [8]. First commercial scale projects would have around 35 – 50 €/tCO2, forecast 30 – 45 €/tCO2 achievable by 2030. Distribution of the total will be 64 – 72% for capture, 11 – 12% in transport and 11 – 24% in storage. Construction of new built coal fired power stations implementing CO2 capture technologies will bring better results than retrofit existing ones [8], mainly because of high efficiencies reached by the implementation of new technology and units. In the year 2010 there are globally 238 active or planed CCS projects [16], with 151 projects integrated. From the total, up to 80 are large-scale integrated projects, presenting the following relation by technology implementing: 33 Pre-combustion, 22 Post-combustion, 13 Gas processing, 2 Oxyfuel combustion, 3 Pre-combustion and Post-combustion, 1 Pre-combustion and Gas processing, 1 Oxyfuel combustion and 5 not specified. On 9 December 2009, the European Commission announced details of the 6 CCS demonstration scale projects (around 300 MWe) which receive funding of 1 billion € under the EEPR [17]. The origin of EEPR (European Energy Program for Recovery) is the global 200 billion € European Economic Recovery Plan presented by the Commission at the end of 2008. The six projects selected are Belchatow (Postcombustion, Poland), Compostilla (Oxyfuel, Spain), Hatfield (Pre-combustion, UK, location/project to be confirmed), Jänschwalde (Oxy + Post-combustion, Germany), Porto
Tolle
(Post-combustion,
Italy)
and
Rotterdam
(Post-combustion,
The
Netherlands). Final Investment Decision for the construction of these projects is scheduled by middle 2012, bearing further funding under the NER300 (New Entrant Reserve), in principal only for selected ones. NER makes funding available for commercial-scale CCS projects, with the funds generated through the sale of 300 million EU ETS allowances for the New Entrant Reserve of Phase 3 of the EU ETS. The European Commission estimates that the sale of these allowances will raise between 15 – 30 €/tCO2, dependent on the carbon price. The UK Demonstration Programme, managed by the OCCS (Office Of Carbon Capture and Storage), part of the UK Government´s DECC (Department of Energy and Climate Change), will fund 4 CCS commercial-scale projects, involving coal and gas fuel, with up to £1b to support the capital cost of the first one ongoing [18]. Main features of the UK Demonstration Programme are: alignment with NER300 schedule and compatibility with other funding programme; projects will receive a fixed strike price per tCO2 abated, and therefore the fund received will be strike price minus EAUs (EU ETS);
Oviedo ICCS&T 2011. Extended Abstract
requirements and specification of CO2 transport will be proposed by the Project Proposer; no onshore storage projects will be funded; only will be funded offshore UK storage in compliance with storage terms of the Energy Act 2008, Storage Carbon Dioxide Regulation 2010 and the EU Directive 2009/31/CE. US Power generation sector produced more than 40% of total US anthropogenic CO2 emissions in 2008, and the majority result from the combustion of coal, about 1.9 billion tCO2 [11]. EPRI and other recent studies result that wide-scale deployment of CCS provides the largest share of potential CO2 reduction. The CCPI (Clean Coal Power Initiative) will begin to demonstrate, by 2015, commercial-scale capture and storage or beneficial reuse technologies that target to achieve 90% capture efficiency for CO2 to enable subsequent commercial deployment in the coal fired utility industry. Under the CPPI Programme [11], 7 CO2 capture demonstration projects in USA are planed and ongoing. 3 of them Pre-combustion technology related, another 3 of them Post-combustion related and only 1 will demonstrate Oxy-combustion deployment. Within the USA framework, the Recovery Act funding is being used for the following CCS related activities [11]: CCPI with a total of $800 million; Industrial Carbon Capture & Storage with a total of $1.5 billion; around $20 million for scale-up a current project; a total of $100 million is being used to characterize about 10 geological formations; a total of $20 million is being used in education related to CCS sector; FutreGen 2.0 with a total of $1 billion for the construction of Oxy-combustion power station to capture 1 million tCO2/y since 2015; Carbon Capture and Storage Simulation Initiative with a total of $40 million is being used to accelerate CCS technology development using advanced simulation and modeling techniques; and the National Risk Assessment Partnership to develop the tools and science base for ensuring longterm storage. In China, the GreenGen Project ongoing will demonstrate the feasibility of 250 MW IGCC in 2009, scale up to 400 MW and 25% CO2 captured by 2015. According to OECD/IEA currently exist more than 20.000 km of CO2 pipeline globally, and forecasts the total CO2 pipeline needs over 200.000 km for the period 2030 – 2050 contingent on the level of optimisation in building common carriage networks able to link multiple sources and storage sites. Currently, a pan-European project called GeoCapacity is underway in order to provide a comprehensive database of European CO2 storage availability.
4. Conclusions Coal reserves are widespread and the international coal market ensures that demand is largely met from the most economic suppliers. Coal will represent, at least, a steady
Oviedo ICCS&T 2011. Extended Abstract
share of the World Primary Energy demand over the next 20 years and the future contribution to power generation relies mostly on how coal's carbon intensity can be coupled with sustainable emission levels. CCS and clean coal technologies are paving the way to meet the most demanding greenhouse-gas emission targets and the design of the future coal power stations will be more capital intensive and more sustainable.
Acknowledgment Special thanks to Javier Pisa for his outstanding support.
References [1] IPCC – CO2 equivalent - IPCC data base [2] IPCC 4Assessment Report 2007 [3] U.S. Census Bureau [4] BP 2011, Energy Outlook 2030 [5] International Energy Outlook 2010 [6] The Role of Coal Power Generation 2008, EURACOAL statistics, official website [7] World Energy Outlook 2009 [8] McKinsey&Company 2008 – Carbon Capture & Storage: Assessing the Economics [9] Air Liquide Technical Papers 2010 – Air Separation Unit for Oxy-coal Combustion [10] Air Liquide / B&W Technical Presentations 2010 – Oxy-coal Combustion for Demonstration Plants [11] DOE/NETL Carbon Dioxide Capture and Storage RD&D Roadmap – December 2010 [12] International Energy Agency, Clean Coal Technologies, 2008 [13] MIT, The Future of Coal, 2007 [14] OECD/IEA, Carbon Capture & Storage Roadmap, 2010 [15] CO2 Capture Project, Annual Report 2010, April 2011 [16] Global CCS Institute, The Status of CCS Projects, Interim Report 2010 [17] CCS European Energy Program for Recovery (EEPR), Energy European Commission website 2010 [18] UK Department of Energy and Climate Change, UK CCS Commercial Scale Demonstration Programme, Delivering Projects 2-4, December 2010
Co-combustion of coals and biomass blends in an entrained flow reactor under oxy-fuel atmospheres J. Riaza, L. Álvarez, M.V. Gil, C. Pevida, F. Rubiera, J.J. Pis
Instituto Nacional del Carbón, INCAR-CSIC, Apartado 73, 33080 Oviedo, Spain [email protected] Abstract Co-combustion of different coals and biomass blends were carried out in an entrained flow reactor under different O2/CO2 atmospheres (21%O2/79%CO2, 30%O2/70% O2 and 35%O2/65%CO2), and the results obtained were compared with those attained in air. Coal burnout for the 21%O2/79%CO2 atmosphere were lower than in air. However, for O2/CO2 conditions when the O2 concentration was increased to 30 and 35%, an improvement on coal burnout was observed. The effect of blending biomass with coals of different rank under oxy-fuel atmospheres resulted in higher burnout for all cases. A reduction of NO emissions was observed when biomass was mixed with coal.
1. Introduction Carbon dioxide capture and storage technologies are necessary to produce energy from fossil fuels in a friendly environmental way. Oxy-fuel combustion is one of the most promising CO2 capture technologies as it could be adapted to both new and existing pulverized coal-fired power stations, and a significant and simultaneous reduction on NOX and CO2 emissions can be achieved. During oxy-fuel combustion a mixture of O2 and recycled flue gas (mainly CO2 and H2O) is used for fuel combustion. Due to higher O2 concentration and differences in gas properties between N2 and CO2, oxy-fuel combustion differs greatly from air combustion in several ways, inclufing coal burnout and pollutant formation [1]. Another approach for reducing CO2 emissions is the use of renewable fuels such as biomass, that is considered a neutral carbon fuel because the CO2 released during its utilisation is an integral part of the carbon cycle. Co-firing coal with biomass offers other environmental advantages due to the reduction of sulphur oxides (SOX), and the potential reduction of nitric oxides (NOX) [2]. The combination of oxy-fuel combustion with biomass could be used as a sink for CO2 emissions. The aim of this work is to asses the effect of cofiring biomass with coal in both air and oxy-fuel conditions.
1
2. Experimental Two coals of different rank were used: a semi-anthracite (HVN) and a high-volatile bituminous coal (SAB). Also two different biomasses were used: olive waste (OR) and pine straw (PI). The samples were ground and sieved to obtain a particle size fraction of 75-150 μm. The results of the proximate and ultimate analyses, and high heating value of the samples are shown in Table 1 . Table1. Proximate and ultimate analyses and high heating values of the samples Sample HVN SAB OR PI
Rank sa hvb -
Proximate Analysis (wt%, db)
Ultimate Analysis wt%, daf)
HHV
Ash
V.M
F.C.a
C
H
N
S
Oa
(MJ/kg,db)
10.7 15,0 7.6 3,8
9.2 29.9 71.9 79,8
80.1 55.1 20.5 16,4
91.7 81,5 54,3 45,9
3.5 5,0 6,6 6,1
1.9 2,1 1,9 0,7
1.6 0,9 0,2 0,0
1.3 10,5 37,0 42,3
31.8 27.8 19.9 18,9
sa: semi-anthracite; hvb: high-volatile bituminous coal. db: dry basis; daf: dry and ash free bases. a Calculated by difference.
Combustion experiments were carried out in an entrained flow reactor (EFR). This reactor is electrically heated to a temperature of 1000 ºC, and it reproduces some of the characteristics found in pulverised coal combustion such as high heating rates and short residence times. The samples were fed in from a hopper and the mass flow was controlled using a mechanical feeding system. The samples were introduced through an air-cooled injector to ensure that their temperature did not exceed 100 ºC before entering the reaction zone. The gases were preheated before being introduced into the reactor through flow straighteners. The flow rates of N2, CO2 and O2 were set to ensure a particle residence time of 2.5 seconds. A water-cooled collecting probe was inserted into the reaction chamber from below. Nitrogen was introduced at the top of this probe to quench the reaction products. Particles were removed by a cyclone and coal burnout was calculated by the ash tracer method. The exhaust gases were monitored using a battery of analysers (O2, CO, CO2, NO, N2O, SO2). A schematic diagram of the entrained flow reactor (EFR) used in the experiments can be seen in Fig. 1.
2
Figure 1. Schematic diagram of the entrained flow reactor (EFR). Four binary mixtures of O2, N2 and CO2 were employed to study the behaviour of the coals and blends. Thus, for the combustion tests, air (21%O2/79%N2) was taken as reference and three binary mixtures of O2 and CO2 were compared: 21%O2/79%CO2, 30%O2/70% CO2 and 35%O2/65%CO2.
3. Results and Discussion Coals and biomass blends were burnt under different levels of oxygen excess for each atmosphere studied. The fuel ratio, defined as the ratio between the coal mass flow rate and the stoichiometric value, was used to asses the excess of oxygen during combustion. The burnout values attained for coals SAB and HVN at different fuel ratios are shown in Fig. 2. It can be seen that the coal burnout decreased as the fuel ratio increased, due to the lower availability of oxygen at higher fuel ratio values. For both coals, the burnout obtained under the 21%O2-79%CO2 atmosphere was lower than that reached under 21%O2-79%N2 conditions. CO2 has a higher specific molar heat than N2, which implies that when N2 is replaced by CO2 the heat capacity of the gases increases, leading to lower flame and gas temperatures. Therefore, the particle temperature during
3
the 21%O2-79%CO2 atmosphere can be expected to be lower, causing the combustion rate of the char and the coal burnout value to fall [3].
100
100
HVN
SAB
90 Burnout (%)
Burnout (%)
90
80 70 21%O2/79%N2 21%O2/79%CO2
60
80 70 21%O2-79%N2 21%O2-79%CO2 30%O2-70%CO2 35%O2-65%CO2
60
30%O2/70%CO2 35%O2/65%CO2
50
50 0.2
0.4
0.6
0.8
1.0
1.2
1.4
0.2
0.4
Fuel ratio
0.6
0.8
1.0
1.2
1.4
Fuel Ratio
Figure 2. Burnout of coals HVN and SAB under different atmospheres at different fuel ratios. Under the 30%O2-70%CO2 and 35%O2-65%CO2 atmospheres, the burnout for both coals was higher than in air, since the higher oxygen concentration produced an increase in the char combustion rate. Increasing the O2 fraction in CO2 up to 30% is still insufficient to match the specific heat capacity of air. However, coal burnout in the 30%O2-70%CO2 atmosphere reached a higher value than in air, which means that another parameter must have changed to offset the negative effect of the specific heat capacity of the gas. Though the gas temperature increases only slightly when the O2 fraction in bulk gas is increased, it is likely that the increase in the mass flow rate of O2 from the bulk gas to the coal surface at higher O2 concentrations promotes the consumption rate of the volatiles, providing extra heat feedback to the coal particle to enhance its devolatilisation, ignition and combustion [4]. For the co-combustion tests, HVN-OR blends were used and their burnouts were also determined at different fuel ratios. Fig. 3 shows the burnout of the HVN-OR blends at a selected fuel ratio of 0.8. The burnout of the blends presents a behaviour similar to that of the individuals coals under the different atmospheres studied, that is, there is a decrease in burnout when N2 is replaced by CO2 for the same oxygen concentration, and an improvement in O2/CO2 atmospheres when the oxygen concentration increases.
4
100
HVN-OR
Burnout (%)
90
80
70
81.7 82.4
79.5
83.5 83.5 80.4 79.7
86.1 85.5 83.0
81.0
77.2
21%O2-79%N2 21%O2-79%CO2
60
30%O2-70%CO2 35%O2-65%CO2
50
HVN
90% HVN-10% OR 80% HVN-20% OR
Figure 3. Burnout of HVN-OR blends at a fuel ratio of 0.8. As shown in Fig. 3, blending biomass with coal has an impact on burnout. For all the atmospheres studied, an improvement on coal burnout was achieved with the increase in the percentage of biomass. The highest improvement on burnout was observed for the combustion of HVN-OR in 30%O2/70%CO2. When the oxygen concentration increased to 35%, no significant differences were observed in comparison with combustion under 30%O2. The concentrations of NO at a fuel ratio of 0.8 during HVN-OR combustion are shown in Fig. 3. The NO concentration (mg NO/ g burned coal) obtained under the 21%O2/79%CO2 atmosphere was lower than that achieved under 21%O2/79%N2. This reduction on NO emissions can be explained due to the higher CO concentrations in oxy-fuel environments [5]. In the 30%O2/70%CO2 and 35%O2/65%CO2 atmospheres, the NO concentration was slightly higher than that in the oxy-fuel atmosphere containing 21% of O2, since higher oxygen concentrations enhance coal nitrogen conversion to NO. However, the NO concentrations still remained lower than in air, and small differences between the three oxy-fuel atmospheres were observed. As seen in Fig. 4, a decrease in NO emissions was observed with biomass blending. This decrease is significantly marked for air-firing conditions.
5
12 HVN-OR mgNO/g burned fuel
10
9.7 8.8
8
6.7 6.9
7.4
7.1 6.5
6.5
6.6
6.1
6.6 6.5
6 4 2
79%N2-21%O2 79%CO2-21%O2 70%CO2-30%O2 65%CO2-35%O2
0
HVN
90% HVN-10% OR 80% HVN-20% OR
Figure 4. NO emissions of HVN-OR blends at a fuel ratio of 0.8.
4. Conclusions The burnout of individual coals and blends was determined under both air and oxyfiring conditions. The burnout of the samples in a 21%O2/79%CO2 atmosphere were lower than those obtained in air-firing conditions due to the higher specific molar heat of CO2 compared to N2, and the lower diffusivity of O2 in CO2 than in N2. However, when the O2 concentration was increased to 30 and 35%, the burnouts increased above those reached under air-firing, due to an increase in the mass flow of O2 to the coal particles. When blending biomass, an improvement on burnout was observed for both air and oxy-firing conditions. NO emissions were lower during oxy-fuel combustion than in air-firing due to a higher reduction of NO by reaction with CO. A slight increase in NO emissions was observed as the O2 concentration increased in O2/CO2 atmospheres. A decrease on NO emissions was observed for both air and oxy-firing conditions when the percentage of biomass in the blends was increased.
Acknowledgements Work carried out with financial support from the Spanish MICINN (Project PS120000-2005-2) co-financed by the European Regional Development Fund. L.A. and J.R. acknowledge funding from the CSIC JAE program, co-financed by the European
6
Social Fund, and the Asturias Regional Government (Severo Ochoa Program), respectively.
References [1] Wall TF, Liu Y, Spero C, Elliot L, Khare S, Rathman R et al. An overview on oxyfuel coal combustion-state of the art research and technology development. Chemical Engineering Research and Design 2009; 87: 1003-16. [2] Demirbas A. Potential applications of renewable energy sources, biomass combustion problems in boiler power systems and combustion related environmental issues. Progress in Energy Combustion Science 2005; 31: 171-92. [3] Berejano PA, Levendis Y. Single-coal particle combustion in O2/N2 and O2/CO2 environments. Combustion and Flame 2008; 153: 270-87. [4] Shaddix CR, Molina A. Particle imaging of ignition and devolatilization of pulverized coal during oxy-fuel combustion. Proceedings of the Combustion Institute 2009; 32: 2091-8. [5] Hu YQ, Kobayashi N, Hasatani M. Effects of coal properties on recycled-NOX reduction in coal combustion with O2/recycled flue gas. Energy Conversion and Management 2003; 44: 2331-40.
7
Effect of the activation temperature and the burn-off degree on the CO2 capture capacity of microporous activated carbons M.V. Gil, M. Martínez, S. García, J.J. Pis, F. Rubiera, C. Pevida Instituto Nacional del Carbón, INCAR-CSIC. Apartado 73. 33080 Oviedo, Spain [email protected] Abstract Phenol-formaldehyde resins and a low-cost biomass residue, olive stones (OS), were used to prepare two activated carbons for the separation of CO2 in post-combustion processes. Two phenol-formaldehyde resins were synthesized: Resol, obtained by using an alkaline environment, and Novolac, synthesized in the presence of an acid catalyst. Carbon precursors were prepared by mixing the Resol resin, R, with potassium chloride, K, and the Novolac resin, Cl, with olive stones, OS. The precursors were carbonized under an inert atmosphere of N2 at 1000 ºC, yielding the RKC10 and ClOSC10 materials. The last stage in the synthesis of the adsorbents involved physical activation with carbon dioxide. Response surface methodology (RSM) was used as a tool for rapidly optimizing the activation parameters in order to obtain the highest CO2 capture capacity of the activated carbons. By means of this methodology it was possible to obtain the optimum values of temperature and burn-off degree during the activation step to maximize CO2 uptake by the activated carbons, within the experimental region considered. The maximum value of CO2 capture capacity for ClOSC10 was obtained when the activation was carried out at 942 ºC, independently of the burn-off value. For RKC10, the maximum value of CO2 capture capacity was obtained with an activation temperature of 722 ºC and a burn-off of 50%. Values of CO2 adsorption capacity of 4.4 and 7.3 wt.% at 35 °C and atmospheric pressure were achieved for the RKC10 and ClOSC10 activated carbons, respectively.
1. Introduction Coal is the most abundant and widely geographically distributed fossil fuel. The stability of its supply and relatively low cost ensure its inclusion in the energy mix in the foreseeable future. The use of coal in power plants generates high amounts of CO2. It is widely accepted that climate change is a global phenomenon influenced by greenhouse gas emissions to the atmosphere, CO2 being the main greenhouse gas contributing to global warming. In the short-to-medium term, carbon capture and storage (CCS) will be
1
necessary in order to reduce CO2 emissions to the atmosphere, CO2 capture being the most costly component of the CCS process [1]. This has led to intensive research aimed at the production of CO2 capture materials with significant levels of CO2 uptake. Different types of adsorbents, such as zeolites and activated carbons, have been used for this purpose. Activated carbons have a high adsorption capacity at ambient pressures and present important advantages over zeolites, such as their hydrophobicity, their significant lower cost and the lower amount of energy needed to regenerate them. Adsorption with activated carbons at atmospheric pressure could be a useful technology for post-combustion CO2 capture. These materials can be obtained from almost any carbonaceous product by a process of carbonization followed by an activation step. However, the use of naturally occurring precursors to produce activated carbons limits the purity, strength and physical form of the end-product materials. This drawback can be overcome by using polymeric precursors, where the reproducibility and purity of the precursor is within the control of the manufacturer, and the physical forms and structures can be tailored by means of the polymer production process [2]. Phenolic resins constitute a family of low-cost polymers, one of the most common being those produced from phenol and formaldehyde [3]. Response surface methodology (RSM) is a multivariate statistical technique used to optimize processes, i.e., to establish the conditions in which to apply a procedure in order to obtain the best possible response in the experimental region studied. This methodology involves the design of experiments and multiple regression analysis as tools to assess the effects of two or more independent variables on dependent variables [4]. It is based on the fit of a polynomial equation to the experimental data to describe the behaviour of a set of data. In the present work, phenol-formaldehyde resins and olive stones were employed as precursor materials for the preparation of microporous activated carbons for use in post-combustion CO2 capture. The CO2 capture capacity of the different activated carbons was optimized in relation to temperature and burn-off degree during the activation stage by means of response surface methodology. The objective of this study was to determine the optimum values of activation temperature and burn-off degree for the activated carbons, which maximize the CO2 capture capacity within a given experimental region.
2
2. Experimental Phenol-formaldehyde resins and a low-cost biomass residue, olive stones (OS), were used as starting materials. Two types of phenol-formaldehyde resins were synthesized. The first one was obtained by basic catalysis using sodium hydroxide (NaOH) and is commonly referred to as Resol. In this case a 2.5:1 formaldehyde-tophenol ratio was used. The second type of resin was synthesized by acid catalysis with hydrochloric acid (HCl) and is known as Novolac. In this case a 1:1.22 formaldehyde-tophenol ratio was used. Then, the resins were cured in a rotary evaporator (40-70 ºC) and a forced-air convection oven (60-100 ºC) and, finally, the cured resins and the olive stones were ground and sieved to obtain a particle size fraction of 1.0-3.35 mm. Two carbon precursors were then prepared by incorporating potassium chloride to the Resol resin, and olives stones to the Novolac resin (80:20 wt. ratio of OS:resin). The precursors were then carbonized in a horizontal furnace under a nitrogen flow at 1000 ºC, which yielded the RKC10 and ClOSC10 carbonized samples. The carbonized materials were physically activated with CO2 in a thermobalance under a 10 mL min-1 flow rate of CO2 at different temperatures. The RSM was used to evaluate the effect of temperature and burn-off degree during the activation stage on the CO2 capture capacity of the activated carbons. The independent variables were activation temperature (T) and burn-off degree attained after the activation process (B), while the dependent variable was the CO2 capture capacity. For ClOSC10, the activation temperature was studied between 900 and 1000 ºC and the burn-off degree between 30 and 50%. For RKC10, the activation temperature was assessed between 600 and 800 ºC and the burn-off degree between 10 and 50%. A three-level full factorial experimental design was used and 13 experiments were carried out, which are shown in Table 1, together with the experimental values of CO2 capture capacity. The mathematical-statistical treatment of the experimental data consisted in fitting a polynomial function to the set of data: y = β0 + β1x1 + β2x2 + β12x1x2 + β11x1x1 + β22x2x2 + ε
(1)
where β0 is the constant term, β1 and β2 represent the coefficients of the linear parameters, β12 represents the coefficient of the interaction parameter, β11 and β22 represent the coefficients of the quadratic parameters and ε is the residual associated with the experiments. Multiple regression analysis was used to fit Eq. (1) to the
3
experimental data by means of the least-squares method, which makes it possible to determine the β coefficients that generate the lowest possible residual. Table 1. Levels of the independent variables (coded levels in parentheses), activation temperature (T) and burn-off degree (B) for the activated carbons, using a three-level full factorial design, and experimental values of CO2 capture capacity Run RKC10 ClOSC10 CO2 capture CO2 capture T (ºC) B (%) T (ºC) B (%) capacity (wt.%) capacity (wt.%) 1 600 (-1) 10 (-1) 0.8 900 (-1) 30 (-1) 6.8 2 600 (-1) 30 (0) 0.8 900 (-1) 40 (0) 7.0 3 600 (-1) 50 (+1) 0.7 900 (-1) 50 (+1) 7.0 4 700 (0) 10 (-1) 2.5 950 (0) 30 (-1) 7.1 5 700 (0) 30 (0) 3.8 950 (0) 40 (0) 7.5 6 700 (0) 50 (+1) 5.0 950 (0) 50 (+1) 7.5 7 800 (+1) 10 (-1) 1.1 1000 (+1) 30 (-1) 6.5 8 800 (+1) 30 (0) 3.2 1000 (+1) 40 (0) 6.7 9 800 (+1) 50 (+1) 3.1 1000 (+1) 50 (+1) 6.6 10 700 (0) 30 (0) 3.7 950 (0) 40 (0) 7.2 11 700 (0) 30 (0) 3.0 950 (0) 40 (0) 7.1 12 700 (0) 30 (0) 3.6 950 (0) 40 (0) 7.3 13 700 (0) 30 (0) 3.2 950 (0) 40 (0) 7.5
Evaluation of the fitness of the models was carried out by applying an analysis of variance (ANOVA) and a lack of fit test. The coefficient of determination adjusted by the number of variables (Adj-R2) and the absolute average deviation (AAD) were calculated in order to check the accuracy of the model. Adj-R2 represents the proportion of variability of the data that is accounted for by the model. The AAD is a direct parameter that describes the deviations between the experimental and calculated values and it is calculated by means of the following equation [5]: AAD (%) = 100 [Σi=1 n (|yi,exp – yi,cal|/yi,exp)]/n
(2)
where yi,exp and yi,cal are the experimental and calculated responses, respectively, and n is the number of experiments. The statistical analyses were carried out by SPSS Statistics 17.0 software. The model obtained can be three-dimensionally represented as a surface (response surface plot) and the best operational conditions inside the studied experimental region can be found by visual inspection. A two-dimensional display of the surface plot generates the contour plot, where the lines of constant response are drawn on the plane of the independent variables. Response surface and contour plots were generated using the software SigmaPlot 8.0. The CO2 capture capacity of the adsorbents at atmospheric pressure was assessed
4
in a Mettler Toledo TGA/DSC 1 thermogravimetric analyzer under a CO2 flow rate of 100 mL min-1 at 35 ºC up to constant weight. The maximum CO2 uptake at atmospheric pressure and 35 ºC was evaluated from the increase in mass experienced by the sample, and it was expressed in terms of mass of CO2 per mass of dry adsorbent (wt.%).
3. Results and Discussion Table 2 shows the results of fitting Eq. (1) to the experimental data by multiple regression analysis, and those obtained from evaluating the fitness of the model by means of ANOVA, together with the Adj-R2 and AAD values. The ANOVA tests showed that the models for CO2 capture capacity were statistically significant at a 95% confidence level (p-value<0.05), whereas their lack-of-fit was found to be statistically non-significant at a 95% confidence level (p-value>0.05). Table 2. Results of multiple regression analysis and ANOVA used to to the CO2 capture capacity experimental data of the activated carbons RKC10 ClOSC10 Coded Sum of Coded DF p-value coefficient squares coefficient Intersection 3.579 74.306 1 0.000 7.334 T 0.850 4.335 1 0.007 -0.167 B 0.733 3.227 1 0.014 0.117 TB 0.525 1.103 1 0.100 -0.025 T2 -1.878 9.737 1 0.001 -0.521 B2 -0.128 0.045 1 0.714 -0.071 Model 20.696 5 0.002 Residual 2.156 7 Total 22.852 12 Lack-of-fit 1.684 3 0.083 Pure error 0.472 4 R2 0.906 0.887 Adj-R2 0.838 0.806 AAD (%) 8.12 1.39
fit the polynomial model Sum of squares 312.009 0.167 0.082 0.003 0.749 0.014 1.233 0.157 1.391 0.029 0.128
DF p-value 1 1 1 1 1 1 5 7 12 3 4
0.000 0.030 0.098 0.748 0.001 0.459 0.003 0.822
Table 2 also shows which of the terms in the models were statistically significant at a 95% confidence level (p-value<0.05), and those that were not statistically significant were eliminated. The Adj-R2 and the AAD values were found to be acceptable, between 0.806-0.949 and 1.4-8.1% respectively. Once the non-significant terms were eliminated and the coded coefficient values were decoded, the polynomial models for the response variables as a function of the true independent variables were as follows: CO2 uptake RKCl0 (wt.%) = -97.8905 + 0.2782 T + 0.0367 B – 0.0002 T2 CO2 uptake ClOS10 (wt.%) = -187.2095 + 0.4129 T – 0.0002 T
2
(1) (2)
5
Fig. 1 presents the response surface plots and the contour plots for the CO2 capture capacity as a function of the independent variables, activation temperature and burn-off, for the activated carbons.
Fig. 1. Response surface and contour plots for CO2 capture capacity as a function of the activation temperature and burn-off corresponding to the RKC10 (a) and ClOSC10 (b) activated carbons.
For the RKC10 carbonized material (Fig. 1a), the CO2 capture capacity increased as the degree of burn-off increased over the entire temperature range studied, since no interaction effect between T and B was detected in the experimental region under study (the T·B interaction term was not statistically significant as shown in Table 2). A curvature was also observed, indicating the achievement of a maximum response. In this case, the highest CO2 capture capacity (4.4 wt.%) was obtained with an activation temperature of 722 ºC and a burn-off degree of 50%. For the No2OS-1000 carbonized
6
material (Fig. 1b) the burn-off degree had no effect on CO2 uptake, whereas a maximum response was obtained in relation to the temperature. The highest CO2 capture capacity (7.3 wt.%) was obtained at 942 ºC, irrespective of the burn-off value.
4. Conclusions The response surface methodology was successfully used to determine the optimum activation conditions (temperature and burn-off degree) that maximize CO2 uptake by the activated carbons. For the ClOSC10 activated carbon, the maximum value of CO2 capture capacity was obtained when the activation was carried out at 942 ºC, independently of the burn-off value, whereas for the RKC10 the maximum value of CO2 capture capacity was obtained with an activation temperature equal to 722 ºC and a burn-off value of 50%. The results showed that the activated carbon derived from Novolac phenol-formaldehyde resin type and olive stones, ClOSC10, presented a higher potential as adsorbent for CO2 post-combustion capture processes, with a value of CO2 adsorption capacity of 7.3 wt.% at 35 °C, which is comparable to that of good commercial activated carbons.
Acknowledgements Work carried out with financial support from the Spanish MICINN (Project ENE2008-05087). M.V. Gil acknowledges funding from the CSIC JAE-Doc Program co-financed by the European Social Fund.
References [1] IPCC. IPCC special report on carbon dioxide capture and storage. Cambridge, United Kingdom and New York: IPCC; 2005. [2] Martín CF, Plaza MG, García S, Pis JJ, Rubiera F, Pevida C. 2011. Microporous phenolformaldehyde resin-based adsorbents for pre-combustion CO2 capture. Fuel 2011;90:2064–71. [3] Knop A, Pilato LA. Phenolic resins. Chemistry. Applications and performance. Berlín: Springer-Verlag; 1985. [4] Myers RH, Montgomery DH. Response surface methodology. USA: John Wiley & Sons; 1995. [5] Baş D, Boyacı İH. Modeling and optimization I: Usability of response surface methodology. J Food Eng 2007;78:836-45.
7
INFLUENCE OF LIGHT OVER THE BIOFIXATION OF CO2 BY THE MICROALGAE CHLORELLA MINUTISSIMA Cristiane Redaelli1,2, Rosane Rech1, Nilson Romeu Marcilio2 1
Federal University of Rio Grande do Sul – Institute of Food Science and Technology Porto Alegre – RS (Brazil) - e-mail: [email protected] 2 Federal Univesity of Rio Grande do Sul – Department of Chemical Engineering Porto Alegre– RS (Brazil) – e-mail: [email protected]
ABSTRACT In the present study, development of a high efficiency process for carbon dioxide biofixation through the use of photosynthetic microalgae is proposed. Flat-plate air-lift photobioreactors have been built aiming a good mass transfer coefficient (kLa) and low power consumption. Growth of the microalgae Chlorella minutissima was tested under light intensities between 2,200 Lux and 24,500 Lux. The biomass production, lipid content and carbon dioxide biofixation rate were evaluated. The light intensity of 17,000 Lux presented the best results, reaching 0.38 g.L-1 of dry biomass with total lipid content of 10.36%. At the exponential growth phase, the maximum specific growth rate was 0.61 d-1 and the carbon dioxide biofixation rate was 16 gCO2.m-3.h-1. Keywords: Chlorella minutissima, photobioreactor, microalgae, light intensity
INTRODUCTION The carbon dioxide released by the burning of fossil fuels is the main responsible for the greenhouse effect. In a certain range, the greenhouse effect is vital for life on Earth, but beyond this can cause global warming, a great environmental problem (Borges et al., 2007). Several studies have been developed to promote reduction or sequestration of the carbon dioxide emitted from stationary sources such as coal thermoelectric. Recently, the development of photobioreactors for microalgae cultures has gained interest in the scientific community due to the ability of these microorganisms to use CO2 as the sole carbon source and produce high-value products such as biofuel, food, feed and bioactive compounds (Chisti, 2007). According Teixeira and Morales (2008), microalgae oils have similar chemical and physicochemical characteristics to vegetables oils, having great potential to be used as raw material for biodiesel production. The advantages of microalgae over higher plants as a source of transportation biofuels are numerous: 1) they synthesize and accumulate large quantities of neutral lipids and grow at high rates; 2) oil yield per area of microalgae cultures could greatly exceed the yield of best oilseed crops; 3) they can be cultivated in saline/brackish water/coastal seawater on non-arable land, and do not compete for resources with conventional agriculture; 4) microalgae utilize nitrogen and phosphorus from a variety of wastewater sources; 5) they grow in suitable culture vessels (photobioreactors)
throughout the year with higher annual biomass productivity on an area basis (Khan et al., 2009). While in the past natural waters (lakes, lagoons, ponds) or artificial ponds were used to grow algae, more recently closed photobioreactors have been employed. Openculture systems have almost always been located outdoors and rely on natural light for illumination. Although they are inexpensive to install and run, photobioreactors, on the other hand, have some advantages such as lower evaporation losses, lower risk of contamination and smaller area required (Xu et al. 2009). The most employed designs of photobiorreactors are: flat plate, tubular and bubble column. Still can be used airlift and modifications of the three configurations mentioned above, as is the case of annular bubble column reactor and the reactor flat plate with shaped dome (Posten, 2009). One advantage of photobiorreactors over artificial ponds is the possibility light intensity control of the cultures. Acording Lima and Sato (2001), the light intensity associated to nitrogen concentration have great influence on lipid production by microalgae. In the present work, a flat-plate airlift photobiorreactor was used to evaluate the influence of light intensity over biomass production, lipid content and carbon dioxide biofixation rate by the marine microalgae Chlorella minutissima.
MATERIAL AND METHODS Microalgae and Growth Media The microalgae Chlorella minutissima used in this study was gently provided by Professor Sérgio Lourenço of the Fluminense Federal University. The cells were maintained in 250 mL Erlenmeyer flasks with 50 mL of f/2 culture medium (Guillard, 1975) in germination camera under controlled temperature (20 °C), with a photoperiod of 12/12 hours (light/dark), provided by fluorescent lamp light. Photobiorreactor and Growth Conditions The cells were grown in flat-plate air-lift photobioreactors with an internal heatexchanger (FPA-IHE) with 2.2 L of f/2 culture medium. The FPA-IHE was designed and characterized by the own research group and is showed in Figure 1 (Kochem, 2010). The photobioreactors were filled with water and sterilized by adding 10 mL of commercial solution of sodium hypochlorite (2.5%). After 15 minutes, 2.5 mL of sodium thiosulfate solution (250 g.L-1) were added to perform the neutralization. After two hours, this solution was discarded and the photobioreactors were filled with 2 liters of sterile culture medium (Andersen, 2005). The photobioreactors were continuous illuminated by a panel of electronic lamps (13 W, white light, Tashibra) each. The panel was divided into sections by wooden supports to isolate each experiment. The number of turned on lamps and the distance between the light panel and the photobioreactors were adjusted to provide average light intensities of 2,200 Lux, 10,000 Lux, 17,000 Lux or 24,500 Lux. The cultures were kept at 30 oC, and air flow rate was 0.5 L.min-1. All experiments were done in duplicate. Pre-cultures were performed by adding 10 mL of algae culture from the germination camera into 100 ml of sterile culture medium in 500 mL conic flasks placed in an orbital shaker at 30 °C. The flasks were continuous illuminated by an electronic circular lamp (30 W, cool light, Avant) corresponding to a light intensity of 2,500 Lux. After 7 days, 100 mL of sterile culture medium were added to the flasks and
the cultures were grown form more 5 days. The photobioreactors were inoculated with 200 mL of the pre-cultures.
Figure 1. Flat-plate air-lift photobioreactors with an internal heat-exchanger (FPA-IHE).
Analytical Procedures Temperature of the cultures was measured by thermometers and light intensity was monitored using digital light meter (MS6610 / Akso). The pH was measured by pH indicator strip (Alkalit ® / Merck), range from 7.5 to 14. Algae growth was monitored by measuring optical density at 570 nm (Lourenço, 2006) and related to cell-dry-weight (CDW) by a calibration curve (Equation 1, Figure 2). X ( g.L−1 ) = 0,304.OD570nm
(1)
CDW (g.L-1 )
0.20 0.15 0.10 0.05 0.00 0
0.2
0.4
0.6
0.8
Optical Density (570 nm) Figure 2. Relationship between biomass in g.L-1 and optical density in 570 nm.
At the end of growth, the biomass of each photobioreactor was separated from
the culture medium by centrifugation and lyophilized for lipid analysis. The total lipid content was determined by Soxhlet. The carbon dioxide biofixation rate was calculated considering that the carbon content of the biomass is 45% (w/w) (Sydney et al., 2010).
RESULTS AND DISCUSSION The results for maximum biomass and maximum specific growth rate, calculated in the exponential phase of growth, are presented in Table 1. The cultures grown at light intensity of 17,000 Lux achieved the highest values of biomass and specific growth rate, while those grown at light intensity of 2,200 Lux showed the lowest values for biomass and specific growth rate. Table 1. Results of biomass and maximum specific growth rate for the light intensities of 2,200, 10,000, 17,000 e 24,500 Lux.
Ilumination (Lux)
Biomass (g.L-1)
Maximum Specific Growth Rate (d-1)
2,200 10,000 17,000 24,500
0.23 0.35 0.38 0.36
0.36 0.55 0.61 0.51
Growth curves are shown in Figure 3.A. All cultures of Chlorella minutissima showed exponential growth phase between 20 and 100 hours of cultivation, independent of light intensity. Statistical analysis of biomass in 114 hours of cultivation, showed that light intensity has a significant influence on biomass production (p <0.001). Tukey's test showed that biomass values have no significant difference within light intensities of 10,000 Lux, 17,000 Lux and 24,500 Lux, and biomass at 2,200 Lux differs from the others. Results for carbon dioxide biofixation rates are shown in Figure 3.B. Cultures grown at light intensity of 2,200 Lux absorbed around 5 gCO2.m-3.h-1 after 48 hours of culture. Cultures grown between 10,000 Lux and 24,500 Lux had similar profiles of carbon dioxide biofixation rate. The curves shown an increase until 48 h of culture, followed by a plateau and a peak at 96 hours, and then a drastic fall. The highest carbon dioxide biofixation rate at the plateau, around 16 gCO2.m-3.h-1, was achieved at 17,000 Lux. 2,200 Lux
10,000 Lux
17,000 Lux
A
0.40
CO2 Biofixation Rate (g.m-3.h-1)
Biomass (g.L-1)
24,500 Lux
25
0.50
B
20
0.30
15
0.20
10
0.10 0.00 0
24
48
72
Time (h)
96
120
144
168
5 0 0
24
48
72
Time (h)
Figure 3. A. Growth curves. B. Carbon dioxide biofixation rate.
96
120
144
The total lipids content in the biomass is shown in Figure 4. The increase in light intensity until 17,000 Lux leads to an increase in total lipid content of the biomass, however a further increase to 24,500 Lux cause a drastic reduction on this variable. Degen et al. (2001) tested light intensities until 29,600 Lux in cultures of Chlorella, reaching a maximum growth rate at the light intensity of 18,500 Lux, equivalent of 0.08 h-1 (1.92 d-1), above this value, a photoinhibition fenomena was observed. 12.00%
17,000 Lux; 10.36%
Total Lipid Content
10.00%
8.00%
6.00%
10,000 Lux; 7.34% 2,200 Lux; 6.00%
4.00%
2.00%
24,500 Lux; 2.06%
0.00%
Figure 4. Results of total lipid content.
CONCLUSIONS The light intensity of 17,000 Lux had the best results, reaching a biomass of 0.38 g/L with a total lipid content of 10.36%. At the exponential growth phase, the maximum specific growth rate was 0.61 d-1 and the carbon dioxide biofixation rate was 16 gCO2.m-3.h-1.
ACKNOWLEDGMENTS Authors are very grateful to CNPq (Brazilian National Council of Research and Development) and to RNC (Brazil Coal National Network, http://www.ufrgs.br/rede_carvao) by the financial support for this research.
REFERENCES ANDERSEN, R.A. (2005), Algal Culturing Techniques. Elsevier, Inglaterra. LIMA, U.A.; SATO, S. (2001) Produção de lipídeos por microrganismos in Biotecnologia Industrial Volume 4, Edgard Blücher Ltda, Brasil. BORGES, L.; FARIAS, B.M.; ODEBRECHT, C.; ABREU, P. (2007), Potencial de absorção de carbono por espécies de microalgas usadas na aqüicultura: primeiros passos para o desenvolvimento de um “Mecanismo de Desenvolvimento Limpo”. Revista Atlântica, Rio Grande, v. 29, p. 35-46.
CHISTI, Y. (2007), Biodiesel from microalgae. Biotechnology Advances, v. 25, p. 294306. GUILLARD, R. L. L. (1975), Culture of phytoplankton for feeding marine invertebrates. Plenum Publishing, p. 29-60. KHAN, S.A.; RASHMI; HUSSAIN, M.Z.; PRASAD, S.; BANERJEE, U.C. (2009), Prospects of biodiesel production from microalgae in India. Renewable and Sustainable Energy Reviews, doi:10.1016/j.rser.2009.04.005. KOCHEM, L.H. (2010), Caracterização de fotobiorreator air-lift para cultivo de microlgas. Chemical Engineering Final Project, Federal University of Rio Grande do Sul, Porto Alegre, Brasil. LOURENÇO, S.O. (2006), Cultivo de microalgas marinhas: Princípios e Aplicações, RiMa, Brasil. POSTEN, C. (2009), Design principles of photo-bioreactors for cultivation of microalgae. Engineering in Life Sciences, v. 9, p. 165-177. SYDNEY, E.B.; STURM, W.; CARVALHO, J.C.; THOMAZ-SOCCOL, V.; LARROCHE, C.; PANDEY, A.; SOCCOL, C.R. (2010), Potential carbon dioxide fixation by industrially important microalgae. Bioresource Technology, v.101, P.5892–5896. TEIXEIRA, C.M.L.L.; MORALES, E. (2006), Microalga como matéria-prima para a produção de biodiesel. Proceedings of the I Congress of the Brazilian Biodiesel Technology, Brasília. XU, L.; WEATHERS, P.J.; XIONG, X.R.; LIU, C.Z. (2009), Microalgal bioreactors: Challenges and opportunities. Engineering in Life Sciences, v. 9, nº 3, p. 178-189. DEGEN, J.; UEBELE, A.; RETZE, A.; SCHMID-STAIGER, U.; TRÖSCH, W. (2001), A novel airlift photobioreactor with baffles for improved light utilization through the flashing light effect. Journal of Biotechnology, v. 92, p. 89-94.
INFLUENCE OF TEMPERATURE AND SALINITY OVER CO2 BIOFIXATION BY THE MICROALGAE DUNALIELLA TERTIOLECTA Nicéia Chies Da Fré1, Rosane Rech2 and Nilson Romeu Marcílio1 1
Federal University of Rio Grande do Sul – Department of Chemical Engineering ZC 90040-040 Porto Alegre – RS – E-mail: [email protected], [email protected] 2 Federal University of Rio Grande do Sul – Food Science & Technology Institute PO Box 15090 – ZC 91501-970 Porto Alegre – RS E-mail: [email protected]
ABSTRACT Microalgae Dunaliella tertiolecta is a unicellular photosynthetic microorganism that can fix CO2 efficiently from the atmosphere and from industrial exhaust gases. In the present work, the effects of medium salinity and temperature over the growth of the marine microalgae D. tertiolecta in airlift photobioreactor were studied. The experiments were performed according to a central composite design. Growth curves, CO2 biofixation rates and lipid content of the resulted biomass were determined for each experiment. The best growth conditions were 0.715 M NaCl and 28 oC. Under these conditions, 0.39 ± 0.03 g/L of biomass were obtained in 93 hours of culture and the lipid content of biomass was 11.75 ± 1.8 %. The greatest values of CO2 biofixation rate (11.34 ± 1.6 g CO2/m3.h) occurred at 28 oC, independent of the medium salinity. Keywords: microalgae, Dunaliella tertiolecta, lipid, airlift photobioreactor.
INTRODUCTION Microalgae can fix carbon dioxide from different sources including CO2 from the atmosphere, from industrial exhaust gases and in form of soluble carbonates. The cultivation of microalgae under controlled conditions promotes carbon dioxide biofixation and yields products of commercial interest as pigments, lipids, proteins, sugars, among others (Sydney et al., 2010). Additionally, it has been estimated that biomass production per hectare may be 10 or more times higher for microalgae than for conventional crops (Chisti, 2007a,b). Dunaliella tertiolecta is a marine microalgae that is simple to cultivate, does not clump on surfaces and is highly salt tolerant, which might be useful in large-scale outdoor cultivation (Elenkov et al., 1996). Salt tolerance of D. tertiolecta has been reported to be in the range of 0.17–1.5 M of NaCl (Olliver et al., 1994). However other investigators working on D. tertiolecta have gone far beyond these limits by quoting NaCl tolerance range of 0.05–3.00 M (Jahnke and White, 2003). In this study, the effects of the medium salinity (0.430 to 1.0 M NaCl) and temperature (21 to 35 °C) over the growth of D. tertiolecta in airlift photobioreactor was evaluated. The CO2 biofixation rates and the lipid and pigment content of the resulted biomass were also measured.
METHODS The experiments were performed with the unicellular green marine microalgae Dunaliella tertiolecta. The cells were maintained in 250 mL Erlenmeyer flasks with 50 mL of culture medium f/2 (Guillard, 1975) in camera germination under controlled temperature (20 °C) with a photoperiod of 12/12 hours (light/dark) provided by fluorescent lamp light. In order to obtain an adequate volume of culture for inoculation of photobioreactors, pre-cultures were done in rotatory shaker. Homogeneous aliquots of 10 ml of culture held in camera germination were inoculated in 500 mL Erlenmeyer flasks containing 200 mL of culture medium f/2 plus 22 g/L (0.43 M NaCl). The flasks were kept in rotatory shaker at 28 °C with continuous light for 12 days. The experiments were performed in flat-panel-airlift photobioreactors with an internal heat-exchanger (FPA-IHE) (Kochem, 2010). The pre-cultures were inoculated in sterile photobioreactors containing 2,000 mL of culture medium f/2 plus NaCl (0.430 M, 0.513 M, 0.715 M, 0.917 M or 1.0 M) according to the experimental design. The photobioreactors were continuously illuminated at 18,000 lux by a panel of electronic lamps (24 × 13 W cool light, Tashibra). The temperature was kept constant (21, 23, 28, 33 or 35 °C) by passing water from a thermostatic bath through the heat exchanger of the photobioreactor. The cultures were performed in triplicate according to the central composite design (CCD) showed in Table 1. Table 1. Central composite design of experiments.
The growth was followed by measures of the optical density at 570 nm three times a day and correlated with cell dry weight (CDW). After 96 hours of culture, the biomass from the photobioreactors was separated from medium by centrifugation, washed with distilled water and lyophilized. The total lipid content of the lyophilized biomass was determined by Soxhlet. RESULTS AND DISCUSSION The growth curves of D. tertiolecta in FPA-IHE photobioreactors are shown in Figures 1, 2 and 3. The highest biomass concentration was achieved at 0.715 M NaCl and 28 °C. The experiments performed at 33 °C and 35 °C reached the decline phase around 70 hours, what did not occur at other temperatures.
Figure 1. Growth of Dunaliella tertiolecta at salinity of 0.715 M and different temperatures.
Figure 2. Growth of Dunaliella tertiolecta at salinity the higher salinities of CCD.
Figure 3. Growth of Dunaliella tertiolecta at salinity the lower salinities of CCD. The response surface for biomass values at 93 hours of culture as function of temperature and salinity is shown in Figure 4. Analyzing the response surface, it is possible to conclude that the best conditions for growth of D. tertiolecta are temperature of 28 oC and salinity of 0.715 M NaCl. Under these conditions, 0.39 ± 0.03 g/L of biomass were achieved at 93 hours of culture.
Figure 4. Response surface for biomass values (g/L) in 93 hours of culture as a function of temperature and salinity. Figure 5 shows the response surface for total lipid content of cells with temperature and salinity. The highest lipid content occurred at 28 °C and 0.715 M NaCl, the same condition of the highest biomass concentration. Under these conditions, a total lipid content of 11.75 ± 1.8% was achieved. This value is very similar to that obtained by Sydney et al. (2010), which was 11.44 ± 1.8%, but at different growth conditions from those used in this work. The CO2 biofixation rates are shown in Figure 6. The highest biofixation rates (11.34 ± 1.6 g CO2/m3.h) were achieved at 28 °C, being independent of medium salinity.
Figure 5. Response surface for total lipid content (%) as a function of temperature and salinity.
Figure 6. Average rates of CO2 fixation in microalgae Dunaliella tertiolecta grown over 100 h for the experiments. CONCLUSIONS In this work, the biomass production, total lipid content and CO2 biofixation rate by D. tertiolecta were optimized by central composite design with temperature and medium salinity. At temperature of 28 °C and salinity of 0.715 M NaCl the highest values of biomass and lipid content were achieved. The highest CO2 biofixation rates occurred at 28 °C, independent of the medium salinity.
ACKNOWLEDGMENTS Authors are very grateful to CNPq (Brazilian National Council of Research and Development) and to RNC (Brazil Coal National Network, http://www.ufrgs.br/rede_carvao) by the financial support for this research.
REFERENCES CHISTI, Y. (2007a), Biodiesel from microalgae. Biotechnology Advances, v.25, p. 294-306. CHISTI, Y. (2007b), Biodiesel from microalgae beats bioethanol. Trends in Biotechnology, v.26, p. 6. ELENKOV, I.; STEFANOV, K.; DIMITROVAKONAKLIEVA, S. and POPOV, S. (1996), Effect of salinity on lipid composition of Cladophora vagabunda. Phytochemistry, v.42, p. 39-44. GUILLARD, R.R.L. (1975), Culture of phytoplankton for feeding marine invertebrates. In: Smith, WL & MH Chanley (Eds.) Culture of Marine Invertebrate Animals. Plenun, New York, p. 29-60.
JAHNKE, L.S. E WHITE, A.L. (2003), Long-term hyposaline and hypersaline stresses produce distinct antioxidant responses in the marine alga Dunaliella tertiolecta. Journal Plant Physiology, v.160, p. 1193-1202. KOCHEM, L.H. (2010), Caracterização de fotobiorreator air-lift para cultivo de microlgas. Trabalho de conclusão do curso de Engenharia Química, Universidade Federal do Rio Grande do Sul, Porto Alegre, Brasil. OLLIVER, B.; CAUMETTTE, P.; GARCIA, J.L. AND MAH, R.A. (1994), Anaerobic bacteria from hypersaline environments. Microbiol. Rev. 58, 27-38. SYDNEY, E.B.; STURM, W.; DE CARVALHO, J.C.; THOMAZ-SOCCOL, V.; LARROCHE; C.; PANDEY, A. E SOCCOL, C.R. (2010), Potential carbon dioxide fixation by industrially important microalgae. Bioresource Technology, v.101, p. 5892-5896.
Carbon Oxides Emission via the Atmospheric Oxidation of Coals: Effect of Coal Rank ,1, 3
Uri Green 123-
,3
1,2
, Zeev Aizenstat and Haim Cohen
Department of Biological Chemistry, Ariel University Center at Samaria, Ariel, Israel. Chemistry Department, Ben-Gurion University of the Negev, Beer Sheva , Israel. Chemistry Department, Hebrew University of Jerusalem, Jerusalem, Israel.
Abstract More than 50% of the electrical power in Israel is produced through bituminous coal combustion. As Israel has no domestic coal reserves it is imported using large ships and is subsequently stored in strategic stockpiles. From the moment the coal is mined it undergoes degradation via atmospheric weathering. This process involves a series of reactions between atmospheric oxygen and the coal surface which result in several gaseous emissions, mainly carbon dioxide and carbon monoxide[1]. The present research aims at expanding our knowledge of the formation of low temperature (30-150C) oxidation products and their dependence on coal rank. The focus of this work is primarily on the functional groups that can serve as precursors to the formation of carbon oxides. The source of the oxygen in the carbon oxides formed can be either atmospheric oxygen gas or inherent oxygen content in the coal macromolecule backbone. The precise relationship and the specific functional groups at the coal surface that are reactive are still unclear. The rank of the coal has shown an appreciable effect[2] on the oxygenated products as such simulation experiments with different coals (lignites, subbitumineous and bituminous) were conducted. From the results of these experiments a comparison of the behavior of variously aged coals has been evaluated. Conclusions drawn from the present research can help to gain a better understanding of the reactions occurring while coal undergoes the oxidation process.
Keywords: Atmospheric oxidation, carbon oxides, coal surface, coal rank
1
a. S. L. Grossman Ph.D. Thesis, “Low Temperature Atmospheric Oxidation of Coal”, Chemistry Department, Ben-Gurion University of the Negev, Beer-Sheva, Israel, (1994). b. S. L. Grossman, S. Davidi and H.Cohen, Fuel, 70, 897 (1991). 2 Zeev Aizenstat, Uri Green & Haim Cohen, "Modes of Formation of Carbon Oxides (COx (x = 1,2)) From Coals During Atmospheric Storage: Part I Effect of Coal Rank", Energy Fuels, 2010, 24 (12), pp 6366–6374
Introduction Recently[2], we have suggested that the reactions of the coal during weathering to yield carbon oxides (CO and CO2) can occur via 2 different types of precursors: They are formed either from inherent oxygen or surface oxides decomposition:
Route 1
Route 2
inherent oxygen 1a
CO
(1)
inherent oxygen 1b
CO2
(2)
surface oxides 2a
CO
(3)
surface oxides 2b
CO2
(4)
It is suggested that the carbon oxides which are emitted from the coal (and it is dependent on coal rank) stem from atmospheric oxygen and also from thermal decomposition of the inherent oxygen content within the coal macromolecular structure. It is of interest to investigate and to try to determine the different modes by which carbon oxides are released during the low temperature oxidation/weathering processes. We have decided to study intensively 3 classes of coal: Bituminous, SubBituminous and Lignite; each containing an increasing amount of inherent O. For this purpose, several coals which are fired either in the Israeli utility plants (bituminous and sub-bituminous coals) or in German power plants (lignite coals) have been chosen.
Experimental All chemicals and gases used throughout the study were of AR grade and supplied by Aldrich, Fluka, Merck or Maxima. The water used throughout this study was purified water (via ion exchange columns). Coal. Experiments in this work were carried out with three classes of coals; bituminous (from South Africa), sub-bituminous (from Indonesia) and lignite (from Grmany). The properties of all the coals are presented in Table 1. The experiments were performed in sealed glass vials (40 ml) charged with coal (particle size 74µ ≤ X ≤ 250µ) in an air, argon or oxygen atmosphere &&ve oven model FT 300. and heated at 55-950C for various periods in a nu All the coals in the present study were prepared by grinding them down and separating them by grain size via sieves. The coal samples were then dried in a Heraeus vacuum oven model VT6060 for 24 hours at 60oC. Gas Chromatography. The amount of the gases (CO2, CO,, N2, O2, hydrocarbons) in the reactors was determined using a gas chromatograph (Varian model 3800) equipped with a thermal conductivity detector & a flame ionization detector connected in series. The gases were separated on a carbosieve B 1/8”, 9’ ss column using a temperature programmed mode. The experimental error in the G.C. determination is ±5%.
The gaseous atmosphere was sampled (1ml samples) after the reaction, with gas tight syringes and measured in the gas chromatograph. As the reactions studied are gas/surface reactions, the reproducibility of the results is not good, thus each experiment has been carried out with duplicate reactors, the error is ±15%, (the main source is the nature of the heterogeneous reactions studied in the experiments). Table 1. Properties of Coalsa type
Ashwf
VMdaf
SA
moisture 1.20
13.77
IN
2.61
HA
b
ultimate analysis (wt %, daf) CV( J ⋅ g
H 4.10
O
S
28.27
C 74.19
5.59
0.46
28,416
10.65
36.38
73.25
4.63
9.05
0.67
28,564
34.53
5.09
52.39
66.12
4.32
23.65
0.16
25,323
10.57
4.16
63.24
4.50
−1
27.24 0.20 24,516 VM = volatile matter; CV = calorific value; daf = dry ash free. wf = water free.
LUS a
proximate analysis (wt %)
52.78
SA = South Africa; IN = Indonesia; HA = Hambach, Germany; LUS = Lusatia, Germany.
Results Kinetics and Rates. The kinetic investigations were carried out in sealed glass vials (40 ml) used as batch reactors charged with 0.5 gr coal (particle size 74µ ≤ X ≤ 250µ) in an air atmosphere. The reactors were heated at different temperatures for various periods of time. In order to enable an accurate measurement of the evolving gasses we sampled the vials at the oven temperature. The results of the experiments to determine the rates of reaction for formation of the carbon dioxide are given in Table 2 and the results for carbon monoxide are presented in Table 3. Table 2. Rate constants for carbon dioxide emission at different temperatures # T (C)
rate constants (mL/g·h) SA IN HA 55 0.006 0.012 0.053 65 0.013 0.034 0.174 75 0.034 0.072 0.374 85 0.054 0.164 0.624 95 0.188 0.208 0.934 # rates measurements were determined in plots from 0.5h R2 =0.95±5%
LUS 0.080 0.252 0.413 0.735 1.247 up to 48h, average
)
Table 3. Rate constants for carbon monoxide emission at different temperatures # T (C)
rate constants (mL/g·h) SA IN HA LUS 55 0.0013 65 0.0014 0.0021 75 0.0009 0.0023 0.0046 0.0054 85 0.0018 0.0031 0.0069 0.0063 95 0.0028 0.0048 0.0104 0.0123 # rates measurements were determined in plots from 0.5h up to 48h, average R2 =0.95±5% Regarding Table 2, CO2 is produced from the first moment of heating. Simulation rates were determined up to 48hr. However, regarding the CO emissions a noticeable induction period was observed. This delay period is coal rank dependent as well as temperature. For example at 95C for IN coal CO emissions were detected only after 1h while for SA coal emissions were detected after 3h. B.Z. DLUGOGORSKI et al [3] suggested that the variation of the ratio of CO2/CO production rates is directly related to the dissociation of unstable intermediates. In the following figure 3, which is based on this suggested parameter, we present the time dependence of the carbon oxide emissions for the four different rank coals. Figure 3
80 70 SA
R = CO2/CO
60
IN
50
HA
40
LUS
30 20 10 0 0
10
20
30
40
50
time (h)
Figure 3. DLUGOGORSKI ratio depicting the emissions of carbon oxides at 950C for various periods of time in simulation reactors containing 0.5g coal. It is clear from figure 3 that the initial behaviour of lower rank coals (HA, LUS) is significantly different than that of the higher rank coals (IN, SA). There is one main point which is fascinating in the above figure. We notice that in the initial stage (<5h) the emissions from the lower rank coals start at a high ratio then fall to a minimum before climbing again until reaching an 3
B.Z. Dulgogorski et al., Pathways for Production of CO2 amd CO in Lowtemperature Oxidation of Coal., Energy & Fuels (2003), 17, 150-158
asymptote. On the other hand both the higher rank coals start from the baseline (R=0) then climb till reaching an asymptote. This significantly different behaviour portrayed by the ratio R might possibly relate to the higher concentration of inherent oxygenated functional groups that are prone to decarbolxylation. Thus, for the lower rank coals the initial high R value is the result of thermal decomposition of the inherent functional groups. The minimum which is reached is the last of the easily released inherent surface functional groups. Thereafter, the climb in the ratio of emissions until the asymptote can be contributed to low temperature atmospheric oxidation alike the plots of the higher rank coals. It is also interesting that in both higher rank coals there is a significant delay stage before the emission of the carbon monoxide. At this initial stage we can clearly see the temperature dependence of the emissions. In order to properly evaluate this effect experiments were performed in sealed glass vials (40 ml) used as batch reactors. The reactors were charged with 0.5 gr coal (particle size 74µ ≤ X ≤ 250µ) in an air atmosphere. They were then heated at different temperatures for various periods of time. The emissions of CO from IN, SA, LUS and HA coals are plotted in figures 4,5,6,7 respectively.
CO emission from IN coal
CO emission from SA coal
CO emission from LUS coal
CO emission from HA coal
It is clear that temperature and rank have an overall effect on the emission of CO. Interestingly, regarding the lignite coals (low rank) one can clearly see that once the 75C threshold is crossed there is no delay in the release of CO from the coal matrix. This corresponds to previous results[2] where we noted that the emission of CO in the case of lignite is mass dependent. The induction period for CO emission seems to indicate that the emission of CO during LTO requires formation of a stable surface oxide which is thereafter is destabilized to produce CO.
Conclusion It is clear at this point that the mechanism responsible for the carbon oxides emission is different in the case of the low rank coals (HA, LUS) as opposed to the high rank coals. Furthermore, as shown above in figure 3 it is clear that these low rank coals show some similarity in their reactivity. At this point it is also apparent that the IN coal reacts differently from the SA coal.
References a. S. L. Grossman Ph.D. Thesis, “Low Temperature Atmospheric Oxidation of Coal”, Chemistry Department, Ben-Gurion University of the Negev, Beer-Sheva, Israel, (1994). b. S. L. Grossman, S. Davidi and H.Cohen, Fuel, 70, 897 (1991). 2. Zeev Aizenstat, Uri Green & Haim Cohen, "Modes of Formation of Carbon Oxides (COx (x = 1,2)) From Coals During Atmospheric Storage: Part I Effect of Coal Rank", Energy Fuels, 2010, 24 (12), pp 6366–6374 3. B.Z. Dulgogorski et al., Pathways for Production of CO2 amd CO in Lowtemperature Oxidation of Coal., Energy & Fuels (2003), 17, 150-158
1.
Oviedo ICCS&T 2011. Extended Abstract
Environmental pollution by migration of gas produced in the Underground Coal Gasification process Authors: Magdalena Ludwik-Pardała 1, Krzysztof Stańczyk Addresses: Główny Instytut Górnictwa, Plac Gwarków , 40-166 Katowice, e-mail address: [email protected] (M. Ludwik-Pardała1) Abstract The Underground Coal Gasification Process (UCG) belongs to Clean Coal Technologies since in comparison to the traditional coal conversion methods it does not generate high environment pollution. Still, during the underground gasification and after the process is terminated the gas and liquid products contain substances harmful to the environment which can migrate to the surroundings of the underground reactor. The synthesis gas obtained in the process is a mixture consisting above all the following components: CO2, CO, H2, CH4, and, depending on the method used in the process, the percentage of the main components of synthesis gas mentioned above is changeable. At the same time in the coal gasification process other chemical compounds of synthesis gas such as: H2S, NH3, metals, BTEX and VOC as well as tar are produced. The migration of the gas and liquid contaminants during the conducted UCG experiments has been repeatedly monitored, still, so far, not all of the contaminant migration mechanisms have been well researched and described - among other the phenomena accompanying the gas transport. Thus it seems crucial to describe the gas migration from the underground georeactor and thus the contaminants migration to the georeactor surroundings. 1. Introduction The comparison of Underground Coal Gasification (UCG) process with conventional coal processing to energy and chemicals presents the latter as the more environmentally friendly due to its lack of air, water and ground emissions. Still, both the gas generated in the process and the liquid products of the gasification process can migrate into the strata surrounding the underground georeactor. Synthesis gas obtained in the gasification process is a mixture consisting of the following components: CO2, CO, H2, and CH4. Depending on the method used in the process, the percentage of the main components of synthesis gas mentioned above is changeable. At the same time in the coal gasification process other chemical compounds of synthesis gas such as: H2S, NH3, metals, BTEX and VOC as well as tar are produced.[1] The underground coal gasification experiments in the world have been carried out at various scales since 1940. The precise process monitoring has shown that the gasification process is not neutral to the strata surrounding the georeactor. The first data on the migration of contaminants from the UCG process come from large scale experiments carried out in 1950-1960 in USSR. In the area of 10 km2 from the georeactor phenols were observed in the ground waters (Klimentov, 1964). The American experiments (Hanna, Hoe Creek, Rawlins, Rocky Hill 1, Rocky Mountain 1, Carbon County UCG – 1976-1995) in Wyoming [2] demonstrated the scale of the migration, especially, of the ammonia, sulphur and carbon dioxide. These compounds, contained in the produced synthesis gas were responsible for the change of the water pH and the contamination of the surrounding strata, which, in turn, was the cause of the decrease of oxygen amount in the ground waters and grounds [3,4]. The underlying reasons of the gas migration from the UCG georeactor can be divided
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1
Oviedo ICCS&T 2011. Extended Abstract
into four groups: • Geological - reservoir and rock pressure gradient, the textural strata properties (porosity, permeability, and factorization), the nature of the reservoir fluids, the temperature, the hydrogeological conditions, the tectonic disorders, etc. • Mechanical - the strata and coal around the cavern cracking resistance in high temperature. • Technological - the scale of the process, temperature, the gasification media (air, steam, oxygen), the georeactor depth, the quality of coal • Technical - the quality of the pipes for transporting the synthesis gas, the resistance of materials to heavy products The monitoring of different experiments has shown the complex character of the environmental problems in UCG since the contaminants are able to migrate as gas and liquid depending on the physical conditions such as pressure and temperature - they are able to change their phases between gas-liquid and liquid-gas. The most important issue is the dependence between the hydrostatic pressure of the surrounding strata and the pressure inside the reactor. If the pressure inside the reactor increases the hydrostatic pressure of the surrounding strata, the gases produced in the process will be pushed from the reactor into the strata in an uncontrolled way. The risk of exceeding the hydrostatic pressure decreases with the depth of the UCG process. If the UCG process proceeds in a controlled and safe way, then the risk of gas migration is lower. There are, however, several physico-chemical phenomena which facilitate gas transport from the reactor into the surrounding strata. In the fire channel behind the front of fire in the cavern, the rapid sorption and diffusion processes may occur. The presence of steam facilitates the hot gas absorption n the porous structure of coal and coke. The high temperature of the process results in textural changes in coal and strata. The surface area and the pore volume increase in the burned (coked) coal which simplifies the sorption of the gases[5]. The sorption properties of the surrounding strata decrease under the influence of temperature due to sintering and vitrification which results in cracking. This may be the underlying reason for the gas transport into the virgin strata. Hot gases can be easily adsorbed in the unconverted strata. They can also condensate and accumulate as liquid contaminants. The strata layers which remain unconverted by the heat can absorb CO, CO2 and H2 [8]. The contaminants from synthesis gas such as tars, BTEX, VOC, H2S, NH3 and metals might be eluted from the strata by ground waters and process waters which will result in their migration into wide area. The CO and CO2 absorbed in coal and coke, or strata can be easily washed out by the water let into the cavern after the gasification process. The CO2 reacts with water forming H2CO3, the acid precipitates the heavy metals from the ash and strata. Also benzene is easily dissolvable in CO2 which increases the risk of its migration to water. [6,7] 2. Experimental section The experiments on the migration of contaminants in the UCG process took place during the simulation of the UCG process in the ex-situ reactor in the “Barbara” Experimental Mine of Główny Instytut Górnictwa (Central Mining Institute) in Katowice. The textural, adsorption and diffusion properties were measured in the samples of coal and strata located in different parts of the ex-situ reactor -situ and influenced by different temperature ranges. The gasification process in the ex-situ reactor was carried out during 170 h applying oxygen (400 m3/h) and steam (250 m3/h) as the gasification media. The coal seam was simulated by two coal blocks from “Bielszowice” Coal Mine (Coal characteristic- Table 2) – its diameters were 2.6 m x 0.5 x 0.6 m, the weight: 850 kg. The strata surrounding Submit before May 15th to [email protected]
2
Oviedo ICCS&T 2011. Extended Abstract
the coal were simulated by clay shale. In the experiment 1200 m3 of synthesis gas were produced - the composition of the gas is presented below (Table 1). The samples of the strata and the coal were taken from different locations in the middle part of the reactor. Figure 1 presents the sampling scheme.
3
Number of samples:
1_3
4 Gasified Coal fire channel
The strata Samales from the bottom of the reactor 2-melted rock under the fire channel 5-strata from the bottom of the reactor 1_1- strata placed between the 2 and 5 Gasified coal Burned (coked_ coal from the fire channel)
2 5
The Strata samples above the cavern: 3-strata from the top of reactor 4-strata from over the gasified coal 1_3-strata placed between the 3 and 4 samples
1_1
Fig.1. The sampling scheme 2.1 The measurement methodology 1) The textural properties measurements included the determination of such parameters as: Specific surface area- SBET, specific surface area of mesopores Smeso and volume of micropores Vmicro from nitrogen physical adsorption isotherm at 77K evaluated applying the modified BET equation; Intrusion volume, Vintr, and apparent density, ρHg, from high pressure mercury porosimetry; True density, ρHe, from helium pycnometer. Sample porosity, ε, was obtained as ε=1−(ρHg/ρHe). The textural characteristics were measured in the Institute of Chemical Process Fundamentals of the ASCR in Praha. 2) Adsorption properties of the materials from the ex-situ reactor were measured in the micro-thermo-weight Sartorius 4436. The measurements took place at different pressure ranges of CO2 and H2: 1, 2, 3 and 4 atm. The weight of each of the samples was 200 mg, the grain size 0.1-0.2 mm, the measurement lasted almost 24 hours. Statics and kinetics of sorption were determined as the result of the calculation of absorption. 3) Diffusion coefficients were measured applying the Inverse Gas Chromatography method. The samples from the ex-situ reactor were placed in a chromatographic column of the length of 50 cm. The chromatographic column was then placed in the oven. The
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Oviedo ICCS&T 2011. Extended Abstract
different carrier gas flow (Ar) rates through the column and the different temperature rates were measured by two tracers: H2 and the inert gas He. The breakthrough curve of the tracer gas was recorded as the response peak by the TCD cathetometer. As the results of the analysis of the peak shape applying the Van Deemeter method the diffusion coefficients such as A = Eddy-diffusion, B = Longitudinal diffusion, C = mass transfer kinetics of the analyses between mobile and stationary phase and u = Linear Velocity have been calculated. 3. Results and discussion The temperature monitoring allowed to control the influence of the temperature on the samples. The same properties were observed in the gasified coal – the area of the mesopores and the volume of the micro pores increased under the influence of high temperature. The figures below present the distribution of temperature in the cavern and the surrounding strata measured during the simulation of the UCG process in the ex-situ reactor. (FIG 2 and 3)
Fig. 2. Temperature distribution in the fire channel
Fig 3 . Temperature distribution in the strata above the cavern. The gas composition was measured every half hour during the experiment by the chromatographic method, the average gas composition is presented table 1. The gasification reagents (media) were oxygen and steam.
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Oviedo ICCS&T 2011. Extended Abstract
Table 1. Average gas composition Reagent Oxygen steam
CO2 C2H6 64.2 0.08 19.41 0.22
H2 11.99 49.65
O2 2.1 1.21
N2 5.53 7.3
CH4 2.37 8.35
CO 13.59 13.67
H2S 0.09 0.23
The gasified coal properties from “Bielszowice” mine are presented below (Table 2). Table 2. Coal characteristic Parametr As recived Total moisture Wtr [%] Ash Atr [%] Total sulphur Str [%] Calorific value Qir [kJ/kg] Analytical Moisture Wa [%] Asha [%] Volatiles Va [%] Heat of combustion Qsa [kJ/kg] Calorific value Qia [kJ/kg] Total sulphur Sa[%] Carbon Cta[%] Hydrogen Hta [%] Nitrogen Na[%]
Value
Standard
1.55 2.21 0.28 33370
PN-G-04511:1980 PN-G-04560:1998; PN-ISO 1171:2002 PN-G-04584:2001; PN-ISO 334:1997 PN-G-04513:1981
1.47 2.21 32.41
PN-G-04511:1980 PN-G-04560:1998; PN-ISO 1171:2002 PN-G-04516:1998; PN ISO-562:2000 PN-G-04513:1981
34572 33399 0.28 83.84 4.94 1.15
PN-G-04513:1981 PN-G-04584:2001; PN-ISO 334:1997 PN-G-04571:1998 PN-G-04571:1998 PN-G-04571:1998
The measured textural properties of the samples placed in different range of temperature in reactor ex-situ are presented below (Tab 3) Table 3. Textural characteristic of various samples [source: ICPF-Institute of Chemical Process Fundamentals of the ASCR] V micro ρHe SBET Smeso ε ρ (mm3/g (g/cm3 Hg 3 Samples (m2/g) (m2/g) (g/cm ) (- ) ) ) 2 strata 1.19 2.68 1.20 0.55 1_1 strata 5 strata 4 strata 1_3 strata 3 strata Burned (coked) coal
18.7 18.8 10.6 12.1 12.2
11.1 11.5 5.9 7.1 7.2
4.1 4.3 2.6 2.8 2.9
2.63 2.63 2.33 2.37 2.37
0.91 0.91 0.97 0,95 0.95
0.65 0.65 0.58 0.60 0.60
12.4
7.3
3.0
2.12
0.92
0.57
Gasified coal
10
5.1
2.6
2.00
0.91
0.55
The coked coal and the gasified coal samples specific surface area and the amount of the micro- and meso- pores have increased in the area of higher temperatures. For example the specific surface area in the coal samples from the cavern is 24% higher than that of Submit before May 15th to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
the coal sample taken from under the strata layer. Also the porosity and the sorption parameters are better in this sample. The analysis of the textural results of the clay shale has shown a decrease of the textural properties in the materials with the increase of the temperature. The comparison between the textural parameters of the strata located above and below the cavern has shown that the specific surface area in the strata from the bottom of the reactor is 43.6% lower than the strata located in the upper part of the reactor. The amount of meso- and micro- pores also decreased under the influence of temperature. The sorption properties of the samples presented table below. The results presented below (tab 4, 5)compares the sorption capacity of the samples for a different range of measured gas pressure. The sorption capacity was measured for a hydrogen and carbon dioxide. Table 4. Sorption capacity of CO2 Sorption capacity [kg/t sample] Sample 1 2 3 4 atm atm atm atm Burned 38.4 47.0 57.2 59.8 (coked) coal Gasified coal 14.2 27.1 29.5 31.9 4 strata 1.0 1.8 2.2 2.5 5 strata 1.2 1.4 1.5 1.6 3 strata 3.9 4.2 8.5 8.9 2 strata 0 1.3 1.5 2.8 Table 5. Sorption capacity of H2 Sorption capacity [kg/t Sample sample] 1 2 3 4 atm atm atm atm Burned (coked) 2.4 3.4 4.9 5.1 coal Gasified coal 5.6 8.2 10.9 12.7 4 strata 0.7 5.5 6.0 6.3 5 strata 1.1 1.2 1.2 1.3 3 strata 0.2 0.5 0.4 2.4 2 strata 0.09 0.2 0.2 0.2 The measurements of the kinetics of CO2 sorption demonstrated that the sorption rate has decreased together with: the strata from the bottom of the reactor, the strata above the cavern, the gasified coal and the coked coal The CO2 sorption rate in the coal samples demonstrated decreasing of the rate of the process together with the increase of the rate of devolatilization of the carbon matter under the influence of the temperature. The CO2 sorption kinetics in the strata demonstrated that the saturation periods are between 120-180 minutes in the strata located above the cavern but in the samples of the strata from the bottom of the reactor the sorption phenomena does not exist. The H2 sorption rate in the coal samples demonstrated rapid growth of the sorption rate in the devolatilized carbon matter. The coke samples located in the highest temperature intervals had the saturation rate of 180-600 minutes. The saturation time of the gasified coal samples was 800-1000 minutes. The H2 sorption kinetics in the strata samples Submit before May 15th to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
demonstrated that the sorption rate decreased under the influence of the temperature. The saturation time in the strata samples located in the temperature interval between 100-1350C was 120-130 minutes. The strata from the bottom of the reactor in higher temperatures were saturated during 450-850 minutes. The sorption capacity presented above (Table 4, 5) for the coal samples was 20-40 times higher than that of the strata samples. The chromatographic measurements have shown that the influence of high temperatures on the strata samples from the bottom and the top of the reactor resulted in the decline of the diffusion parameters. The increase of the temperature resulted in the decline of the diffusion processes and the displacement of the equilibrium into the adsorption processes. 4. Conclusions Based on experimental results the migration processes of synthesis gas from UCG process by means of the absorption and diffusion methods are the result of several elements. The most important are the textural properties of the georeactor surroundings and the impact of temperature on the adsorption and diffusion processes. The temperature influence on the coal and strata materials plays significant role in textural remodeling. The pore volume in the strata decreases, which results in weaker adsorption properties. In contrast to the strata the influence of temperature has increased the specific surface area of coal and its pore volume which improved its adsorption properties. The textural properties of the measured strata and coal samples show the same temperature dependence for the diffusion properties of the strata and gasified coal. The rate of the diffusion processes became slower in the samples of molten strata. The gas transport through the cracking in the cavern and the surrounding strata into the raw part of the rock layers might be the reason of synthesis gas adsorption in the downstream area. The measurements of the strata samples placed in the lower range of temperature showed that their textural properties are different than those of the samples placed in the areas of higher temperature. The specific surface area and the pore size of strata samples unconverted by the heat are greater which cause better adsorption processes and diffusion rate of the synthesis gas than in the strata samples to a greater extent transformed by the heat . The environmental aspect of UCG process seems to be of a comprehensive character due to the dependence on many factors. This relationship between the temperature of the process, the geological mechanisms and the gas transport properties by means of adsorption and diffusion is a subject which is not fully recognized. Taking this relationship into consideration the measurements of the adsorption and diffusion process in UCG seem to be significant. Depending on the scale of the venture, the stage of the process, the depth and the geological and hydrological characteristics of the bed, as well as on the applied safety measures, process parameters, the gas utilization technologies, the emitted contaminants and their concentrations vary, various is also the emission risk. Thus, from the point of view of the preparation of the experiment the potential risks should be presented in order to minimize the contamination and the potential negative impact on the surroundings of the underground georeactor.
Acknowledgements The works presented in this paper have been performed in the frame of HUGE project (Hydrogen Oriented Underground Coal Gasification for Europe) and supported by the
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Oviedo ICCS&T 2011. Extended Abstract
RFCS under the Contract no. RFCR-CT-2007-00006 and Polish Ministry of Science and Higher Education
References [1] Burton, E., Friedman, J. and R. Upadhye, 2005 Best Practices in Underground Coal Gasification. Lawrence Livermore National Laboratory. Submitted to U.S. DOE. Contract No. W-7405-Eng-48; [2] Wabash River Coal Gasifiaction Repowering Project Final Technical Report DEFC21-92MC29310 Development of Underground Coal Gasification for Power Generation, Contract No.: C/07/00378/00/00, Cranfield University, May 2009 [3] Boysen, J. E., C. G. Mones, J. R. Covell, S. Sullivan and R. R. Glaser. 1988. Underground Coal Gasification Contaminant Control Program: Simulation of Postburn UCG, Contaminant Production. GRI 88/0170. [4] Boysen, J.E., C.G. Mones, R.R. Glaser and S. Sullivan. 1987. Physical and Numerical Modeling Results for Controlling Groundwater Contaminants Following Shutdown of Underground Coal Gasification Processes. DOE/FE/60177-2392. [5] Review of underground coal gasification technological advancements, Report No. COAL R211 DTI/Pub URN 01/1041 by D P Creedy, K Garner S Holloway, N Jones, T X Ren, 2001, Rozdział 5, Appendix 43 [6] Review of the Feasibility of Underground Coal Gasification in the UK, Cleaner Fossil Fuels Programme, DTI /Pub URN 04/1643 1 [7] Review of Environmental Issues of Underground Coal Gasification, Report No. COAL R272 DTI/Pub URN 04/1880, November 2004, Martin Sury et al., rozdział 5 i 6 [8] O. Šolcova, K. Soukup, J. Rogut, K. Stanczyk, P. Schneider, Gas transport through porous strata from underground reaction source; the influence of the gas kind, temperature and transport-pore size, Fuel Processing Technology 90 (2009) 1495–1501
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8
Fly Ash as a Potential Scrubber for Low Activity Radioactive Waste
Roy Nir Liberman1, Gyora Segev,1 ,2, Eyal Elish3, Eitan J. C. Borojovich,3 and Haim Cohen1,4 1-
Department of Biological Chemistry, Ariel University Center at Samaria, Ariel, 40700 Israel, phone: 00972-52- 4306878, fax: 00972-8-9200749, email: [email protected] ; [email protected]
2-
Israel Atomic Energy Commission, Tel-Aviv, Israel.
3-
Department of Analytical Chemistry , NRCN, Box 9001, Beer-Sheva, Israel.
4-
Chemistry Department, Ben-Gurion University of the Negev, Beer Sheva, Israel email: [email protected]
Abstract Recently it has been observed that Class F fly ash, which is a byproduct of pulverized bituminous coal combustion in power utilities, can serve as an efficient scrubber and fixation reagent for acidic wastes containing a variety of trace elements. The fly ash produced in Israel has a highly basic reaction when exposed to water (pH >10), which is a result of the very low sulfur phosphorus content due to restrictions imposed on coal imports because of environmental regulations. Thus, it is feasible to use the ash as a chemical scrubber for acidic wastes. Furthermore the fixation product can serve as a partial substitute to sand and cement for concrete production. Bricks produced using the aggregate as sand substitute, have proved to be strong enough according to the concrete standards and the fixation of the trace element in the concrete is excellent (checked via the improved TCLP1311 and the CAL WET methods). Thus, the fly ash might be a potential fixation reagent for radionuclides. The possibility to use the fly ash as a potential fixation reagent has been studied through simulation experiments in aqueous solutions containing cesium cations, Cs+, divalent strontium, Sr2+ and Ce3/4+ as a simulation reagent to the radionuclides. It has been found that indeed the fly ash serves as a good fixation reagent to these metal ions. The mechanism of the fixation processes is discussed. Keywords: fly ash, radioactive wastes, chemical scrubber, trace elements, cesium, strontium, cerium Introduction Today in Israel, most of the electricity is produced by 4 coal fired power stations (>60% in 20081), the amount of bituminous coal burnt in the utilities, is ~13 Mtons per year. The coal (which is imported mainly from South Africa but also from Columbia, Australia, Indonesia and Russia2) contains ~10% of inorganic mineral materials. The fly ash particle size is between 3-250 μm and contains different types of spheres. The first are the cenospheres (image 1A) which are "glass bubbles" which includes mainly aluminum and silicon oxides and also carbon dioxide or nitrogen which give the ash its light weight properties. The second are the
plerospheres (image 1B) which are "hollow glass bubbles" with small particles inside filling it. The fly ash has large surface area which gives it the possibility to act as a potential fixation reagent. The total amount of coal fly ash which is produced in Israel is ~1.3Mtons per year. Due to Israel's strict environmental regulations the coals imported to Israel contain low content of sulfur and phosphorus. Thus, the fly ash produced in Israel has a highly basic reaction when exposed to water (pH >10), because of high lime (CaO) content and is defined as Class F. Of course the high base content makes the fly ash feasible as a chemical scrubber for acidic wastes. The chemical composition of the South African and Columbian fly ash (SAFA, COFA) is given in Table 1. Table 1: the major components and Minor elements in the SAFA and COFA of the South African and Columbian fly ashes COFA*
SAFA*
54.4 20.8 1.05 6.18 4.65 2.05 0.12 0.05 0.75 0.13 9-7
40.9 31.4 1.75 3.05 8.35 2.45 0.05 0.02 1.95 0.35 5-4
Component %weight SiO2 Al2O3 TiO3 Fe2O3 CaO MgO K2O Na2O P2O5 SO3 C
Element ppm Ag As Ba Be Cd Co Cr Cu Mn Ni
COFA** SAFA** 9.5 <10 1,150 5.07 2 27 133 60 375 70
13.6 < 10 2,350 9.43 <2 40 150 77 360 68
Image 1: SEM images of the A- Cenosphere and B- Plerosphere
The fact that the fly ash has very good adsorption properties, due to the high surface area, created different types of utilization such as: structural filling for road construction and stabilization of soil in agriculture. Fly ash as a neutralization and fixation reagent for acidic waste Recently it was shown that the fly ash can act as an effective neutralization and fixation regent for dangerous acidic waste3. Two wastes have been tested: 1) Acidic Organic waste of regeneration processes of used motor oil. This waste contains more than 10M of acid waste per liter. The waste also contains high concentration of heavy and toxic metals.
2) Acidic waste from the phosphorus industry. This waste is a byproduct of melting the phosphate mineral with sulfuric acid. Results of these tests have shown that the fly ash is very effective as a neutralization and fixation reagent. The product created from these tests was a grey aggregate (sand like) which fixate the heavy metals effectively. The fixation quality was tested by two types of washing processes: the improved TCLP13114 and the CALWET5 All results showed that the concentration of leached heavy metals was under the drinking limit (DL). These results indicate that the fly ash might act as a potential scrubber for low activity radioactive waste.6 Radioactive wastes We had decided to study three different types of radioactive wastes: Cs137 which is one of the nuclear fission byproducts of U235 7 and has a half-life of 30.17 years. During the decay it emits β rays (0.19MeV) to form metastable nucleus of Barium (137m), which further decays fast (2.6 minutes) with emission of γ rays (0.60 MeV) to the stable isotope of Barium. Sr90 The strontium radionuclide is also one of the major nuclear fission byproducts of U235.8 It's half-life is 28.90 years, which during this time it emits β irradiation (0.546MeV) and decays to a stable isotope of Yttrium (Y90). Actinides The Actinides which are main byproducts during the fission are all radioactive. They are a derivative of the 6f shell and usually decay as an α emitters. A typical actinide is the plutonium – Pu239 which is created in nuclear plants7. Simulants of the radionuclides Working with radioactive material is very expensive, due to the high cost of protective measures needed to be taken. Thus we decided to work with simulants of the radionuclides. The simulants are: Cs133 (A stable isotopes of Cesium), Sr88 (A stable isotopes of Strontium) and Ce as a simulant to the actinide elements. Experimental section Materials and Equipment All water used in the study were DDW with aresistivty large than 10 MOhms/cm. Chemicals: The chemicals we used were of analytical grade: Sigma/Aldrich, Riedel de Haen, B.D.H and Merck. Fly ashes South Africa (SAFA) an Columbian (COFA) flyashes were supplied by the Utilities. The density of SAFA – 0.98gr/cm3 and that of COFA – 0.85gr/cm3. Gases- Nitrogen (99.95%, dry) and air (dry) were supplied by MAXIMA Ltd. Equipment The following instruments were used throughout this study: pH-meter – of El-Hama company model Cyberscan510. Calibration of this device was made by three buffer solutions : pH=4; pH= 7; pH=10.
ICP-OES – Induced Chemical Plasma Atomic Emission spectroscopy – from Varian, model 710-ES. Some of the cesium samples was analyzed in ICP-OES from Varian model VISTAPRO and some cesium samples were analyzed in ion chromatograph of Dionex model DX-500. Orbital Shakers from Cocono, model TS-400 Teflon filers for PET syringe from Tamar LTD, model FP-030/0.2μ. Analytical scales from Mettler Experimental method The experiment was carried in 6 steps (as shown in scheme 1) : 1) 2) 3) 4) 5) 6)
Preparing a simulant solutions (Cerium, Strontium or Cesium) in a known volume Mixing the solution with fly ash in a PET bottle. Mixing the mixture in the orbital shakers at 250rpm for different periods of time. Taking a sample and filter it. Measuring the pH and acidify the sample to prevent sedimentation. Measuring the concentration of the simulation element studied in the ICP-OES
Scheme 1 – work method
Results and Discussion Fixation of Ce3+ and Ce4+with SAFA Solutions of 20ppm of Cerium were prepared from Ce(NO3)3 or Ce(SO4)2.2(NH4)2SO4.2H2O salts and acidified with HCl to pH2. The mixture of the fly ash with the solution was mixed for 24 hours (Solid/Liquid ratio – 1/20, 20 gr fly ash + 400ml Cerium solution). The mixing was carried out in 500 ml PET bottles in orbital shakers at 250 rpm. The acidity at pH2 was necessary to prevent precipitation of Ce3+ or Ce4+in the solution. At the end of the mixing the solution was filtered and was acidified to a pH1 with 65% HNO3. Finally [Ce] was measured in the ICP-AES (as shown in tables 3, 4): Table 3: fixation of Ce3+ with SAFA in a 1:20 ratio (S/L) [Ce]f ppm
[Ce]0 ppm
SAFA weight (gr)
Shakin Exp. g number period <0.1 19 10 24 1 <0.1 19 10 24 2 19 19 0 0 3 * [Ce]0 concentration before mixing with flyash, [Ce]f concentration after mixing with flyash Table 4: fixation of Ce4+ with SAFA in a 1:20 ratio (S/L)
[Ce]f ppm <0.05 <0.05 <0.05
[Ce]0 SAFA Mixing Exp. ppm weight (gr) period number 20 20 24 4 20 20 96 5 1 + 0.1M 20 30 6 HCl * [Ce]0 concentration before mixing with flyash, [Ce]f concentration after mixing with flyash From the results in Tables 3,4 it is concluded that: There is a very efficient fixation of Ce3+ and Ce4+ in the fly ash, thus the fly ash can be a very effective fixation reagent for actinides (assuming Ce3+ and Ce4+can simulate actinide). The fixation mechanism is probably Coordination bonding. Fixation of Cesium Solution of 20ppm of Cesium was prepared from CsNO3 Salt, mixed with 20 gr of fly ash (1:20 S/L ratio) for different periods of time: 1–24 hrs. The mixture was in a 500 ml PET bottle on an orbital shaker at 250 rpm. At the end of the mixing the samples was filtered and acidify with 65% HNO3 to prevent precipitation. Analysis of [Cs+] was carried out in the solutions with ICP-AES (results shown in Table 5). Table 5: fixation of Cs with SAFA/COFA in a 1:20 ratio (S/L) Fly Ash Source
[Cs]f ppm
[Cs]0 ppm
SAFA Mixing Exp. weight time number (Gr) (hrs) COFA 15.1 20 20 1 7 COFA 14.1 20 20 24 8 SAFA 14.9 20 20 1 9 SAFA 14.4 20 20 5 10 N.A 20 20 0 0 11 * [Cs]0 concentration before mixing with flyash, [Cs]f concentration after mixing with flyash
From the results in Table 5 it is concluded that: There is an appreciable fixation of Cs+ ions in the fly ash thus that is why the fly ash can act as a fixation reagent for Cesium ions. The fixation mechanism is probably that the flyash behaves as a cation exchanger, The anionic groups on its surface are aluminate, silicates. Fixation of Strontium Solutions of 20ppm of Strontium were prepared with Sr(NO3)2 salt. Mixture of solution and SAFA was introduced into a 500 ml PET bottle with ratio of 1/20 (S/L). The mixture was mixed for different period of times: 5–96 hrs in an orbital shaker at 250 rpm. In one set of experiments Ex12-13, 0.5gms of Na2CO3 were also added to the solution and in the second set Ex 14-15, air was bubbled through the solution in order to dissolve carbon dioxide from the air into the basic solution. After mixing was completed the flyash was filtered from the solution and a sample from the solution was taken and acidified to pH1 with 65% HNO3 to prevent precipitation. Finally the concentration of the strontium was analyzed with ICP-AES (as shown in Table 6):
Table 6: fixation of Sr2+ with SAFA in a 1:20 ratio (S/L) [Sr]f ppm
[Sr]0 ppm
SAFA Mixing weight period EX. (Gr) (Hr) 0.15 10 20 5 12 0.15 10 20 5 13 0.16 10 20 5 14 0.12 10 20 5 15 9.86 10 20 0 16 * [Sr]0 concentration before mixing with flyash, [Sr]f concentration after mixing with flyash From the results in Table 6 it is concluded that: There is an efficient fixation of Sr in the fly ash. The fixation mechanism is probably via a formation of SrCO3 solid which interacts electro statically with the fly ash surface. The insoluble strontium salt is precipitated due to the carbonate anions present in the solution,. as a electrostatic precipitate, thus why we can assume that the strontium is fixed on the fly ash surface as SrCO3 . Conclusions The fly ash can act as a very good fixation reagent in three types of bonding mechanism. All of the simulants showed very effective fixation in the fly ash surface and much more detailed experiments have to be carried out to continue checking the effectiveness of the fixation. Acknowledgement We would like to thank the Israel Coal Ash Administration for funding the research, And to Ariel University center and Dr. Hanan Teller for supplying us the equipment and work space. Also acknowledgement to the Israel Geological Survey for chemical analysis and Ms. Natali Litbek for the SEM analysis. References [1] Ravina, D.; Lulav, O.: Coal Ash in Israel: AshTech Conference in Birmingham: 2006. [2] Internet site: http://www.coal-ash.co.il. [3] H. A. Foner, T.L. Robl, J.C. Hower and M. U. Graham, Characterization of Fly Ash from Israel with Reference to Its Possiblre Utilization, FUEL, Vol. 78, 215-223, 1999. [4] U.S.-Environmental Protection Agency (EPA): Toxicity characteristic leaching procedure (TCLP). Part 261, Appendix II-method 1311, 40 CFR Ch. 1,1990 [5] California Waste Extraction Test (CAL WET), Environmental Health Standards - Hazardous Waste, 66261.126 Appendix II. [6] T. Nokumura, The Material Flow of Radioactive Cesium 137 in the USA, 2000. [7] Internet site: www.epa.gov/rpdweb00/docs/source-nagement/csfinallongtakeshi.pdf. [8] Internet site: Strontium 90,k28 www.absoluteastronomy.com/topics/Strontium-90. [9] Internet site: http://www.wci-Coal.com.World Coal Institute. [10] Eli Lederman, "Coal Fly Ash as a Chemical Reagent for Scrubbing Industrial Acidic Waste", Ben Gurion University of the Negev, Israel, 2008.
Mobility of major and minor species in fly ash-brine co-disposal systems: upflow percolation test O.O. Fatoba1,*, W.M. Gitari3, L.F. Petrik1 and E.I. Iwuoha2 1
Environmental and Nano Sciences Research Group, Chemistry Department, University of the
Western Cape, Private Bag X17, Bellville, 7535, South Africa. Tel: +27 (0) 21-959-3878/3304, Fax: +27 (0) 21-959-3878/3304 2
SensorLab, Chemistry Department, University of the Western Cape, Private Bag X17, Bellville, 7535, South Africa 3
Department of Ecology and Resources Management, School of Environmental Sciences, University of Venda, Private Bag X5050, Thohoyandou, 0950, Limpopo, South Africa * Corresponding author: [email protected]
Abstract Apart from the generation of fly ash, brine (hyper-saline wastewater) is also a waste material generated in South African power stations as a result of water re-use. These waste materials contain major and minor species such as Na, K, Ca, Mg, Cl, SO4, Fe, Mn, Mo, Cr, As, Pb, Ni and Cu. The co-disposal of fly ash and brine has been practiced by some power stations in South Africa with the aim of utilizing the fly ash to capture the salts in brine. The effect of the chemical interaction of the species contained in these waste materials when co-disposed on the mobility of species in the fly ash-brine systems is the focus of this study. The up-flow percolation test was employed to determine the mobility of species in the fly ash-brine systems. The results of the analyzed eluates from the up-flow percolation tests revealed that some species such as Ca, Ba, Cr and Mo were released from the fly ash into the systems while some species such as Co, Cu, Zn, Pb, Mg, Mn, Na, Cl and SO4 were removed from the brine solution during the interaction with fly ash. The pH of the up-flow percolation systems was observed to have a significant effect on the mobility of major and trace species from the fly ash-brine systems. The study showed that some contaminant species such as Mg, B, Cl, SO4, Co, Pb and Zn were immobile and could be significantly removed from brine solution using fly ash.
Keywords: Fly ash; brine; co-disposal; mobility of species; interaction of species; up-flow percolation tests
1. Introduction The combustion of coal for power generation is on the increase due to the increase in demand for electricity in some countries in the world. Huge amounts of fly ash are produced as a result of the increase in coal combustion. According to US Coal Combustion Product (CCP) Production & Use Survey Report in 2008, US power plants produced about 70 million tons of fly ash annually, out of which nearly 45 % was beneficially used [1]. In South African, one of the power utilities produces about 40 million tons of fly ash annually [2], out of which less than 10 % is utilized. The fact remains that, despite its beneficial use for agricultural purposes, waste stabilization, additive to cement, road construction among others [3-6], significant amounts of fly ash are being disposed of in ash dump. Fly ash contains major and minor species such as Ca, Na, Mg, K, SO4, Cl, As, Pb, Cu, Cr, Fe, Mo, Mn etc., and these species could leach out in significant quantity from fly ash when in contact with aqueous solution [7-10]. The disposal of fly ash has been a major concern to coal-fired power stations due to the possible release of contaminants to the surrounding soils, surface and groundwater. Apart from the fly ash generated in the power stations in South Africa, brine (a hyper-saline wastewater), is also generated in significant quantities. Brine contains some major and trace species such as Na, Mg, K, SO4, Cl, As, Pb, Cu, Se, Cr, in significant quantities. The regulation guiding the disposal of brine has regulated its disposal due to its chemical composition thereby causing some power stations to co-dispose fly ash and brine with the aim of using fly ash to capture the species in brine. Despite several studies on the release of species from fly ash when in contact with water and acid mine drainage (AMD) [8-9, 11], studies have not been carried out on the mobility of species when fly ash is in contact with brine solution. This study is aimed at understanding the chemical interactions and mobility of species in fly ash-brine systems. The up-flow percolation tests were employed in this study and the eluates were analyzed for major and minor species.
2. Material and method
2.1 Up-flow percolation tests The test was carried out according to European standard method prEN14405 [12]. This method involves packing the material to be tested (fly ash) in a column in a standardized manner, and the brine is percolated in an up-flow direction through the column at a specified flow-rate to attain a fixed liquid/solid (L/S) ratios. This method was used to determine the release of constituents from fly ash packed in a column after a specific volume of brine had percolated through it. A continuous vertical up-flow was used in this study so that the column would be properly saturated with the brine and preferential percolation of the leachant would also be avoided.
Table 1: Concentration of species released from Secunda fly ash-brine up-flow percolation systems as a function of L/S ratio (concentration in mg/L) L/S ratio Element (mg/L)
UB
0.1
0.2
0.5
1
2
5
10
20
35
55
80
B
1.984
0.311
0.256
0.200
0.508
0.546
0.510
0.378
2.321
6.117
3.698
3.998
Ba
0.058
24.713
7.690
7.435
0.322
0.269
0.192
0.289
0.291
0.198
0.118
0.074
Ca
91.012
163.586
216.093
262.890
487.030
471.740
368.249
221.451
423.391
378.602
502.700
377.808
Co
0.008
BDL
BDL
0.000
BDL
BDL
BDL
0.001
0.003
0.006
0.002
0.006
Cr
0.015
1.150
1.968
2.895
1.199
0.164
0.024
0.015
0.031
0.027
0.013
0.016
Cu
0.189
1.303
1.104
0.962
0.550
0.219
0.148
0.091
0.033
0.027
0.023
0.026
0.116
0.074
0.100
0.103
0.093
0.155
0.114
BDL
0.148
0.059
0.042
Fe
0.102
Mg
147.500
0.213
0.065
0.054
0.072
0.062
0.092
0.167
0.351
0.801
1.418
10.662
Mn
0.002
0.005
0.003
0.001
0.002
0.001
0.002
0.002
0.002
0.002
0.001
0.002
Mo
0.038
1.306
1.352
1.486
0.432
0.229
0.193
0.088
0.059
0.057
0.173
0.075
Na
4323.220
4966.308
4491.683
4537.628
4415.768
4577.120
4342.918
4260.288
4225.258
4132.132
4283.775
4342.352
Ni
0.120
0.014
0.007
0.011
0.012
0.082
0.110
0.128
0.128
0.117
0.134
0.132
Pb
0.007
0.321
0.049
0.054
0.077
0.059
0.029
0.002
0.001
0.001
0.004
0.001
Zn
0.110
1.619
0.977
0.856
0.689
0.419
0.239
0.042
0.005
0.007
0.139
0.060
Cl
2424.000
1945.800
2019.233
1945.817
2217.500
2092.667
2173.450
1970.783
2028.050
2120.567
1778.647
1688.656
SO4
8858.000
49.250
199.233
639.500
3542.750
3779.000
3928.817
6854.350
7764.600
7761.817
7299.228
7256.240
7.89
13.21
13.6
13.67
11.95
11.56
12.73
11.34
10.79
10.61
9.38
9.24
14.63
34.75
34.25
33
26.5
27.65
25.65
18.62
18.13
15.48
17.1
17.01
pH EC (mS/cm)
Table 2: Concentration of species released from Tutuka fly ash-brine up-flow percolation systems as a function of L/S ratio (concentration in mg/L) L/S ratio Element (mg/L)
UB
0.1
0.2
0.5
1
2
5
10
20
35
55
80
B
1.984
0.378
BDL
0.087
0.733
1.695
2.081
2.547
3.364
4.204
2.915
2.855
Ba
0.058
0.346
0.283
0.280
0.246
0.200
0.159
0.186
0.115
0.057
0.050
0.038
Ca
91.012
92.995
194.811
203.465
291.754
335.976
339.748
323.169
227.742
143.986
115.942
115.013
Co
0.008
0.005
0.001
0.001
0.002
0.002
0.003
0.004
0.007
0.010
0.020
0.019
Cr
0.015
9.827
0.971
0.957
0.214
0.241
0.229
0.239
0.126
0.112
0.013
0.013
Cu
0.189
0.105
0.037
0.044
0.049
0.055
0.059
0.051
0.052
0.081
0.092
0.089
Fe
0.102
0.115
0.298
0.170
0.131
0.100
0.102
0.128
0.083
0.091
0.040
0.028
Mg
141.961
0.313
0.118
0.109
0.104
0.108
0.492
1.634
76.146
114.629
170.435
169.940
Mn
0.002
0.002
0.002
0.002
0.002
0.002
0.001
0.002
0.002
0.003
0.005
0.002
Mo
0.038
4.085
0.320
0.299
0.076
0.087
0.085
0.093
0.067
0.062
0.042
0.041
Na
4323.220
1322.633
3937.013
3962.138
4445.447
4429.410
4478.548
4268.860
4348.297
3942.667
4499.293
4462.698
Ni
0.120
0.053
0.110
0.114
0.118
0.129
0.117
0.129
0.125
0.113
0.132
0.130
Pb
0.007
0.005
0.001
0.002
0.003
0.001
0.003
0.002
0.002
0.001
0.002
0.001
Zn
0.110
0.091
0.052
0.047
0.042
0.048
0.026
0.027
0.029
0.075
0.148
0.147
Cl
2424.000
998.617
1820.983
1820.983
1821.000
1901.767
1894.417
1947.283
1951.683
2149.933
1817.323
1753.438
SO4
8858.000
1707.367
6561.683
6688.483
7358.783
7380.233
7483.617
7804.200
7767.650
7661.950
7347.974
7283.133
7.89
7.83
12.44
12.49
10.86
11.62
11.59
10.55
9.55
9.36
8.71
8.74
14.63
5.6
13.23
13.73
13.97
14.36
14.57
17.43
17.67
15.45
17.33
17.05
pH EC (mS/cm)
3. Results and discussion
3.1 Major elements in the fly ash-brine leachates The concentration of Ca in UB was about 91 mg/L while SO4 levels were exceptionally high, being nearly 9000 mg/L (Tables 1 and 2). The concentration of Ca in leachates from both Secunda and Tutuka up-flow percolation systems (Tables 1 and 2) slightly increased after equilibration at the beginning of the tests compared with its concentration in the unreacted brine solutions. The concentration of Ca gradually increased until a maximum of about 500 mg/L and 350 mg/L was reached at L/S 55 and 5 in Secunda and Tutuka systems respectively. The trend of Ca released from the fly ash-brine systems showed that apart from the initial dissolution of readily soluble Ca-rich phases such as CaO, the dissolution of Ca-rich phases locked in the fly ash matrix contributed to the gradual increase in Ca concentration. The concentration of SO4 in the leachates from both Secunda and Tutuka systems showed a very significant decrease after equilibration at the beginning of the up-flow percolation tests compared with the concentration of SO4 in unreacted brine solutions (UB) (Tables 1 and 2). This indicates that SO4 was removed from the brine during and after the equilibration. The trends of SO4 showed that SO4 was removed from the brine solutions by the fly ashes as a result of precipitation of SO4 bearing phases. The removal of SO4 from the systems especially at the beginning of the tests could be attributed to the formation of Ca-SO4-rich mineral phases such as gypsum [13]. Na concentration was very high in UB, being approximately 4323 mg/L while Cl level was high being 2424 mg/L. Two different trends were observed for the release of Na into the leachates from Secunda (Table 1) and Tutuka (Table 2) up-flow percolation systems. A slight increase in the concentration of Na (approximately 5000 mg/L) was observed in Secunda systems (Table 1) after equilibration while initially, at L/S 0.1, a significant decrease to about 1000 mg/L in Na concentration was observed in Tutuka (Table 2) systems immediately after equilibration and at the beginning of the up-flow percolation tests. After the initial increase in the concentration of Na in Secunda system, a slight and slow decrease (slightly lower than the levels in UB) was observed and this decrease in concentration (>4000 mg/L) was maintained throughout the period of the tests. The slight increase of Na at L/S 0.1 in Secunda systems when compared with the
concentration in UB, could be attributed to the leaching of Na from the fly ash during the three days equilibration period while the initial decrease in Na concentration in Tutuka systems immediately after the equilibration period may be as a result of the formation of transient Na-rich phase upon contact with fly ash due to the super-saturation of Na in the brine containing systems. The nearly immediate dissolution of the transient Na-rich phase or the reduction in the capacity of the fly ash to capture more Na from the brine due to continuous flow of brine may account for the increase in Na concentration in Tutuka system at L/S 0.2. These trends indicate that after the initial removal of Na, the capacity of the fly ashes to remove more Na from the brine solutions during the tests reduced as a result of continuous flow of fresh brine solution. The trend of Cl was similar to that observed for Na in the leachates. However, the concentration of Cl in the leachates after contact with fly ash was lower than the concentration in UB throughout the period of the up-flow percolation tests (Tables 1 and 2). This indicates that some Cl was removed from the brine solution throughout the period of the up-flow percolation tests. Cl was initially removed from the brine solution in Tutuka systems after equilibration. The removal patterns of Cl in Tutuka systems corresponded to those of Na indicating that the removal of Cl from the systems could be controlled by the formation of a transient halite (NaCl) phase at the beginning of the tests. Due to the sufficient concentration of Ca and Al in the upflow percolation systems as a result of matrix dissolution of fly ash, the probable formation of Friedel’s salt [Ca4Al2Cl2(OH)12·4H2O] could also account for the reduction in the concentration of Cl after the initial stages in the systems [14-16].
3.2 Minor elements in the fly ash-brine leachates Low levels of Fe and Mn were present in UB, 0.1 mg/L and 0.002 mg/L respectively. The trend of Fe and Mn released into the leachates of the up-flow percolation tests was almost the same as no significant difference was observed in the concentrations of these species in the leachates when compared with their concentrations in UB (Tables 1 and 2). The slight increase in the concentrations of Fe and Mn at the beginning of the tests could be attributed to the partial dissolution of their oxides from the fly ash surface when in contact with the brine solutions. The slight fluctuation observed in the release of Fe and Mn could be attributed to the dissolution and
precipitation of Fe and Mn-oxyhydroxides in the systems due to the high pH of the systems. It was suggested that the concentrations of the Fe and Mn in fly ash leachates are controlled by precipitation of their hydroxides [17]. The increase in the concentration of Mn in Tutuka system over time could be attributed to the decrease in the pH of the systems over time which may have caused the dissolution of Mn precipitates from the matrix of the fly ash. Although present in minor concentration, the levels of Pb and Ni were an order of magnitude higher than Co concentration in the unreacted brine (UB). Pb being present in concentration of approximately 0.3 mg/L, Ni levels was around 0.1 mg/L but Co concentration being approximately 0.009 mg/L (Tables 1 and 2). Co was immediately removed from the brine solutions after contact with Secunda fly ash (Table 1) but increased after L/S 35 to around 0.006 mg/L. In the case of Tutuka systems, Co levels decreased from approximately 0.01 mg/L to 0.005 mg/L. Co removal in the leachates of the up-flow percolation systems was initially significant compared to the concentration of Co in UB. This indicates that Co was removed to levels below 0.001 mg/L from the brine solutions by interactions with the fly ashes at L/S 0.1-10 during the interactions tests. Thereafter, the concentration of Co gradually increased in the leachates of Secunda systems at L/S 10, and at L/S 1 in the case of Tutuka systems. The initial removal of Co from the brine solutions during the percolation tests despite the continuous inflow of brine solution indicates that Co was removed as a result of transient precipitation of Cocontaining phases or co-precipitated with other phases in the systems. Ni was removed from the brine solution during the initial stages of the up-flow percolation systems. The concentration of Ni in the leachates was lower than what was observed in the unreacted brine solutions (UB) from the beginning of the tests until L/S 10 in the case of Secunda systems and until L/S 2 in the case of Tutuka systems. This implies that Ni was initially removed from the brine solutions by the fly ashes. Ni has been observed not to be easily hydrolyzed [18] which could suggest that its removal at the beginning of the tests may not be as a result of precipitation of transient Ni-containing mineral phases in the systems. However, the removal of Ni from the systems at the beginning of the test could be due to adsorption of Ni on the negatively charged surface of the fly ash particles [19]. The fact that Ni is weakly adsorbed on the surface of fly ash particles [18] could lead to easy desorption of Ni over time thereby
increasing its concentration in the leachates. The lowering of the pH (≈9) of the up-flow percolation systems towards the end of the tests could lead to the leaching of adsorbed Ni into the systems.
4. Conclusion The removal and subsequent release of some species in the systems could be as a result of precipitation followed by dissolution of the transient secondary mineral phases formed during the interactions. Ca, Ba, Sr, Cr and Mo showed a continuous leaching throughout the percolation tests, which indicates that these species leached continuously from the fly ashes by contact with brine flows. However, the study has shown that some contaminant species such as Mg, B, Cl, SO4, Co, Pb and Zn were immobile and could be significantly removed from brine solution using fly ash. The pH of the up-flow percolation systems has been shown to have a significant effect on the mobility of major and trace species from the fly ash-brine systems.
References [1]. American Coal Ash Association, 2008: website: http://www.acaa-usa.org. [2]. ESKOM Annual Report, 2009: http://www.eskom.co.za/annreport09/annreport09/ [3]. Iyer, R.S. and Scott, J.A. 2001: Power station fly ash- a review of value added utilization outside of the construction industry. Resources Conservation and Recycling 31: 217-228. [4]. Kumpiene, J., Lagerkvist, A. and Maurice, C. 2006: Stabilization of Pb- and Cucontaminated soil using coal fly ash and peat. Environmental pollution 145: 365-373. [5]. Campbell, A.E. 1999: Chemical, physical and mineralogical properties associated with the hardening of some South African fly ashes. Unpublished MSc Thesis, University of Cape Town, South Africa. [6]. Foner, A.H., Robl, L.T., Hower, C. and Graham, M.U. 1999: Characterization of fly ash from Israel with reference to its possible utilization. Fuel 78: 215-223. [7]. Ilic, M., Cheesman, C., Sollars, C. and Knight, J. 2003: Mineralogy and microstructure of sintered lignite coal fly ash. Fuel 82: 331-6.
[8]. Baba, A. and Kaya, A. 2004: Leaching characteristics of solid wastes from thermal plants of western Turkey and comparison of toxicity methodologies. Journal of Environmental Management 73: 199-207. [9]. Polettini, A. and Pomi, R. 2004: The leaching behaviour of incinerator bottom ash as affected by accelerated ageing. Journal of Hazardous Materials 113: 209-215. [10]. Adriano, D.C., Page, A.L., Elseewi, A.A., Chang, A.C., and Straughan, I. 1980: Utilization and disposal of fly ash and other coal residues in terrestrial ecosystems: A review. Journal of Environmental Quality 9: 333-344. [11]. Gitari, W.M., Petrik, L.F., Etchebers, O., Key, D.L., Iwuoha, E. and Okujeni, C. 2006: Treatment of acid mine drainage with fly ash: Removal of major contaminants and trace elements. Journal of Environmental Science and Health, Part A 41: 1729-1747 [12]. prEN14405. 2003: Leaching behaviour test – Up-flow percolation test - Horizontal standard. [13]. Fatoba, O.O. 2008: Chemical compositions and leaching behaviour of some South African fly ashes. Unpublished MSc Thesis, University of the Western Cape, South Africa. [14]. Suryavanshi, A.K. and Swamy, R.N. 1996: Stability of Friedel’s salt in carbonated concrete structural elements. Cement and Concrete Research 26 (5): 729-741. [15]. Bothe, J.V. and Brown, P.W. 2004: PhreeqC modeling of Friedel’s salt equilibria at 23±1 degrees C. Cement and Concrete Research 34 (6): 1057-1063. [16]. Hyks, J., Astrup, T. and Christensen, H. 2009: Long-term leaching from MSWI airpollution-control residues: Leaching characterization and modelling. Journal of Hazardous Materials 162: 80-91. [17]. Roy, W.R. and Griffin, R.A. 1984: Illinois basin coal fly ashes. 2. Equilibria relationships and quantitative modelling of ash-water reactions. Environmental Science and Technology 18 (10): 739-742. [18]. Rio, S., Delebarre, A., Hequet, V., Le Cloirec, P. and Blondin, J. 2002: Metallic ion removal from aqueous solutions by fly ashes: multicomponent studies. Journal of Chemical Technology and Biotechnology 77: 382-388. [19]. Ram, C.L., Srivastava K.N., Tripathi C.R., Thakur K.S., Sinha K.A., Jha K.S., Masto E.R. and Mitra S. 2007: Leaching behaviour of lignite fly ash with shake and column tests. Environmental Geology 51: 1119-1132.
Oviedo ICCS&T 2011. Extended Abstract
Removal of Arsenate from Water by Adsorption onto Lignite
Yuda Yürüm* and Z. Özlem Kocabas Faculty of Natural Science and Engineering, Sabanci University, Orhanlı 34956 Tuzla, Istanbul/Turkey Phone: 90 216 4839512, Fax: 90 216 4839550 *Corresponding author ([email protected] Abstract Carbonaceous material has been extensively tested in environmental applications, especially in separation technologies. In the present work, a laboratory study was conducted to investigate the ability of lignite for the removal of arsenate (As(V)) from water. Lignite was characterized by x-ray diffraction and scanning electron microscopy (SEM) and batch adsorption experiments were performed with respect to pH, contact time, kinetics, initial arsenic concentration and adsorption isotherms. The maximum As(V) uptake were found 1.01 mg/L at pH 3.26 when 0.2 g/L lignite was used at an initial arsenic concentration of 1 mg/L. According to kinetic sorption data, the higher regression coefficients (R2) were obtained after the application of pseudo-second order to the experimental data of As(V) initial concentrations. The result of the sorption experiment, which takes into consideration the effects of equilibrium concentration on adsorption capacity, was analyzed with two popular adsorption models, Langmuir and Freundlich models. According to the regression coefficient values for As(V) adsorption onto lignite, the Langmuir model (R2 = 0.998) described the isotherm better than the Freundlich model (R2 = 0.745) at pH 5, respectively. The maximum arsenic uptake (qmax) value computed from slope of the linearized Langmuir plot was 2.865 mg/g for the adsorption of As(V) onto lignite. This study proposed the use of lignite as a potential adsorbent material for waters contaminated with As(V) ions. Keywords: lignite; arsenic removal; kinetics; adsorption isotherm
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1
Oviedo ICCS&T 2011. Extended Abstract
1. Introduction Arsenic is one of the most toxic contaminants found in the environment and management of the hazardous element has become a major public concern in the twentieth century [1]. The significantly high contamination level of arsenic has been reported for many countries as USA, China, Chile, Bangladesh, Taiwan, Mexico, Argentina, Poland, Canada, Hungary, Japan, and India [2]. Due to its toxic and carcinogenic effects on human beings, the contamination level of arsenic in water has taken serious consideration by environmental authorities. According to World Health Organization (WHO), 10 µg/L has been adopted as the new contamination level of arsenic in drinking water [3]. In natural water, arsenic occurs both in organic and inorganic forms. Although organic arsenic is detoxified by methylation process, inorganic arsenic is needed a wellestablished treatment. Inorganic arsenic exits in -3, 0, +3 and +5 oxidation states in aquatic systems. The elemental state 0 and -3 are quite rare as compared to +3 and +5 oxidation states. Especially, As(III) has greater toxicity and mobility than As(V) [4]. There are various technologies for removal of arsenic from water, such as coagulation, membrane processes, ion exchange and adsorption [5]. Nowadays, removal of arsenic by adsorption is acquired importance due to its technical simplicity and applicability in rural areas, where people are more subjected to polluted drinking water with arsenic [6]. The main objective of this study is to investigate an effective and inexpensive water purification system by using lignite. With this aim, experimental study evaluates the efficiencies of arsenic removal from water by adsorption process under various experimental conditions including adsorbent to arsenic ratio, pH and contact time to determine the optimum process conditions.
2. Experimental
Materials and Reagents Lignite supplied from Soma Basin in Turkey was washed with distilled water and boiled for 30 min, then dried at 100oC in oven for 32 h and stored in a desiccator for further analysis and experiments. Analytical grade hydrochloric acid, sodium hydroxide was purchased from Merck. All the solutions were prepared by deionized water using QSubmit before 31 May 2011 to [email protected]
2
Oviedo ICCS&T 2011. Extended Abstract
H2O, Millipore Corp. deionizer until 18.2 MΩcm of resistivity. Na2HAsO4.7H2O salt (99.9 % from Sigma-Aldrich) was dissolved in deionized water for 50 mg/L As(V) stock solution. The working solutions were prepared by diluting these stock solutions with deionized water. All stock solutions were prepared weekly and frozen to prevent oxidation.
Adsorption Experiments Experiments for the effect of varied pH on adsorption efficiency were carried out by adjusting pH values of 1 mg/L As (V) solution from 2.0 to 9.0 with the addition of 0.2 g/L adsorbent material. The pH of solutions was adjusted to specified values with diluted HCl or NaOH, and the mixtures were shaken in an incubator shaker at 180 rpm mixing rate for 24 h at 25oC. At the end of experiment, the solution was separated from the solid adsorbent by using 0.45 µm PVDF membrane filter. In kinetic experiments, 100 mL of 1 mg/L As(III) or As (V) solution with 0.02 g/L adsorbent material at pH 5 were used and sampled at different time intervals. A series of batch adsorption tests were conducted by using a known amount of adsorbent with 50 mL aqueous solution of As (V) of desired initial concentrations of 0.4-8 mg/L at pH 5 and 7 to develop adsorption isotherms. Each adsorption experiment in the present study was repeated three times and average values were reported below.
Arsenic Analyses Arsenic concentrations of the solutions were measured with a Varian, Vista-Pro CCD simultaneous inductively coupled plasma ICP-OES spectrophotometer. Samples before and after adsorption experiments were analyzed to obtain residual arsenic concentration.
3. Results and Discussion
3.1.
Effect of pH
The pH of aqueous solution can affect considerably the removal of As(V) ions, and the formation of surface charge groups onto adsorbent. The adsorptions of an initial As(V) concentration of 1 mg/L onto lignite has been conducted using the initial pH range (pHinitial) of 2.0-9.0. After completion of reaction, equilibrium pH (pHfinal) values as well as the arsenic uptake, qe (mg/g) were measured and results were shown in Fig. 1.
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3
Oviedo ICCS&T 2011. Extended Abstract
10
3.0 As(V) pHFinal
2.5
9 8 7 6
1.5
5 1.0
4 3
0.5 0.0
pHFinal
qe (mg/g)
2.0
2 1
2
3
4
5
6
7
8
9
10
pHInitial
Figure 1. The pHfinal and arsenic uptake, qe as a function of pHinitial for adsorptions of As(V) onto lignite (Initial arsenic concentration: 1mg/L, adsorbent amount: 0.2 mg/L, temperature: 25oC and contact time: 24 h). The predominant As(V) species show variations with pH values: H3AsO4 (pH 0.0-2.0), H2AsO4- (pH 3.0-6.0), and HAsO42- (pH 7.0-11.0) [7]. The maximum arsenic uptakes were found 2.74 mg/g for As(V) at pH 5.87.
3.2.
Adsorption Kinetics
The kinetic studies were conducted in order to understand the adsorption behavior of the lignite by taking subsamples at different time intervals. The experiments were performed with 0.02 g of adsorbent material and 100 mL of 1 mg/L As(V) solutions 1, 5, 10, 20, 40, 60, 100, 140, 180, 300, 450, 720 and 1080 min reaction times at pH 5. The plot of arsenic uptake versus adsorption time t for As(V) was presented in Fig. 2. Since a large number of adsorption sites at the initial stage of adsorption, the arsenic uptake increased fast in followed by a slower increase before achieving a plateau.
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4
Oviedo ICCS&T 2011. Extended Abstract
2.8
qe (mg/g)
2.4
2.0
1.6
1.2
0
200
400
600
800
1000
1200
Time (min)
Figure 2. Arsenic uptake on lignite versus adsorption time (Initial arsenic concentration: 1 mg/L, adsorbent amount: 0.2 mg/L, pH: 5, temperature: 25oC). The kinetic data obtained from this study were first analyzed by employing the pseudofirst order (1) and the pseudo-second-order (2) equations [8, 9]; (1)
log(qe - qt)= log(qe) – (kad.t) / 2.303 t/qt = (1/k2qe2) + t/qe
(2)
where, qt is the amount of arsenic adsorbed (mg/g) at time t, qe is the maximum adsorption capacity (mg/g) for the pseudo first-order adsorption and pseudo second order adsorption, kad is the pseudo-first-order rate constant for the arsenic adsorption process (1/min), k2 is the pseudo second order rate constant (g/mg.min). The experimental data were fitted with both the linearized pseudo first order and linearized pseudo second order models. The parameters calculated from the pseudo first order and pseudo second order linearized equations were listed in Table 1. Table 1. Parameters of pseudo first order and pseudo second order kinetic models for adsorption of As(V) on lignite at pH 5.
Pseudo first order
Pseudo second order
qe (mg/g)
kad (min-1)
R2
0.995
0.072
0.963
q2 (mg/g)
k2 (g/mg min)
R2
2.834
0.023
0.999
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5
Oviedo ICCS&T 2011. Extended Abstract
3.3.
Adsorption Isotherms
The adsorption of arsenic, initially ranging from 0.4 to 8 mg/L concentration onto lignite has been performed at pH 5 and pH 7 using 0.2 g/L adsorbent material in the batch mode. A time of 720 min. was used to reach equilibrium in the experiments to construct the adsorption isotherms. The relationship between the adsorption capability, (qe) and equilibrium concentration, (Ce) of arsenic onto lignite was shown in Fig. 3. It was apparent that the extent of As(V) adsorption at pH 5 was higher than that at pH 7 over the entire range of equilibrium concentration values.
pH 5 pH 7
qe (mg/g)
3
2
1
0
1
2
3
4
5
6
7
8
9
Ce (mg/L)
Figure 3. Adsorption isotherm of As(V) on the lignite (Initial arsenic concentration: 0.4-8 mg/L, adsorbent amount: 0.2 mg/L, pH: 5 and 7, temperature: 25oC and contact time:12 h). In this study, the Langmuir and Freundlich isotherm models were investigated to evaluate adsorption patterns of As(V) on lignite with respect to its concentration of equilibrium in solutions. The linearized forms of the Langmuir, (equation (3)) and Freundlich, (equation (4)), isotherms are; Ce / qe = 1 / (qmax L) + Ce / qmax ln qe = ln Kf + 1/n ln Ce
(3) (4)
where qe is the amount adsorbed on solid (mg/g), Ce is the equilibrium solution concentration (mg/L), qmax is adsorption capacity (mg/g), L is a constant related to enthalpy of sorption which should vary with temperature (L/mg), Kf, (mg/g) is related to the adsorption capacity of the adsorbent and 1/n is a constant known as the heterogeneity factor is related to surface heterogeneity.
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6
Oviedo ICCS&T 2011. Extended Abstract
3.0 pH 5
pH 7
10
2.5
8
1.5
Ce/qe
Ce/qe
2.0
1.0
4
y=0.349x+0.045 2 R =0.998
0.5
0
1
2
3
4
5
6
7
y=1.226x+0.354 2 R =0.997
2
(a)
0.0
6
(b) 0
8
0
1
2
3
4
5
6
7
8
9
Ce (mg/L)
Ce (mg/L)
Figure 4. Linearized Langmuir isotherm of As(V) on lignite for a) pH 5 b) pH 7 According to the regression coefficient values for As(V) adsorption onto lignite for the two different pH values, the Langmuir model described the isotherm better than the Freundlich model. Additionally, the maximum arsenic uptakes were obtained for both pH values after the application of linearized form of Langmuir isotherm to the experimental data of As(V) initial concentrations (Fig. 4). The maximum adsorption capacity of lignite at pH 5 was about approximately 2.79 mg/g which was higher than that at pH 7 for As(V). Table 2 summarized the results of the adsorption parameters. Table 3. Calculated isotherm parameters for As(V) adsorption onto lignite at pH 5 and pH 7 Isotherm models
Isotherm parameters qmax (mg/g)
L (L/mg)
R2
pH 5
2.865
7.756
0.998
pH 7
0.816
3.462
0.997
Kf (mg/g)
n
R2
pH 5
2.195
5.081
0.745
pH 7
0.550
4.643
0.764
Langmuir
Freundlich
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Oviedo ICCS&T 2011. Extended Abstract
4. Conclusions In the current study, evaluation of the As(V) removal efficiency of lignite was performed considering effect of pH and contact time. According to kinetic sorption experiments, the higher regression coefficients (R2) were obtained after the application of pseudo-second order to the experimental data of As(V) initial concentrations. The most important observation was that the higher adsorption capacities achieved from Langmuir isotherm model was found for As(V) at pH 5 as compared to at pH 7 and the calculated maximum adsorption capacity values was 2.865 mg/g for As(V) pH 5.
Acknowledgement We greatly acknowledge the support of Prof. Ismail Çakmak and ICP Laboratory head technician Mr. Veli Bayır of the Biological Sciences and Bioengineering Program at Sabanci University for providing us the measurement facilities of inductively coupled plasma (ICP) spectrometer.
References [1] Henke KR. Arsenic: Environmental Chemistry, Health Threats and Waste Treatment, John Wiley & Sons, Ltd Publication, 2009. [2] Jain CK, and Ali I. Arsenic: occurrence, toxicity and speciation, Water Resources Res 2000; 34:4304–12. [3] Agency, U.E.P. Office of Ground Water and Drinking Water, 2002. [4] Maiti A, DasGupta S, Basu JK, De S. Batch and Column Study: Adsorption of Arsenate Using Untreated Laterite as Adsorbent, Ind Eng Chem Res 2008;47:1620-9. [5] USEPA, Capital Costs of Arsenic Removal Technologies: U.S. EPA Arsenic Removal Technology Demonstration Program Round 1, (EPA, 2004b). EPA, Editor. 2004. [6] Song S, Lopez-Valdivieso A, et al. Arsenic removal from high-arsenic water by enhanced coagulation with ferric ions and coarse calcite, Water Res 2006;40:364-72. [7] Elizalde-Gonzalez MP, Mattusch J, Einicke WD, Wennrich R. Sorption on natural solids for arsenic removal. J. Chem Eng 2001;81: 187-95. [8] Yurum Y, Dror Y, Levy M. Effect of Acid Dissolution on the Mineral Matrix and Organic-Matter of Zefa Efe Oil-Shale. Fuel Process Technol 1985;11:71-86. [9] Kannan N, Sundaram MM. Kinetics and mechanism of removal of methylene blue by adsorption on various carbons - a comparative study. Dyes Pigments 2001;51:25-40. Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
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Oviedo ICCS&T 2011. Extended Abstract
Biomass and Mineral Coal in South Brazil: Potential Use for Energy Generation in Bubbling Fluidized Bed Gomes G.M.F.1, Dalla Zen L.2, Vilela A.C.F.1, Osório E.1 1
LASID. Universidade Federal do Rio Grande do Sul (UFRGS). Av. Bento Gonçalves, 9500. 03306-000 - Porto Alegre - RS – BRASIL. [email protected] 2 Fundação de Ciência e Tecnologia do Estado do Rio Grande do Sul (CIENTEC). Av. das Indústrias, 2277 – Distrito Industrial – Laboratório de Combustão – Cachoeirinha – RS – Brasil. Abstract Due to the large variety of available natural resources, Brazil can be considered a country that relies on several and different options for primary energy generation. From this perspective, biomass has been under-utilized within the available energy matrix. The State of Rio Grande do Sul, one of the 27 federal units located in southern Brazil possesses, per its agricultural characteristics and soil use for silviculture, an excellent potential for the use of biomass as a decentralized energy source. Some examples are the availability of wood and rice husk in specific regions and during all the year. Yet, the mineral coal exploitation is located in places near the availability of those biomasses, making the use of fluidized bed a good option for energy generation. This work makes an analysis of the possibilities of using those kinds of biomass with high availability in the State of Rio Grande do Sul as primary fuels. The data obtained shows that there is an availability of 155, 488 and 466 MWth of availability of rice husk, wood waste and mineral coal, respectively.
1. Introduction Due to the need to reduce greenhouse gas emissions and the world dependence on fossil fuels, the use of biomass as an energy source has increased during recent years, being accompanied by parallel technology development and a consequent cost reduction. Latin America has shown the greatest increase in biomass use for energy generation worldwide, with an average of 2.28% per year during the period between 1999 and 2003. For comparison, the world’s annual average increase was 0.57% during the same period [1]. In Brazil, the energy matrix has always been internationally distinct due to the widespread sharing of renewable energy sources; initially a consequence of hydroelectric enterprises and later of the use of sugar cane alcohol as an automotive fuel.
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Oviedo ICCS&T 2011. Extended Abstract
Over a long-term period, it has been necessary to prepare the country for a change from experiencing predominant hydroelectric energy expansion to realizing an expansion with totally different characteristics, which must recruit the increased use of renewable energy sources, such as eolic and biomass energy [2]. In 2007, the energy availability in Brazil achieved a capacity of 238.8 tep, having a portion of 45.9% of renewable energy, far above the world’s renewable energy average – 12.9%. The biomass portion, in turn, had the contribution of 15.9% of sugar cane products and 12.0% of firewood and charcoal [3]. Each year, about 330 millions of metric tons of biomass residues are generated in the country. However, apart from its use as fuel source, the use of biomass residues as energy source has not had a considerable attention, partly due to the lack of an adequate program and part due to the inherent difficulties in the use of these residues – they have poor energy characteristics causing high costs during transportation, handling and storage [3]. Among the options for the Brazilian energy matrix, biomass is rather interesting due to its diversification throughout the vast Brazilian territory and the possibility of its use near its place of generation. The State of Rio Grande do Sul, located in the southern region of Brazil, is one of the Federative Units and is the southern-most state of Brazil, lying between the latitudes 27° and 34°S and the longitudes 50° and 57°W. Its geography has contributed to its farming and cattle production, especially the production of grains and livestock. In recent years, the use of soil for forestry, mainly Eucalyptus and Pinus, has gained considerable importance. From the above considerations, the generation of agricultural and forestry waste in the State of Rio Grande do Sul may possess a high potential for use in primary energy generation. In the State of Rio Grande do Sul, 80% of the thermal power plant projects produce between 300 and 600 kWe. Still, according to Mayer (2009) [4], in this region of Brazil, there is a great coincidence of geographical distribution between the production of biomass waste (rice husk and wood) capable of being employed for auto generation and repressed demand for electricity, possibly caused or determined by the low index of industrialization and an aggregation of value. Moreover, there is, through the concessionaries, interest in buying surplus electrical energy in regions of such repressed demand [5]. Based on these considerations, this article rovides an analysis of the possibilities of
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Oviedo ICCS&T 2011. Extended Abstract
using two types of biomass in the State of Rio Grande do Sul as primary fuel: rice husk and wood. Furthermore, as the national mineral coal mines are located in southern Brazil, we also describe the potential use of coal for energy— in particular, due to the possibility of cofiring the biomasses described with mineral coal using fluidized bed technology for energy generation.
2. Availability of solid fuels in southern Brazil 2.1 Rice husk The state of Rio Grande do Sul is, traditionally, an agricultural state, taking a large stake in the brazilian primary production. As a result, the potential for the exploitation of agricultural wastes is extremely significant, as shown in Figure 1, where one can observe the energy potential of the three southern states of Brazil. As can be seen, Rio Grande do Sul (RS) has a distinguished territory, with high potential for the generation of energy from farm waste across almost all its southern half.
Figure 1: Estimate of potential energy from the use of agricultural waste in southern Brazil [6]. The State of Rio Grande do Sul is the main producer of rice in Brazil, having produced 7.9 million tons in 2009, equivalent to 63% of the national production [7]. It is important to highlight that Rio Grande do Sul has 24 municipalities among the largest producers of rice in the country, where a yield of 7.5 kg/ha has been reached, which is
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Oviedo ICCS&T 2011. Extended Abstract
higher than the southern regional average of 5.85 kg/ha and the national average of 3.38 kg/ha. A great disadvantage of using rice husk as a fuel is the high cost of transport due to its low density. However, the decentralized use of this residue for energy generation could eliminate such disadvantage. When used on a scale exceeding 1 MW, it is possible to achieve economic feasibility with the use of rice husk [5].
2.2 Forest wastes In 2008, primary forestry production in Brazil added US$ 7.5 billion. Of this total, 69.3% was from the forestry segment, and 30.7% was from extractivism. In the production of wood from silviculture, the State of Rio Grande do Sul was the main producer, providing 33.9% of the more than 42 million cubic meters produced in the country. In paper and cellulose derived from forest plants, the State had a national contribution of 5.0%, adding almost 3 million cubic meters, and it still had a significant production of wood in torah, firewood for charcoal and other uses [8]. Considering that all of these activities produce waste upon holding and milling, we can observe the high potential for generating biomass from forest residues for energy use. Moreover, the State of Rio Grande do Sul (RS) has an area comprising 6% and 10% of the total Eucalyptus and Pinus, respectively, planted in Brazil [9]. Also, taking into account the density and the heating value, there is an energetic availability of 15.4x106 GJ/yr of wastes from Eucalyptus in RS out of 287.6 x106 GJ/yr for all the country [3]. The success of reforestation programs that have been carried out in Brazil may be attributed to the relatively high productivity of forests, the low cost of manual labor and the low price of soils. High yields have been reached due to the planting of rapidly growing exotic species of Eucalyptus and Pinus [10]. 2.3 Mineral Coal In Brazil, lignite and sub-bituminous coals can be found. However, the mines that can be exploited are located in the southern region, in the States of Rio Grande do Sul, Santa Catarina and Parana. Approximately 88% of the resources are located in Rio Grande do Sul. The most important ones, listed geographically from southwest to northeast, are Candiota, Capané, Irui, Lion, Charqueadas, Morungava/Chico Lomã, Santa Terezinha, as well as mines located in the State of Rio Grande do Sul, as shown in Figure 2 [11]. The coal varies, by rank, from southwest to northeast, from bituminous
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Oviedo ICCS&T 2011. Extended Abstract
high volatile C to bituminous high volatile A [12]. The main source is Candiota, located in the southern half of the state, which is also the largest Brazilian ore bed. In this place, the layer of coal has great continuity and small coverage; it presents very good quality and can be extracted from the mouth of the mine without the need of milling. In the same state, in its central part, there are deposits that allow some milling and short-distance transport. Due to these characteristics, the mineral coal from Brazil should be consumed where it is found [2].
Figure 2: Location of the main mineral coal deposits in the States of Rio Grande do Sul and Santa Catarina [11]. Considering a 60% mine recovery factor, a 50% rate of use, an average capacity factor of 55% and an efficiency of 35%, the Brazilian coal reserves would be sufficient to supply thermal plants with 28.000 MW over the next 100 years [13].
3. Energy availability from rice husk, wood waste and mineral coal From the data obtained in Section for rice production and using the coefficient proposed from Felfli et al. [4], it is possible to estimate the amount of rice husk produced in the State of Rio Grande do Sul. Yet, if one considers the fuel
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Oviedo ICCS&T 2011. Extended Abstract
characterization previously obtained [14], the energy availability from rice husk can be obtained, as shown in Table 1. Also, comparing with the data which shows the thermal plants using rice in the State of Rio Grande do Sul [4], a utilization of almost 44% of the energy available from rice husk is already achieved. Taking into account the availability already mentioned, it is possible to estimate the total energy availability for rice husk, wood waste from Eucalyptus and mineral coal in the State of Rio Grande do Sul, as shown in Table 2. From the table, there is the confirmation of high energy availability from wood, furthermore if we consider that the data accounts only Eucalyptus wastes from planted forests – there are also the Pinus planted forests and a considerable area of Acacia along with industrial wastes. The mineral coal availability accounts the thermal power plants in operation in the State of Rio Grande do Sul.
Table 1: Energy availability from rice husk. 7.9 Total rice production (106 x t) Coefficient 0.18 1.42 Total rice husk production (106 x t) LHV (kJ/kg) 3440 Total energy availability (MW) 155 From the mineral coal use, the total energy availability from power thermal plants in operation in Brazil is 1415 MWth. There is also a prevision for an increment in the mineral coal participation for primary energy production from two new plants located in Rio Grande do Sul of 350 MWth each. Table 2: Energy available from rice husk, Eucalyptus wastes and mineral coal in the State of Rio Grande do Sul. Fuel Energy availability (MWth) Rice husk 155 Wood waste (Eucalyptus) 488 Mineral coal 466 To sum up, the possibility of using rice husk and wood wastes as biomass fuels for energy generation on a decentralized way in the southern region of Brazil is real. Also, from the coal mines available, this fuel could have an increased use by cofiring in fluidized boilers, as the technical viability has been proven in this paper.
4. Conclusion
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Oviedo ICCS&T 2011. Extended Abstract
The results show that there is a big possibility of installing small thermal power plant (STPs) in places where the various fuels studied in this work are available in south Brazil, and as a consequence, optimize costs related to transport. Also, this optimization of resources utilization provides the possibility of introducing independent producers and autoproducers of energy. If we consider the generation of energy from rice husk, the State of Rio Grande do Sul has a capacity of 155 MWth and, for energy from wood, there is a capacity of 488 MWth if we consider only Eucalyptus wastes from silviculture. Yet, it is possible to use the mineral coal vastly available in the region to increment the capacity. At this moment, oxyfuel tests, form the same project, are been carried out on the same pilot plant on order to show the possibility of adapting combustion plants to this technology. Moreover, such tests are showing the possibility of using the same fuels in the oxyfuel technology.
Acknowledgements We would like to thank National Council of Scientific and Technological Development – CNPq – from Brazil for the financial support and Foundation of Science and Technology of the State of Rio Grande do Sul (CIENTEC/RS) supplying the Campus’ plants for combustion and oxyfuel tests.
References [1] Cortez, LAB, Lora, EES, Gómez, EO. Biomassa para energia. Campinas: Editora da UNICAMP; 2008. Portuguese. [2] BRASIL. Matriz Energética Nacional 2030. Ministério de Minas e Energia & Empresa de Pesquisa Energética; 2007. 254 p. Portuguese. [3] BRASIL. Plano Decenal de Expansão de Energia 2008-2017. Ministério de Minas e Energia & Empresa de Pesquisa Energética; 2009. 734 p. Portuguese. [4] Mayer, FD. Aproveitamento da casca de arroz um uma micro central termelétrica – Avaliação dos impactos econômicos e ambientais para o setor arrozeiro do Rio Grande do Sul [dissertation]. Programa de Pós-Graduação em Engenharia de Produção: Universidade Federal de Santa Maria; 2009. [5] Felfli, FF, Mesa, JM, Rocha, JD, Filippetto, D, Luengo, CA, Pippo, WA. Biomass briquetting and its perspectives in Brazil. Biomass and Bioenergy 2011; 35: 236-242.
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Oviedo ICCS&T 2011. Extended Abstract
[6] CENBIO. Centro Nacional de Referência em Biomassa; 2010. Available from: http://www.cenbio.iee.usp.br. Portuguese. [7]
IRGA.
Instituto
Riograndense
do
Arroz;
2011.
Available
from:
http://www.irga.rs.gov.br. [8] IBGE. Instituto Brasileiro de Geografia e Estatística; 2011. Available from: http://www.sidra.ibge.gov.br. Portuguese. [9] ABRAF [Internet]. Associação Brasileira dos Produtores de Florestas Plantadas. Anuário Estatístico 2009; 2011 Available from: http://www.abraflor.org.br/. Portuguese. [10]
SBS.
Sociedade
Brasileira
de
Silvicultura;
2010.
Available
from:
http://www.sbs.org.br. Portuguese. [11] SPG. Atlas Sócio-Econômico do Rio Grande do Sul [electronic version]. Secretaria de
Planejamento
e
Gestão;
2011.
Available
from
http://www.scp.rs.gov.br/atlas/default.asp. Portuguese. [12] Gomes, AM, Ferreira, JA, Albuquerque, LF, Süffert, T. Carvão fóssil. Estudos Avançados 1998; 12: 89-106. Portuguese. [13] BRASIL. Plano Nacional de Energia 2030. Ministério de Minas e Energia & Empresa de Pesquisa Energética; 2006. 324 p. Portuguese. [14] Zen, LD, Gomes, G, Kalkreuth, W, Garavaglia, L, Sampaio, C. Evaluation of the use of Co-Combustion (Coal-Biomass) in Small Fluidized-Bed Thermal Power (STPP) in Southern Brazil. In: Internacional Conference on Coal Science and Technology, 2007, NottiNgham, England, 2007.
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Oviedo ICCS&T 2011. Extended Abstract K.S. Milenkova and A.G. Borrego
Effect of particle size in the devolatilization behaviour of coal chars of different rank K.S. Milenkova and A.G. Borrego Instituto Nacional del Carbón, INCAR-CSIC. Ap 73, 33080 Oviedo, Spain
Abstract The effect of particle size on the devolatilization behaviour of coal chars has been studied for coals ranging in rank from high volatile bituminous to anthracite and having moderate to high inertinite content. The conditions tested were similar to those occurring in pulverized coal combustion. The temperature of char devolatilization was 1300 ºC and the residence time was around 30 ms. Five size factions within 150 and 20 microns, which are within the range typical for pulverized fuel combustion were tested for each coal. Both the aspects related with morphology/optical texture and those related to specific surface area and reactivity have been considered.
Some morphological and optical textural aspects of the coal chars exhibiting plastic behaviour and swelling during devolatilization were significantly affected by the particle size. The high volatile bituminous coal yielded chars with incipient anisotropic walls. The main differences between the size fractions were that the smallest particles exhibited thinner char walls and higher specific surface areas resulting in a progressively higher reactivity. A similar behaviour was observed for the semianthracite that under the present conditions fused and formed well-developed anisotropic domains. The smaller chars with smaller anisotropic domains and higher specific surface areas were more reactive than the large ones. This behaviour was not observed in the case of the inertinite-rich semianthracite which showed similar reactivity for most of the fractions except the smallest one. This can be explained in relation to the differences in plastic behaviour between vitrinite and inertinite, the latter less sensitive than vitrinite to differences in swelling. Only the smallest fraction showed a significantly higher reactivity than the others. The anthracite, which did not fuse upon devolatilization, enhanced the anisotropy of the particles in the reactor. The particles did not exhibit significant differences in morphology except for a higher number of cleats in the smallest particles. The reactivities were also similar for the different fractions.
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Oviedo ICCS&T 2011. Extended Abstract K.S. Milenkova and A.G. Borrego
The main conclusion of the present study is that the effect of particle size on the devolatilization behaviour of coal chars is not the same for any coal and it is very much related to coal rank and maceral composition of the parent coal.
1. Introduction This study deals with the influence of particle size in pulverised coal char reactivity in relation to the characteristics of the parent coals. It is well-known that particle size has a negative effect on the combustibility of pulverised coals at high temperature because the reaction is either controlled by external diffusion (Regime III) or internal pore diffusion (Regime II) [1]. It is not only the size of the particle but the distribution of vacuoles, the wall-thickness, the porosity and the intrinsic reactivity of the char-forming material what finally determines the combustibility of the particles. The char structure is achieved after fast volatile release. The particle size is also relevant for the final char structure because larger particle sizes restrain the transport of volatile matter out of a particle and promote secondary reactions [2], which increase the yield of light gases and decrease the yield of tar [3]. The devolatilization behaviour will therefore have a key role on the morphology of the char acquired after re-solidification of the carbonaceous residue [4] and on the optical texture of the char walls [5], both relevant variables for the combustion rate of the devolatilized char. Despite that the influence of particle size in the devolatilization/combustion behaviour of pulverized coal has been a topic of interest for many years [6, 7], overall there are not systematic studies dealing with this issue and they are particularly scarce for low volatile coals. In this study four coals commonly burned in Spanish power plants ranging in rank from high volatile bituminous to anthracite have been considered for a detailed study of the influence of particle size in the devolatilization behaviour of the particles. The results will be discussed in relation to the characteristics of the parent coals.
2. Experimental The coals are three Palaeozoic coals from Northern Hemisphere ranging from high volatile bituminous rank to anthracite with low to moderate inertinite content (in increasing rank order MO<SO
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Oviedo ICCS&T 2011. Extended Abstract K.S. Milenkova and A.G. Borrego
inertinite content of NI compared with SO [8] and to the 15 vol% of higher rank anthracite in NI. Table 1. Proximate ultimate and petrographic analyses of the coals. Ash V.M. C H N O S Rr V L I db % daf % % mmf- vol % MO 23.6 38.0 85.7 5.1 1.7 8.0 1.2 0.84 71.0 11.9 13.1 SO 33.4 17.2 90.0 3.5 1.5 7.6 0.7 2.06 76.2 23.8 NI 15.5 14.4 88.0 4.1 2.3 5.6 1.5 2.00 60.8 39.2 NN 30.8 9.0 91.6 2.6 1.0 5.6 1.6 3.73 95.7 4.3 VM=Volatile matter content, Rr=vitrinite random reflectance, V=vitrinite, I=inertinite, L=liptinite, db=dry basis, daf=dry-ash-free basis, mmf=mineral matter free, vol=volume.
Each coal ground typical pulverised coal size (80 % below 75 µm) was separated into 6 granulometric fractions (>150, 150-100, 100-75; 75-45, 45-20, and <20 µm). Each fraction was characterized by temperature programmed thermogravimetric analysis (from 40 to 900 ºC at heating rate 20 ºCmin-1 and flow rate 35 cm3min-1 ) to produce a combustion profile and to determine the ash content. Each fraction was also fed in a drop tube reactor (DTR) described elsewhere [9]. The DTR was operated at 1300 ºC, with a flow rate of 600 L h-1 for the body of the reactor and 300 L h-1 for the feeder and a 5% of oxygen in the reacting gas (below stoichiometry). The estimated residence time of the particles in the reactor was 0.3 s. Under these conditions volatiles are burned away and the coal is devolatilized but not completely burned generating a char that can be further characterized.
The conversion of the samples in the DTF was estimated using the ash-tracer expression ashcoal Conversion(%) = 1 − 100 − ashcoal
100 − ashchar ashchar
× 100
Chars were embedded in polyester resin and polished for petrographic examination under incident polarised light with 1 λ retarder plate. Reactivity of the char was measured in a thermobalance under air at 550ºC with a flow rate of 35 cm3min-1 after heating the char under nitrogen to the selected temperature. Aparent reactivity was calculated as R=1/mo(dm/dt) where mo is the initial ash-free mass of coal. CO2 adsorption isotherms were measured in a micromeritics ASAP 2020 adsorption equipment at 0 ºC temperature in the interval of pressure 0.035-0.0001 torr. Prior to gas adsorption experiments the chars were outgassed under vacuum at 5 ºC min-1 with holding temperatures of 90 ºC (1 h) and 350 ºC (4 h). The Dubinin-Radushkevich (D-R)
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Oviedo ICCS&T 2011. Extended Abstract K.S. Milenkova and A.G. Borrego
equation [10] was applied to the adsorption data to obtain the specific surface area (SCO2). Normalized reactivity was obtained dividing the apparent reactivity by SCO2 referred to an ash-free-basis.
3. Results and Discussion The combustion profiles of the samples indicated that mineral matter content hardly varied for the different size fractions except for the fraction <20 µm, which was not fed to the reactor. In addition the maceral composition did not change for the fractions analysed. The combustion profiles also showed the best combustibility for the hvb coal, intermediate for the semi-anthracites and the lowest for the anthracite [11]. In addition, they showed improved combustibility as the particle size decreased and this was mainly reflected in the reaction rate and the burnout temperature whereas ignition temperature and temperature of maximum reaction rate were less sensitive to the variation of particle size. This result is not unexpected because combustion in the thermobalance progress following a shrinking core model [12] and smaller particles burn away faster [13]. 8E-04
70 60 Rmax-ap(s-1)
Conversion(%)
6E-04
50 40 30 20
4E-04
2E-04
10
MO
SO
NI
MO
NN
SO
NI
NN
0E+00
0 >150
300
150-100 100-75 75-45 Size fraction (µm)
>150
45-20
150-100
100-75
75-45
45-20
Size fraction (µm)
5.E-06
200
-1
2 -1
SCO2 (m g )
-2
Rmax-in (s m )
4.E-06
100
3.E-06 2.E-06 1.E-06
MO
SO
NI
MO
NN
SO
NN
NI
75-45
45-20
0.E+00
0 >150
150-100
100-75
75-45
45-20
Size fraction (µm)
>150
150-100 100-75
Size fraction (µm)
Figure 1. Conversion, apparent reactivity measured in air at 550 ºC in a thermobalance (Rmaxap), Specific surface area measured by CO2 adsorption (SCO2) and reactivity normalized to surface area (Rmax-in) for the various coal chars and size fractions.
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Oviedo ICCS&T 2011. Extended Abstract K.S. Milenkova and A.G. Borrego
The conversion in the reactor calculated using the ash-tracer is shown in Figure 1. It must be stressed that the oxygen content/residence time in the reactor were not enough to complete combustion of the coals (5%O2). The burnout for the various coals followed the trend expected from the chemical analysis, the higher the volatile matter content of the parent coal, the higher the conversion. Nevertheless, Conversions of NI and NN varied within narrow limits despite the higher volatile matter content of NI. The semi-anthracites and the anthracite exhibited similar conversions for the various size fractions whereas for the hvb coal conversion of the largest particles was slightly higher than that of smallest particles.
The apparent reactivity (Rmax-ap) of MO chars measured under kinetic control in a thermobalance was the highest of the set and increased as particle size decreased. Rmax-ap was similarly low for the large fractions of anthracite and semi-anthracites chars. As particle size decreased Rmax-ap increased for SO and NN chars whereas NI char reactivities maintained very low. The reactivity of the coal chars measured in a thermobalance and the specific surface area measured by CO2 adsorption allow accounting separately for the factors involved in the apparent reactivity.
The specific surface area (Figure 1) was the highest for the anthracite. Anthracite is known to have slit-shape porosity due to the spatial arrangement of the ordered structural polyaromatic units. The structure of anthracite can be visualized as graphitic-like units with pseudo-parallel arrangement with C-C crosslinks. Upon fast heating in the DTF the structure suffered contraction and limited realignment of aromatic units favoured by volatile release but mobility is so low that porosity annihilation did not occur. Therefore the specific micropore surface area of anthracite chars remains high. Values around 300 m2g-1 have been reported for anthracite chars prepared at temperatures similar to that of the present study [5]. A drop in surface area has been observed as the particle size decreased for anthracite chars (Figure 1). This might be related to the highest peak temperature reached by smallest particles which would have caused a reduction in porosity [5]. This is confirmed by the reflectance of the chars which slightly increased for the smallest particles [14] indicating a more severe re-organization of the anthracite structure as the particle size decreased. A similar behaviour was observed for the inertinite-rich semianthracite (NI) in Figure 1. The rational for NI behaviour is not so clear. This char is very heterogeneous consisting on Submit before January 15th to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract K.S. Milenkova and A.G. Borrego
anisotropic well-ordered material derived from vitrinite, anthracite-like material (around 15 volume %) and inertinite-derived material. The latter is expected to have a carbonrich disordered structure with relatively high internal surface area [5] whereas the material derived from the vitrinite would have among the lowest surface areas as observed for SO char. Vitrinite in SO char generated well—ordered anisotropic domains which were larger in the largest particles as consequence of enhanced secondary reactions (Figure 2). Increase in crystallite size favoured by secondary reactions (large particles) and improved devolatilization favoured by faster heating rate (small particles) may have opposite effects on the development of internal surface area that may account for the variation observed in Figure 1. SCO2 increased with particle size for the hvb coal char (MO) as previously reported for size fractions of lignite char [15]. As char becomes more ordered with decreasing heating rate for any given temperature [16], the largest particles will exhibit larger crystallite size, which is translated into larger anisotropic domains in the optical micrograps (Figure 2). This can be observed by the development of incipient anisotropy in the thick walls of the MO char cenospheres, whereas walls appear to be isotropic in the smallest particles. Similarly well-developed anisotropic domains were observed in the walls of the largest particles in SO whereas the smallest ones were brighter and exhibited smaller domains (Figure 2).
MO (20-45 µm)
MO (>150 µm) 50 µm
SO (20-45 µm)
SO (>150 µm)
Figure 2. Optical micrographs showing the improved anisotropy development observed in the large particles compared to the small ones for hvb coal chars (MO) and semi-anthracite chars (SO).
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6
Oviedo ICCS&T 2011. Extended Abstract K.S. Milenkova and A.G. Borrego
Figure 3 shows images of the chars generated in the DTR from the different coals. The smallest size fraction has been selected to show the appearance of the chars because various particles can be shown in a single image. Overall the high volatile bituminous coal (MO) generated highly swollen chars dominated by isotropic cenospheres. The isotropic material indicates disordered material which typically exhibits high reactivity [5] and is associated to high specific surface areas and abundance of active sites. The vitrinite-rich semianthracite (SO) generated multi-vesiculated particles with welldeveloped anisotropic domains, which were formed during a short plastic stage. During this stage aromatic units grew and stacked yielding well-ordered anisotropic domains with low reactivity and low specific surface areas (Figure 1). Indeed the intrinsic reactivity of largest fraction with the largest anisotropic domains was the lowest one.
SO
NI
50 µm MO
NN
Figure 3. Appearance of the chars from the smallest size fraction (20-45 µm) of the coals considered in this study. MO=hvb coal, SO and NI=semi-anthracite and NN=anthracite.
The main difference of SO chars with the inertinite-rich semianthracite NI chars was the predominance of massive particles and/or abundance of unfused material in NI chars (Table 2). In addition the 15% anthracite contained in NI char was not evenly distributed among the different size fractions but concentrated in the largest ones (Table 2).
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7
Oviedo ICCS&T 2011. Extended Abstract K.S. Milenkova and A.G. Borrego
The data in Table 2 may bring some light to the interpretation of the results observed in Figure 1. For example the similar SCO2 of NI and NN chars both having different devolatilization behaviour resulting in different char morphologies can be explained by the high amount of both inertinite and anthracite particles in the large fractions of NI responsible for the higher SCO2 of these chars compared with those of the vitrinite-rich semianthracite (SO). The anthracite did not fuse under the actual operating conditions and therefore NN chars were formed by massive angular particles without signs of devolatilization except for the enhancement in vitrinite reflectance observed in the particles [17]. They show different colour in Figure 3 as consequence of the random orientation of the strongly anisotropic particles. Table 2. Petrographic analysis of the chars applying an adapted 1996 ICCP system for description of char optical texture. Different fractions of NN chars were not analysed becaused consisted of over 95% of unfused vitrinite-derived anisotropic material. Sample
Size fraction µm
>150 100-150 chMO 75-100 75-45
Anisotropic Isotropic Anisotropic Fragments Particles fused inertinite- inertinitefrom derived vitrinitederived accompanyin derived material material g coal material
24. 3 22.7 20.2 22.9
33.2 46.4 50.0 43.1
19.5 17.2 15.2 9.4
17.1 11.7 10.1 12.1
6.0 1.9 4.5 12.4*
-
20-45
29.3
23.1
6.2
18.5
22.8*
-
>150
-
49.3
15.9
26.8
8.0
-
100-150
-
35.5
14.2
40.4
7.8
2.1
-
34.5 41.9 36.1
17.9 11.8 12.4
36.3 35.5 24.3
7.7 7.0 23.1*
3.6 3.8 4.1
-
7.7 9.6 8.2 6.2 8.6
32.9 37.4 34.1 22.9 19.0
20.7 26.1 34.9 43.6 41.2
4.5 3.5 5.2 7.5 24.9*
34.2 23.5 17.7 19.8 6.3
chSO 75-100 75-45 20-45
chNI
Isotropic vitrinitederived material
>150 100-150 75-100 75-45 20-45
*fragments plus unidentified particles in the smallest size fractions. In SO the accompanying coal is a hvb coal, in NI the accompanying coal is an anthracite consisting of unfused anisotropic vitrinite particles..
4. Conclusions A detailed study of char morphology, optical texture, specific surface area and reactivity of different size fractions of coal chars of different rank has shown the effect of the Submit before January 15th to [email protected]
8
Oviedo ICCS&T 2011. Extended Abstract K.S. Milenkova and A.G. Borrego
devolatilizing behaviour in different parameters relevant for char combustibility under pulverised fuel combustion conditions. Overall the larger the size of the particles, the lesser the volatile release and the lower the heating rate of the particles resulting in lower fluidity upon devolatilization. Significant differences occurred among the different materials forming the chars. For material with caking but non-coking properties the enhanced devolatilization of the smallest particles increased the surface area resulting in an increased apparent reactivity of the chars. The surface areas were rather low for the material passing through a plastic stage. In this case the smaller size of the anisotropic domains in the smaller particles can be responsible for their enhanced apparent reactivity. In the case of the anthracite chars with no ability to melt during devolatilization a drop in surface area occurred for the smaller particles due to a better realignment of aromatic domains. This caused a certain contraction in the structure and could make more accessible the more reactive active sites.
Acknowledgement. The research leading to these results has received funding from the Research Programme of the Research Fund for Coal and Steel (Grant Agreement number ECSC
7220-PR/121).
References [1] Zhangfa W. Fundamentals of pulverised coal combustion CCC/95, IEA London 2005, 36 pp. [2] Devanathan, N., Saxena, S.C., A transport model for devolatilization of large nonplastic coal particles: The effect of secondary reactions, Ind. Engng Chem. Res. 1987:26, 539-548. [3] Saxena S. C. Devolatilization and combustion characteristics of coal particles Prog. Energy Combust. Sci. 1990:16, pp. 55-94 [4] Yu, J., Lucas, J.A., Wall T.F. Formation of the structure of chars during devolatilization of pulverized coal and its thermoproperties: A review. Prog. Energy Combust. Sci. 2007:33, 135–170. [5] Alonso, MJG, Borrego, AG, Alvarez, D, Parra, JB, Menendez, R. Influence of pyrolysis temperature on char optical texture and reactivity. J Anal Appl Pyrolysis 2001:58–59, 887–909. Barranco et al. [6] Fletcher, T.H. Swelling properties of coal chars during rapid pyrolysis and combustion. Fuel 1993: 72, 1485-1495 [7]Cloke M., Lester E., Thompson A.W. Combustion characteristics of coals using a drop-tube furnace. [8] Borrego A.G., Marbán, G., Alonso, M.J.G., Álvarez D., Menéndez R. Maceral effects in the determination of proximate volatiles in coals. Energy & Fuels 2000: 14, 117-126
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9
Oviedo ICCS&T 2011. Extended Abstract K.S. Milenkova and A.G. Borrego
[9] Borrego, A.G., Alvarez D. Comparison of chars obtained under oxy-fuel and conventional pulverized coal combustion atmospheres. Energy Fuels 2007:21, 31713179. [10] Dubinin, M.; Radushkevich, L. Proc. Acad. Sci. USSR 1947, 55, 331-335. [11] Cumming JW. Fuel 1984;63:1436–40. [12]. Alonso M.J.G., Borrego A.G., Alvarez D., Kalkreuth W., Menéndez R. Physicochemical transformations of coal particles during pyrolysis and combustion. Fuel, 2001:80, 1857-1870 [13] Méndez L.B., Borrego A.G.,. Martinez-Tarazona M.R, Menéndez R. Influence of petrographic and mineral matter compositionof coal particles on their combustion reactivity. Fuel 2003:82, 1875–1882 [14] Milenkova K.S. Origen de los inquemados en cenizas volantes de centrales térmicas. Ph D. Thesis. University of Oviedo, 2004. [15] Field, M.A. Rate of combustion of size-graded fractions of Char from a low-rank coal between 1200 K and 2000 K. Combustion and flame 1969:13, 237-252. [16] Lu, L., Kong, C., Sahajwalla, V., Harris D. Char structural ordering during pyrolysis and combustion and its influence on char reactivity. Fuel 2002: 81, 1215-1225. [17] Borrego A.G., Martín, A.J. Variation in the structure of anthracite at a fast heating rate as determined by its optical properties: An example of oxy-combustion conditions in a drop tube reactor. : Int. J. Coal Geol., 2010:81, 301-308.
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10
Program topic: Coal Combustion
The element distribution during the co-combustion of coal with wood and wood wastes Z. Klika and L. Bartoňová VŠB – Technical University of Ostrava, Tř. 17. listopadu 15, 708 33 Ostrava – Poruba, Czech Republic, +420 596991111, [email protected], [email protected]
Abstract Combustion of coal with limestone, co-combustion of coal with limestone and wood matter and combustion of wood matter were performed in CFB in 7 different combustion regimes. Inorganic matter composition and properties of all input and output materials were characterized. For this, chemical composition of solid samples was determined. Balance calculations were performed for Cl, Zn, As, Se, Hg and Pb and all results were based on the 1 GW (=103 MW) of circulating fluidised-bed boiler output. The results show that combustion of wood and/or co-combustion of lignite with wood waste brings about significant environmental benefits.
1. Introduction Worldwide, there are many reasons for utilization of biofuels for energy purposes. First of all it is significant environmental benefit [1-4] and reduction of the dependency on imported oil. Using the biomass brings about a reduction of acid rains, CO2 as a greenhouse gas and reduction of hazardous elements that can get into the environment as a volatile species or can be dissolved from solid state products of coal combustion.. Biomass is a renewable fuel available for heat and power production in relatively substantial amounts [57]. In this contribution the comparison of the trace elements (Cl, Zn, As Se, Hg and Pb) in input and output streams of CFB were performed for combustion of lignite, wood and/or cocombustion of lignite with wood wastes.
1
2. Experimental section In power station situated in the Mondi Packaging Paper-Mill Štětí area near Mělník (Czech republic) lignite combustion and co-combustion of lignite/wastes were performed in circulating fluidised-bed boiler (K11) at 870°C. Waste wood coming from the cellulose production was used. Simplified diagram of the combustion facility is given in Fig, 1.
Fig. 1. Simplified diagram of the combustion facility Lignite was combusted and/or co-combusted with wood wastes in seven combustion regimes (I – VII). In reference regimes I and IV lignite was combusted and in regimes II, V, VI and VII lignite with wood wastes and in regime III only wood waste (wood, sawdust and wood chips) were combusted. Data to individual regimes are given in next paragraph.
3. Results and discussion Mass flows of lignite, wood wastes and other alternative fuels are given in Table 1. They relate to original (not dried) samples. Except input flows of fuel also mass flows of bottom ash (BA) and fly ash (FA) were measured. Chemical analyses of selected elements (Cl, Zn, As, Se, Hg and Pb) were performed in all input and output materials. Mean values and estimated standard deviations ( c x ,i ± 2s) for lignite, limestone and wood calculated from regimes I – VII are given in Table 2. Mean values for S, B, WCh, SS, ASF and So (see Table 1) have not been calculated because they were used for co-combustion only in one (or two) regims. Anyway, each sample of fuel in each regime I – VII was collected for analyses as average sample from about 16 sample portions obtained during the measurement in each regime. Table 1. Mass flows of lignite, wood wastes and other alternative fuels in regimes I -VII. 2
Lignite (C) Limestone (L) Sawdust (S) Tree-bark (B) Wood (W) Wood chips (WCh) Sewage sludge (SS) Alternative solid fuel (ASF) Soap (So)
I 25920 2630 -
II 11840 970 5220 5620 -
III 5690 27220 27220 -
-
-
-
-
mass (kg/h) IV 29980 3490 -
V 34960 4320 19728 -
VI 26750 3240 15160 15160
VII 28550 2630 24260 -
-
-
-
940
-
-
-
-
-
200
Table 2. Concentrations and estimated standard deviations ( c x ,i ± 2s) of selected trace elements in lignite, wood wastes and other alternative fuels in regimes I - VII. c x ,i ± 2s (ppm) Lignite Limestone Sawdust Tree-bark Wood Wood chips Sewage sludge Alternative solid fuel Soap
Cl 407 (±150) 733 (±300) 4 16 258 (±37) 11 22 128
Zn 26 (±7) 34 (±3) 5.6 41 59 (±8) 3 3.4 208
As 87 (±40) 1.9 (±1) 0.24 0.18 0.32 (±0.3) 1.2 0.08 0.48
Se 1 (±0,5) 0.6 (±0.25) 0.013 0.038 0.05 (±0.01) 0.05 0.02 0.2
Hg 0.16 (±0.04) 0.01 (±0.001) 0.05 0.06 0.05 (±0.05) 0.05 0.003 0.7
Pb 11 (±0.5) 8 (±4) 0.19 1.7 4(±1) 0.43 0.6 37
364
49
0.18
0.11
0.001
0.8
For solid output streams BA and FA the concentrations of elements differs more than those in input materials (Table 2) and therefore the mean concentration for each combustion regime have not been calculated. The concentrations of elements are given for BA (cBA) and FA (cFA) in Tables 3A and 3B, respectively. Table 3A. Concentrations of selected trace elements in BA in regimes I - VII . cBA (ppm)
Regime I II III IV V VI VII
Cl 569 836 566 1568 1315 1141 490
Zn 188 251 223 151 139 94 170
As 248 140 298 97 97 39 158
Se 1,4 1,3 1,0 0,6 0,3 0,2 0,4
Hg 0,002 0,005 0,003 0,002 0,002 0,001 0,002
Table 3B. Concentrations of selected trace elements in FA in regimes I – VII. 3
Pb 41,2 56,0 48,5 33,4 33,3 41,1 32,9
cFA (ppm)
Regime I II III IV V VI VII
Cl 303 375 440 1163 1003 975 650
Zn 172 243 268 225 261 196 264
As 375 438 291 535 375 268 586
Se 12,5 7,9 7,8 7,4 5,0 5,1 7,3
Hg 0,605 0,280 0,010 0,779 0,202 0,672 0,537
Pb 58,9 65,0 57,4 36,1 45,0 46,6 41,8
In order to compare the flow of elements the boiler outputs (Qoutput) for regimes I – VII were calculated. Data used for the Qoutput calculation and calculated Qoutput are given in Table 4. The ratio (R) is defined for each regime I – VII as ratio between mass of organic wood wastes (OrgF) and total fuel (F). It is calculated using the formulae (Eq. 1): R = 102.(mOrgF/mF)
(1)
where mOrgF is sum of mass of all fuels except lignite (i.e S, B, W, WCh, SS, ASF and So, see Table 1) and mF = mOrgF + mC (C is for lignite, see Table 1). This ratio varies from 0 (for regimes I and IV) to 1 (for regime III). Table 4. Data used for Qoutput calculation. mF Qcal η Qoutput R (kg/h) (MW..) (MW) (-) I 17315 15 0.89 64 0 II 15441 14 0.90 54 48 III 46750 12 0.90 140 100 IV 16040 15 0.89 59 0 V 26403 12 0.90 79 36 VI 25940 12 0.90 78 54 VII 30673 12 0.90 92 46 Note: mF is mass flow of dry and ash free fuel (kg/h); Qoutput is boiler output (MW), Qcal is fuel calorific value related to dry and ash free basis (MJ/kg), and η is boiler efficiency. The total input mass flows of the element (Cl, Zn, As, Se, Hg and Pb) min (kg.h-1.GW-1) have been calculated for each combustion regime I – VII. The data are plotted against the ratio R for each combustion regime (Fig. 2).
4
Fig. 2. The plot of the logarithm of total input mass flows of element min (Cl, As, Zn, Pb, Se and Hg) and ratio R. The results show that the order of total input mass flows of the elements (min) ranging from Cl, As, Zn, Pb, Se to Hg do not change with changes of ratio R (ratio between organic wood wastes and total fuel) that varies from 0 to 100 % (Fig. 2). The total input mass flows of elements were calculated using data given in Tables 1 and 2. However the input mass flows of all elements decrease with increasing portion of wood waste on total combusted fuel. From our results it can be supposed that decreasing of these elements in input flows for combustion of wood wastes and/or co-combustion of coal with wood wastes bring about also essential decrease of these elements in flue gas. For circulating fluidised-bed boiler output 1 GW there are following decrease of total input mass flows of the elements between regimes combusting lignite (regimes I and IV; R=0) and regime IV combusting 100 % of wood wastes (R = 100 %): For Cl: from 219 to 36 kg/h For As: from 40 to 0.3 kg/h For Zn: from 13 to 11 kg/h For Pb: from 5 to 0.8 kg/h For Se: from 0.5 to 0.03 kg/h For Hg: from 0.10 to 0.02 kg/h The calculation of the element volatilities from difference between mass of input and output streams was burdened by errors due to variable composition of fuel (namely wood wastes) and therefore they are not presented. Conclusion This work was focused on the comparison of some hazardous mass element flows in the circulating fluidised-bed power station in Štětí during the combustion of lignite and wood 5
waste materials. Seven combustion regimes with different ratio of wood waste to total fuel mass were performed. The total input mass flows of the elements (Cl, Zn, As, Se, Hg and Pb) in all combustion regimes have been based on the 1 GW (=106 MW) of circulating fluidisedbed boiler output. It was found that: Using higher ratio wood waste to total fuel (lignite with wood waste) bring about the essential decrease of mass elements (Cl, Zn, As, Se, Hg and Pb) in process of co-combustion of lignite with wood wastes and therefore also lower volatilized mass of elements.
Acknowledgment The authors would like to express their acknowledgement to the Grant Agency of the Czech Republic (grant GA 105/08/0913) for the financial support of this work. References [1] Kajikawa, Y., Takeda, Y. Technological Forecasting and Social Change, 2008, 75, 1349. [2] Haykiri-Acma, H., Yaman, S. Waste Management, 2008, 28, 2077. [3] Van Loo, S., Kopperjan, J. (Editors). Handbook of Biomass Combustion and Co-firing, Vol. 2. Earthscan, London, UK, 2008, pp. 1–442. [4] Ogaji, S., Probert, D. Applied Energy, 2009, 86, 2272. [5] Escobar, J.C., Lora, E.S., Venturini, O. J., Yánez, E. E., Castillo, E. F., Almazan, O. Renewable and Sustainable Energy Reviews, 2009, 13, 1275. [6] Yilgin, M., Pehlivan, D. Applied Energy, 2009, 86, 1179. [7] Pettersson, A., Amand, L. E., Steenari, B. M. Fuel, 2009, 88, 1758.
6
The influence of particle size and density on the combustion of Highveld coal. G.W. van der Merwea, R.C. Eversona*, H.W.J.P. Neomagusa, and J.R. Bunta,b a
School of Chemical and Minerals Engineering, North-West University, 2520, Potchefstroom, South Africa. b
Sasol Technology (Pty) Ltd, Box 1, Sasolburg, 1947, South Africa
*Corresponding author. Tel.: +27 18 299 1986; E-mail address: [email protected]
Abstract Coal from the Highveld seam 4 deposit in South Africa was studied to determine and understand the influence of density and particle size on the high temperature combustion characteristics. All of the charring, as well as the combustion experiments were conducted in a high temperature horizontal tube furnace where charring was done at 1100 °C and combustion at 1000 °C. The furnace was equipped with carbon monoxide and carbon dioxide analysers to monitor carbon conversion during experimentation. The parent coal was characterised in terms of proximate, ultimate, calorific value, and FTIR analysis. The chars were investigated by FTIR analysis. The coals were categorized according to density from 1.4 g.cm-3 to 2.0 g.cm-3 using 0.2 g.cm-3 intervals and the particle size effect was studied using particles of 20, 30 and 40 mm diameter.
The characterization of the different coal samples showed that the parameters were generally significantly different for the different density fractions but did not vary significantly across the particle size ranges. FTIR analysis showed that the properties of the density separated fractions varied significantly; this was most prominent in the substitution structures occurring on the aromatic structures. It was found that the chemical composition of the size separated particles remained relatively constant. The properties of the char particles having different densities tended to become more similar, especially with regard to their aromatic substitution structures.
It was also found that significant shattering occurred at heating rates of 50 °C/min, while very little shattering was observed at 15 °C/min. The low density particles formed porous ash residues, while the high density particles formed very hard and
solid ash residues that could contain unreacted carbon in the core. Combustion studies showed that particle density and size had a significant influence on the time required for complete conversion of the chars. The low density particles required less time for full conversion than for the high density particles and the smaller particles reacted faster than the larger particles. Modelling of the experimental data showed that the shrinking unreacted core model can be used to describe the combustion characteristics of both the size and density separated particles. The controlling mechanism was found to be a combination of internal and external diffusion. From the modelling results it was found that effective ash layer diffusion became more prominent as the density increased and that the obtained mass transfer coefficients correlated well with values found in literature.
1. Introduction Energy resources and energy availability are critical to any civilization, and access to cheap fossil fuels has been the backbone of our society since the industrial revolution [1]
. The major fossil fuels fuelling this growing and ever more demanding society
include coal, oil and natural gas.
Along with the increasing use of coal as major fuel stock, public policy is aiming to create, develop and implement clean coal technologies to achieve higher efficiency power generation and coal conversion
[2]
. The main increase in public demand for
clean coal technologies and more efficient processes is due to the fact that the use of fossil fuels has some negative effects on the environment. The energy supply processes can be improved by way of both the more efficient use of coal in existing technology, and the designing of new systems that use coal as optimally as possible.
To design better technologies or to make current processes more efficient it is necessary to understand the influence of parameters like particles size density
[4]
, temperature
[5]
, pressure
[6]
and coal composition
[7]
[3]
, particle
as these parameters
greatly influence performance characteristics like conversion rate, product quality and unwanted emissions. Given that as coal resources diminish the utilisation of low grade coal will become unavoidable, making it essential that the performance of low grade coal be understood in order to make suggestions regarding system upgrades.
2. Experimental
The ROM (Run of mine) coal obtained from the Highveld region was submitted for preparation to the laboratories at the South African Bureau of Standards (SABS). Prior to density separation, the entire 250 kg sample was sieved to remove the – 0.5 mm size fraction, as these fines would significantly increase the time for sink / float separation, and could affect the reference density of the separating liquid. The remaining sample was then density separated into the density cuts ranging from F 1.4 g.cm-3 to S 2.0 g.cm-3 at 0.2 g.cm-3 intervals, using sink / float techniques with TBE (tetrabromoethane) as medium. The density separated samples were then left overnight in a convection oven at 250 °C to allow the TBE to evaporate. The dried density cuts were then sieved to obtain the desired sizes (20, 30, 40mm). Sieves of appropriate sizes were stacked on top of each other, with a collection drum at the bottom to collect the smallest size.
Representative samples were then sent to the analytical laboratories at the South African Bureau of Standards (SABS) for characterization analyses. These representative samples were also used to determine the calorific values of the respective coals as well as to do analytical work in terms of FTIR analysis. Due to the fact that conventional density measurement techniques like mercury intrusion porosimetry are only capable of using small particle sizes, an apparatus that was able to measure the density of large coal particles without destroying or damaging the particles was developed. This apparatus can handle samples up to 60 mm in diameter.
The samples selected according to the required size and density was firstly charred inside the horizontal tube furnace. The particles were charred at 1100°C under a nitrogen atmosphere. The particles were heated at 15°C/min to minimize particle breakage as a high heating rate leads to significant cracking of the larger particles. After the particles reached the required temperature, they were left for 45 minutes. The furnace was then turned off to allow the particles to cool down under a nitrogen atmosphere. As soon as the furnace temperature reached 100°C the chars were removed and weighed. The nitrogen flow rate during charring was set to 5 Nl/min. The horizontal tube furnace setup used can be seen in Figure 1.
Figure 1: Horizontal tube furnace setup
After the charring of the different particles was completed, the charred particles were combusted inside the horizontal tube furnace. The particles were heated to 1000 ˚C at a heating rate of 15°C/min and were kept at this set point for an additional 20 minutes to ensure isothermal conditions inside the particle. The gas feed was then changed to an air flow rate of 10 Nl/min so that the particle combustion could commence. Both the carbon dioxide and carbon monoxide production rate were recorded using the custom logging software. As soon as gas concentration reached no significant reading anymore, it was assumed that combustion was complete and the reactor was shut down and allowed to cool. As soon as the furnace temperature reached 100°C the remaining ash weight was measured
3. Results and Discussion
As the various density fractions contain different amounts of ash it can be expected that particles having different densities will behave differently during combustion. Some typical ash forms produced are shown in Figure 2.
(a) F 1.4 g.cm-3 (b) F 1.8 g.cm-3 (c) S 2.0 g.cm-3
Figure 2: Various ash forms obtained from different densities
The first type (a) is a porous ash that easily disintegrates upon touching and is typically found for the low density coals (F 1.4). The second type (b) is also porous but has a stronger structural integrity with some cracking and is typically found for the middle density coals (i.e. F 1.8). The third type (c) is a solid ash with little to no cracking that needs considerable force to break, and is typically found for the most dense coals (S 2.0). The combustion of 30 mm particles (density = 1.6 g.cm-3) at 1000 °C was stopped at various intervals into combustion. When these particles were opened it could be seen that the combustion followed shrinking unreacted core behaviour. This behaviour is shown in Figure 3. It can be seen that the core stays unreacted as the reaction progresses and that the central unreacted core shrinks as the reaction progresses.
Figure 3: Photographic representation of the shrinking unreactive core model at various stages of conversion (0-100%)
Remiarova et al.
[8]
and Holikova et al.
observed this behaviour in 5 mm particles combusted at 650 °C, [9]
also observed it for 10 mm particles at the same temperature.
They found that the reaction immediately consumes the oxygen at the sharp ash-core interface, and concluded that combustion at 650 °C is controlled by either external diffusion, or ash layer diffusion or a combination of the two controlling mechanisms. On the basis of these findings the shell-progressive or shrinking core type models were used for modelling purposes.
The influence of particle density on the combustion reaction at 1000 °C can be seen in Figure 4. All the runs for density variation were done on particles with an equivalent spherical radius of 30 mm.
Figure 4: Effect of particle density on the combustion of 30 mm particles at 1000 °C
It was observed that the density of the particle has a significant influence on the combustion rate of coal. It can clearly be seen that the lower density particles (density = 1.354 g.cm-3) react the fastest, while the higher density particles take longer to reach full conversion. It can also be noted that the highest density (density = 1.859 g.cm-3) does not follow the trend as would be expected. This deviation might be due to the hard ash properties of the residual combustion matter presenting high diffusion resistance; the different shape of the particles; or maceral disproportionation occurring in the coal samples.
The effect of particle size on the combustion reaction can be seen in Figure 5 which shows the effect as the particle radius is increased from 20 mm to 30 mm to 40 mm.
Figure 5: Effect of particle size on combustion at 1000 °C
From this Figure it can be observed that the larger particles take longer to reach full conversion. Work done by Tevrucht and Griffiths [10] showed a similar trend when the oxidation reaction was controlled by diffusion mechanisms rather than reaction controlling mechanisms. Diffusional effects were also observed for particles as small as 1-4 mm at 1500 K [11] and it was shown by other researchers that air combustion of 24 mm carbonaceous particles at 800 K was ash layer diffusion controlled
[3]
. It was
also shown for lignite cylinder (2.5 cm diameter, 6 cm long) combustion in air that the burning rate varied inversely with ash layer thickness, and that the burning rate can be predicted by assuming effective ash layer diffusivity
[12]
. During diffusion controlled
coal conversion processes smaller particles will oxidize more rapidly due to a higher surface area to volume ratio [10].
4. Conclusions From the qualitative observations made it was determined that extensive fragmentation occurred for large particles at heating rates of 50 °C/min, and that the fragmentation behaviour was only observed for the medium to high density particles when the heating rates were reduced to 25 °C/min. When the heating rate was further
reduced to 15 °C/min very little fragmentation was observed. No fragmentation was observed when the particles were reheated after charring.
It was observed that the low density coal particles produced very porous ash residues with poor structural integrity, while the high density particles produced a very dense ash residue that required significant force to break. High density particles can pose a significant resistance to ash layer diffusion effects as it was observed that some high density particles contained unreacted particle cores after several hours of combustion.
The quantitative results also showed that density has an effect on the time to full conversion of 30 mm particles, and that the lowest density particles required less time for full conversion than the higher density particles. The smaller particles reacted faster than the larger particles and the reaction is diffusion controlled.
It was found that temperature did not affect the conversion time required for the 30 mm particles and that the shrinking unreacted core model can be used to model combustion behaviour. The mass transfer coefficients were found to be in accordance with values found in literature and the effective ash layer coefficient was further found to become more prominent as the particle density increased. It was also found that the fitted values for external mass transfer coefficients followed the trend predicted in the literature.
Acknowledgement The authors would like to thank Sasol for the funding of this study. References [1] Hopkins, R. (2006). Energy Descent Pathways: evaluating potential responses to peak oil. England. [2] Beer, J. M. (2000). Combustion technology developments in power generation in response to environmental challenges. Progress in Energy and Combusiton Science. 26, pp.301-327. [3] Wang, S. C. and Wen, C. Y. (1972). Experimental evaluation of nonisothermal solid-gas reaction model. AIChE Journal. 18(6), pp.1231-1238.
[4] Koekemoer, A. F. (2009). The influence of minerals content and petrographic composition on the gasification of inertinite-rich high ash coal. Masters thesis. Potchefstroom: North-West University. [5] Oberholzer, A. (2009). Characterization and steam gasification of large, low grade coal particles using a specially designed thermo gravimetric analyser. Masters thesis Potchefstroom: North-West University. [6] Niska, S., Liu, G. S. and Hurt, R. H. (2003). Coal conversion submodels for design applications al elevated pressures. Part 1. Devolatilization and char combustion. Progress in Energy and Combustion Science. 29, pp.524-477. [7] Mendez, L. B., Martinez-Tarazona, M. R., Borrego, A. G., and Menendez, R. (2003). Influence of petrographic and mineral matter composition of coal particles on their combustion reactivity. Fuel. 82, pp.1875-1882. [8] Remiarova, B., Markos, J., Zajdlik, R. and Jelemensky, L. (2004). Identification of the mechanism of coal char particle combustion by porous structure characterization. Fuel Processing Technology. 85, p.303. [9] Holikova, K., Markos, J. and Jelemensky, L. (2004). Mechanism of coal char burning at a low oxygen content in the feed stream. Journal of Thermal Analytic Calorim. 76(1), p.237. [10] Tevrucht, M. L. and Griffiths, P. R. (1989). Activation energy of air-oxidized bituminous coals. Energy and Fuels. 3, pp.522-527. [11] Essenhigh, R. H. (1991). Rate equations for the carbon-oxygen reaction: an evaluation of the Langmuir adsorption isotherm at atmospheric pressure. Energy Fuel. 5, pp.41-46. [12] Park, K. Y. and Edgar, T. F. (1987). Modelling of Early Cavity Growth for Underground Coal Gasification. Ind. Eng. Chem. Res. 26, pp.237-246.
Oviedo ICCS&T 2011. Extended Abstract
Effect of Blending Waste Materials with Coal on Minerals and Reactivity of Char and Coke A.M. Fernández1, C. Barriocanal1, S. Gupta2, D. French3 1 Instituto Nacional del Carbón, CSIC, POB 73, 33080 Oviedo. Spain. E-mail address: [email protected] 2 SMaRT@UNSW, School of Materials Science & Engineering, UNSW, Sydney, Australia 3 CSIRO Energy Technology, North Ryde, NSW Australia. Abstract Three types of carbon bearing waste materials were blended with a typical medium rank coking coal in order to improve the coke quality and increase the recycling of waste materials for ironmaking. The additives included were: coke fines, two types of waste tyres, a bituminous residue from the distillation column of benzol and a commercial coal tar pitch. It was concluded that the presence of additives did not significantly modify the mineral phases of the cokes. However, the reactivity of the blended cokes showed a slightly increase. Keywords: Coke; tyre, pitch, mineralogy, reactivity.
1. Introduction Coke is essential for producing iron via the blast furnace route. However, due to rising costs, a variety of additives are commonly used in cokemaking in order to decrease the consumption of coking coal, to recycle waste products “in situ”, and also to improve the coke’s properties or to reduce costs. High quality coke in a blast furnace is essential to achieve a low rate of coke consumption, high productivity and a cheaper production of hot metal [1-3]. For these reasons, the interaction of such additives with coal and their influence on coke quality is of growing interest [4-9]. The aim of the present study is to investigate the influence of carbon bearing waste materials on coke quality and their impact on coke reactivity.
2. Experimental section A typical medium rank Australian coal (A) was chosen as the base coal of the coke blends. The additives blended with this coal were coke fines (CF), residues obtained from a tyre recycling plant (Ty, TxTy); a bituminous residue (RP) from a coke oven gas treatment installation and a commercial tar pitch (CTP). The base coal and their blends
1
Oviedo ICCS&T 2011. Extended Abstract
were carbonized in a semi-pilot coke oven [10]. The additives were pyrolyzed separately in a horizontal oven at 1000 ºC. A custom-made fixed bed reactor was used to measure the CO2 reactivity of the char and coke specimens as described elsewhere [11]. In order to elucidate the elemental composition of inorganic matter present in each sample, an XRF analysis was conducted by means of a SRS 3000 Bruker spectrometer as per ASTM D4326-04 standards [12]. Coke/char samples were examined using a Field Emission Scanning Electron Microscope (Model QUANTA FEG 650).
3. Results and Discussion Table 1 shows the proximate and ultimate analysis of the base coal and additive samples used in this study. Table 1 Proximate and ultimate analysis of the coal and additives Coal and additive composition A CF Ty TxTy CTP RP Proximate analysis, db Moisture, wt % 0.9 1.0 0.7 0.7 Ash yield, wt % 9.8 10.7 9.3 8.4 0.9 0.9 Volatile matter, wt % 22.3 1.1 63.0 65.7 65.7a 69.7 a Ultimate analysis, daf Carbon, wt % 88.7 96.2 87.1 83.1 92.8 87.6 Hydrogen, wt % 5.0 0.6 7.6 7.2 4.2 5.1 Nitrogen, wt % 2.1 1.5 0.3 0.3 1.1 3.6 Sulfur, wt % 0.55 0.53 2.00 1.74 0.62 1.72 Oxygen, wt % 3.6 1.2 3.1 7.7 1.3 2.0 a values obtained from the thermogravimetric analysis The chemical composition of the coal and additive ashes are shown in Table 2. Coal A does not have any significant amount of alkalis but it does contain some iron, as well as acidic components. Coke fines have slightly higher iron and calcium contents. The tyre waste samples are rich in zinc and have high calcium and sodium concentrations. The ash yields of the two pitch samples are low. Both of them have high iron contents, especially the residual pitch sample which also has the lowest silica content.
2
Oviedo ICCS&T 2011. Extended Abstract
Table 2. Oxide analysis of the raw coal and additives (Oxide wt.%) Ash yield wt % SiO2 Al2O3 TiO2 Fe2O3 CaO MgO K2O Na2O P2O5 ZnO a a
Samples A CF Ty TxTy 8.93 10.27 9.36 8.06 5.64 6.02 5.57 4.39 2.46 3.11 0.14 0.14 0.17 0.18 0.03 0.03 0.38 0.44 0.06 0.12 0.12 0.22 0.66 0.60 0.03 0.06 0.07 0.06 0.09 0.12 0.04 0.04 0.03 0.05 0.67 0.63 0.01 0.07 0.03 0.03 0.00 0.00 2.09a 2.02a
CTP RP 0.79 0.88 0.33 0.09 0.10 0.03 0.01 0.00 0.19 0.74a 0.10 0.01 0.02 0.00 0.01 0.01 0.02 0.00 0.01 0.00 0.00 0.00
based on atomic absorption analysis
Figure 1 compares the variation of the apparent reaction rate of CF, coke A and coke blend A10CF (10 wt % coke fines + 90 wt % coal A). It shows that the reaction rate of coke A, after 10 % conversion, is less than 10 units which is within the range of the reaction rates of typical high CSR cokes. The reaction rate of coke fines is more than two times the former rate. Consequently, the reaction rate of the base coal increases with
A pparent R ate x 10 -6 (g/g/s)
the addition of 10 % coke fines. 13 11 9 7 CF Coke-A10CF Coke-A
5 3 0.00
0.02
0.04
0.06
0.08
0.10
Fractional wt loss
Figure 1 Variation of CO2 apparent reaction rates of coke A, the coke blend and CF with carbon conversion. Coke A10CF (Figure 2) contains a considerable amount of iron that can catalyse the gasification reaction.
3
Oviedo ICCS&T 2011. Extended Abstract
Figure 2 SEM image of coke A10CF Coke A10CF Figure 3 compares the variation of the apparent reaction rate of tyre chars, coke A and coke blends A5Ty and A5TxTy (5 wt % waste tyre + 95 wt % coal A). The
9 a) 8 7 6 Coke-A5TxTys Coke-A5TY Coke-A
5 4 0.00
0.02
0.04 0.06 Fractional wt loss
0.08
0.10
A p parent R ate x 10 -6 (g /g/s)
A p p aren t R a te x 1 0 -6 (g /g /s)
reactivity of the coke is enhanced by the addition of the waste tyre material. 40
b)
36 32 28 Char-TxTy Char-Ty
24 20 0.00
0.02
0.04 0.06 Fractional wt loss
0.08
0.10
Figure 3 Variation of the CO2 apparent reaction rates of a) coke, coke blends and b) tyre char samples with carbon conversion. SEM/EDS analysis of coke A5Ty (Figure 4) shows that zinc is present in the coke in association with sulphur, indicating that either zinc sulphide or sulphate is present in the amorphous phase. It is most likely that the zinc phases are catalysing the gasification reaction.
4
Oviedo ICCS&T 2011. Extended Abstract
Figure 4 a) SEM image of coke A5Ty Figure 5 compares the variation of the apparent reaction rate of both types of pitch semicokes with that of coke A with fractional carbon conversion. The semicokes show a higher reactivity than the coke A. Consequently, the reactivity of the coke blends
A pparent R ate x 10 -6 (g/g/s)
containing pitch may be higher than that of the single coke (A). 40 35 30 25 20 15 10 5 0 0.00
Semicoke-RP Semicoke-CTP Coke-A
0.02
0.04
0.06
0.08
0.10
Fractional wt loss
Figure 5 Variation of CO2 apparent reaction rates of coke A, the residual pitch (RP) and the commercial tar pitch (CTP) semicokes with carbon conversion. The SEM images of coke A5RP (5 wt % residual pitch + 95 wt % coal A) are shown in Figure 6. Although the pitch additives are characterized by high iron contents (Table 2), especially the residual one, the increase in iron in coke A5RP is not important due to the low percentage of ash from the residual pitch.
Figure 6. a) SEM images of semicoke RP, b) semicoke CTP and c) coke A5RP.
4. Conclusions The apparent reaction rates of the coke blends made of coke fines and waste tyre was higher than that of the base coke. This can be attributed to the high reactivity of these
5
Oviedo ICCS&T 2011. Extended Abstract
additives and the modification of the base coke mineralogy. The iron oxides found in the coke fines and zinc minerals present in waste tyres could have catalysed the gasification reaction of the blend cokes. The reactivity of the residual pitch semicoke was found to be higher than that of the commercial one because of its higher iron content. Pitch blending did not produce any appreciable change in the mineral composition of the coke blends.
Acknowledgement AMF thanks the Government of The Principality of Asturias for the award of a predoctoral grant with funds from the PCTI-Asturias. AMF would also like to thank Dong-Min Jang from UNSW for his assistance in conducting the reactivity measurements.
References [1]
Arendt P. CRI and CSR - An assessment of influential factors. Cokemaking
International 2000; 12:62-68. [2]
Patrick JW, Wilkinson HC. Coke reactivity. The yearbook of the coke oven
manager’s association; 1983, p. 191. [3]
Ariyama T, Sato M. Optimization of ironmaking process for reducing CO2
emissions in the integrated steel works. ISIJ International 2006; 46:1736-1744. [4]
Elliot MA (Ed.). Chemistry of Coal Utilization, 2nd Supplementary Volume.
Wiley Interscience, New York; 1981. [5]
Loison R, Foch P, Boyer A. Coke. Quality and Production, Butterworths,
London; 1989. [6]
Alvarez R, Barriocanal C, Díez MA, Cimadevilla JLG, Casal MD, Canga CS.
Recycling of Hazardous Waste Materials in the Coking Process. Environmental Science and Technology 2004; 38 (5): 1611-1615. [7]
Diaz MC, Steel KM, Drage TC, Patrick, JW., Snape, C.E. Determination of the
effect of different additives in coking blends using a combination of in situ hightemperature 1H NMR and rheometry. Energy & Fuels 2005; 19 (6): 2423-2431. [8]
Menendez JA, Pis JJ, Alvarez R, Barriocanal C, Canga CS, Diez MA.
Characterization of petroleum coke as an additive in metallurgical cokemaking. Influence on metallurgical coke quality. Energy & Fuels 1997; 11: 379-384. [9]
Best MH, Burgo JA, Valia HS. Effect of coke strength after reaction (CSR) on 6
Oviedo ICCS&T 2011. Extended Abstract
blast furnace performance. Proceedings - Ironmaking Conference 2002; pp. 213-239. [10]
Casal MD, Canga CS, Diez MA, Alvarez R, Barriocanal C. Low-temperature
pyrolysis of coal with different coking pressure characteristics J. Anal. Appl. Pyrolysis 2005; 74: 96–103 [11]
Grigore M, Sakurovs R, French D, Sahajwalla V. Influence of mineral matter on
coke reactivity with carbon dioxide. ISIJ International 2006; 46(4):503-512. [12]
Vassilev SV, Menendez T, Alvarez D, Diaz-Somoano M, Martinez-Tarazona
MR. Phase-mineral and chemical composition of coal fly ashes as a basis for their multicomponent utilization. 1. Characterization of feed coals and fly ashes. Fuel 2003; 82: 1793–1811. [13]
Grigore M, Sakurovs R, French D, Sahajwalla V. Coke gasification: The
Influence and Behaviour of Inherent Catalytic Mineral Matter. Energy & Fuels 2009; 23:2075-85.
7
An investigation into the catalytic potential of coal ash constituents on the CO2 gasification rate of high ash, South African coal. B.B. Hattingh1*, R.C. Everson1, H.W.J.P. Neomagus1 and J.R. Bunt1,2 1
Energy Systems, School of Chemical and Minerals Engineering, North-West University,
Potchefstroom Campus, Private Bag X6001, Potchefstroom 2520, South Africa. 2
Sasol Technology (Pty) Ltd, Box 1, Sasolburg, 1947, South Africa
*Corresponding author. Tel.: +27 18 299 1545; E-mail address: [email protected] ________________________________________________________________________ Abstract CO2 gasification experiments were conducted on three South African coals with similar carbon-structural properties in order to evaluate the catalytic potential of inorganic species contained within the ash. Gasification experiments were performed in a high pressure thermogravimetric analyser on coal particles measuring 200 μm in size. A response surface methodology was followed to assess coal reactivity at temperatures ranging from 900-1000°C, pressures ranging from 1-10 bar and CO2 composition ranging from 10-30 mol.%. For all three coals, reactivity decreased with ash content and was found to be dependent on the ash composition. Specifically, it was found that reactivity showed an increase with increasing amounts of CaO and MgO in the ash as well as the Alkali Index (AI) and Al2O3/SiO2 ratio Keywords: South African coal, characterisation, CO2 gasification, catalytic effect, ash constituents
1. Introduction
The importance of coal today, apart from its role as an important energy source, lies in its ability to be converted into useful chemical products as well as liquid fuels [1]. The conversion of coal to petrochemical products involves the initial step of gasification which primarily consists of initial rapid pyrolysis (devolatilization) of the coal to form char, tar and gasses; and the subsequent gasification of the formed char to produce “syngas” [2-4]. Coal type and coal characteristic properties such as carbon content, volatile matter, maceral content and rank have shown to play an important role during
gasification behaviour and has generally been attended to in literature [5-7]. Furthermore it has been established that (1) maceral- and microlithotype composition [5-7], (2) concentration of active sites or reaction surface area (micropore surface area) [8-10], (3) accessibility of a particular reactant gas to the active sites (porosity) [11] and (4) the catalytic effect of the inorganic constituents (the presence of catalytically active species such as Ca-, K-, Na- and Mg) [12-16] are some of the major factors affecting gasification rates. Currently a particular interest has been shown in assessing the catalytic activity of inorganic species on coal gasification. A large number of these studies however consist of removing the original mineral matter by accustomed demineralisation techniques and impregnating/loading the demineralised coal/char with inorganic additives such as K2CO3 and Na2CO3 [15-18]. The purpose of this investigation is to report the inherent (already contained minerals within the coal) catalysis on CO2 gasification based on the comparison between the effects of the different parent properties of the coal.
2. Experimental section
Three coals: XA, XB and XC all originating from the South African Highveld no. 4 seam were used in this investigation. The choice of the three coals was based on their similarity in rank and carbon-structural properties (elemental, structural and petrographical properties). All three coals were therefore fully characterised on a chemical, petrographical and physical basis according to standard procedures and methods.
CO2 gasification rate measurements were performed on screened coal samples with an average particle size of 200 µm. The gasification experiments were conducted in a Bergbau-Forshung GMBH7 high pressure thermogravimetric analysis system (TGA) [19], which consisted of lowering the desired amount of sample (45 mg in this case) into a well mixed gaseous region that was controlled at a specific reaction temperature and reactant partial pressure. The effect of total pressure (1, 5.5 and 10 bar), temperature (900, 950 and 1000°C) and reactant partial pressure (10, 20 and 30 mol.% CO2) was evaluated by implementation of a response curve methodology (RCM) as a statistical method for combining variables [20-21]. For this particular case a face-centred central composite design method (CCD) was employed to conduct the experiments [20-21] in a random order as to prevent any systematic bias.
2
3. Results and Discussion
A full characterisation study confirmed that all three coals contained similar elemental, surface and petrographical features, with the largest differences, from a coal characteristic perspective, attributed to the mineralogical properties. The three coals were therefore characterised as Bituminous Medium rank C-D, inertinite-rich coals (71.8-75.9 vol.% m.m.f.b.) containing relatively large amounts of ash with values of 21.9 wt.%, 29.2 wt.%, 44.4 wt.% (d.b) respectively for coals XA, XB and XC. The mineral- and ash analyses, as obtained from XRD- and XRF (on a loss of ignition free basis) analysis, are presented in Table 1 and 2 respectively.
Table 1: Mineral composition (graphite-free basis) of coals XA, XB and XC (wt %) Species Anatase Calcite Dolomite Kaolinite Muscovite Pyrite Quartz Rutile Siderite
XA
XB
XC
0.18 6.32 13.95 54.32 2.21 2.04 19.64 1.05 0.28
1.22 3.22 9.30 48.61 8.52 1.27 27.01 0.78 0.07
0.79 2.86 6.88 50.40 10.60 2.43 24.90 1.14 0.00
Table 2: Ash composition of coals XA, XB and XC as wt.% and on an ash mass basis (g/100g coal). Species SiO2 Al2O3 Fe2O3 P2O5 TiO2 CaO MgO K2O Na2O SO3
XA wt.% 44.00 25.95 1.32 1.71 1.64 15.28 4.31 0.33 0.91 4.53
XB a.m.b (g) 9.46 5.58 0.28 0.37 0.35 3.29 0.93 0.07 0.20 0.97
wt.% 52.80 26.04 1.89 0.25 1.90 8.80 3.04 0.42 0.56 4.26
XC a.m.b (g) 14.52 7.16 0.52 0.07 0.52 2.42 0.84 0.11 0.15 1.17
wt.% 64.12 23.96 3.25 0.17 1.14 2.38 1.19 1.51 0.28 2.05
a.m.b (g) 28.47 10.64 1.44 0.08 0.51 1.05 0.53 0.67 0.12 0.91
All three coals contained substantially large amounts of kaolinite and quartz, with coal XA having the least amount of these two species. This also holds true for what was observed for Al2O3 and SiO2 present in the ash (both on wt.% and a.m.b. basis). Anatase and siderite was only present in relative small amounts within all three coals. No significant differences (wt.% and a.m.b) were observed in the amounts of Fe2O3 and K2O 3
present in the ash of coals XA and XB. The amounts of these species were however slightly higher for coal XC. This slightly higher amount of Fe2O3 present in the ash of coal XC can be related to the higher amount of pyrite present in this coal in comparison to coals XA and XB. A large difference was however found between the amounts of CaO (mainly derived from calcite and dolomite) present in the ash of each coal. Of all three coals the largest amount of this species was confined to coal XA (15.3 wt.%).
From the CO2 reactivity experiments of the three coals it was established that coal reactivity increased for a respective increase in temperature, pressure and CO2 composition. In addition for all the different reaction conditions (CCD experimental combinations) it was found that coal XA showed the fastest reaction rate, followed respectively by coals XB and XC. In order to explain the substantial difference between the reaction rate of the three coals, the influence of the different coal properties on reactivity was assessed. This was done by constructing qualitative plots of the dimensionless relative reactivity (Rrel) (as defined in Equations 1 and 2) and the respective coal properties. The relative reactivity for each coal was calculated from its initial reactivity (R0) relative to the coal with the highest initial reactivity (in this case coal XA). R0 =
dX dt
Rrel =
Eq. (1) t =0
dX / dt t =0 dX / dt
× 100
Eq. (2)
XA,t = 0
The average relative reactivity of each coal was used in the construction of the qualitative plots. The values of initial- and relative reactivities for all three coals at the respective CCD experimental conditions are summarised in Table 3. No systematic trends (in terms of catalytic effect) could be observed between the relative reactivities of the three coals and operating conditions. In all cases coal XA has the largest initial- and relative reactivity followed by coals XB and XC. Coal reactivity (initial and relative) was found to decrease with increasing ash value, which differed however from what was observed by Samaras et al. [22] and Skodras and Sakellaropoulos [23].
4
Table 3: Summary of initial reactivities of coals XA, XB and XC at different CCD conditions. Experimental conditions P yCO2 T (°C) (bar) (mol.%) 1000 1 30 1000 1 10 1000 5.5 20 1000 10 30 1000 10 10 950 1 20 950 5.5 30 950 5.5 20 950* 5.5 20 950* 5.5 20 950 5.5 10 950 10 20 900 1 30 900 1 10 900 5.5 20 900 10 30 900 10 10 Average Rrel
Coal initial reactivity x 102 (min-1) and relative reactivity XA XB XC R0 Rrel R0 Rrel R0 Rrel 4.34 2.26 1.32 100 52 30 1.90 1.14 0.67 100 60 35 4.29 2.02 1.45 100 47 34 6.30 2.57 2.04 100 41 32 2.08 1.60 1.17 100 77 56 1.46 0.65 0.48 100 45 33 2.48 1.28 0.89 100 52 36 1.64 0.92 0.73 100 57 45 1.70 0.89 0.65 100 52 38 1.83 0.95 0.71 100 52 39 1.07 0.62 0.51 100 58 48 1.84 0.90 0.78 100 49 42 0.75 0.32 0.28 100 43 38 0.48 0.26 0.19 100 54 40 0.85 0.49 0.35 100 58 41 1.10 0.64 0.56 100 58 51 0.52 0.36 0.30 100 69 58 100 54(±4.3) 41(±3.9)
* Repeatability runs This inconsistency between the effects of ash on reactivity possibly indicates that the reactivity of a specific coal is not solely ash amount dependent, but more dependent on the inherent constituents of the ash. Therefore the abundance of a certain constituent can enhance the inherent catalytic effect of the ash. A comparison between relative reactivity and mineral contents present in the ash revealed that relative reactivity increased for increasing CaO and MgO contents, which indicated an enhanced catalytic effect for coal XA, which contained the largest amount (wt.% and a.m.b) of these species. This corresponded to the catalytic effects of Ca and Mg observed by numerous other authors [12-15,22,23,25,26]. Two indices have however been proposed in an attempt to describe the overall catalytic effect of inorganic species within coal. The AI [27] is a parameter frequently used to describe the overall influence of catalytically active species within the ash and is defined as the ratio of the sum of the fraction of the basic compounds in the ash (CaO, MgO, K2O, Na2O and Fe2O3) to the fraction of the acidic compounds (SiO2 and Al2O3) in the ash, multiplied by the ash value (Eq. (3)): ⎛ CaO + K 2 O + MgO + Na 2 O + Fe2 O3 ⎞ ⎟⎟ AI = ash% × ⎜⎜ Al 2 O3 + SiO2 ⎠ ⎝
5
Eq. (3)
Another predictive indicator used for assessing the influence of ash components on coal reactivity, is the Al2O3/SiO2 ratio [28]. The influence of the AI and the Al2O3/SiO2 ratio
120
120
100
100
Average Rrel
Average Rrel
on the average relative reactivity is depicted in Figure 1.
80 60 XA 40 XB 20
80 60 XA 40 XB 20
XC
XC
0
0 0.37
0.49
4.2
0.59
5.2
6.5
Alkali Index
Al2O3/SiO2 ratio
Figure 1: Influence of Al2O3/SiO2 ratio and Alkali Index on coal relative reactivity.
From the figure it is clear that coal reactivity increases for both an increase in Al2O3/SiO2 ratio and AI value. The increased reactivity of coal XA can therefore be explained by the large amount of catalytic species such as CaO and MgO etc. and lower concentrations of SiO2 present in the ash, thus confirming the relatively strong catalytic effect of the ash of coal XA. Similar trends were observed by Skodras and Sakellaropoulos [23], Sakawa et al. [27] and Zhang et al. [29]. Based on the above observations the predominance of inherent catalytic active species such as Ca2+ can enhance the reactivity of a particular coal. Equally an increase in the presence of non-active ash constituents can significantly reduce the heating value and available reactive surface area, which can consequently lead to a lower reactive nature of the coal.
4. Conclusions
The reactivity of the coals with a particle size between 150 and 250 μm was determined in a thermogravimetric analyser. The gasification reactivity was evaluated at temperatures between 900 and 1000 oC, pressures between 1 and 10 bar, and mole fractions of CO2 between 10 and 30%. For the selected coals, the reactivity decreased with ash content, and was found to be dependent on the composition of the ash. Specifically, the increase in gasification rate could be well correlated with calcium and magnesium content, Al2O3/SiO2 ratio and alkali index respectively.
6
References:
[1]
K. Miura, Mild conversion of coal for producing valuable chemicals. Fuel Processing Technology 2000:62: 119-135.
[2]
J.R. Bunt, F. Waanders, Identification of the reaction zones occurring in a commercialscale Sasol-Lurgi FBDB gasifier. Fuel 2008:87:1814-1823.
[3]
J.C. Van Dyk, M.J. Keyser, J.W. Van Zyl, Suitability of feed stocks for the Sasol-Lurgi Fixed Bed Dry Bottom gasification process. Proceedings of 2001 conference on gasification technologies, San Francisco, USA, 2001.
[4]
J.C. Van Dyk, M.J. Keyser, M. Coertzen, Syngas production from South African coal sources using Sasol-Lurgi gasifiers. International Journal of Coal Geology 2006:65(34):243-253.
[5]
M. Cloke, E. Lester, Characterization of coals for combustion using petrographic analysis: A review. Fuel 1994:73(3):315-320.
[6]
J.C. Crelling, N.M. Skorupska, H. Marsh, Reactivity of coal macerals and lithotypes. Fuel 1988:67:781-785.
[7]
R.H. Hurt, J.R. Gibbens, Residual carbon from pulverized coal fired boilers: 1. Size distribution and combustion reactivity. Fuel 1995:74(4):471-480.
[8]
H.N.G. Siauw, D.P.C. Fung, S.D. Kim, Study of the pore structure and reactivity of Canadian coal-derived chars. Fuel 1985:67:700-706.
[9]
S. Kajitani, N. Suzuki, M. Ashizawa, S. Hara, CO2 gasification rate analysis of coal char in entrained flow coal gasifier. Fuel 2006:85:163-169.
[10]
G.Ö. Çakal, H. Yücel, A.G. Gürüz, Physical and chemical properties of selected Turkish lignites and their pyrolysis and gasification rates determined by thermo gravimetric analysis. J. Anal. Appl. Pyrolysis 2007:80:262-268.
[11]
H. Slaghuis, T.J. Van der Walt, D. Glasser, D. Hildebrandt, Anomalies in gasification kinetic behaviour of some South African coal chars, Proceedings of 1993 international Pittsburgh coal conference on coal-energy and environment, Pittsburgh, USA, 1993.
[12]
D. Cazorla-Amoros, A. Linares-Solano, C. Salinas-Martinez De Lecea, A temperatureprogrammed reaction study of calcium-catalyzed carbon gasification. Energy & Fuels 1992:6(3):287-293.
[13]
C. Nattermann, K.J. Hütinger, Correlations between coal reactivity and inorganic matter content for pressure gasification with steam and carbon dioxide. Fuel 1994:73(10):16821684.
[14]
R. Meijer, H.J. Muhlen, J.A. Moulijn, Burn-off behaviour in alkali-catalyzed CO2 gasification of bituminous coal char: A comparison of TGA and Fixed-bed reactor. Fuel Process. Technol. 1991:28:5-17. 7
[15]
A. Molina, A. Montoya, F. Mondragón, CO2 strong chemisorption as an estimate of coal char gasification reactivity. Fuel 1999:78:971-977.
[16]
J. Adanez, L.F. De Diego, Reactivity of lignite chars with CO2: influence of the mineral matter, Int. Chem. Eng. 1993:33(4):656-662.
[17]
Q. Sun, W. Li, H. Chen, B. Li, The CO2 gasification and kinetics of Shenmu maceral chars with and without catalyst. Fuel 2004:83:1787-1793.
[18]
G. Jaffri, J. Zhang, Catalytic activity of the black liquor and calcium mixture in CO2 gasification
of
Fujian
anthracite,
Chinese
Journal
of
Chemical
Engineering
2007:15(5):670-679. [19]
R.C. Everson, H.W.J.P. Neomagus, R. Kaitano, The modeling of the combustion of highash coal–char particles suitable for pressurised fluidized bed combustion: shrinking reacted core model. Fuel 2005:84:1136-1143.
[20]
D.C. Montgomery, Design and analysis of experiments. 4th ed., New York: John Wiley & Sons; 1997.
[21]
R.H. Myers, D.C. Montgomery, Response Surface Methodology: Process and product optimization using designed experiments. 2nd ed., New York: John Wiley & Sons; 2002.
[22]
P. Samaras, E. Diamadopoulos, G.P. Sakellaropoulos, The effect of mineral matter and pyrolysis conditions on the gasification Greek lignite by carbon dioxide. Fuel 1996:75(9):1108-1114.
[23]
G. Skodras, G.P. Sakellaropoulos, Mineral matter effects in lignite gasification. Fuel Process. Technol. 2002:77-78:151-158.
[24]
D.P.C. Fung, S.D. Kim, Chemical reactivity of Canadian coal derived chars. Fuel 1983:63:1197-1201.
[25]
S. Su, J.H. Pohl, D. Holcombe, J.A. Hart, A proposed maceral index to predict combustion behaviour of coal. Fuel 2001:80:699-706.
[26]
L.R. Radovic, P.L. Walker Jr, R.G. Jenkins, Effect of lignite pyrolysis conditions on calcium oxide dispersion and subsequent char reactivity. Fuel 1983:62:209-212.
[27]
M. Sakawa, Y. Sakurai, Y. Hara, Influence of coal characteristics on CO2 gasification. Fuel 1982:61:717-720.
[28]
D. Erincin, A. Sinağ, Z. Misirlioğlu & M. Canel, Characterisation of burning and CO2 gasification of chars from mixtures of Zonguldak (Turkey) and Australian bituminous coals. Energy conversion and management 2005:46:2748-2761.
[29]
L. Zhang, J. Huang, Y. Fang, Y. Wang, Gasification reactivity and kinetics of typical Chinese anthracite chars with steam and CO2. Energy & Fuels 2006:20:1201-1210.
8
Catalyst Recovery using by Support in Steam Gasification of Lignite at Low Temperature Young-Kwang Kim1, Joo-Il Park1, Jin Miyawaki2, Isao Mochida3, Seong-Ho Yoon1,2* 1
Interdisciplinary Graduated School of Engineering Sciences, Kyushu University 2 3
Institute of Materials Chemistry and Engineering, Kyushu University
Research and Education Center of Carbon Resources, Kyushu University
Abstract Low temperature catalytic gasification of lignite was studied using novel alkali metal carbonate (K2CO3) supported perovskite oxide catalyst. The novel developed catalysts of K2CO3 supported on perovskite oxides showed higher activity than K2CO3 supported on γ– alumina and K2CO3 only under gasification conditions of less than 800 oC. Furthermore, when using the novel catalysts, the syngas composition showed higher H2 and lower CO2 ratios compared to the gasification without any catalyst. More hopefully, tar formation sharply decreased to less than 50 wt% of the non–catalytic gasification. Also, after the gasification reaction, less loss of K2CO3 and no coke formation were confirmed on the surface of used catalysts.
1. Introduction Coal gasification is currently being widely examined as an effective procedure to convert the coal for power generation and CO/H2 production which can reduce the carbon emission at the use of coal as energy [1]. Hydrogen as the main product of coal gasification is expected to be an important secondary energy in the 21st century in the CO2 free society [2]. The gasification with steam is considered as a method of increasing H2 yield through the reduction of H2O during the gasification. The steam gasification of coal is a highly endothermic slow reaction, requiring high temperature above 1000 oC under some pressure [3]. Commercially acceptable gas production rate at lower temperature below 800 oC needs the catalyst to enhance the steam gasification. Among the catalysts proposed, the most favored catalysts are alkali metal salts especially K2CO3 [4-6]. However, large loss has been reported from the reaction region by evaporation during the reaction at the temperature around 800 oC [7-11]. The supporting
material is a method to solve this issue. Yamaguchi et al. [12] introduced alumina as a support of K2CO3. K2CO3 supported on alumina showed higher catalytic activity than K2CO3 itself. However, recycling ability of the catalyst supported on alumina was not sufficient. In a previous study [11], we found an interesting support for effective catalytic activation and keeping loss of K2CO3 in the coal combustion. We reported that K2CO3 supported on several perovskite type oxides showed catalytic ability and good recovery of K2CO3 on the support after the combustion. In the present study, we studied the catalytic activity and catalyst holding of K2CO3 / perovskite system in the steam gasification of an Indonesian lignite in the fluidized bed gasifier. Among of supports, we selected LaMn0.8Cu0.2O3 system which showed excellent performance in the preliminary study. The catalyst of K2CO3 supported on LaMn0.8Cu0.2O3 which carries the active form of potassium to be transferred from the perovskite surface through the reduction of K2CO3 on the surface is expected to activate H2O to gasify the coal. The active species can be vapored from the coal surface, but being captured by the perovskite surface in the fluidized bed. 2. Experimental method 2.1. Sample properties An Indonesian lignite, Adaro coal, was selected in the present study. Raw coal was ground to the size of 100~250 um, then dried at 80 oC for 24 hr in the heating oven. Table 1 shows the several properties of Adaro coal. Adaro coal carried a very low level of ash. Table 1. Properties of lignite (Adaro coal) Proximate analysis (wt%)
Coal Adaro Coal
Elemental analysis (wt%, dry & ash free base)
VM
Ash
C
H
N
43.3
2.84
70.91
5.075
1.026
O
diff.
22.989
Calorific value cal/g 6010
2.2. Catalyst preparation Perovskite–type oxide, LaMn1–xCuxO3 (LMC82, x = 0.2), was prepared according to sol–gel method [13, 14]. Quantitative amounts of lanthanum nitrate, manganese nitrate, and copper nitrate which were mixed and stirred for 8 h in citric acid water solvent. Water was evaporated from the mixed solution by vacuum evaporator at 45 °C, until a viscous gel was obtained. The gel was kept at 110 °C over–night and the resultant mass was then ground, and
calcined at 750 °C for 5 h. K2CO3 (0–20 wt%) was impregnated onto perovskite support (also, on γ–alumina) through immersing in K2CO3–ethanol/H2O. The solvent were removed at 110 o
C in vacuum for the wet impregnation (WI) method [11].
2.3. Experimental setup Catalytic steam gasification of the lignite (Adaro coal) was carried out in a bubbling type fluidized reactor and the fluidized was zone was separated by nickel mesh, as shown in scheme 1. Coal and catalyst (ca. 2 g, 50:50 wt %) were fluidized for gasification at a temperature from 600 to 800 °C. The reactor was heated up to the target temperature at about 200 °C min–1 of a ramping rate in N2 / H2O vapor flows (400 / 100 ml min–1 of flow rate) and kept at the target temperature. The produced syngas including steam was sent to a three stage–condenser in the first stage to collect heavier tar, the second to collect steam at –10 oC, and the last one to collect lighter product through the quartz filter. The product syngas (H2, CO, CO2, and CH4) was passed through 2 traps. One is for collecting the heavy tar, and another one is for collecting of light tar and cooling of gas. After then, gas was filtered through the quartz separator. The cleaned gas was analyzed every 5 min using gas chromatograph (J–science GC 7000). The lines between the steam generator and the reactor as well as between the reactor and the first stage of the tar collector were heated at 200 oC to avoid the condensation of tarry products. The collected tar was dissolved in THF solvent. After a solvent evaporation at 100 oC, the condensed tars were weighed to quantify the condensable tars. At the end of each experiment, the catalysts were sampled and separated from char or ash by size separation (ash < ~100 μm). The amount of potassium on the catalyst sampled was determined by ICP–Mass through the water extraction. The amount of coke included in the catalyst was quantified by TGA.
Scheme 1. Bubbling type fluidized bed reactor
3. Results 3.1. Catalytic activities of the catalysts Catalytic activities of developed perovskite supporting and reference catalysts were evaluated with Adaro coal in the steam gasification. Fig. 1 (a) shows the carbon conversion profiles for Adaro coal with and without the catalyst at 700 oC and 20 vol% of steam in N2. The sample, including of coal and catalyst, was added in the reactor and mixed by the bubbling gas flow. In gasification reaction, the reaction products were about 90–95 % of carbon in syngas and 5–10 % carbon in tar. It must be noted that tar component was not completely collected, because some light tar was evaporated when the drying of water for collecting heavy tar. This error appeared about 1 % in the carbon balance. The carbon conversion was calculated on dry, ash and catalyst free basis by the following equation:
(1) The equation (1) includes only the carbon conversion to syngas. Then, it was not reach 100 % of carbon conversion, because tar was not included in this calculation, as shown in Fig. 1 (a). The gas component including carbon in syngas was about 99 % being consisted of CO, CO2 and CH4. 100
(b)
10 wt% K2CO3 / LMC82 10 wt% K2CO3 / Al 10 wt% K2CO3 LMC82 Al coal only
90 80 70 60
100 90 80
CH4 CO2
70
Syngas (%)
Carbon conversion (mol%)
(a)
50 40 30
60
40 30
20
20
10
10
0
CO
50
H2
0 0
30
60
90
120
150
180
Time (min)
co al on ly
Al
LM C8 2
10 wt%
K2 CO 3
10 wt%
10 wt%
K2 CO 3/
Al
K2 CO 3/
LM C8 2
Fig. 1. (a) Carbon conversion ratios of steam gasification for Adaro coal at 700 oC: without and with catalysts (γ–alumina, LMC82, 10 wt% K2CO3, 10 wt% K2CO3 on γ–alumina, and 10 wt% K2CO3 on LMC82), (b) syngas compositions The initial weight loss in the first 10 min in Fig. 1 (a) was due to the volatile release.
A part of volatile was converted to syngas, while another part of heavier volatile was captured by two traps and a quartz filter. The amounts of the volatiles were similar in the non-catalytic and catalytic reactions. In coal gasification without catalyst, the curve shows very slow gasification rate after the initial weight loss of volatile release. Carbon conversion increased very slowly without any catalysts, indicating very slow gasification under present conditions. Γ–alumina and LMC82 increased slightly the gasification rates. K2CO3 did not increase significantly the gasification rate. 10 wt% K2CO3 supported on γ–alumina provided the same conversion with that of K2CO3. Alumina appears no enhancement of K2CO3 activity. 10 wt% K2CO3 supported on LMC82 gave much higher activity by 20 times than the same amount of K2CO3, achieving 75 % conversion within 80 min but the increase became slow to achieve 85 % by 170 min. The major gases were H2: 56 vol%, CO: 9 vol%, and CO2: 31 vol% in the product from coal alone, whereas H2: 61 vol%, CO: 11 vol%, and CO2: 26 vol% were produced over 10 wt% K2CO3 impregnated on LMC82. Catalyst enhanced not only the conversion of fixed carbon but also tar decomposition. In the initial step of steam gasification, volatile material was released from coal by heat. After then, the volatile material underwent steam reforming, and inactivated component remained as tar. The catalyst steam-reformed the tar into syngas effectively. Table 2 shows quantities of the condensed tar by the catalytic reforming. Table 2. Effect of tar decrement by catalytic steam gasification o
Catalyst
Amount of condensed tar (g), 700 C
w/o catalyst
0.0797 (100%)
10 wt% K2CO3
0.0640 (80.3%)
10 wt% K2CO3 / γ–alumina
0.0411 (52.6%)
10 wt% K2CO3 / LMC82
0.0469 (58.8%)
10 wt% K2CO3 / LMC82 x 2 g
0.0343 (43.0%)
( ) : relative to amount without catalyst
3.2. Catalytic steam gasification of Adaro coal at other temperatures Fig. 3 shows catalytic activities of K2CO3 supported LMC82 in the steam gasification at 600 to 800 oC. The catalytic activity was very dependent on the reaction temperature. Steam gasification of Adaro coal alone at 600 to 800 oC was very slow, but the catalyst increased the reactivity extremely above 700 oC. The higher temperature was the
initial conversion rate was faster. This tendency appears with melting temperature (891 oC) of K2CO3. Because, K2CO3 can be evaporated easily from the surface of support near melting point, and reacted with coal. However, catalyst loss is also increased at the high temperature, because K2CO3 mobility increased enough to run out from the reaction region. Though catalyst loss was about 20 % in catalytic gasification at 700 oC, it increased to about 80 % at 800 oC. Thus, 700 oC was the best temperature for the catalytic performance in this catalytic system. 100
o
800 C o 750 C o 700 C o 650 C o 600 C
Carbon conversion (mol%)
90 80 70 60 50 40 30 20 10 0 0
30
60
90
120
150
180
Time (min)
Fig. 3. Carbon conversions of catalytic steam gasification for Adaro coal with 10 wt% K2CO3 supported on LMC82 at 600–800 oC 3.3. Effects of catalyst amount The effects of K2CO3 impregnation amount on LMC82 in the steam gasification of Adaro coal at 700 oC is shown in Fig. 4. Impregnation amounts were 0, 5, 15, 20 wt% by LMC82 weight. At a given reaction time, the carbon conversion rate increased by increasing K2CO3 impregnation amount from 0, 5 to 20 wt%. However, the catalyst of more impregnation amounts of 15 and 20 wt% K2CO3 showed similar activities to that of 10 wt% K2CO3. When two times amount of 10 wt% K2CO3 on LMC82 was applied, catalytic gasification was more effective than one par of 20 wt% K2CO3 on LMC82 to increase the rate of carbon conversion. The complete conversion was achieved in about 120 min under the conditions.
100
Carbon conversion (mol%)
90 80 70 60 50 40 30
10 wt% K2CO3 / LMC82 x 2g 20 wt% K2CO3 / LMC82 15 wt% K2CO3 / LMC82 10 wt% K2CO3 / LMC82 5 wt% K2CO3 / LMC82 LMC82
20 10 0 0
30
60
90
120
150
180
Time (min)
Fig. 4. Effect of K2CO3 impregnation amount for steam gasification of Adaro coal at 700 oC: K2CO3 impregnation amount (0–20 wt%), and catalyst amount of 2 times of 10 wt% K2CO3 on LMC82 3.4. Analysis of the catalyst condition during coal gasification The size separation of the catalyst recovered after the gasification about 90 wt%, and another 10 wt% was impossible to recover from the ash. Fig. 5 shows potassium amount on the recovered catalysts during the catalytic gasification reaction. The support of γ–alumina released K+ continuously the during gasification reaction. After the gasification reaction, about 65 % of potassium was lost from γ–alumina support. The LMC82 lost potassium up to 40 % by the carbon conversion of 30 % like γ–alumina, however the amount of potassium on the support increased after the minimum to gradually to reach of 80 % when the carbon conversion reached 80 %. The support may release potassium once on the coal and char, but appears to recover potassium on the surface after the gasification. The main reasons for the loss are loss of potassium through the reaction region and trap on the ash. 100 90 80
K+ ion (%)
70 60 50 40 30 20 10 wt% K2CO3 / LMC82 10 wt% K2CO3 / Alumina
10 0 0
20
40
60
Carbon conversion (%)
80
100
Fig. 5. K+ ion amounts on supports during steam gasification at 700 oC Fig. 6 shows coke formation on the two kinds of the catalysts monitored during the catalytic gasification. The catalysts were contaminated by coke when the volatile material was released. The catalyst supported γ–alumina showed large coke formation about 16 wt% by 40 % of carbon conversion. Then, the coke produced was decomposed completely by 5 hr later when carbon conversion reached 85 %. In contrast, the catalyst supported LMC82 suffered very small amount coke below 2 wt% at largest which disappeared by the end of the gasification.
Coke formation (wt %, coke / catalyst)
25 10 wt% K2CO3 / LMC82 10 wt% K2CO3 / Alumina
20
15
10
5
0 0
20
40
60
80
100
Carbon conversion (%)
Fig. 6. Coke formation on catalysts during steam gasification at 700 oC
4. Discussion In the present study, we studied the catalytic activity and catalyst holding of K2CO3 / LMC82 system in the steam gasification of an Indonesian lignite in the fluidized bed gasifier. Lignite has a large portion of the volatile matter. Hence the catalytic gasification of this coal has two aspects, tar reforming and characterization. Although these volatile matter has high reactivity, it causes some problems of deactivation by the formation of tar and coke on the catalyst as well as the reactor wall. The catalytic reforming of the volatile over the same catalyst was found very effective, leaving no coke on the catalyst. Alumina support showed the significant trap of tar over 45 %, however, this effect appears the capture about tar by its large surface. The captured tar formed the coke on its surface to be slowly gasified. Large surface is favor for tar removal, however high surface reactivity for reforming into syngas is also required. Thus the perovskite surface is an excellent catalyst system to reforming the tar without any significant
formation of coke from the tar. After the release of volatile matter, coal char reactivity was depended on potassium on the support which much interact solid char. The large part of potassium was produced to be migrated from support to the char during the gasification. Since potassium species must contact the char surface to gasify the surface carbon, alkali carbonate may be reduced on the perovskite surface to migrate on the char surface during the fluidized mixing. The difference of activities of catalyst depending upon the support means that the activities of potassium depended on very much the supports. The activated potassium must be recovered on LMC82 at the end of the gasification. The reaction path of K2CO3 supported on perovskite oxide is schematically illustrated as shown in scheme 2. First, potassium active species on the surface of supports moves to coal. Second, the active species on coal gasified carbon, and then moves to neighboring carbon. Finally, potassium was recovered on the surface of support.
Scheme 2. Schematic diagram for shuttle mechanism of catalyst (K2CO3) on support Potassium is the active species and should be recovered for the recycle of catalyst. However, the 100 % of potassium recovery was not achieved so far. The evaporation from the reaction region and reaction with ash cause 20 % of loss of potassium in this gasification. From these results, a participation support of LMC82 showed several favorable effects on the catalytic performance of K2CO3 as below : First, the increment of reforming the volatile matters and tar removal Second, the activation of potassium species to migrate as gasify the char Third, the effective recovery of potassium species after gasification
References [1] Ekins P. Hydrogen Energy: Economic and Social Challenges: Earthscan 2009. [2] Shoko E, McLellan B, Dicks AL, da Costa JCD. Hydrogen from coal: Production and utilisation technologies. International Journal of Coal Geology. 2006;65(3-4):213-22. [3] Molina A, Mondrag F. Reactivity of coal gasification with steam and CO2. Fuel. 1998;77(15):1831-9. [4] Wang J, Sakanishi K, Saito I, Takarada T, Morishita K. High-Yield Hydrogen Production by Steam Gasification of HyperCoal (Ash-Free Coal Extract) with Potassium Carbonate: Comparison with Raw Coal. Energy & Fuels. 2005;19(5):2114-20. [5] Wang J, Jiang M, Yao Y, Zhang Y, Cao J. Steam gasification of coal char catalyzed by K2CO3 for enhanced production of hydrogen without formation of methane. Fuel. 2009;88(9):1572-9. [6] Douchanov D, Angelova G. Effect of catalysis and inlet gas on coal gasification. Fuel. 1983;62(2):231-3. [7] Meijer R, Weeda M, Kapteijn F, Moulijn JA. Catalyst loss and retention during alkali-catalysed carbon gasification in CO2. Carbon. 1991;29(7):929-41. [8] Huhn F, Klein J, J?tgen H. Investigations on the alkali-catalysed steam gasification of coal: Kinetics and interactions of alkali catalyst with carbon. Fuel. 1983;62(2):196-9. [9] Marsh H, Mochida I. Catalytic gasification of metallurgical coke by carbon dioxide using potassium salts. Fuel. 1981;60(3):231-9. [10] Shadman F, Sams DA, Punjak WA. Significance of the reduction of alkali carbonates in catalytic carbon gasification. Fuel. 1987;66(12):1658-63. [11] Miyazaki T, Tokubuchi N, Inoue M, Arita M, Mochida I. Catalytic Activities of K2CO3 Supported on Several Oxides for Carbon Combustion. Energy & Fuels. 1998;12(5):870-4. [12] Yamaguchi T, Wang Y, Komatsu M, Ookawa M. Preparation of New Solid Bases Derived from Supported Metal Nitrates and Carbonates. Catalysis Surveys from Japan. 2002;5(2):81-9. [13] Slagtern A, Olsbye U. Partial oxidation of methane to synthesis gas using La-M-O catalysts. Applied Catalysis A: General. 1994;110(1):99-108. [14] Doggali P, Kusaba S, Teraoka Y, Chankapure P, Rayalu S, Labhsetwar N. La0.9Ba0.1CoO3 perovskite type catalysts for the control of CO and PM emissions. Catalysis Communications. 2010;11(7):665-9.
Evaluation of coal gasification reaction from composition of gases produced (2) M. Kaiho1, O. Yamada1, H. Yasuda1, S. Shimada2 and M. Fujioka3 1
National Institute of Advanced Industrial Science and Technology 16-1 Onogawa, Tsukuba, Ibaraki 305-8569 JAPAN E-mail [email protected] 2 The University of Tokyo 5-1-5 Kashiwanoha, Kashiwa, Chiba 277-8563 JAPAN 3 Japan Coal Energy Center 9F, MeijiYasuda Seimei Mita Bldg., 3-14-10 Mita, Minato-ku Tokyo 108-0073 JAPAN
Abstract Estimation of chemical reaction process occurred in the reactor of coal gasification is subject of importance to evaluate and improve performance of plant operations. We have previously proposed methods to evaluate reaction formula as follows; CHmOn+αO2+βH2O→γH2+δCO+εCO2+ηCH4 where CHmOn is coal, based on the results of ultimate analysis of coal and chemical analysis of gas produced. We have prepared the mathematically derived reaction formula whose α ~ η were written as functions of the amount of O2 reacted, H2 burned, CO consumed by shift reaction and CH4 formation to elucidate a chemical reaction mechanism. In this study, an extended method which allows us to deal with the gas that contains C2 compounds and tar was proposed. The reaction formula is described as CHmOn＋αO2＋βH2O→γH2＋δCO＋εCO2＋ηCH4＋ζC2H4＋θC2H6＋λCHm’On’, where CHm’On’ is tar. Values of α ~ λ were calculated from the concentration of each gas produced and molar ratio of tar to total gas using equations that were derived based on the stoichiometry of above reaction formula. Reaction formula was also derived mathematically from the standard reaction formula by taking account of the variation of CHmOn+0.5(1-n)O2→0.5mH2+CO O2 and effects of shift reaction, and formation of CH4, C2H4, C2H6 and tar. Chemical process of Rocky Mountain-1 underground coal gasification project performed in USA was investigated using the method mentioned above.
1.
Introduction In order to improve performance of a plant operation in coal gasification, chemical reaction process occurred in a gasifier should be sufficiently clarified. In the fundamental studies of coal gasification, rates of water gas reaction and Boudouard reaction have been focused. Although knowledge on kinetics of these two reactions is variable to design various gasification processes, it alone is not sufficient to understand the chemical process in gasifier from gas composition because of two reactions out of more than 15 elementary reactions are not enough. In addition, it is difficult to select an appropriate reaction path scientifically from numbers of possible combinations of various elementary reactions which is assumed for the composition of gas produced. Equilibrium constants have been often used to understand gas composition produced by gasification. Since an argument base on chemical equilibrium is only available to the apparently stable state, it should be inapplicable to the analysis of transient chemical state in the gasifier during the actual operation which is constantly fluctuating in various conditions. We developed a method to estimate the reaction formula described by following manner, i.e., CHmOn+αO2+βH2O -> γH2+δCO+εCO2+ηCH4, from the composition of gas produced and reported previously1. Since the formula obtained is depending exactly on the material balance of gasification, it should be quite useful to evaluate carbon conversion, cold gas efficiency, heat of reaction, reaction temperature, and estimation of reaction process. This method, however, only allows an evaluation of high temperature gasification process such as reaction in an entrained bed gasifier, because we limited the products to H2, CO, CO2, and CH4 in order to simplify the mathematical procedure. To remove this disadvatage, we have modified the method applicable to gasification at lower temperature, such as fixed bed, fluidized bed, and underground coal gasification processes, by the addition of C2H4, C2H6, and tar to the reaction formula mentioned above. The equations to calculate quantity of each component in the formula can be derived by the same mathematical process that we have presented previously1,2. 2. Estimation of reaction formula We propose a method that allows us to deal with the gas that contains C2 compounds and tar. The reaction formula is described as follows, where CHm’On’ is tar; CHmOn+αO2+βH2O -> γH2+δCO+εCO2+ηCH4+ζC2H4+θC2H6+λCHm’On’ (1)
Increase in numbers of product included has enabled us to apply our method to the process that is operated at lower temperature than entrained bed gasification, such as fixed bed gasification and the underground coal gasification. Elemental balance for C, H and O in eq. (1) is described as in eqs. (2), (3) and (4) respectively. 1 + δ + ε + η + 2ζ + 2θ + λ m + 2β = 2γ + 4η + 4ζ + 6θ + m’λ n+ 2α + β = δ + 2ε + n’λ
(2) (3) (4)
The total moles of gas yield in eq. (1) is expressed as Σ. Σ=γ+δ+ε+η+ζ+θ
(5)
When the concentrations of H2, CO, CO2, CH4, C2H4, and C2H6 (dry and N2 free) are represented by p, q, r, s, t, and u respectively, the molar quantity of each gas is described as follows; H2 ; γ = pΣ CO ; δ = qΣ CO2 ; ε = rΣ CH4 ; η=sΣ C2H4 ; ζ=tΣ C2H6 ; θ=uΣ
(6) (7) (8) (9) (10) (11)
In practical process, a definite amount of sample gas has been separated from gasification plant by thin pipe, cooled to remove condensable water vapor and tar, and to analyze the volume and the gas composition. Yield of tar has been usually evaluated from the quantity of condensed matters in above sampling system. The mole ratio of tar yield to total gas is presented by v. Then λ in eq. (1) is described as eq. (12). λ = vΣ
(12)
Consequently, it will be possible to obtain ten solutions, α ~ η and Σ, mathematically because we could prepare the eleven equations from (2) to (12). (2) is rewritten as in
eq. (13) by employing (7) ~ (12). 1 = qΣ + rΣ + sΣ + 2tΣ + 2uΣ + vΣ
(13)
Eq. (13) is rearranged to eq. (14). 1 Σ=
─────────────── q + r + s + 2t + 2u + v
(14)
Yields of gas and tar, γ ~ λ in eq. (1) therefore can be expressed by eqs. (15) ~ (21). p H2 ;
γ = ─────────────── q + r + s + 2t + 2u + v
(15)
q CO
; δ =─────────────── q + r + s + 2t + 2u + v
(16)
r CO2 ;
ε = ─────────────── q + r + s + 2t + 2u + v
(17)
s CH4 ;
η = ─────────────── q + r + s + 2t + 2u + v
(18)
t C2H4 ;
ζ = ─────────────── q + r + s + 2t + 2u + v
(19)
u C2H6 ;
θ = ─────────────── q + r + s + 2t + 2u + v
(20)
v Tar
;
λ = ─────────────── q + r + s + 2t + 2u + v
(21)
Eq. (3) is rearranged to express the amount of H2O, β as follows; β = γ + 2η + 2ζ + 3θ + 0.5m’λ– 0.5m When γ, η, ζ, θ, and λ in above equation are substituted by using eqs. (15), (18), (19), (20), and (21) respectively, eq. (22) will be obtained. p+ 2s + 2t + 3u + 0.5m’ β = ──────────────── – 0.5m q + r + s + 2t + 2u + v
(22)
Eq. (4) is also rearranged to following equation to lead the expression for α. δ + 2ε – β + n’λ – n α = ───────────── 2 When δ, ε, β, and λ are substituted by using eqs. (16), (17), (21), and (22) respectively, equation (23) will be obtained.
(-p+ q + 2r - 2s - 2t -3u ) - 0.5m’v + n’v α = ──────────────────────── + 0.25m - 0.5n 2 (q + r + s + 2t + 2u + v)
(23)
Equations from eqs. (13) to (23) are distinctively complicated compared with those presented previously1 in which the numbers of species of gas product was limited only four. 3.
Feature of this method Since every equation employed to evaluate α ~ λ in eq. (1) was derived from the elemental balance of eq. (1) without any arbitrary assumption and approximation, it
may be applicable to any practical process regardless of a type of reactor or a rank of coal used. This method is available for the investigation of a chemical reaction process occurred in actual gasifier. Eq. (1) seems to be obtained from the flow rates of coal, gasifying agent, and each product. As can be seen from the fact that carbon conversion for practical gasification process has been usually found to be lower than 100%, accuracy of industrial instruments used in the plant to control flow rate of raw materials (coal, O2, and H2O) and to measure that of product (gas, tar, drain, ash, and residual char) is insufficient to obtain the formula that satisfies the law of conservation of mass. In another word, our method that estimate the formula without flow rate data mentioned above would be an quite unique way to estimate a perfect formula from the gas composition. Although the formula obtained by this method is written in the most simplified form, we can say that it is more desirable form to estimate the reaction process compared with simple composition of gas. Since the formula obtained here expresses exactly the material balance of gasification, product in molar ratio of H2, CO, and CH4 can be readily estimated from γ, δ, and η to judge whether molar ratio of O2 or H2O supplied to coal is appropriate from α or β. We can also estimate carbon conversion, thermal efficiency, heat of reaction, and temperature of gasification during the operation based on the material balance. In this way, this method will be applicable to actual gasification process to maintain operation conditions at the optimum state. 4.
Analysis of reaction formula obtained Reaction formula shown in eq. (1) is composed of various kinds of chemical processes including pyrolysis of coal, decomposition of tar, oxidation of char, combustion of gas, secondary reactions and others. Approach to analyze the chemical process of such a reaction formula has scarcely been known. We tried to prepare procedure to elucidate the contribution of particular reaction process concealed in the reaction formula. Firstly, following formula was defined as a standard formula. CHmOn + 0.5( 1 - n )O2 ->
0.5mH2 + CO
(24)
Since eq. (24) is determined singly by the value of m and n, it is considered to be a standard to make any approach to clarify the chemical process. Our approach began from making up a gap between eqs. (1) and (24) by mathematical way. We assumed that gasification is performed in two steps, namely partial-oxidation and secondary
reaction. In partial-oxidation step, carbon in coal is reacted with O2 and H2O and coal is converted into CO, CO2, and H2O. In secondary reactions, CO produced by partial oxidation may cause shift reaction and FT (Fisher-Tropsh) synthesis. Thus, eq. (24) mathematically changed according to the reaction formula of both steps mentioned above and finally eq. (1) was derived. Further details of derivation are described in following part. 4. 1 Partial oxidation step The amount of O2 in eq. (24), 0.5(1- n), is regarded as a stoichiometric amount of O2 for gasification. α in eq. (1) is written as followes, where Oex means a deviation of O2 from standard formula; α＝0.5(１- n) + Oex
(25)
Eq. (25) is very useful expression to estimate reaction mechanism of the partial oxidation step. Mechanism of partial oxidation step is estimated in two ways according to the sign of Oex. In the case of Oex > 0, coal will be first gasified with 0.5(1- n) mol of O2 along with eq. (24), and then 2Oex mole of CO and H2 in total will be burned with Oex mole of O2 to produce 2Oex mole of CO2 and H2O in total. In the case of Oex < 0, it can be assumed that -2Oex mole of carbon remain unconverted because the amount of O2 is not enough to convert all of C in coal into CO. The residual C will react with -2Oex mole of H2O supplied to produce -2Oex(H2+CO). Therefore, the final product will be CO+(0.5m-2Oex)H2. In this manner, the partial oxidation step may be expressed mathematically by using Oex. 4.
2 Secondary reaction step CO produced in a partial oxidation step causes shift reaction and hydrocarbon formation. In the case of shift reaction, CO + H2O -> H2 + CO2, we can easily express the variation of yields of CO, H2O, H2, and CO2 numerically based on the chemical reaction formula. Hydrocarbons such as CH4, C2H4, C2H6, and tar are produced by pyrolysis of coal and hydrogenation of carbon in coal as well as FT synthesis reaction. There is no question that FT synthesis is classified as secondary reaction step. The idea that pyrolysis and hydrogenation are also classified in the secondary reaction step, however, would be doubtful. It is necessary, we consider, to investigate the difference in the material balance of following hydrocarbon formation.
CH4
Pyrolysis ; C + 4H + H2O -> CH4 + H2O Hydrogenation ; C + 2H2 + H2O -> CH4 + H2O FT synthesis ; CO + 3H2 -> CH4 + H2O
(26) (27) (28)
C2H4
Pyrolysis ; 2C + 4H+ H2O -> C2H4+ H2O Hydrogenation ; 2C + 2H2+ H2O -> C2H4+ H2O FT synthesis ; 2CO+4H2+H2O -> C2H4+ H2O
(29) (30) (31)
C2H6
Pyrolysis ; 2C + 6H + 2H2O -> C2H6 + 2H2O Hydrogenation ; 2C + 3H2+ 2H2O -> C2H6+ 2H2O FT synthesis ; 2CO+10H2 -> C2H6 + 2H2O
(32) (33) (34)
Pyrolysis ;
(35)
CHm’On’
C+m’H+n’O+(1-n’)H2O ->
CHm’On’ +(1-n’)H2O
Hydrogenation ; C+0.5m’H2 +n’O+(1-n’)H2O -> CHm’On’+(1-n’)H2O (36) FT synthesis ; CO+{(1-n’)+0.5m’}H2 -> CHm’On’ + (1-n’)H2O (37)
Comparison of the material balance of pyrolysis and hydrogenation was made for the first step. We assumed that the pyrolysis and hydrogenation proceeds under an atmosphere containing moisture, or H2O. Pyrolysis is considered as a reaction that hydrocarbon is formed from C, H, and O atoms in coal and represented by eqs. (26), (29), (32), and (35). When H2 molecule is formed from H atom in the left side of eqs. (26), (29), and (32), the formula changes to express hydrogenation reactions between C atom in coal and H2 gas as shown in eqs. (27), (30), (33), and (36) respectively. That is to say, the difference between both formula that produce hydrocarbon produced by pyrolysis and by hydrogenation are no longer distinguishable from the viewpoint of material balance. In the next step, we investigate the difference between hydrogenation and FT synthesis. Possible reactions in FT synthesis is expressed in eqs. (27), (29), (31), (34), and (37). When water gas reaction is occurred in the left side of hydrogenation formula, it turns into FT synthesis in each component. Therefore, we cannot distinguish whether the path of hydrocarbon production is hydrogenation or FT synthesis so far as investigating the material balance of gasification. In another word, it means that the effect of formation of hydrocarbon on the yields of other inorganic components can be evaluated mathematically according to eqs. (28), (31), (34), and (37).
4. 3 Construction of mathematical reaction formula 1) In the case of Oex > 0 In the case of Oex>0, it can be assumed that coal is gasified with 0.5(1 - n)O2 and converted into product gas, i.e., 0.5mH2 + CO in the initial stage of gasification. Then 2Oex mole of product gas is burned with Oex mole of O2 to generate 2Oex mole of H2O and CO2. When mole of H2 burned is taken as x, that of CO is presented as (2Oex-x) and reaction formula after partial oxidation step was written as follows; CHmOn + { 0.5( 1– n) + Oex }O2 -> (0.5m–x)H2 +(1–2Oex + x)CO +(2Oex–x)CO2 - xH2O At the secondary reaction step, when y mole of CO caused shift reaction with H2O, y mole of H2 and CO2 are formed. The reaction formula after shift reaction will be as follows; ChmOn+{0.5( 1 - n) + Oex }O2+ｙH2O -> (0.5m+ｙ - x)H2 + (1 - 2Oex -ｙ+ x )CO+(2Oex +ｙ- x)CO2 + xH2O When z mole of CH4 is formed, molar amounts of H2 and CO will decrease by 3z and z respectively and those of H2O and CH4 will increase by z according to eq. (28). Reaction formula after CH4 formation is written as follows; CHmOn + {0.5( 1 -n) + Oex }O2 +(ｙ- x- z )H2O -> (0.5m+ｙ- x - 3z )H2 +(1-2Oex-ｙ+ x -z )CO +(2Oex +ｙ- x)CO2 + zCH4 When w mole of C2H4 is formed, molar amounts of H2 and CO will decrease by 4w and 2w respectively and those of H2O and C2H4 will increase by 2w and w respectively according to eq. (31). Reaction formula after C2H4 formation is written as follows; CHmOn + {0.5(1- n) + Oex }O2 +(ｙ- x -z -2w)H2O -> (0.5m+ｙ- x - 3z - 4w)H2 +(1-2Oex -ｙ+ x – z -2w )CO +(2Oex +ｙ- x )CO2 + zCH4 + wC2H4 When u mole of C2H6 is formed, molar amounts of H2 and CO will decrease by 5u and 2u respectively and those of H2O and C2H6 will increase by 2u and u respectively according to eq. (34). Reaction formula after C2H6 formation is written as follows;
CHmOn + { 0.5( 1 -n) + Oex }O2 +(ｙ- x - z -2w-2u)H2O -> (0.5m +ｙ -x - 3z - 4w-5u)H2 +(1-2Oex-ｙ+ x -z –2w -2u )CO +(2Oex +ｙ- x)CO2 + zCH4 + wC2H4 + uC2H6 When v mole of CHm’On’ is formed, molar amounts of H2 and CO will decrease by { (1-n’) + 0.5m’} v and v respectively and those of H2O and C2H6 will increase by (1– n’)v and v respectively according to eq. (37). Reaction formula after CHm’On’ formation is written as follows; CHmOn + {0.5( 1 -n) + Oex }O2 +{ｙ- x -z -2w -2u - (1- n’)v }H2O -> [0.5m +ｙ- x - 3z - 4w -5u - {(1-n’) + 0.5m’}v]H2 +(1-2Oex -ｙ+ x -z -2w -2u - v)CO +(2Oex +ｙ- x)CO2 + zCH4 + wC2H4 + uC2H6 + vCHm’On’
(38)
2)
In the case of Oex < 0 In the case of Oex < 0, the amount of O2 is insufficient to complete the standard reaction in formula (24). Coal is reacted with αO2 which is less than standard amount of O2 at the initial stage of gasification by following reaction formula, where -2OexC is residual carbon; CHmOn + { 0.5( 1- n ) + Oex }O2 -> 0.5mH2+(1+ 2Oex)CO -2OexC. Since there is no carbon in eq. (1), residual carbon in above formula should be gasified with H2O as follows; -2OexC - 2OexH2O -> -2OexCO - 2OexH2 The final formula after partial oxidation step is expressed as follows; CHmOn +αO2 - 2OexH2O -> (0.5m -2Oex)H2+CO At the secondary reaction step, y mole of CO causes shift reaction with H2O. Reaction formula will be as follows; CHmOn+{ 0.5(1- n ) +Oex }O2+(-2Oex +ｙ)H2O -> (0.5m -2Oex +ｙ)H2+(1 -ｙ)CO +ｙCO2 When z mole of CH4 is formed according to eq. (28), reaction formula is as follows;
CHmOn + {0.5 ( 1- n ) + Oex }O2 + ( -2Oex +ｙ-z )H2O -> (0.5m-2Oex +ｙ- 3z )H2 +(1 -ｙ-z )CO +ｙCO2 + zCH4 When w mole of C2H4 is formed according to eq. (31), reaction formula will be as follows; CHmOn + { 0.5( 1- n ) + Oex } O2 + ( -2Oex +ｙ- z -2w)H2O -> (0.5m -2Oex +ｙ- 4w )H2 +(1 -ｙ-z -2w )CO +ｙCO2 + zCH4 + wC2H4 When u mole of C2H6 is formed according to eq. (34), reaction formula will be as follows; CHmOn + {0.5( 1- n) + Oex }O2 + ( -2Oex +ｙ–z - 2w-2u)H2O -> (0.5m -2Oex +ｙ- 4w - 5u)H2 +(1-ｙ-z -2w -2u )CO +yCO2 + zCH4 + wC2H4+ uC2H6 When v mole of CHm’On’ is formed according to eq. (37), following reaction formula will be is obtained as in eq. (39). CHmOn + { 0.5( 1- n ) + Oex }O2 + { -2Oex +ｙ-z - 2w-2u - (1- n’)v }H2O -> [0.5m-2Oex +ｙ- 4w -5u -{(1-n’) + 0.5m’}v]H2 +(1 -ｙ-z -2w -2u - v)CO +ｙCO 2 + zCH4 + wC2H4+ uC2H6+ vCHm’On’ (39) 3)
Evaluation of practical value of each variable The coefficient of each component of eq. (1) and its numerical expression in eqs. (38) and (39) is shown in Table 1. Table 1 Comparison of each coefficient in eq. (1) with eqs. (38) and (39) (1) Α β γ δ ε η ζ θ λ
(39) : Oex<0
(38) : Oex>0 0.5(1- n ) + Oex ｙ- x - z -2w -2u - (1- n’)v
0.5( 1 - n ) + Oex -2Oex +ｙ- z - 2w - 2u - (1- n’)v
0.5m + y - x - 3z - 4w-5u-{(1-n’) + 0.5m’}v 1 – 2Oex - y + x – z - 2w -2u - v 2Oex + y - x
0.5m -2Oex + y- 4w - 5u - {(1-n’) + 0.5m’}v 1 -ｙ-z -2w -2u - v
z w u v
y z w u v
Quantities of inorganic components, β, γ, δ, and ε are expressed as functions of valuables such as Oex, x, y, z, w, u, and v. When compare the value of each coefficient in eq. (1) with the expression in eqs. (38) or (39), we can evaluate the values of above valuables, Oex, x, y, z, w, u, and v, that affect the composition of gases produced in actual gasification reactions. In the case of Oex>0, the value of Oex is calculated by eq. (25) and that of z, w, u, and v is found to be equal to η, ζ, θ, and λ respectively. Following equation is obtained corelating to the value of x and y. ε = 2Oex + y – x
(40)
Since the value of x, i.e. quantity of H2 burned with excess O2, cannot be determined only from the results of ultimate analysis and gas analysis, we assume the reasonable value of x based on the rate of combustions of H2 and CO. When gasification temperature is considerably low to maintain combustion rate of CO in adequate level, only H2 was virtually burned because of its enormous combustion rate. In this case, x becomes 2Oex and eq. (41) is valid. x = 2Oex
(41)
Eq. (41) is modified with eq. (40) to obtain eq. (42). (42)
ｙ＝ε
When coal is gasified at relatively high temperature, the product gas in eq. (24) is burned with OexO2 because CO and H2 are considered to be burned at comparable rate actually. The ratio of H2 to whole gas in the right side of eq. (24) is 0.5m/(1+0.5m) and x is calculated by 2Oex{0.5m/(1+0.5m)}. Thus, eq. (43) is obtained. x = mOex／(1+0.5m)
(43)
The equation to calculate the value of y is obtained as in eq. (44) by modifying eq. (42) and eq. (40). ｙ = ε-2Oex +{ mOex/(1+ 0.5m) }
(44)
In the case of Oex<0, relation, y = ε, can be found from Table 1. Since the quantity of O2 is insufficient to satisfy the standard reaction formula, CO2 cannot be produced by the combustion of CO. Consequently, CO2 is considered to be produced only by shift reaction and ｙ= ε is valid. Significant contribution of shift reaction in gasification has been often predicted. Role of shift reaction in the formation of H2 or CO2, however, has never been evaluated. Above approach, although requires reasonable assumption, can account for the contribution of shift reaction to composition of gas generated in actual gasifier. The method presented in this study will offer a novel point of view to understand practical reaction state in actual gasifier and will help plant operation at optimum conditions. 5.
Conclusion It is generally accepted that gasification consists of more than five chemical processes including pyrolysis, partial oxidation of char, decomposition of tar, secondary reactions of gases, and combustion of char and gas. Since the each process may be further divided into various elementary reactions, it is rather difficult to construct a detailed reaction model based on many elementary reactions. In addition, it is actually impossible to collect all necessary kinetic data of every elementary reaction by fundamental experiments. Simulation that can be applied for scientific analysis of coal gasification seems to be very difficult due to a very complicated phenomenon both from experimental and theoretical point of view. Chemical process cannot be fully understood by an accumulation of kinetic knowledge. The method presented in this study may be the first one that can discuss the reaction process scientifically based on gas composition obtained at a practical gasification plant. This method is simply based on stoichiometry of reaction formula without any arbitrary assumption and approximation. It would be applicable to any gasification process regardless of a type of gasifier or a rank of coal used. This study can be summarized to be one of the first attempt that a material balance of coal gasification, which has only used to evaluate carbon conversion and cold gas efficiency, is applied to elucidate the gasification phenomena. References [1] Kaiho M, Yasuda H, Yamada O, Proceeding, 2009 International Conference on Coal Science and Technology, Cape Town, in CD-ROM, 14-4, 2009. [2] Kaiho M, Yamada O, Shimada S, Proceeding, 2008 International Pittsburgh Coal Conference, in CD-ROM, 32-5, 2008.
Effect of calcium on the interaction of CO2 with carbonaceous materials during coal gasification Juan D. González, Fanor Mondragón, and Juan F. Espinal* Institute of Chemistry, University of Antioquia, Medellín, Colombia A.A. 1226 *
Corresponding author
E-mail address: [email protected] Abstract A computational chemistry study was carried out to evaluate the effect of CaO adsorption on the reaction of CO2 with a model structure of carbonaceous materials by using density functional theory DFT at B3LYP/6-311G(d) level of theory. It was found that the presence of CaO increase the electron density of free active sites making them thermodynamically more favorable for oxidation reactions. On the other hand, CO desorption is an endothermic process and it becomes less favorable when the chemisorbed oxygen atom has the ability to interact with a neighbor Ca atom. 1. Introduction Coal is expected to play an increasingly important role in this century both as energy source and as a raw material for obtaining chemical products [1, 2]. This tendency will continue at least until renewable energy sources are developed at large scale. In this context, gasification is considered to be an efficient and environmentally promising option for the use of coal. The use of catalysts during coal gasification has a long history in scientific studies [3-5]. Using active catalysts, carbon conversion at a given temperature can be significantly increased. Alternatively, the process can be operated at substantially lower temperatures than the ones that are required for non-catalytic gasification [6]. Several catalysts are known to be effective at increasing the gasification rate of coal char with CO2. Among others potential catalysts, calcium is attractive due to its superior catalytic properties like low agglomeration and volatilization during gasification and high activity at low loadings [7]. Even though, there has been intensive experimental research devoted to the catalytic effect of calcium, the mechanisms of calcium-catalyzed gasification are still not clear [8-10]. Due to the complexity of the
system, molecular modeling has been used as a powerful tool for improving our understanding of the catalyzed gasification process at molecular level. Nevertheless, most of the theoretical studies that can be found in the scientific literature are limited to the use of semi-empirical methods which are not quite accurate [11, 12]. In this work we evaluated the effect of calcium oxide on CO2 coal gasification by using density functional theory (DFT) and a possible mechanism including thermodynamic information was proposed. 2. Computational details Chars obtained at high temperatures from coal are carbonaceous materials made up of macrostructures formed mainly by aromatic clusters of different sizes. For char, it has been reported from solid-state 13C-NMR experiments that it has structures of randomly connected graphene clusters consisting of 3-7 benzene rings [13]. Since the gasification reactions take place at high temperatures, the active sites are simulated as unsaturated edge carbon atoms of a graphene layer. We selected as char model a graphene layer formed by 5 fused aromatic rings, which has proved to give good results in terms of bond lengths and bond angles compared with experimental data [14]. The computational method B3LYP/6-311G(d) represents a suitable method in terms of energy and computational cost. First we evaluated several interactions of the CaO molecule with the active sites and basal plane of carbonaceous models, then the most stables structures (CaO-Carbonaceous models) were chosen to perform Natural Bond Orbital (NBO) analysis, in order to visualize the most probable active sites for oxidation reactions such as CO2 dissociative chemisorption. We calculated the change in enthalpy and Gibbs free energy
of all reactions by using equations 1
and
2.
Where
is the sum of total electronic energy at 0 K and thermal correction
energy to the enthalpy and
is the sum of total electronic energy at 0 K and
thermal correction energy to the Gibbs free energies, respectively. In addition, changes
in enthalpy and Gibbs free energy for CO desorption were calculated and compared between the catalyzed reaction and the non-catalyzed reaction. All molecular systems, reactants and products, were optimized in its ground state and it was confirmed that the geometries correspond to stable structures that do not have imaginary vibration frequencies. Gaussian 09 package which include NBO version 3.1 was used for all calculations [15, 16]. 3. Results and Discussion 3.1.
Interaction of calcium oxide with carbonaceous models
The most stable systems (CaO-Carbonaceous models) after calcium oxide interaction with carbonaceous models are shown in the Figure 1. Note that atoms have been labeled in order to facilitate the NBO analysis and that the
-bond network is implied in these
figures.
+ ΔH298 = -158.4 kcal/mol ΔG298 = -146.9 kcal/mol
II-CaO-a
+ ΔH298 = -147.6 kcal/mol ΔG298 = -136.5 kcal/mol
II-CaO-b
Figure 1. Most stable structures after CaO interaction with carbonaceous models. Green atom represents calcium, red oxygen, grey carbon and white hydrogen. In spite of the different possible interactions of CaO with active sites and basal plane of the carbonaceous models, which are exothermic processes, the most stable structures are
those where CaO interacts with two actives sites in a horizontal approach. It is probably due to the stabilization of two reactive sites and the fact that in this structures oxygen and calcium atoms are present in a configuration that is close to their most stables valences. Structure II-CaO-a is more stable than II-CaO-b by about 11 kcal/mol. It is worth to note that both structures have an available active site where one oxygen atom could be chemisorbed after CO2 dissociation. 3.2. NBO analysis In order to explore the effect of CaO adsorption on the carbonaceous models, Natural Bond Orbital (NBO) analysis over CaO-carbonaceous model structures was carried out. Chemical bond data and atomic charge changes of some selected atoms are listed in Tables 1 and 2. Table 1. Bond length and bond order (Wiberg index) for structures II-CaO-a and II-CaO-b model
bond
length (Å)
bond order
II-CaO-a
C(1)-O(6)
1.29
1.33
O(6)-Ca(7)
2.36
0.15
Ca(7)-C(3)
2.22
0.42
C(1)-Ca(7)
2.33
0.44
Ca(7)-O(6)
2.07
0.30
O(6)-C(3)
1.33
1.14
II-CaO-b
Table 2. Net NBO atomic charges before and after CaO adsorption model
atom
before CaO adsorption
after CaO adsorption
Δ charge
II-CaO-a
C(5)
0.21
-0.27
-0.48
II-CaO-b
C(5)
0.21
0.19
-0.02
It can be seen from Table 1 that the C-O interactions (see length and bond order) are stronger than Ca-O and Ca-C. This result could be related to the stabilization of the C-O interaction by resonance with the π-bond network of the aromatic system. Table 2
shows that largest change in the atomic charge of the active site C(5) after CaO adsorption occurs in structure II-CaO-a. This result could be explained considering the fact that the Ca atom can donate electron density whereas the O atom attracts electron density due to its more electronegative character. This is consistent with the results reported by Chen and Yang [11] who found that the presence of electropositive atoms can increase the electron density of carbon atoms making them more negatively charged. 3.3.
Reaction with CO2 and CO desorption
The reaction of CO2 with the active site involves a dissociative chemisorption and subsequent CO desorption for both catalyzed and non-catalyzed reactions, which are shown in Figure 2.
+CO2 -CO ΔH= -40.7
-CO ΔH= 44.9
II-CaO-a
Δ
+CO2 -CO
-CO
ΔH= -26.3
ΔH= 19.6
+CO2 -CO
-CO
II-CaO-b
ΔH= -22.1 ΔH= 18.7
II
Figure 2. CO2 dissociative chemisorption and CO desorption for both catalyzed and non-catalyzed reactions. Enthalpies are in kcal/mol
Several inferences can be made from the results shown in Figure 3. Dissociative chemisorption of carbon dioxide is an exothermic process for both catalyzed and noncatalyzed reactions being more exothermic (almost twice) on II-CaO-a compared with II-CaO-b and II. It is in good agreement with the results of the NBO population analysis (see Table 2), where it was found that free active site (C(5)) experiments an increment in its electron density, making this edge carbon atom suitable for bonding with an oxygen atom. On the other hand, CO desorption is an endothermic process with positive reaction enthalpies. Note that CO desorption after dissociative chemisorption of CO2 on II-CaO-a present the most positive ΔH value. This result might be related to the fact that for this structure the chemisorbed oxygen interacts strongly with the neighbor Ca atom, making the CO desorption process less favaroble, from a thermodynamic point of view. Finally, the ΔH values for CO2 dissociative chemisorption and CO desorption are in good agreement with those reported in the scientific literature for the non-catalyzed process [17]. 4. Conclusions We have found that CaO interaction with carbonaceous models change the electron density of free active sites making the oxygen chemisorption more favorable on these negatively charged carbon atoms from a thermodynamic point of view. It has been shown that CO desorption is an endothermic process being less favorable when the oxygen atom can interact with neighbor atoms like Ca. Acknowledgements The authors would like to thank COLCIENCIAS and the University of Antioquia for financial support of the project 1115-452-21242. The authors also acknowledge partial support from the “Sostenibilidad” Program (2009-2011) of the University of Antioquia
References [1] I. Mochida, K. Sakanishi, Catalysts for coal conversions of the next generation, Fuel, 79 (2000) 221-228. [2] E. Shafirovich, A. Varma, Underground Coal Gasification: A Brief Review of Current Status, Industrial & Engineering Chemistry Research, 48 (2009) 7865-7875. [3] D.W. McKee, Mechanisms of the alkali metal catalysed gasification of carbon, Fuel, 62 (1983) 170-175. [4] F. Kapteijn, G. Abbel, J.A. Moulijn, CO2 gasification of carbon catalysed by alkali metals: Reactivity and mechanism, Fuel, 63 (1984) 1036-1042. [5] W.-Y. Wen, Mechanisms of Alkali Metal Catalysis in the Gasification of Coal, Char, or Graphite, Catalysis Reviews: Science and Engineering, 22 (1980) 1-28. [6] Y. Huang, X. Yin, C. Wu, C. Wang, J. Xie, Z. Zhou, L. Ma, H. Li, Effects of metal catalysts on CO2 gasification reactivity of biomass char, Biotechnology Advances, 27 568-572. [7] Y. Zhang, M. Ashizawa, S. Kajitani, S. Hara, A new approach to catalytic coal gasification: The recovery and reuse of calcium using biomass derived crude vinegars, Fuel, 89 (2010) 417-422. [8] A. Linares-Solano, E.J. Hippo, P.L. Walker Jr, Catalytic activity of calcium for lignite char gasification in various atmospheres, Fuel, 65 (1986) 776-779. [9] Y. Ohtsuka, K. Asami, Ion-Exchanged Calcium from Calcium Carbonate and LowRank Coals: High Catalytic Activity in Steam Gasification, Energy & Fuels, 10 (1996) 431-435. [10] Y. Zhang, S. Hara, S. Kajitani, M. Ashizawa, Modeling of catalytic gasification kinetics of coal char and carbon, Fuel, 89 (2010) 152-157. [11] S.G. Chen, R.T. Yang, The Active Surface Species in Alkali-Catalyzed Carbon Gasification: Phenolate (C---O---M) Groups vs Clusters (Particles), Journal of Catalysis, 141 (1993) 102-113. [12] S.G. Chen, R.T. Yang, Unified Mechanism of Alkali and Alkaline Earth Catalyzed Gasification Reactions of Carbon by CO2 and H2O, Energy & Fuels, 11 (1997) 421-427. [13] S.T. Perry, E.M. Hambly, T.H. Fletcher, M.S. Solum, R.J. Pugmire, Solid-state 13C NMR characterization of matched tars and chars from rapid coal devolatilization, Proceedings of the Combustion Institute, 28 (2000) 2313-2319.
[14] N. Chen, R.T. Yang, Ab initio molecular orbital calculation on graphite: Selection of molecular system and model chemistry, Carbon, 36 (1998) 1061-1070. [15] M. J. Frisch, H. B. Schlegel, G. E. Scuseria, M. A. Robb, J. R. Cheeseman, G. Scalmani, V. Barone, B. Mennucci, G. A. Petersson, H. Nakatsuji, M. Caricato, X. Li, H. P. Hratchian, A. F. Izmaylov, J. Bloino, G. Zheng, J. L. Sonnenberg, M. Hada, M. Ehara, K. Toyota, R. Fukuda, J. Hasegawa, M. Ishida, T. Nakajima, Y. Honda, O. Kitao, H. Nakai, T. Vreven, J. A. Montgomery, Jr., J. E. Peralta, F. Ogliaro, M. Bearpark, J. J. Heyd, E. Brothers, K. N. Kudin, V. N. Staroverov, R. Kobayashi, J. Normand, K. Raghavachari, A. Rendell, J. C. Burant, S. S. Iyengar, J. Tomasi, M. Cossi, N. Rega, J. M. Millam, M. Klene, J. E. Knox, J. B. Cross, V. Bakken, C. Adamo, J. Jaramillo, R. Gomperts, R. E. Stratmann, O. Yazyev, A. J. Austin, R. Cammi, C. Pomelli, J. W. Ochterski, R. L. Martin, K. Morokuma, V. G. Zakrzewski, G. A. Voth, P. Salvador, J. J. Dannenberg, S. Dapprich, A. D. Daniels, Ö. Farkas, J. B. Foresman, J. V. Ortiz, J. Cioslowski, and D. J. Fox, Gaussian 09, Revision A.1: Gaussian Inc., Wallingford CT 2009. [16] A.E.R. E. D. Glendening, J. E. Carpenter, and F. Weinhold, NBO Version 3.1. [17] L.R. Radovic, The mechanism of CO2 chemisorption on zigzag carbon active sites: A computational chemistry study, Carbon, 43 (2005) 907-915.
Interaction of calcium with carbonaceous materials: A DFT study
Juan D. González, Fanor Mondragón, and Juan F. Espinal* Institute of Chemistry, University of Antioquia, Medellín, Colombia A.A. 1226 *
Corresponding author
E-mail address: [email protected]
Abstract A systematic theoretical study using density functional theory (DFT) was carried out to provide molecular-level understanding on the interaction of calcium oxide with different carbonaceous models. It was found that B3LYP/6-311G(d) represents a good balance between accuracy and computational cost. Several configurations for the CaO interaction with carbonaceous models were evaluated, finding that its adsorption is an exothermic process and the most favorable interaction occurs when calcium oxide interacts with the active sites of a zigzag edge.
1. Introduction It is expected that coal will be the major energy source among fossil resources in the coming decades because of its abundant reserves, which are expected to last for about 200 years at current consumption rates[1]. Currently, the most common use of coal is through its combustion. However, this process has low efficiencies and releases large amounts of pollutants. Therefore, better alternatives for using coal must be implemented. Coal gasification seems to be the best technology for the production of chemicals and energy from coal. This is a process that takes place at high temperatures. However, from an economic point of view, low temperatures are desirable. The use of catalysts has been proposed to achieve low temperatures in the gasification and speed up the reaction of coal with gasifying agents such as H2O and/or CO2. Several catalysts are known to be effective for increasing the gasification rate of coal in steam and CO2 [2-4]. Among other potential catalysts, calcium is one of the most interesting metals due to its low agglomeration and volatilization during gasification and high activity at low
loadings [5]. There is abundant information in the scientific literature about the catalytic effect of calcium in the gasification of coal, mainly through experimental techniques [4, 6, 7]. However, the way in which calcium interacts with carbonaceous materials and the mechanisms of calcium catalyzed gasification is still poorly understood. This is mainly due to the complexity of coal structure and the limitations of available experimental techniques. Molecular modeling is a powerful tool and it is a complement to experimental methods for improving our understanding of the catalyzed gasification process. Despite of its utility, there have been a limited number of studies on the catalytic mechanisms of coal gasification by metals and most of them employ less accurate semi-empirical methods for their calculations [8, 9]. In the present work we used density functional theory (DFT), which has proved to yield a reasonable balance between reliable results and computation cost modeling non-catalyzed coal gasification [10-13]. In this study, we first selected the model chemistry (Theory level/Basis set) and then we used it to evaluate the interaction of calcium oxide with the active sites of clean carbonaceous models and its effect over carbon atoms. This interaction is a first step to get a better insight on the effect of calcium in coal gasification with CO2.
2. Computational details Carbonaceous materials are macrostructures formed mainly by aromatic clusters. For char, it has been reported from solid-state
13
C-NMR experiments that it has
structures of randomly connected graphene clusters consisting of 3-7 benzene rings [14]. Since the gasification reactions take place at high temperatures, the active sites are simulated as edge carbon atoms of a graphene layer that have lost a hydrogen atom. Initially, the search for an appropriate computational method (theory level/basis set) was carried out in order to obtain reliable energy values at a reasonable computation time. Adsorption energies were calculated as: ΔEAds(0 K) = ∑ EProducts - ∑ EReactants for a small system represented by one molecule of calcium oxide and two fused aromatic rings with two active sites (See Figure 1). Note that the -bond network is implied in all figures. The calcium oxide molecule was placed close to the active sites of the two rings system, then this structure was optimized to a minimum on the potential energy surface by combining the B3LYP, B3PW91, MPW1PW91 and M06-2X functionals with the following basis sets: 3-21G, 6-31G(d), 6-31G(d,p), 6-31+G(d), 6-31+G(d,p),
6-31++G(d),6-31++G(d,p), 6-311G(d), 6-311G(d,p), 6-311+G(d), 6-311+G(d,p), 6-311++G(d), and 6-311++G(d,p). A single point energy calculation on a geometry optimized at B3LYP/6-311++G(d,p) was performed with the highly correlated method CCSD(T)/6-311++G(d,p). This adsorption energy was used as reference to select an appropriate computational method (functional/basis set). The selected computational method was used to evaluate several interactions of the CaO molecule with the active sites of carbonaceous models; we also evaluated the interaction on the basal plane. All molecular systems, reactants and products, were optimized in their ground state and it was confirmed that the geometries correspond to a minimum on the potential energy surface by means of frequency calculations. Gaussian 09 package[15] was used for all calculations.
3. Results and Discussion 3.1. Theory level and basis set selection The molecular model that was used to calculate the adsorption energy in the search of an appropriate computational method is shown in Fig. 1. Considering that the computational cost of calculations increases exponentially with the number of atoms, we used a small model to represent the structure of carbonaceous materials in order to find an appropriate computational method at a lower computational cost. It is worth to mention that it has been reported that this model (C13H7) is the smallest model that can be used to obtain parameters in excellent agreement with experimental data such as bond lengths and bond angles [11].
Calcium Oxygen Carbon Hydrogen
Figure 1. Molecular model used to simulate CaO adsorption and to calculate ΔEAds.
The adsorption energies obtained in the simulations and the computational cost by combining different functionals with different basis sets are shown in Figures 2(a) and 2(b), respectively. The reference value for the adsorption energy obtained at CCSD(T)/6-311++G(d,p)//B3LYP/6-311++G(d,p) level is also plotted in Figure 2(a).
a
b
Figure 2. Search of appropriate computational method: (a) ΔEAds, (b) average calculation time per optimization cycle at different functional/basis set combinations.
All of the functionals show a similar tendency for the adsorption energy. For the same theory level, adding p polarization functions on the basis set does not have a significant effect on ΔEads, but adding the first diffuse function makes the adsorption energy less exothermic by about 7 kcal/mol. In the case of computational cost, adding one diffuse function increases significantly the calculation time, while adding p polarization functions has a minimal effect. From Figure 2 (a) and (b) we conclude that the most suitable method in terms of energy and computational cost is B3LYP/6-311G(d). The difference between the value of ΔEads at this level and the reference CCSD(T)/6311++G(d,p)//B3LYP/6-311++G(d,p) is only 5.8 kcal/mol (3.6%), which is acceptable.
3.2. Interaction of CaO with carbonaceous models Having selected the computational method (B3LYP/6-311G(d)), we studied several possibilities for the interaction of CaO with different graphene layers used as models of char (see Figure 3). Structure I has only one active site, while structure II has three active sites and there are no active sites in structure III.
I
II
III
Figure 3. Single graphene layers used as models of char.
The most stable systems after calcium oxide interaction with carbonaceous models are shown in the Figure 4. It was found that this interaction is an exothermic process and that the most favorable interaction occurs when calcium oxide interacts with two active sites on the zigzag edge through a horizontal approach (II-CaO-h) instead of a vertical approach. This result might be explained by the fact that in this structure calcium and oxygen atoms are present in a configuration that is close to their most stables valencies. On the other hand, even though the interaction of CaO with the basal plane of carbonaceous models (structure III) is exothermic, it is the less thermodynamically favorable process. Probably, due to a repulsion between the oxygen atom and the system of the aromatic structure. It can be noted that CaO interacts with the basal plane by means of the calcium atom which carries a partial positive charge.
+ ΔEAds= -105,8 kcal/mol I-CaO
+ ΔEAds= -157,9 kcal/mol
II-CaO-h
+ ΔEAds= -74,2 kcal/mol
II-CaO-v
+ ΔEAds= -28,3 kcal/mol III-CaO
Figure 4. Interaction of CaO with different carbonaceous models.
4. Conclusions
We have performed a systematic search for an appropriate computational method to study the interaction of calcium oxide with different carbonaceous models. We found that B3LYP/6-311G(d) is an appropriate method in terms of accuracy and computational cost. It has been shown that the interaction of CaO with carbonaceous models is an exothermic process and that the most favorable interaction occurs when calcium oxide interact with two active sites of a zigzag edge in a horizontal approach.
Acknowledgements
The authors would like to thank COLCIENCIAS and the University of Antioquia for financial support of the project 1115-452-21242. The authors also acknowledge partial support from the “Sostenibilidad” Program (2009-2011) of the University of Antioquia
References [1] International Energy Agency, world energy outlook, 2008. [2] D.W. McKee, Mechanisms of the alkali metal catalysed gasification of carbon, Fuel, 62 (1983) 170-175. [3] H. Ohme, T. Suzuki, Mechanisms of CO2 Gasification of Carbon Catalyzed with Group VIII Metals. 1. Iron-Catalyzed CO2 Gasification, Energy & Fuels, 10 (1996) 980-987. [4] J. Wang, M. Jiang, Y. Yao, Y. Zhang, J. Cao, Steam gasification of coal char catalyzed by K2CO3 for enhanced production of hydrogen without formation of methane, Fuel, 88 (2009) 1572-1579. [5] Y. Zhang, M. Ashizawa, S. Kajitani, S. Hara, A new approach to catalytic coal gasification: The recovery and reuse of calcium using biomass derived crude vinegars, Fuel, 89 (2010) 417-422. [6] Y. Zhang, S. Hara, S. Kajitani, M. Ashizawa, Modeling of catalytic gasification kinetics of coal char and carbon, Fuel, 89 (2010) 152-157. [7] A. Linares-Solano, E.J. Hippo, P.L. Walker Jr, Catalytic activity of calcium for lignite char gasification in various atmospheres, Fuel, 65 (1986) 776-779. [8] S.G. Chen, R.T. Yang, The Active Surface Species in Alkali-Catalyzed Carbon Gasification: Phenolate (C---O---M) Groups vs Clusters (Particles), Journal of Catalysis, 141 (1993) 102-113. [9] S.G. Chen, R.T. Yang, Unified Mechanism of Alkali and Alkaline Earth Catalyzed Gasification Reactions of Carbon by CO2 and H2O, Energy & Fuels, 11 (1997) 421-427. [10] A. Montoya, F. Mondragón, T.N. Truong, First-Principles Kinetics of CO Desorption from Oxygen Species on Carbonaceous Surface, The Journal of Physical Chemistry A, 106 (2002) 4236-4239. [11] N. Chen, R.T. Yang, Ab initio molecular orbital calculation on graphite: Selection of molecular system and model chemistry, Carbon, 36 (1998) 1061-1070. [12] N. Chen, R.T. Yang, Ab Initio Molecular Orbital Study of the Unified Mechanism and Pathways for Gas−Carbon Reactions, The Journal of Physical Chemistry A, 102 (1998) 6348-6356. [13] Z.H. Zhu, J. Finnerty, G.Q. Lu, R.T. Yang, A Comparative Study of Carbon Gasification with O2 and CO2 by Density Functional Theory Calculations, Energy & Fuels, 16 (2002) 1359-1368.
[14] S.T. Perry, E.M. Hambly, T.H. Fletcher, M.S. Solum, R.J. Pugmire, Solid-state 13C NMR characterization of matched tars and chars from rapid coal devolatilization, Proceedings of the Combustion Institute, 28 (2000) 2313-2319. [15] M. J. Frisch, H. B. Schlegel, G. E. Scuseria, M. A. Robb, J. R. Cheeseman, G. Scalmani, V. Barone, B. Mennucci, G. A. Petersson, H. Nakatsuji, M. Caricato, X. Li, H. P. Hratchian, A. F. Izmaylov, J. Bloino, G. Zheng, J. L. Sonnenberg, M. Hada, M. Ehara, K. Toyota, R. Fukuda, J. Hasegawa, M. Ishida, T. Nakajima, Y. Honda, O. Kitao, H. Nakai, T. Vreven, J. A. Montgomery, Jr., J. E. Peralta, F. Ogliaro, M. Bearpark, J. J. Heyd, E. Brothers, K. N. Kudin, V. N. Staroverov, R. Kobayashi, J. Normand, K. Raghavachari, A. Rendell, J. C. Burant, S. S. Iyengar, J. Tomasi, M. Cossi, N. Rega, J. M. Millam, M. Klene, J. E. Knox, J. B. Cross, V. Bakken, C. Adamo, J. Jaramillo, R. Gomperts, R. E. Stratmann, O. Yazyev, A. J. Austin, R. Cammi, C. Pomelli, J. W. Ochterski, R. L. Martin, K. Morokuma, V. G. Zakrzewski, G. A. Voth, P. Salvador, J. J. Dannenberg, S. Dapprich, A. D. Daniels, Ö. Farkas, J. B. Foresman, J. V. Ortiz, J. Cioslowski, and D. J. Fox, Gaussian 09, Revision A.1: Gaussian Inc., Wallingford CT 2009.
Oviedo ICCS&T 2011. Extended Abstract
USE OF A WASTE GENERATED IN THE CEMENT INDUSTRY AS ADDITIVE IN PROCESS OF COAL AND BIOSOLID GASIFICATION IN FLUIDIZED BED Elizabeth Rodríguez Acevedo 1, Farid Chejne Janna 1, William Jurado 2 1
Universidad Nacional de Colombia. Cr. 80. # 65-223 Of. M3 – 207 Medellin, Colombia. E-mail: [email protected] 2 Investigación y Desarrollo de Tecnologías de Proceso, CEMENTOS ARGOS E-mail: [email protected]
Abstract
This work shows the effect of a waste rich in carbonates and calcium oxides produced in the cement industry used as additive in the process of coal gasification in fluidized bed for capture of sulfur. The coal and coal/additive which had a 10, 20 y 30% weight in excess of additive concentration was gasified at constant conditions of temperature and steam flow. Coal and additive were physico-chemically characterized, syngas composition was analyzed during gasification by on-line chromatography and the characterization of bottom ashes were characterized was focused on the percentage of sulfur retained. Also the syngas composition is presented for biosolid gasification to compare the percentage of CO2 retained.
1. Introduction
Coal is used currently for power generation due to the extensive knowledge, availability and low cost comparative to others fuels. The measured reserves in Colombia are 6.648 billion tons [1] Coal contains sulfur in its composition than during thermal processes become potential sulfur contaminants. The use of carbonates and calcium oxides during thermal processes is a method for the capture of polluting sulfur compounds [2, 3, 4] Cement industry generates waste which contains carbonates and calcium oxides; For one cement company produced 22480 Ton/year (to 2008) of this waste, it’s generates expenses of transportation and space for disposal of this material. Due to the composition of waste can be used as an additive in the gasification process for the removal of sulfur.
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Oviedo ICCS&T 2011. Extended Abstract
Another important residue is biosolid produced in wastewater treatment plant after stabilization process, daily 100 TON is produced in Medellín. This waste has a percentage of organic matter that can be energetically harnessed through gasification. The syngas produced has a high percentage of H2 and CO2 which can be enriched when CO2 is captured. The main pollutants of sulfur are sulfur oxide and hydrogen sulfides [5, 6]. The following chemical reactions play an important role in the sulfur dioxide capture in gasification in fluidized bed (between 800 y 900°C).
S + O2 SO2 + ½ O2
SO2 + 296 Kj/g mol SO3
CaO + ½ O2 + SO2 SO3 + H2O
(1) (2)
CaSO4 + 486Kj/g mol
H2SO4
(3) (4)
The limestone (CaCO3) and dolomite (CaCO3.MgCO3) are main sorbents used for sulfur oxide capture [6] (Equation 5).
CaCO3 + SO2 + ½ O2
CaSO4 + CO2
(5)
But this reaction not carried out in one step. The first step is the calcination through an endothermic reaction, as follows:
CaCO3
CaO + CO2 - 183 Kj/ g mol
(6)
After calcination, the carbon dioxide creates and expands many pores in the limestone. This represents a greater surface area for the subsequent sulfidation which is second step (Equations 3 and 7).
SO2 + ½ O2
SO3
CaO + SO3
CaSO4
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(7)
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Oviedo ICCS&T 2011. Extended Abstract
The calcination reaction in equation 6 is relatively fast, for example the calcination reaction of a particle of 0,5 mm would take about 50 seconds to be completed. The sulfidation reaction in equation 3 is relatively slow, 1200 seconds is required for a calcined particle of 0,5 mm [6].
In some cases the calcium sulfate can be reduced to free SO2 (Equation 8). This reaction is favored at high temperatures but can occur at a relatively low temperature around 850°C, if there is sufficient CO [6].
CaSO4 + CO
CaO + SO2 + CO2
CaSO4 + 4 CO CaO + SO2
(8)
CaS + 4 CO2
(9)
CaS + 3/2 O2
(10)
The removal of H2S can be performed at low or high temperature by solid sorbents such as zinc ferrites or titanates (500-700 °C) or calcium sorbents (800 – 1000 ºC) for producing CaS.
Calcination:
CaCO3
Calcined Sulfidation:
CaO + H2S
Direct sulfidation:
CaCO3 + H2S
CaO + CO2 + 178 Kj / mol
(11)
CaS + H2O – 59 Kj / mol
(12)
CaS + H2O + CO2 + 119 Kj / mol
(13)
The origin of sorbent and porous system characteristics of calcined matter (porosity and pore size distribution) strongly affects reaction rates and conversions of sulfidation.
2. EXPERIMENTAL SECTION
The initial characterization was performed for coal, additive and biosolid. The materials were gasified in fluidized bed to laboratory scale. Coal and mixtures of coal/additive were gasified in fluidized bed at constant conditions of temperature and steam flow. Also the syngas composition is biosolid gasification is presented for compare the percentage of CO2 retained.
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Oviedo ICCS&T 2011. Extended Abstract
2.1 Experimental process description The Figure 1 shows the schematic of the gasification system, which is operated in batch.
Figure 1. Schematic of the systems of gasification The gasification system is comprised following elements:
-
Reactor, which has an internal diameter of 0,14 m and a height of 1,5 m. Electrical resistances were used for heating.
-
Boiler.
-
Steam superheater, which increases the temperature of steam at 400°C.
-
Gas cleaning system: Filters for particulate matter, condenser and silica gel for remove moisture and tars.
-
Gas chromatograph for analysis of the syngas composition on-line. H2, CO2, CO, N2, O2 and CH4 are the analyzed compounds.
El coal or mixture is introduced into the reactor. The system is heated and the superheated steam is introduced when the temperature is 300°C. The chromatograph
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Oviedo ICCS&T 2011. Extended Abstract
begins to take records of the syngas composition on-line after 800°C. The system is stabilized in 1000-1060°C and reacts for four hours. The Table 1 shows the constant conditions of operating for the four tests.
Table 1. Operating conditions Condition Temperature (°C) Steam (lpm) Coal (kg)
Value 1000-1060 1 2
The additional additive concentration in the mixtures are 10, 20 y 30% (percentage weight).
2.2 Material characterization The additive is a wasted generated in the clinker production process. The composition of additive is shown in the Table 2. Additive composition The highest percentages correspond to calcium oxide and silicates. Carbonates are analyzed independently which represent 73%. Table 2. Additive composition
COMPOUND Al2O3 SiO2 Fe2O3 CaO MgO SO3 Na2O K 2O TiO2 P2O5 LOI (Lost of Ignition)
% 2,35 6,63 1,72 57,56 0,10 2,06 0,02 0,29 0,13 0,07 29,47
The initial characterization of the additive, coal and biosolid are presented in the
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Oviedo ICCS&T 2011. Extended Abstract
Table 3, which presents the proximate analysis, elemental analysis and others important characteristics in the gasification process. The coal is from Amagá-Antioquia. Table 3. Characteristics of solids
Coal
SAMPLE Additive
Proximate analysis Moisture (%) Ash (%) Volatile matter (%) Fixed carbon (%)
9,87 9,56 38,63 41,94
0,15 68,16 16,2 15,49
5,2 65,2 25,2 4,4
Total sulfur (%)
0,58
0,32
0,64
Ultimate analysis C (%) H (%) N (%) O (%)
58,30 4,42 0,18 25,49
9,95 0,38 1,65 21,01
14,06 2,41 3,85 13,84
Others Heating Value (kJ/kg)) Mean particle size (mm) Apparent density (kg/m3)
22399 5 637
100,5 1-0,25 873
6163 5 619
Biosolid
3. Results and Discussion
The sulfur retention is presented for coal and mixtures gasification and syngas composition is presented for coal, mixtures and biosolid gasification.
3.1 Sulfur retention
The gasification time was not enough for the samples are completely consumed, so there results of retention of sulfur are partial. The percentage of unburned is 70% approximately. The
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Oviedo ICCS&T 2011. Extended Abstract
Table 4 presents the analysis of sulfur in the bottom ash (The unburned is included in ash) and additive after gasification.
Table 4. Percentage of sulfur retained
Gasification 0% additive Gasification 10% additive Gasification 30% additive
Sulfur percentage (%) Bottom ashes Additive after gasification 0,58 0,55 0,4 0,53 0,45
The coal contains 0,58% of sulfur initially. The sulfur percentage in ash for gasification for 0% of additive is constant (0,58%), but decreases to 0,55% for 10% of additive (decrease of 5,2%) and to 0,53% for 30% of additive (decrease of 8,6%).
The initial additive contains 0,32% of sulfur. The sulfur percentage in the additive at the end of the gasification for 10% of additive increases from 0,32% to 0,4% (increase of 25%) and 0,45% (increase of 40,6%) for 30% of additive.
3.2 Syngas composition
The composition of the syngas is shown in Figure 2 for the four test with coal. The H2, CO2, CO and CH4 are reported. Ethylene and Ethane are observed but are not reported. The time zero corresponds to 800°C in the system, which in one hour arrive at 1000°C where is stabilized. The syngas composition is very similar after one hour for the four test, the percentage of H2 is approximately 60%, CH4 between 6 and 8 % and 4% for CO.
The tests are carried out only with steam, the composition (Figure 2) shows that the shift reaction plays an important role, equation 14, which reduces the CO to produce H2 and CO2.
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Oviedo ICCS&T 2011. Extended Abstract
CO + H2O
CO2 + H2
(14)
The CO2 behavior is shown in Figure 3 for coal gasification.
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Oviedo ICCS&T 2011. Extended Abstract
0% of additive
10% of additive
20% of additive
30% of additive
Figure 2. Syngas composition. Coal Gasification
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Oviedo ICCS&T 2011. Extended Abstract
Figure 3. Comparison of CO2 in the syngas for coal gasification The CO2 obtained for coal gasification increases by shift reaction but after the composition is stabilized. Coal gasification for 10% of additive has an increase in first two hours, after is stabilized at 33 % of CO2, gasification for 20% of additive is stabilized at 34% of CO2 approximately after three hours and gasification for 30% of additive is less slope, after four hour the percentage of CO2 is 28 %.
The Figure 4 shows a comparison of CO2 for biosolid gasification for 0% and 30 % of additive. The gasification was carried out under different operating condition at with respect to coal gasification conditions, in this case was performed with samples of 200 g, steam 2 lpm and temperature 850°C. The concentration of CO2 is 70% (0% additive) and 10 – 28% (30% additive), a biggest difference with respect to coal gasification.
Figure 4. Comparison of CO2 in the syngas for biosolid gasification
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Oviedo ICCS&T 2011. Extended Abstract
3 Conclusions
Retention of sulfur was observed in the coal gasification; however the coal sample was not consumed completely which show a partial retention of sulfur. The capture CO2 is significant for coal gasification at 30% additive, but the value is low; while the biosolid gasification presented a high percentage of capture for different operational condition. Therefore, the later indicates that both results given by the samples of coal and biosolid cannot be directly compared.
The use of the cement residues as additive in the
gasification process allowed the reduction of pollutant releases a low-cost
Acknowledgement
Authors acknowledge the financial support of INCARBO, ARGOS, EPM, COLCIENCIAS, CIIEN and Universidad Nacional de Colombia.
References
[1] Unidad de Planeación Minero Energética. Colombia. www.upme.gov.co . 2011 [2] GOYAL A., BRYAN B.G., REHMAT A., PATEL J. G. In Situ Desulfurization in a Fluidized-Bed Coal Gasifier. Energy Sources. 1990: 19: 161 – 179
[3] AZNAR M. CABALLERO A. CORELLA J. TOLEDO J. M. Hydrogen Production by Biomass Gasification with Steam-O2 Mixtures Followed by a catalytic Steam Reformed and a CO-Shift System. Energy and Fuel. 2006:
[4] THE EUROPEAN COMMISSION. The Integrated Gasification Combined Cycle. Information WebPage. An Advanced Clean Coal Technology. 2000.
[5] ATTIA, Y. A. Processing and Utilization of High Sulfur Coals. COAL SCIENCE AND TECNOLOGY 9. Elsevier Scientific Publishing Company. 1982
[6] ABAD SECADES, A. Eliminación de H2S de Gases de Gasificación en Reactores de Lecho Móvil a Presión. Tesis Doctoral. Universidad de Zaragoza. España. 2003
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Oviedo ICCS&T 2011. Extended Abstract
CALCIUM LOOPING TECHNOLOGIES FOR CO2 CAPTURE Juan Carlos Abanades Spanish Research Council, CSIC-INCAR (Spain) [email protected] Abstract There is a range of CO2 capture technologies sufficiently developed today to be demonstrated at large scale in the next few years, providing that a small economic incentive for CCS becomes available (0.02-0.03 c€/kWhe or 30-50 €/ tCO2avoided). There is also a wide range of R&D developments that can deliver substantial reductions in CO2 capture cost and efficiency penalties. Calcium looping is one of such promising technologies, and its recent progress is briefly reviewed in this presentation. Particular emphasis is given here to the latest developments in three research lines in this field: the postcombustion Ca-looping technology, which is to be demonstrated in 2011 at a pilot plant rated at 1.7 MWt (the largest of this type the world); the use of Ca-looping combustion technology combined with biomass combustion for “in situ” CO2 capture, (negative emissions concept); and the use of Ca-looping for a novel hydrogen production process. The different status of development of these technologies is discussed and key gaps of knowledge highlighted. 1. Introduction Responsible policy for climate change mitigation must consider realistic transition scenarios between today’s dominance of fossil fuels towards a decarbonised energy system for the future. CO2 capture and storage technologies (CCS) can be one of the major players in such a transition (IPCC SRCCS) [1]. This is because this is actually a technology option much more mature than many want to believe. It can deliver truly deep cuts in emissions from power generation at reasonable cost (30-50 €/tCO2 avoided or just between 0.02-0.03 c€/kWhe in the electricity sector), it can be applied to other large stationary sources, it can be applied to transport and other sectors trough decarbonised energy vectors (like electricity or low C fuels) and it can open the way to negative emission concepts when applied to biomass. Although the lack of economic incentives has so far limited large scale demonstration, it is well known that the complete CCS chain only requires technology blocks that have been already developed
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Oviedo ICCS&T 2011. Extended Abstract
and optimized at large scale for other applications [1].
Therefore, with adequate
incentives, CCS could be deployed very rapidly in the next two decades and contribute to a substantial fraction of the mitigation effort to be undertaken in the XXI century. Despite the maturity of major CCS technologies, there is a great interest worldwide in developing new CO2 capture technologies aimed at further reducing the energy penalty and the cost of “existing” capture equipment. As an example, this lecture is concerned with a family of gas separation technologies that have been growing in interest in recent years (see IEAGHG High Temperature Solid Looping Network website and recent reviews on the subject in references [2-5] )
Figure 1. (left) Evolution of the number of published papers on Ca-looping in the WoK database (search terms: “carbonation and calcination” or “Ca-looping” or “carbonate looping”). A search for “CO2 capture ” reveals 2098 papers in the same period.(right)IEAGHG 2nd High Temperature Solid Looping Network meeting . ECN (NL) Sept 2010.
As indicated by Figure 1, the number of peer-reviewed publications has been growing exponentially in recent years. But this is much more important: major utility companies, equipment manufactures, cement and other energy industries are supporting and conducting their own R&D work for these technologies. The next few years are likely to be critical to see if this can be a realistic technology development path or, as it is the case too often in the energy research field, we are condemned to be another foreverpromising scientific curiosity. The only way for a technology to be successful in the CO2 capture field is to deserve and to reach large scale demonstration. A great deal of exciting scientific and technical developments are still needed in the Ca-looping field to Submit before May 15th to [email protected]
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achieve that objective. 1. Recent progress in the development of Ca looping technologies Ca-looping technologies have in common the following high-temperature, reversible gas-solid reaction of CO2 with CaO: CaO(s) + CO2 ↔ CaCO3(s)
ΔH298=-178.8 kJ/mol
The basic concept of a CO2 separation using this reaction goes back to the XIX century. Process configurations for CO2 capture (involving calcination in pure atmosphere of CO2) were pioneered by Prof. Harrison group in the 90s [6], focused on H2 production by Sorption Enhanced Reforming, SER, using CaO. For postcombustion applications, the first proposal was published by Prof. Shimizu et al in 1999 [7]. In this process [Figure 1 and ref 7] there are two interconnected fluidized bed reactors. One of the reactors operates as a CO2 absorber, between 600-700ºC,
and the second reactor
operates in oxyfired mode at much higher temperatures (over 870ºC) to allow the decomposition of CaCO3 into CaO and CO2. Other configurations have been proposed to provide the necessary heat for calcination from the power plant, minimizing energy penalties to just the CO2 purification and compression penalty [8] but the first generation of Ca-looping systems will follow the scheme of Figure 2 and are therefore linked to the successful demonstration of Oxyfired CFB combustion [9]. Flue gas “without” CO2
POWER PLANT
Flue Gas
CARBONATOR
CO2
Concentrated CO2
CALCINER
CaCO3
O2 CaO
Air
ASU Coal
Air
CaCO3 CaO (F0) Purge
Coal
N2
Figure 2. General scheme of a postcombustion Ca-looping system. Power is generated from the “existing” power plant and the “new” capture system, which includes a CFB Oxyfuel Combustor to allow both continuous calcination of CaCO3 formed in the carbonator.
As indicated in Figure 2, the CO2 captured from the flue gases as CaCO3 and the CO2 produced by the oxy-fired combustion of coal in the calciner are recovered in concentrated form from the calciner gas. A considerable fraction (40-55%) of the total energy entering the system is used in the calciner. Most of this energy leaves the system Submit before May 15th to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
in mass streams at high temperature (at T>900ºC) or is recovered as carbonation heat in the carbonator (at around 650ºC). Thus, the large energy input into the calciner comes out of the system as high quality heat that can be used in a highly efficient steam cycle (several groups have estimated this penalty to be between 6-8 net points, including compression). Post combustion Ca-looping is the only capture system that introduces repowering to the existing power plant [10] because the calciner is indeed very similar to a new oxyfired fluidized bed power plant. Some of the main hurdles for the process (sorbent deactivation trends and make up flow requirements, heat integration issues, attrition etc.) are now much better understood and quantified [see reviews 2-5]. I think it is fair to claim for CSIC in Spain, together with our colleagues of CANMET in Canada and IFK in Germany, a substantial contribution to the rapid development and credibility of this postcombustion option. We published a first set of results operating the carbonator fluidized bed [11] that has been expanded in recent years with a data base of experiments in different pilot plant configurations, including continuous CFB operation of interconnected fluidized bed pilots of several 10s of kW (see [12-14] and Figure 2). During 2011, there will be a new pilot of 1.7 MWt entering operation in la Pereda, Asturias, Spain (see for example Sanchez et al in a different communication to this conference, and Figure 3) and a different 1MWt pilot in Germany has already reported some preliminary results [15]. 1.00 Eequil 0.90 0.80
E c arb
0.70 0.60
Tcarb (ºC)
Ecarb CANMET 0.50
ug (m/s)
Ecarb IFK (BFB)
0.40
Ecarb INCAR‐CSIC
0.30 0
10
Operating conditions CSIC IFK CANMET 653 655 669
20
2
1.27
0.90
NCaO/FCO2 (h)
0.29
0.35
2.45
FCaO/FCO2 (‐)
7.86
17.95
‐
vCO2
0.15
0.13
0.08
30
40
50
60
Runni ng ti me (mi n)
Figure 3. Example of experimental capture efficiencies (Ecarb) in three different reactors set-ups and conditions. (see details in [12-14]).
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Oviedo ICCS&T 2011. Extended Abstract
Figure 4. Overview of the 1.7 MWt “La Pereda” pilot plant (Mieres-Spain) and its interconnected CFB reactor system
There are several important variations respect to the general scheme of Figure 2, that can exploit improved thermal integration options, synergies with the cement industry (that is natural consumer of large quantities of CaO purge) or synergies with desulfurization systems (also large consumers of CaO sorbents). Also, general improvements for other postcombustion CO2 capture options (like higher steam conditions) would be beneficial for the Ca-loop. The same applies to future improvements in oxygen separation efficiencies and cost or CO2 purification and compression. One of such variations is of particular strategic importance, because it deals with CO2 capture from biomass firing. Even opponents of CCS as a climate change mitigation technology understand the potential of the negative emission concept when applied to biomass. In this context, one option for the Ca-looping technologies involves merging the two main blocks of Figure 2 into a single “combustor-carbonator” reactor. The high reactivity of biomass allows for effective combustion in air at a sufficiently low temperature (below 700ºC) to allow capture “in situ” by CaO of the CO2 generated during the biomass combustion. This step is followed by the calcination of CaCO3 in an interconnected oxyfuel combustor as above. We have recently reported (Alonso et al [16]) experimental results from a 30 kWt pilot facility and a second test facility of 300
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Oviedo ICCS&T 2011. Extended Abstract
kWt is being commissioned to validate results (see presentations at this conference by Chamberlain et al). Preliminary economic studies have shown that the concept can be competitive if current incentives for biomass use are combined with expected credits from the ETS (CO2 market in Europe). Finally, the following paragraphs will briefly discuss the status of Ca-looping processes for H2 production. As indicated in the introduction, there are several recent reviews providing a comprehensive introduction to the fundamentals of the different technology options under this category. In these processes, CO2 is removed “in-situ” by CaO, so that the reforming of the fuel gas (note that the fuel gas can also come from an “in situ” gasification step) and the water gas shift reaction occur simultaneously together with CO2 capture. For natural gas: CH4(g)+ CaO(s) + 2H2O(g) ' CaCO3(s) + 4H2(g) ΔH298K= -13.7 kJ/mol The presence of a CO2 sorbent like CaO, shifts the equilibrium to the right, with the result that the almost complete conversion of methane and CO is achieved. This leads to a higher hydrogen yield under relatively mild conditions of pressure and temperature. In addition, the process can be simplified and reduced to one step, as the overall reaction is slightly exothermic so that no supplemental energy is required for the production of H2. Other benefits of these processes have been referred to by several authors in the literature (see review [3]). There is an equivalent concept for coal, originally developed in the 1960s (the “Acceptor gasification process” Curran et al [17]) that was successfully demonstrated in a large pilot fluidized bed gasifier. The main hurdle for successful scaling up of these technologies for CCS may not be in the H2 generation step, but in the CaCO3 regeneration step within a CO2-rich atmosphere. If the reactions are carried out in fluidized bed at high pressure, the calcination in oxyfuel conditions is very challenging because the equilibrium dictates operating temperature higher than 1000ºC (material issues and rapid sorbent deactivation over 950ºC do not favour these conditions). If the reactions are carried out in fixed bed systems, it is also very challenging to supply the heat for calcination to the fixed bed of solids. We have recently proposed a novel process to overcome these difficulties that is represented in Figure 5.
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H2 (CO2 free flue gas)
Fuel gas+H2O
CaO → CaCO3 (Cu)
(A)
CO2+H2O
(C)
N2
(B)
CuO → Cu CaCO3 → CaO
Cu → CuO (CaCO3)
Air Fuel gas
Fig. 5 - General scheme of the Cu/Ca three step chemical loop.
The reaction at the top (step A) is a conventional sorption enhanced reforming in the presence of CaO. However, the CaO required in this reaction is coming from a previous step (C) where the heat for calcination is supplied by the exothermic reduction of CuO using a fuel gas (natural gas or other). This produces a gas stream rich in CO2 and H2O suitable for CO2 dehydration and compression for transport and storage. An intermediate step (B in Figure 5) is required for the oxidation of Cu to CuO at conditions so that the calcination of CaCO3 is minimal; in order to avoid the loss of CO2 in the stream of O2 depleted air leaving the Cu oxidation reactor. The sequence of reactions in Fig. 5 can be implemented as a series of steps carried out in fixed bed reactors changing gas atmospheres and pressures. A full conceptual design of the process is now completed [17] considering ideal models for the reactors and incorporating thermodynamic and rate limitations known from the application of similar Ca-looping and chemical looping combustion systems reported in the literature. 4. Conclusions Calcium looping embraces a wide variety of technologies for CO2 separation at high temperature using the reversible carbonation and calcination reactions of CaO/CaCO3. The separation of CO2 at high temperature allows for a very effective heat integration in the CO2 capture system so that most of the additional energy used for regeneration (calcination) is effectively recovered for power generation or hydrogen production. The practical implementation of specific Ca-looping technologies is progressing at different pace. Postcombustion CO2 capture pilot plant experiments in the MWt range are expected in 2011 and a fast scaling up process can take place if the similarities with
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Oviedo ICCS&T 2011. Extended Abstract
existing CFBC equipment, including oxyfuel CFB combustion, can be exploited. An important variation of this process can be applied to biomass combustion with “in situ” capture of CO2 by CaO, leading to a negative emission CCS concept. Precombustion Ca-looping system is inherently more complex but offer higher efficiencies and compact H2 production schemes. As a research topic, Ca-looping technologies are deserving a fast growth in interest. This must translate in the next few years in the validation of the key technologies at increasing scales...
Acknowledgements The postcombustion work at CSIC is funded by Hunosa and Endesa and by the EC FP7 project “CaOling”. The biomass combustion work is founded by Gas Natural Fenosa under the MENOS CO2 project. The Cu-Ca loop work is founded by the Spanish Ministry of Industry and Commerce and MICINN under the project ENE2009-11353.
References [1] Intergovernmental Panel on Climate Change (IPCC). Special Report on Carbon Dioxide Capture and Storage. 2005 [2] Anthony EJ. Solid Looping Cycles: A New Technology for Coal Conversion. Ind. Eng. Chem. Res. 2008; 47: 1747-1754. [3] Harrison DP. Sorption-Enhanced Hydrogen Production: A Review. Ind. Eng. Chem. Res. 2008; 47: 6486-6501. [4] Blamey J.et al. The calcium looping cycle for large-scale CO2 capture. Progress in Energy and Combustion Science 2010; 36: 260-279. [5] Dean et al. The calcium looping cycle for CO2 capture from power generation, cement manufacture and hydrogen production Chemical Engineering Research and Design 2011; 89; 836–855 [6] Han C et al. Simultaneous shift reaction and carbon dioxide separation for the direct production of hydrogen. Chem Eng Sci. 1994; 49: 5875-5883. [7] Shimizu T., et al. A twin-bed reactor for removal of CO2 from combustion processes. Trans. IChemE. 1999; 77 (part A): 62-68. [8] Martínez et al. Conceptual design of a three fluidised beds combustion system capturing CO2 with CaO Int. J. of Greenhouse Gas Control 5 (2011) 498–504 [9] Myöhänen K. et al. Near Zero CO2 emissions in coal firing with oxyfuel CFB boiler. [10] Romeo L.M. et al. Oxyfuel carbonation/calcination cycle for low cost CO2 capture Submit before May 15th to [email protected]
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in existing power plants. Energy Conversion Management 2008; 49: 2809-2814. Chem. Eng. Technol. 2009; 3: 355-363. [11] Abanades JC, Anthony EJ et al. Capture of CO2 from combustion gases in a fluidized bed of CaO. AIChE J. 2004; 50: 1614-1622. [12] Lu D.Y. et al Ca-based sorbent looping combustion for CO2 capture in pilot-scale dual fluidized beds. Fuel Processing Technology 2008; 89: 1386-1395. [13] Rodríguez N. et al. Experimental investigation of a circulating fluidized bed reactor to capture CO2 with CaO. AIChE Journal 2010; 57: 1356-1366. [14] Charitos A. et al. Experimental validation of the calcium looping CO2 capture process with two circulating fluidized bed carbonator reactors. Industrial Engineering Chemistry Research, 2011 (in press) [15] Galloy A., Epple B et al. CO2 capture in a 1MWt fluidized bed reactor in batch mode operation. CCT2011, 8-12 May, Zaragoza-Spain [16] Alonso M. et al. Biomass Combustion with in Situ CO2 Capture by CaO. II. Experimental Results Ind. Eng. Chem. Res., 2011, 50 (11), 6982–6989 [17] Curran G.P., Fink C.E., Gorin E. CO2 acceptor gasification process. Studies of acceptor properties. Adv. Chem. Ser. 1967; 69: 141-161. [18] Fernández et al. Conceptual design of a hydrogen production process from natural gas with CO2 capture using a Ca-Cu chemical loop. 2011. Int. J. of Greenhouse Gas Control (accepted for publication).
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N2O and CO emission increase during oxy-char combustion in fluidized bed Astrid Sánchez1, Yuli Betancur1, Eric Eddings2, Fanor Mondragón*1 1
Chemistry Institute, University of Antioquia, A.A. 1226. Medellín-Colombia. Dept. of Chemical Engineering, University of Utah. Salt Lake City. Utah-USA *[email protected] 2
Abstract Oxy-coal combustion is one of the most promising alternatives for CO2 capture. In this process the combustion reaction is carried out in a higher oxygen concentration than the one used in conventional air combustion. In general, CO2, produced in the combustion is recirculated to the burner. Thus, combustion is carried out in O2/CO2, instead of O2/N2. Suppression of N2 from air avoids dilution of CO2 in N2 which facilitates its capture; also this process diminishes formation of thermal- and prompt-NO. Given that, fuel-N is the main source of N in oxy-combustion systems, it is very important to understand how the simultaneous increasing of O2 and CO2 partial pressures affects the evolution of Nspecies. With this in mind, a synthetic char with high N content was prepared and subjected to combustion under different oxidizing atmospheres. For example, increasing the O2 concentration in an Ar gas or in CO2 and also by keeping constant the O2 partial pressure, while CO2 was gradually increased. It was found that increasing the O2 partial pressure promotes the formation of NOx, and more interesting, it was observed that the presence of CO2 affects the behavior of N2O evolution, giving rise to an increase of these emissions. At the same time it was observed that CO production rises with the char conversion. Reactions were carried out in a fluidized bed reactor at 800 °C. FTIR with a gas cell was employed to continuously monitoring of the gases produced in the combustion reactions.
1. Introduction
Coal has been one of the most employed sources of energy [1] especially by means of combustion processes. One of the main problems associated with the use of this abundant resource is obtaining the maximum energy producing the minimum amount of
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Oviedo ICCS&T 2011. Extended Abstract
pollutant emissions [2]. Emissions from combustion can be classified into nitrogen oxides, fine particulates and greenhouse gases [3]. In coal combustion, the main emitted compounds are CO2 and H2O; additionally, coal contains sulfur and nitrogen in its structure, that are converted to sulfur dioxide (SO2) and nitrogen compounds (nitrogen oxides -NOx-, molecular nitrogen -N2-, and nitrous oxide -N2O) in the process [4]. SO2 and NOx are gases that can cause acid rain and, at the same time NO contributes to the formation of photochemical smog; CO2 and N2O are greenhouse gases [5]; and N2O is implicated indirectly in the depletion of the ozone layer [6-7].
Different technologies have been studied and developed to mitigate such emissions, such as low NOx burners, fluidized bed combustion employing limestone, among others. In the case of reduction of CO2 emissions different options have been proposed. These alternatives makes part of a general technology known as Carbon Capture and Storage (CCS), in which the main objective is the separation and storage the CO2 produced by the fuel combustion. Among the options for CCS we have chemical looping combustion, pre-combustion process, post-combustion process and oxy-combustion among the most popular.
One of the problems with the separation of CO2 from conventional coal combustion is the low concentration of CO2, being around 15% by volume [8],[9]. Oxy-combustion technology offers a solution to this difficulty by firing fossil fuel with high O2 partial pressure [10], rising the levels of CO2 as high as 95 % [8]; however, the use of high O2 concentration will cause a very high temperature in furnaces, for example, the temperature of a lignite char particle in combustion in air can be around 2000 K, while in the combustion in pure oxygen the temperature can increase up to around 2750 K [10]. High temperatures can favor the melting of ashes, formation of NOx and cause damages in furnaces materials being hard to retrofit current boilers.
To moderate the combustion temperatures, the reactant O2 is mixed with recycled flue gas, RFG, (containing mostly CO2 and H2O) [8, 10], RFG can be dried by condensation of water before it enters to the furnace. Thus, coal in the boiler burns in an O2/CO2 environment, instead of an O2/N2 atmosphere [10], as in conventional air combustion.
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Oviedo ICCS&T 2011. Extended Abstract
This technology offers several advantages as the possibility to capture CO2 and the diminishing of NOx emissions, since N2 coming from the air is very low, reducing the formation of thermal NOx. However, coal contains N in its structure (fuel-N) that is split into volatile-N and char-N during the first stages of combustion [11]. Both, volatile- and char-N suffer transformations through homogeneous and heterogeneous reactions giving rise to the formation of NOx (NO and NO2), N2O and N2 [7]. Many differences have been determined in the comparison between conventional air combustion and oxy-fuel combustion: burner stability, heat transfer, flame ignition and propagation, gas temperature, char burnout, radiation properties of the flame, efficiency of the boiler, evolution of pollutants, and ash properties [8, 10, 12]. Several studies have reported the increasing in CO emissions and differences in NOx formation, particularly due to the low N2 coming from air and NO reduction when gas recycling is implemented. However, it is necessary to understand in more detail the conversion of char-N under these conditions and the effect on the gaseous equilibriums involving CO 2, CO and N-species.
2. Experimental section
A stainless steel tube with a 2.6 cm internal diameter and a 76 cm length electrically heated was employed as the combustion reactor. 100 g of sand from the Merck Company®, with particle size 0.1-0.21 mm was used as the fluidizing medium. Sand bed height was approximately 11 cm; while the free board was around 32 cm. Combustion temperature was 800 °C, controlled by a type K thermocouple placed directly into the sand bed. Total gas flow rate was 500 mL/min STP, controlled by means of mass flow controllers, giving rise to a mean gas velocity of 0.94 m/s. A synthetic char with high N content obtained from the pyrolysis of polyacrylonitrile (PAN) at 800 °C, identified as PAN-8, was used to monitoring the evolution of Ncomplexes to nitrogen species under different oxidizing atmospheres. Gases, produced in the combustion were continuously analyzed by means of Fourier Transform Infrared (FTIR) spectroscopy with a gas cell. Spectra were taken every 2 seconds with the aim to obtain high detail in combustion reactions.
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Oviedo ICCS&T 2011. Extended Abstract
3. Results and Discussion
Combustion of PAN-8 was carried out with different oxidizing atmospheres, increasing both the O2 and CO2 partial pressures. Experiments increasing the O2 partial pressures indicate that the combustion reaction is faster with the increase in O2 concentration, and at same time that instantaneous concentrations of all emitted species (NO2, N2O and CO2) are higher with the increasing in O2 concentration, particularly, emission of NO2 is favoured.
When both, O2 and CO2 concentrations rise simultaneously similar conclusions can be achieved. However, when CO2 concentration is increased with a fixed O2 partial pressure more interesting observations can be obtained. Figure 1 shows the comparison for the emission of N2O employing 30% O2 with and without CO2 in the balance gas. From this figure it is possible to see that N2O formation is enhanced by the presence of CO2, especially at the beginning of the reaction. This behaviour could be explained by several homogeneous reactions, although heterogeneous reactions can be affected as well these evolution profiles.
450
N2O concentration (ppm)
400
0%-CO2 60%-CO2
350 300 250
200 150 100 50 0
0
50
100
150
200
250
Reaction time (s)
Figure 1. N2O evolution for PAN-8 combustion employing 30% O2 – 70% Ar () and 30% O2 – 60% CO2 – 10% Ar ()
Figure 2 shows the evolution of CO2, CO, NO, NO2 and N2O for several O2 concentrations when 75% of conversion has been reached. As mentioned before, the
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Oviedo ICCS&T 2011. Extended Abstract
emission of NO2 is favoured by the increasing in O2 concentration; on the other hand, the formation of CO is diminished, probably due to the homogeneous oxidation to produce CO2. This oxidation process can take place by means of different radicals or stable molecules, such as N2O that is reduced to N2 by the reaction: N2O + CO N2 + CO2. This reaction has been confirmed by reactions with homogeneous mixture of CO and N2O. However, when high O2 concentration is present, as in oxy-combustion systems, this reaction is inhibited by the preferential reaction of CO and oxygen.
NO2 75O2/25Ar
CO
NO
N2O
50O2/50Ar 40O2/60Ar 21O2/79Ar
Figure 2. Spectra corresponding to 75% of conversion for combustion of PAN-8 with several O2 concentrations
CO2 has an important effect on CO gaseous phase equilibriums, when CO2 is present in the reactant atmosphere CO emission is high, as can be seen in Figure 3, where it is possible to see that when there is O2 in the reacting mixture all CO is oxidized to CO2. If CO/Ar is the reacting mixture, some CO2 can be detected due to the gas phase equilibrium between these species; however, when CO2 is added, the CO signal becomes bigger indicating that CO2 has displaced the equilibrium.
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CO2 CO
Figure 3. 20CO/20O2/60Ar (blue), 20CO/80Ar (black), 20CO/20CO2/60Ar (red)
4. Conclusions
One of the main changes in oxy-combustion technologies consists in the increasing in O2 and CO2 partial pressures, causing several important differences with respect to conventional combustion. This work was focused in the emission of N-species and CO when both O2 and CO2 partial pressures were systematically increased. High O2 concentration promotes the formation of NO2 and diminishes the evolution of CO; however, the increasing in CO2 concentration enhances the production of N2O and increases the CO emission.
Acknowledgements The authors gratefully acknowledge support from the “Programa Sostenibilidad 20092011” of the “Universidad de Antioquia. A.S. thanks Colciencias and the “Universidad de Antioquia” for her Ph.D. scholarship and the U.S. Department of Energy (DOE) through the Utah Clean and Secure Energy (CASE) Program (FC26-08NT0005015) for her internship.
References
[1] Graus W., Worrell E. Methods for calculating CO2 intensity of power generation and consumption: A global perspective. Energy Policy In Press, Corrected Proof.
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[2] Mastral A.M., Callén M.S., García T. Polycyclic Aromatic Hydrocarbons and Organic Matter Associated to Particulate matter Emitted from Atmospheric Fluidized Bed Coal Combustion. Environmental Science & Technology 1999;33:3177-84. [3] Gupta R. Advanced Coal Characterization: A Review. Energy & Fuels 2007;21:45160. [4] Perry M.B., Clean Coal Technology, in: J.C. Cutler (Ed.) Encyclopedia of Energy, Elsevier, New York, 2004, pp. 343-57. [5] Ren Q., Zhao C., Wu X., Liang C., Chen X., Shen J., et al. Effect of mineral matter on the formation of NOx precursors during biomass pyrolysis. J. Anal. Appl. Pyrolysis 2008;85:447-53. [6] Zhu J., Lu Q., Niu T., Song G., Na Y. NO emission on pulverized coal combustion in high temperature air from circulating fluidized bed – An experimental study. Fuel Processing Technology 2009;90:664-70. [7] Thomas K.M. The release of nitrogen oxides during char combustion. Fuel 1997;76:457-73. [8] Rathnam R.K., Elliott L.K., Wall T.F., Liu Y., Moghtaderi B. Differences in reactivity of pulverised coal in air (O2/N2) and oxy-fuel (O2/CO2) conditions. Fuel Processing Technology 2009;90:797-802. [9] Seepana S., Jayanti S. Steam-moderated oxy-fuel combustion. Energy Conversion and Management 2010;51:1981-8. [10] Bejarano P.A., Levendis Y.A. Single-coal-particle combustion in O2/N2 and O2/CO2 environments. Combustion and Flame 2008;153:270-87. [11] Glarborg P., Jensen A.D., Johnsson J.E. Fuel nitrogen conversion in solid fuel fired systems. Progress in Energy and Combustion Science 2003;29:89-113. [12] Wall T., Liu Y., Spero C., Elliott L., Khare S., Rathnam R., et al. An overview on oxyfuel coal combustion--State of the art research and technology development. Chemical Engineering Research and Design 2009;87:1003-16.
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Oviedo ICCS&T 2011. Extended Abstract
Chemical Looping Combustion of Coal-derived Synthesis Gas containing H2S over Supported Fe2O3 - MnO2 Oxygen Carriers Ewelina Ksepko1, Ranjani V. Siriwardane2, Hanjing Tian2,Thomas Simonyi2, Marek Sciazko1 1 Institute for Chemical Processing of Coal, 1 Zamkowa, 41-803 Zabrze, POLAND, E-mail: [email protected], Tel: 00 48 32 271-00-41 ext. 228 2 U.S. Department of Energy, 3610 Collins Ferry Road, Morgantown, WV 26507-0880, USA, E-mail: [email protected], Tel: 001 304-285-4513
Abstract The paper contains results of research work on novel, promising combustion technology so called as chemical looping combustion (CLC). The objective of our work was to prepare supported bimetallic oxygen carriers and to evaluate the performance of these for the CLC process with synthesis gas/air. Thermo gravimetric analysis (TGA) and low pressure bench scale flow reactor tests were conducted to evaluate the performance. Multi cycle tests were conducted in TGA with oxygen carriers utilizing simulated synthesis gas with and without H2S impurities. Effect of H2S contaminations on both the stability and the oxygen transport capacity was evaluated. Multi cycle CLC tests were also conducted in the bench scale flow reactor at 800 °C with selected samples. Chemical phase composition was investigated by X-Ray diffraction (XRD) technique. Five Cycle TGA tests at 800 °C indicated that all oxygen carriers had a stable performance at 800 °C. It was interesting to note that there was complete reduction/oxidation of the oxygen carrier during the 5-cycle test. The fractional reduction, fractional oxidation and global reaction rates were calculated from the data. It was found, that support had a significant effect on both fractional reduction/oxidation and the reaction rate. The reaction profile was changed by the presence of H2S but there was no effect on the reaction rate due to presence of H2S in synthesis gas. Low pressure bench scale flow reactor data indicated stable reactivity, consumption of oxygen from oxygen carrier and complete combustion of H2 and CO. XRD data of samples after multi-cycle test showed stable crystalline phases and complete regeneration of the oxygen carriers after multi-cycle tests.
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Oviedo ICCS&T 2011. Extended Abstract
1. Introduction The objective of the paper was to prepare three mixed-metal oxide oxygen carriers, consisting of Fe2O3 and MnO2 supported on Sepiolite (Mg4Si6(OH)2*6H2O), ZrO2, and Al2O3, and to evaluate the performance of these carriers during the CLC process in the presence of synthesis gas/air. The advantages of using mixed-metal oxides and the effect of the support on their performance will be discussed. Sulfur is the major impurity in coal synthesis gas, as well as in natural gas. Depending on the coal type and resource, coal-derived synthesis gas may contain 200–8000 ppm H2S, which may interact with the metal-oxide oxygen carrier during combustion reaction, thus affecting the performance of the CLC system. There are only a few experimental papers that report studies on the effect of contaminants [1–3]. Additionally, thermodynamic analyses indicated that there is an interaction between H2S and the metal oxide of an oxygen carrier which may lead to carrier deactivation [4].
2. Experimental section
Preparation of Fe2O3-MnO2 supported on ZrO2, Sepiolite, and Al2O3. The oxygen carriers with a composition of 60 wt % Fe2O3 and 20 wt % MnO2 supported on ZrO2, Sepiolite, and Al2O3 were prepared using the solid-state mixing method. Thermogravimetric
analyses.
TGA
experiments
were
conducted
in
a
thermogravimetric analyzer (TA Instruments Model 2050) in which the weight change of the various metal-oxide oxygen carriers was measured isothermally as a function of time during reduction-oxidation cycles. Five reduction-oxidation cycles were conducted at atmospheric pressure to determine the stability of the carriers. An approximately 20 mg sample was heated in nitrogen in a quartz bowl to the reaction temperature. A simulated coal-derived synthesis gas mixture of 4042 ppm H2S, 13% CO2, 38% CO, 17.8% He, and 30.8% H2 balanced by nitrogen was used for the reduction segment, while zero air was utilized for the oxidation segment. All reaction gas flow rates were 45 sccm. For all experiments, reduction time was 120 min, and oxidation reaction time was 60–90 min. To avoid the mixing of reduction gases and air, the system was flushed with nitrogen for 15–45 min before and after each reduction reaction. For comparison purposes, tests were also conducted with a synthesis gas composed of 12% CO2, 36%
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Oviedo ICCS&T 2011. Extended Abstract
CO, 25% He, and 27% H2 without H2S. To understand the effect of temperature, TGA experiments were carried out both at 800 °C and 900 °C.
3. Results and Discussion Effect of support. The results of the five-cycle CLC TGA data of three bimetallic oxygen carriers—Fe-Mn oxides/Sepiolite, Fe-Mn oxides/ZrO2, and Fe-Mn oxides/Al2O3 at 800 °C are shown in Figure 1. It is interesting to note that all the supported carriers showed stable performance at 800 °C, which indicated that Sepiolite, Al2O3, and ZrO2 are suitable supports for Fe-Mn, which contributes to stable reactivity during cyclic tests. Moreover, Fe-Mn oxides/Sepiolite has the highest oxygen capacity (18.24 wt %) at 800 °C. Fe-Mn oxides/ZrO2 had 17.46 wt %. Fe-Mn oxides/Al2O3 had the lowest oxygen capacity at 15.10 wt %.
Figure 1 Five-cycle reduction-oxidation TGA data for (a) Fe-Mn oxides/ZrO2, (b) Fe-Mn oxides/Al2O3, and (c) Fe-Mn oxides/Sepiolite at 800 ˚C.
22
(b)
(a)
800
800
18
0
20 400
200
200 18
16 0
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800
1000
1200
0 1400
0
200
400
600
800
1000
1200
0 1400
Time (min)
Time (min)
20
(c)
0
Weight (mg)
600
Temperature ( C)
800
18
400
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16
0
200
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600
800
1000
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0 1400
Time (min)
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Temperature ( C)
400
600
Weight (mg)
600
0
Weight (mg)
20
Temperature ( C)
22
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Oviedo ICCS&T 2011. Extended Abstract
When a five-cycle CLC test was conducted with syngas at 900 °C, the reaction profiles were very similar to those achieved at 800 °C with all Fe-Mn oxygen carriers. All FeMn oxygen carriers showed stable performance during the five-cycle test, both at 800 °C and 900 °C. It is assumed that the presence of manganese contributed to better stability and rate of Fe-Mn carriers specifically at higher temperature, while comparing to unstable supported Fe oxide. Effect of H2S impurities. TGA data of the five-cycle reduction-oxidation test for Fe-Mn oxides/Sepiolite with simulated synthesis gas containing H2S are shown in Figure 2a. For comparison, data with pure synthesis gas are also shown. Figure 2b contain expanded views of the third reduction-oxidation cycle.
Figure 2 Thermogravimetric analysis of Fe-Mn oxides/Sepiolite with coal-derived synthesis gas at temperature 800 °C: (a) five-cycle test and (b) third cycle data.
(a)
synthesis gas (b)
1.0 1.0
N2
N2
Air
synthesis gas/H2S 1.0
W/Wox
W/Wox
0.8
Synthesis gas with H2S
0.8
Synthesis gas 0
300
600
900
1200
0.8
Time (min)
600
700
800
Time (min)
The TGA profile in the presence of H2S is different from that without H2S. However, a stable reduction-oxidation profile was obtained during the five-cycle test in the presence of H2S. This could be attributed to an excellent regenerability in and the stable physical structure of bimetallic oxygen carriers even in the presence of the H2S, which was supported by XRD data. Powder diffraction pattern did not revealed any sulfide/sulaftes formation in regenarated oxide simple. Moreover, oxygen transport capacity was was not affected by presence of H2S contaminations (Figure 2b).
4. Conclusions
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Oviedo ICCS&T 2011. Extended Abstract
There are many advantages of using Fe and Mn oxides as oxygen carriers. Both oxides are widely available at low cost, and they create minimal health and environmental concerns. They also contribute to high oxygen transport capacity from 15.10 wt % to 18.24 wt % for aluminia and sepiolite, respectively. Additionally, they lend to significant physical strength to the prepared carrier. Mixing a monometallic iron oxygen carrier with another metal oxide such as manganese oxide improved physical stability over multiple reduction-oxidation cycles. The presence of H2S did not appear to affect the stability during the five-cycle test which may indcate the successful regeneration was achieved.
Acknowledgement. This study was partially financed by the Polish Ministry of Higher Education and Science, project No. 685/N-USA/2010/0.
References [1] Adanez J, Garca-Labiano F, Gayan P, Diego LF, Abad A, Dueso C, Forero C. Effect of gas impurities on the behavior of Ni-based oxygen carriers on chemical looping. Energy Procedia 2009; 1 (1), 11–18. [2] Tian H, Simonyi T, Poston J, Siriwardane R. Effect of hydrogen sulfide on chemical looping combustion of coal-derived synthesis gas over bentonite supported metal oxide oxygen carriers. Ind. Eng. Chem. Res. 2009; 48 (18), 8418–8430. [3] Ksepko E, Siriwardane R.V, Tian H, Simonyi T, Sciazko M. Comparative investigation on chemical looping combustion of coal-derived synthesis gas containing H2S over supported NiO oxygen carriers. Energy & Fuels 2010; 24 (8), 4206–4214. [4] Jerndal E, Mattisson T, Lyngfelt A. Thermal analysis of chemical-looping combustion. Trans IchemE, Part A, Chemical Engineering Research and Design 2006; 84 (A9), 795–806.
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Oviedo ICCS&T 2011. Extended Abstract
Drying Behavior of Brown Coal under the Various Temperature Conditions with Halogen Heat Source and its Formulation Y. Matsushita1, N. Mitsuhara1, T. Harada2 1 Research and Education Center of Carbon Resources, Kyushu University, 6-1 Kasuga-koen, Kasuga, Fukuoka, 816-8580 Japan 2 Research Laboratory, Kyushu Electric Power Co., Inc, 12-1-47 Shiobaru, Minami-ku, Fukuoka, 815-8520 Japan Abstract In this study, using HR83 Halogen moisture analyzer (METTLER TOREDO), as-mined, 5-20 g, Victorian Brown coal in Australia (Loy Yang coal) is heated and dried under the various temperature conditions from 50 to 150 °C with Halogen heat source to measure the drying rate. Furthermore, the experimental results are analyzed to understand the fundamental drying behavior and also to evaluate the formulation that can predict the drying rate under the various temperature conditions, constant temperature conditions as well as some heating conditions. As a result, the drying rate gradually decreases with an increase in drying ratio. The behavior is quite different from wood’s one, that is one of the most famous materials for drying and shows two drying periods, so-called constantrate drying period and decreasing drying period. Drying rate constant can be expressed as a linear relationship or second-order polynomial equation with respect to temperature. The evaluated formulation can predict the drying rate and drying termination time not only in constant temperature conditions but also in some heating conditions with high accuracy.
1. Introduction Advanced utilization technology of low rank coal with high ash or high moisture content is necessary with an increase in coal consumption in developing countries. Victorian Brown coal in Australia is well known Brown coal with high moisture content around 60%. To enhance the energy conversion efficiency, the drying process of Brown coal with low temperature heat source can be considered as the key technology. Some research can be found to analyze the water in Brown coal. For example, Norinaga et al. [1] have investigated the water in Brown coal by using DSC. The study of Brown coal is summarized in Li [2]. In this study, the fundamental drying behavior is investigated by Halogen moisture analyzer and the drying rate is formulated by the simple drying rate equation Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
2. Experiment Loy Yang coal, one of the most famous Brown coal in Australia is used in this study, which is provided by HRL Technology. Proximate and ultimate analysis of Loy Yang coal and Newlands coal, one of the most famous bituminous coal for the comparison are listed in Table 1. Volatile matter in Loy Yang coal is more than in Newlands coal and fixed carbon in Loy Yang coal is less in Newlands coal in proximate analysis, and oxygen atom is much involved in Loy Yang coal in ultimate analysis. Drying rate is calculated by measuring change in weight of water in Brown coal due to drying by Halogen moisture analyzer HR83 (METTLER TOLEDO) which has Halogen heat source. The advantage of the apparatus is weight of the sample can be measured continuously in the constant temperature or the constant heating rate conditions. The disadvantage of the apparatus is the atmosphere gas and humidity in the apparatus cannot be controlled. The measurement conditions are summarized in Table 2. The sample is as-mined, involving 56 wt. % water, Loy Yang coal whose weight and diameter are approximately under φ2 mm and 10 g, respectively. Note that the sample weight is varied only in investigating the effect of the sample weight. The gas introduced in the apparatus is air with humidity 40 RH % at temperature 25 °C (the experimental room condition).
Table 1 Proximate and ultimate analysis of Loy Yang coal and Newlands coal d.b. [wt. %]
d.a.f. [wt. %]
VM
FC
ASH
C
H
O
N
S
LY
53.31
44.80
1.89
79.66
3.97
15.48
0.67
0.22
NL
29.20
56.67
15.13
84.63
4.69
8.69
1.70
0.29
3. Results and discussion 3.1 Drying behavior of Brown coal Figure 1 shows change in weight of water in Brown coal for drying time in the constant temperature condition, , T = 90°C. The weight decrease due to drying in both cases. The drying rate shows its maximum value and then decrease gradually during drying. The reason drying rate increases in the first period of drying is that the sample temperature increase due to heating in the period. However, the drying rate seems to decrease
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Oviedo ICCS&T 2011. Extended Abstract
continuously if the heating time is enough small and can be ignored. The behavior of Brown coal drying is quite different from the one of woody drying that shows the constant-rate drying period and the decreasing-rate drying period. Furthermore, the drying rate is high and the drying time is short with the higher temperature condition (not shown here).
Table 2 Measurement conditions Sample
as-mined Brown coal (Loy Yang coal)
Water content
56 wt. %
Weight of sample
10 g (5–20 g)
Sample diameter
≤ φ2 mm
Atmosphere gas
Air with 40 RH % at 25°C
Constant temperature condition
50–150 °C
Constant heating rate condition
1–100°C/min
Figure 1 Change in weight of water in Brown coal and its drying rate for drying time
3.2 The effect of the sample weight Figure 2 shows the drying rate dw/dt for (1−X) with the sample weight 5, 10, and 20 g in the constant temperature condition, T = 90°C. The drying rate shows its maximum value in the first period of drying, and then decrease gradually like an arc. In case of the sample weight 5 g, the drying rate is large compared with the other cases. This is
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Oviedo ICCS&T 2011. Extended Abstract
because temperature to control overshoots in the first period of drying. However, the drying rate shows almost the same independent of the sample weight.
Figure 2 Drying rate dw/dt for (1−X) with the sample weight
3.3 Evaluation of drying rate equation In the previous section, the relationship between drying rate and drying time cannot be found though drying rate for drying time in the constant temperature conditions is investigated. Therefore, the relationship between drying rate and drying fraction X, defining drying fraction X = 1 − w/wi, in the constant temperature conditions is investigated. Figure shows drying rate, dw/dt for (1−X). The drying rate shows its maximum value at the first period of drying around (1−X) = 0.95, and then decreases gradually like an arc independent of temperature. The hehavior of Brown coal drying does not corresponds with the one of wood drying that shows constant-rate drying period and decreasing-rate drying period. Furthermore, the behaviour is almost the same in the various temperature conditions. In the previous sections, it is already found that there seems to be relationship between dw/dt and (1−X), and the drying rate is independent of weight of the samples. Therefore, the drying rate equation will be evaluated base on the relationship between dw/dt and (1−X)n. dw/dt = k⋅(1−X)n
(1)
Figure 3 shows dw/dt⋅(1−X)0.75 for (1−X) for the various constant temperature
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Oviedo ICCS&T 2011. Extended Abstract
conditions, respectively, where, k is drying rate constant, and n is the order of the drying rate equation and is determined as n = 0.25 by try and error. The drying rate constant, k is evaluated with respect to temperature as k(T) since it can be described as first order (50–90°C) or second-order (50–150°C) polynomial function. Figure 4 shows the drying rate constant, k(T) for the various constant temperature conditions. The drying rate constant, k(T) are evaluated with respect to temperature since drying rate constant, k(T) increases with an increase in temperature. Above all, the drying rate can be evaluated as a function of drying rate constant, k(T) and (1−X)0.75. dw/dt = k(T)⋅(1−X)n
(2)
k(T) = 2.356×10−5⋅T2 + 9.178×10−4⋅T − 3.445×10−1 (50 ≤ T ≤ 150 °C)
(3)
Figure 3 dw/dt⋅(1−X)0.75 for (1−X)
Figure 4 k(T) for temperature
3.4 The validation of the drying rate equation The validation of the drying rate equation evaluated in the previous section is investigated by comparing change in weight of water in Brown coal due to drying under the constant temperature and the constant heating rate conditions. Figures 5 shows predicted and measured change in weight of water in Brown coal due to drying under the constant temperature and the constant heating rate conditions. The drying rate equation describes the measured results well in the constant temperature condition. However, the equation underestimates the termination time of drying compared with the measured one. This is because change in weight of water in Brown coal is larger than the measured one, in other words, the equation overestimates the drying rate in the late period of drying. It is well known that there are some different kinds of water in Brown coal, the one behaves as water in which water molecules bind each other, and another does not in which water molecules bind to the functional group of the Brown coal surface. The
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Oviedo ICCS&T 2011. Extended Abstract
bonds of former one are stronger than those of latter one and it seems to take more time to remove such water. Figure shows predicted and measured change in weight of water in Brown coal due to drying in the constant heat rate condition. The drying rate equation also well predicts the measured results in the first and the middle periods of drying, and overestimates the ones in the last period of drying. Therefore, the evaluated drying rate equation can be applied not only to the temperature conditions but also to the variable temperature conditions.
(a) in constant temperature
(b) in constant heating rate
Figure 5 Predicted and measured change in weight of water in Brown coal
3.5 Drying time Figure shows the drying time needed for 1 g sample to achieve drying fraction 0.25, 0.50, 0.75, or 0.98, respectively with the one estimated by the dry rate equation evaluated in the previous section. The drying time becomes shorter with an increase in drying temperature to achieve the all drying fraction. The drying time drastically increases with a decrease in temperature in case of high drying fraction. Comparing the predicted results with the measured ones, the drying time estimated by the drying rate equation is in good agreement with the measured one in case of lower drying fraction, however, the predicted time shows shorter than the measured one in case of higher drying fraction. The reason the drying rate equation underestimates the drying time in case of higher drying fraction is that the drying rate equation predicts the drying rate well in the first and the middle periods of drying, however overestimates the one in the last period as discussed in the previous section.
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Oviedo ICCS&T 2011. Extended Abstract
Figure 6 Drying time needed for 1 g sample
4. Conclusions In this study, the fundamental drying behavior is investigated by Halogen moisture analyzer and the drying rate is formulated by the simple drying rate equation. This study can be summarized as follows: ⋅ Drying behavior of Brown coal is quite different from that of wood. ⋅ Drying rate of Brown coal decrease gradually with an increase as drying progresses. ⋅ Drying rate of Brown coal can be evaluated by the simple drying rate equation with respect to temperature and drying fraction.
Acknowledgement This research is partially supported by international advanced utilization of brown coal in Victoria state based on the gasification technology for fundamental international collaboration research for clean coal technology in international clean coal technology development project in the energy innovation program presented by New Energy and Industrial Technology Development Organization (NEDO).
Nomenclature k(T) n
[g/s]
drying rate constant order in the drying rate equation
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Oviedo ICCS&T 2011. Extended Abstract
T
[°C]
drying temperature
t
[s]
drying time weight of water in Brown coal
w X
[−]
drying fraction
Subscript i
initial
References [1] Norinaga et al., Evaluation of Drying Induced Changes in the Molecular Mobility of Coal by Means of Pulsed Proton NMR, Energy Fuels, 12(5), pp. 1013-1019 (1998) [2] Chun-Zhu Li, Advances in the Science of Victorian Brown Coal, Elsevier, 2004
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Oviedo ICCS&T 2011. Extended Abstract
MODELLING AND SIMULATION OF A COAL GASIFICATION PROCESS IN PRESSURIZED FLUIDIZED BED
F. Chejne1 , E. Lopera1, C. A. Londoño1, C. A. Gómez1
Universidad Nacional de Colombia, Facultad de Minas. Grupo de Termodinámica Aplicada y Energías Alternativas, TAYEA. Medellín, Colombia. [email protected] Abstract A coal gasification mathematical model that can predict temperature, converted fraction and particle size distribution for solids have been developed for a high pressure fluidized bed. For gases in both emulsion and bubble phase, it can predict temperature profiles, gas composition, velocities and other fluidynamic parameters. In the feed zone, it could be considered a Gaussian distribution or any other distribution for the solid particle size. Experimental data from literature have been used to validate the model. Finally, the model can be used to optimize the gasification process changing several parameters, such as excess of air, particle size distribution, and coal type and reactor geometry.
1. Introduction Many models used for describing pressurized gasification were developed for low pressure processes and then validated with experimental data obtained at moderated pressure. Yan et al. [1] proposed a one-dimensional model incorporating a distinctive characteristic: a net flow from emulsion to bubble phase in the conservation equations, generated by devolatilization, homogeneous and heterogeneous reactions. Further improvements were made to this model [2], incorporating an energy balance in order to predict operation temperature, and to use the new model to simulate coal gasifiers at both pilot and industrial scale. Later Ross et. al. [3] considers non-isothermal behavior of gases taking into account the heat transfer mechanisms in the fluidized bed. Thus, Yan et al. [3] state that the improved model can accurately predict global coal conversion; Hamel et al. [4] presented a gasification model in pressurized fluidized bed reactors. Their model included bed and freeboard fluid-dynamics, kinetic for drying, devolatilization and chemical reactions. De Souza-Santos [5-8] developed a comprehensive mathematical model and a computer program, to use as a tool for engineering design and operation optimization, by predicting the behavior of a real unit during steady-state operation. The basis of the model developed in this paper carried out by Chejne et al. [9]. This model is based in the two-phase theory, which considers bubble and emulsion phase, and can accurately predict solids temperature, gas temperature profiles, Submit before January 15th to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
converted coal fraction, gas composition among other important parameters. The main modifications to the model of Chejne et al. [9] consist in incorporating high pressure effects in transport phenomena, bed fluid-dynamics, physical properties, kinetic models for heterogeneous reactions and other aspects that could change with pressurized operation of gasification process. The model structure is the same, due to mass and energy balance equations do not change with operational conditions. Kinetics of coal reactions in combustion and gasification processes, especially at low pressures, is vast [10-20]. Liu et al. [14, 15] propose an integrated mechanism of reaction, incorporating three combustion and five gasification reactions with CO2, H2O, CO and H2. Although, the purpose of this model is for quantitative design at high pressure, kinetic parameters of coal oxidation, proposed by Hurt et al. [11], were determined at atmospheric pressure. The hydrodynamic behaviour of pressurized fluidized beds is not completely known. Pressure and temperature have strong influence on some characteristics of a gas-solid fluidized bed, such as minimum fluidization velocity, bubble size, velocity and stability. Yates [21] have undertaken a review of theoretical and experimental studies of gas-solid fluidization at elevated pressures and temperatures, for both low velocity operation (bubbling fluidized beds) and high velocity operation (circulating beds). Yates [21] references a methodology for calculating minimum fluidization velocity in fluidized beds at super-ambient conditions, proposed by Yang et al. [22]. However, the mathematical complexity and lack of clarity in the calculation method, do not justify its use in the model developed in this paper, as it was presented by Lopera et al. [23, 36]. Chiester et al. [24] studied experimentally some hydrodynamic parameters at high pressures for several types of particles, and proposed empirical correlations for calculating minimum fluidization velocity and porosity, and bed expansion. These equations accurately represent the experimental data and the effect of pressure on those parameters. For Geldart [25] Group B and D powders, minimum fluidization velocity decreases as pressure increases, while minimum fluidization porosity remains constant. Gogolek et al. [26] reviewed fluid-dynamics in pressurized fluidized bed combustors. In addition to fundamental hydrodynamics, solids mixing, heat and mass transfer, elutriation and entrainment are also treated. Llop et al. [27] presented experimental results for high pressure bed expansion using particles belonging to Groups B and D, and developed a semi-empirical model which allows the prediction of that parameter as a function of the expansion coefficient. They also refer the equations for bubble diameter, according to Mori and Wen [28] for Group B solids, and to Darton [29] for Group D solids. The main objective of this work is to develop a phenomenological model to predict the gasification behaviour in a fluidized bed reactor operating at high pressures, starting from an existing model of the same process at low pressure.
2. Theoretical section: Development of Mathematical model The most important characteristics and advantages of the coal gasification model described in this paper are: One-dimensional, it means any variation in bed occurs in axial direction, steady-state, it is feed and product streams are constant, two phases, they
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Oviedo ICCS&T 2011. Extended Abstract
are emulsion and bubble phase. The emulsion is formed by gas and solids. The bubble is considered free of solid particles; therefore it is formed only by gas, the solid is considered isothermal and the consumption uniform through the bed height. It is justified by high recirculation of solids inside the bed; the consumed particles are quickly replaced by new ones, attrition, elutriation and drag are included for solids, bed fluid-dynamics is described by empirical equations, in order to avoid the solution of highly complex momentum equations. It is very difficult to treat these balance equations properly in a one-dimensional model, devolatilisation and drying are considered instantaneous in the feed zone and the energy equations for both phases are coupled by convection phenomena on the surface of the particles. Inside each equation, the mass and heat transfer coefficients are also calculated. 2.1. Basic equations The transference phenomena that must occurr inside reactor (figure 1) is resumed in figure 2. These transference phenomena are necessary to do the mass balance for the gas in the emulsion, bubbles and for the coal particle.
Figura 1. Schema of reactor in fluidized bed at high pressure
The mathematical model requires getting the variation of each component along the axial direction and the generation (or consumption) term from the heterogeneous and homogeneous reaction and the exchange by convection with the bubble (see details in F. Chejne et al [31] and E. Lopera [32]).
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Oviedo ICCS&T 2011. Extended Abstract
Figura 2. The transference phenomena inside reactor
Due to the homogeneity and isothermal conditions for the solids (coal, limestone and/or inert material), the mass balance is global and is integrated over the volume of the reactor. For the coal, the difference between the inlet and outlet flow is equivalent to the oxygen and gasification reactions in the reactor. The mass balance for the limestone requires that the rate of generation (or consumption) due to sulphur reactions is equivalent to the differences between the inlet and outlet flow. Finally, the inlet and outlet flow for the inert material are equal. 2.2. Fluid Dynamics The fluid-dynamic phenomena involved in the process are too complicated to be studied in analytic form. Therefore, hydrodynamic parameters in the model are determined by means of experimental correlations for high pressure operation of fluidized beds, obtained from different sources. The most important parameters in the fluidization process are the velocities and diameters of different phases (see details in F. Chejne et al [31] and E. Lopera [32]). In order to maintain a fluidized bed, the conditions of minimum fluidization must be satisfied. This means that the drag force of the fluid in motion has to overcome the weight of solid particles, so levitation of the solids is achieved. The minimum gas velocity for fluidization is defined in the point when drag force and weight of solid particles are equivalent. This is the limit between a fixed and a fluidized bed. To find this velocity, the minimum fluidization Reynolds number was calculated using the correlation developed by Chiester et al. [24] for high pressure operation. 2.3. Particle size distribution The present model considers that inside the reactor the average diameter changes due to attrition, elutriation, drag and consumption. The average diameter at the feed point and inside the reactor is calculated as,
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Oviedo ICCS&T 2011. Extended Abstract
d p,M =
1 W ∑l d l l, M n-1
(1)
where the mass fraction of each level in the size distribution is:
Wl =
ml ∑ mi
(2)
i i≠l
and the average diameter of each level in the distribution is calculated as:
d l,ave =
d l+1 + d l 2
(3)
At the feeding point the initial diameter and mass of each level in the distribution are given, and with equation (2) the mass fraction of each level is found. The model manipulates these three vectors (diameter, mass and fraction) with the elutriation fraction, and new values are obtained. From these vectors, the diameters and masses are affected by the consumption fraction, while the fractions remain constant. In other words, the mass fraction of each level is not affected by the consumption, but its value will be affected by attrition and drag. The mass fraction of each level in the distribution inside the reactor is obtained by, 2.4. Chemical Reactions
At the feeding point, it is considered instantaneous devolatilization and drying processes presented simultaneously, for the remaining coal combustion and gasification reactions are considered, that is reactions with oxygen, steam and carbon dioxide. For devolatilization and drying, the kinetic coefficients are calculated from the mass balance principle, while for the combustion, gasification and limestone reactions the kinetic coefficients are function of temperature and composition. For limestone and homogeneous reactions, the kinetic coefficients were calculated with different expressions from several references and it used two different model for solid-gas reaction as you can see in figure 3 (see details in F. Chejne et al [31] and E. Lopera [32]).
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Oviedo ICCS&T 2011. Extended Abstract
Figura 3. (a) Non- exposed core model (b) exposed core model [10]
For combustion and gasification reactions at high pressure, the kinetic model used is the proposed by Roberts et al. [12]. This model uses expressions which have the Arrhenius form to calculate kinetic coefficients of the reaction rates.
R i = k i Pgn
(4)
−E ⎞ k i = A i exp ⎛⎜ i RT ⎟⎠ ⎝
(5)
The exposed core model was chosen for combustion and gasification reactions, while for the limestone reaction the unreacted core model was used. The rates include reaction and diffusion resistance. The rates of combustion, gasification and limestone reactions are then calculated as follows [7]: 3. Results and dsicussion.
In Figure 4, comparison between experimental data and the model results is made. Six different experiments (Kabawata, [30]) and numerical results are compared, and the results are satisfactory. The dashed lines represent ±20% calculation error. It can be observed that most of the results are within the 20% range. The higher differences are presented in the CO and CO2 molar composition, due to the facts previously mentioned. In spite of these differences, the model can predict the tendencies of molar composition obtained in the process.
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Oviedo ICCS&T 2011. Extended Abstract
Figura 4. Participation of gas production to different pressure at 1,8 stem/coal ratio.
Other important advantage of pressurized fluidized beds is the enhancement of convective heat and mass transfer. Despite transfer coefficients change along the process, they were compared in last 3 cm of bed. It was found that at higher pressures heat transfer coefficients are considerable bigger than those at atmospheric pressure, but mass transfer is not so different. 4. Conclusions
We have developed a two-phase mathematical model for description of a high pressure gasification process in a fluidized bed reactor, using the basic conservation equations in compact form in order to obtain an adequate and accurate numerical solution. The high pressure model was developed starting from an existing model for atmospheric pressure and modifying the affected parameters affected by pressure. The numerical results presented here agreed very well with the experimental data taken from literature, thus proving the efficiency and accuracy of both the mathematical and numerical model. The presented model can be used for studying the process behavior at high pressures and for observing the effect of variation of several parameters in coal gasification, as well as for helping in the design of new equipments of this kind. Acknowledgements
There are a lot of institutions and people the authors wish to thank.COLCIENCIAS (Instituto Colombiano para el desarrollo de la Ciencia y Tecnología Francisco José Caldas), ISAGEN, Universidad Nacional de Colombia – Medellín Campus, Universidad Pontificia Bolivariana and Universidad de Antioquia for their logistic and economic
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Oviedo ICCS&T 2011. Extended Abstract
support, and special thank to INCARBO (Centre of Investigation and Innovation of Coal in Guajira, Colombia).
References [1] Yan H, Heidenreich C, Zhang D. Fuel 1998;77:1067. [2] Yan H, Heidenreich C, Zhang D. Fuel 1999;78:1027. [3] Ross DP, Yan H, Zhong Z, Zhang D. Fuel 2005;84:1469. [4] Hamel S, Krumm W. Powder Technol 2001;120:105. [5] de Souza-Santos ML. Fuel 1989;68:1507. [6] de Souza-Santos ML. Fuel 1994;73:1459. [7] de Souza-Santos ML. Solid fuels combustion and gasification. Modeling, simulation, and equipment operation, USA; 2004. ISBN: 0-8247-0971-3. [8] de Souza-Santos ML. Fuel 2007;86:1648. [9] Chejne F, Hernandez JP. Fuel 2002;81:1687. [10] Hurt R, Sun J, Lunden M. Combust Flame 1998;113:181. [11] Hurt RH, Calo JM. Combust Flame 2001;125:1138. [12] Roberts DG, Harris DJ. Energ Fuel 2000;14:483. [13] Roberts D, Harris D. Fuel 2007;86:2672. [14] Liu G, Tate AG, Bryant GW, Wall TF. Fuel 2000;76:1145. [15] Liu G, Niksa S. Prog Energ Combust Sci 2004;30:679. [16] Wall TF, Liu G, Wu H, Roberts DG, Benfell KE, Gupta S, et al. Prog Energ Combust Sci 2002;28:405. [17] Mann M, Knutson R, Swanson M, Erjavec J. In: 12th International conference on coal science, Australia; 2003. [18] Hong J, Hecker W, Fletcher T. Proc Combust Inst 2000;28:2215. [19] Liliedahl T, Sjöström K. Fuel 1996;76:29. [20] Monson C, Germane G, Blackham A, Smoot D. Combust Flame 1995;100:669. [21] Yates JG. Chem Eng Sci 1995;51:167. Review Article Number 49. [22] Yang WCh, Chiester DC, Kornosky RM, Keairns DL. AIChE J 1985;31:1086. [23] Lopera E, Chejne F, Londoño C. VII Congreso Nacional de Ciencia y Tecnologı´ a del Carbón y Combustibles Alternativos. Colombia: Valledupar; 2007. [24] Chiester DC, Kornosky RM, Fan LS, Danko JP. Chem Eng Sci 1984;39:253. [25] Geldart D. Powder Technol 1973;7:285. [26] Gogolek P, Grace J. Prog Energ Combust Sci 1995;21:419. [27] Llop M, Casal J, Arnaldos J. Powder Technol 2000;107:212. [28] Mori S, Wen CY. AlChE J 1975;21(1):109–15. [29] Darton RC. Trans Inst Chem Eng 1979;2:134. [30] Kawabata J, Yumiyama M, Tazaki Y, Honma S, Takeda S, Yamaguchi H, et al. Chem Eng Commun 1981;1981:335. [31] Farid Chejne *, Eliana Lopera, Carlos A. Londoño, Fuel 90 (2011) 399–411. [32] Lopera E: “Modelamiento y simulación de un proceso de gasificación de carbón en lecho fluidizado a alta presión”Tesis de Maestría, Facultad de Minas, Universidad nacional de Colombia; 2009
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8
Oviedo ICCS&T 2011. Extended Abstract
Preliminary Studies on Ash-Free Coal Gasification at Mild Condition J. Yoo, S. Jin, H. Choi, Y. Rhim, J. Lim, D. Chun, S. Kim, S. Lee Clean Coal Research Center, Korea Institute of Energy Research, 102 Gajeong-ro, Yuseong-gu, Daejeon, Korea 305-343 Addresses: [email protected]
Abstract Highly efficient IGCC have adapted the coal to gaseous fuel conversion route. The commercial gasification reactors are being operated at harsh condition (T > 1300 °C and p > 30 bar), putting much burden in economy and safety respect. The kinetics of the gas conversion can be significantly increased by introducing the catalyst. However, the catalytic activity is commonly deactivated by its irreversible interaction with mineral matters in coal. This work addresses how to achieve the coal gasification at the mild condition. Following the production of ash-free coal by thermal extraction using 1methylnaphthalene solvent, preliminary experiments on its pyrolysis and CO2/H2O gasification were carried out. The ash-free coal made of a low rank coal (Samhwa, lignite) was a little less susceptible to devolatilization during pyrolysis than its raw coal. In the absence of the catalysts and at T < 800 ˚C, no evident gasification of the ash-free coal was observed, showing very similar TGA profile to its pyrolysis run. With a further increase in temperature, the gasification reactions kicked in, and which produced CO by CO2 gasification (Boudouard reaction), and H2/CH4/CO/CO2 by steam gasification (coal-water and water-gas shift reaction). As a next step to achieve faster kinetics at the mild condition, a mixture of the gasification catalyst and the ash-free coal will be prepared and evaluated in terms of the advantage over its raw coal.
1. Introduction Coal is one of the most important energy sources, currently accounting for ~25% of energy consumption around the world1, 2. The use of coal is expected to increase as a result of its abundance and economic advantage, even though there are various problems coupled with the combustion of coal such as emission of greenhouse (GHG) and pollutant gases. Especially, a coal–fired power sector is responsible for ~35% GHG
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Oviedo ICCS&T 2011. Extended Abstract
emission and has been continuously blamed as a main culprit of global warming3. The Korean legislature enacted Low Carbon Green Growth Act in 2009, targeting 30% reduction of GHG emissions, compared to BAU, by 20204. A series of measures are required in order to meet the goal without damaging the energy security and the economy. In the way of this challenge, Korean government listed an integrated gasification combined cycle (IGCC) in the action plan, aiming at the development of the commercial scale plant. Many of IGCC demo plants (50 − 600 MW scale) have been operated successfully, mostly in USA, EU, and Japan5. The thermal efficiency was estimated to be above 48%, much improved from conventional coal-fired power plants (~28% for subcritical and ~43% for ultra supercritical class)6. IGCC is now outlooked as a realistic and energyefficient alternative and also as low carbon next generation technology. In the foreseeable future, IGCC is very likely to play a key role for coal utilization in Korea. Coal gasification is one of the most critical processes in highly efficient IGCC7. The majority of the gasification processes scaled up for commercialization adopted entrained-flow, slagging gasifiers, which are practically meaningful only at harsh operating condition (> 1400 °C and 20 − 70 atm)8. This severity puts much burden in economy and safety respect. At the lower temperature (T < 800 °C), the conversion kinetics is generally slow and therefore of no practical use, unless the catalyst-aided coal gasification is performed. The coal to gas conversion can be significantly increased by using alkali, alkali earth, and Ni/Fe based active catalyst 9 . However, the catalytic activity is commonly not repeatable due to deactivation by irreversible interaction of the catalyst with the mineral matters in coal9. The ash in coal is ill-natured, decreasing the power efficiency and also being discharged as an air pollutant10. A lot of works have concentrated on the development of efficient methods to prepare ash-free coal 11 . Among them, thermal extraction with organic solvents has produced ash-free coal, namely “Hypercoal,” most successfully, potentially solving the ash problem12. This work addresses how to realize the coal gasification at the mild condition. The ash in coal was removed by thermal extraction method and the resultant ash-free coal was gasified under H2O and CO2 flow. The pyrolysis and gasification behavior of the ash-free coal was discussed.
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Oviedo ICCS&T 2011. Extended Abstract
2. Experimental A soluble carbon component in Samhwa coal (brown coal) was extracted using 1methylnaphthalene (1-MN) solvent. The extraction was processed via extraction, filtration, and drying step. The raw coal was ground, meshed to < 74 μm, and dried in a vacuum oven at 100 °C. Coal slurry was prepared by adding 20 g coal into 200 g 1-MN. The slurry was added to an extractor (0.5 liter volume), which was then purged with N2 gas. While stirring, the mixture was heated to 370 °C and held for an hour under 30 bar. To separate the solvent extract from the residual matter, a filtering was done at the stainless steel filter unit. The solvent in the extract was removed by being kept in a vacuum oven (~300 °C) under N2 atmosphere for 3 − 4 hr. Finally, extracted coal (EC) was obtained. Properties of Samhwa coal and its solvent extract are tabulated in Table 1. Table 1. Proximate/ultimate analysis and calorific value of Samhwa coal (raw coal) and its solvent extract. Sample (wt%) Raw Extract
Moisture
Volatile matter*
Ash (dry)
6.7 0.9
52.3 46.5
2.2 0.2
Fixed carbon* 38.6 52.3
C
H
N
O
70.5 86.3
5.1 5.2
0.9 0.9
23.5 7.6
Heat value (kcal/kg) 4,554 8,066
*daf: dry & ash-free
Preliminary results on catalytic gasification reactions of ash-free coal were obtained using thermogravimetric analysis {TGA, SDT Q600 (TA Instrument) and Setsys (Setaram)}. In this work, no catalysts were introduced to the system. In the beginning, both the raw coal and its solvent extract were pyrolyzed under 100 cc/min N2 or Ar flow. Temperature increased from 20 to 950 ˚C at 10 ˚C/min ramp rate and then stayed at 950 ˚C for 30 or 60 min. The effluent gases were analyzed using either FT-IR (Nicolet 6700, Thermo Scientific) or quadrupole mass spectrometer {(QMS, E5CN (Omnistar)}. Then, the ash-free coal was gasified with 100 cc/min of CO2 and H2O supply. The analysis of the effluent gases was repeated using the same instruments.
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Oviedo ICCS&T 2011. Extended Abstract
3. Results and Discussion Pyrolysis profile of the ash-free coal was compared with that of the raw coal and shown in Fig. 1. No significant difference was found until 600 ˚C. The raw coal lost more weight than the extract at 600 − 800 ˚C, likely due to higher content of the volatile matter (Table 1). During a further increase of temperature (800 to 950 ˚C) and isothermal at 950 ˚C, the lines were parallel with each other, indicating that the labile nature of the extracted portion was similar to that of the raw coal. 120
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0
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Fig. 1. Pyrolysis profile under N2 flow for Samhwa coal and its ash-free solvent extract. The CO2 gasification of the ash-free coal was performed and compared with its pyrolysis reaction (Fig. 2). A wt% profile of the CO2 gasification was about the same as that of the raw coal at T < 850 ˚C, which pointed out that CO2 behaved like an inert N2 gas and no CO2 gasification occurred in this region (Fig. 2(a)). In the analysis of the effluent gases using FT-IR, CH4 peaks were similarly positioned at 300 − 550 ˚C (Fig. 2(b)). Again, this indicated that CO2 was non-reactive with the coal at the lower temperature. A gap between the two lines was formed from ~850 ˚C and became bigger with increasing temperature. A continuous loss of the weight was shown for CO2 gasification at 950 ˚C isotherm, while there was negligible weight change with N2 flow at the same condition. As shown in Fig. 2(c), a Boudouard reaction (CO2 + Ccoal Æ CO) kicked in and produced CO gas under CO2 flow, but only after temperature reached above ~700 ˚C. A faster kinetics is expected with an introduction of the gasification catalysts, which will be investigated later on.
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Oviedo ICCS&T 2011. Extended Abstract
120
6x 10 -3
(b) CH4
(a) wt%
CO2
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o
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Fig. 2. Comparison of pyrolysis under N2 and CO2 gasification reaction of the ash-free coal. (a) weight change profile, and evolved gas profile for (b) CH4 and (c) CO. The ash-free coal was gasified in the presence of 10% H2O in N2 and pyrolyzed under Ar gas (Fig. 3). The general trend of weight change for the steam reaction was very similar to that for the CO2 gasification, such that the reactions of coal with water at T < 850 ˚C were unnoticeable and mainly pyrolysis products were obtained (Fig. 3(a)). The evolved gases were analyzed as a function of temperature using QMS. The peaks listed below corresponded to the gas evolution by pyrolysis; a H2 peak (the maximum at ~750 ˚C, Fig. 3(b)), a CH4 peak (the maximum at ~520 ˚C, Fig. 3(c)), and a CO2 peak (the maximum at ~500 ˚C, Fig. 3(e)). A pyrolyzed CO peak was not identified due to strong m/e = 28 background. However, a significant increase of H2 and CO production was detected during the temperature ramp of 800 to 950 ˚C and the isotherm at 950 ˚C, as shown in Fig. 3(b) and (d), respectively. This increase happened as a consequence of coal-water (Carbon + H2O Æ CO + H2) reaction. In addition, a water-gas shift reaction (CO + H2O Æ CO2 + H2) resulted in an increase of a CO2 peak at T > 800 ˚C. As a next step, a mixture of up to Submit before May 15th to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
20% catalyst (K, Na, Ca, Mg, Fe, and Ni) and the ash-free coal will be prepared and evaluated in terms of the advantage over its raw coal. 120
1x 1 0
-10
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(b) H2
Ar 10% H2O
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wt%
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Fig. 3. Comparison of pyrolysis under N2 and H2O gasification reaction of the ash-free coal. (a) weight change profile, and evolved gas profile for (b) H2, (c) CH4, (d) CO, and (e) CO2 profile.
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6
Oviedo ICCS&T 2011. Extended Abstract
4. Conclusions Preliminary experiments were performed on the pyrolysis and CO2/H2O gasification of the ash-free coal at the mild condition using TGA, FT-IR, and QMS. The pyrolysis profile of the ash-free coal was similar to that of its raw coal, showing a little less devolatilization. In case of the CO2/H2O gasification without adding catalysts, the gases were evolved mainly from the pyrolysis at T < 800 ˚C. Production of CO by CO2 gasification (Boudouard reaction) and H2/CH4/CO/CO2 by steam gasification (coalwater and water-gas shift reaction) were enhanced by a further temperature increase (800 − 950 ˚C) and an isotherm at 950 ˚C. To obtain faster kinetics at mild condition, the gasification catalysts will be mixed with the ash-free coal and evaluated in terms of the advantage over its raw coal.
References 1. World Energy Outlook 2007 China and India Insights. Paris: IEA; 2007, p663. 2 . Kurose R, Ikeda M, Makino H, Kimoto M, Miyazaki T. Pulverized Coal Combustion Characteristics of High-Fuel-Ratio Coals. Fuel 2004;83:1777-85. 3. Climate analysis indicators tool (CAIT). version 2009. world resources institutel; 2009. 4. Overview of the Republic of Korea’s National Strategy for Green Growth. UNEP; 2010. 5. Liu H, Ni W, Li Z, Ma L. Strategic thinking on IGCC development in china. Energy Policy 2008;36:1-11. 6. Henderson C. Clean coal technologies. IEA clean coal centre report: CCC/74. London: Graham & Trotman; 2003. 7. Pérez-Fortes M, Bojarski A, Velo E, Nougués J, Puigjaner L. Conceptual model and evaluation of generated power and emissions in an IGCC plant. Energy 2009;34:1721-32. 8. Higman C, van der Burgt M. Gasification. 2nd ed. Amsterdam: Gulf professional publishing; 2008. 9. Corella J, Toledo JM, Molina G. Steam gasification of coal at low-medium (600 − 800 ˚C) temperature with simultaneous CO2 capture in fluidized bed at atmospheric pressure. Ind Eng Chem Res 2006;45:6137-6146. 10. Steel K, Patrick J. The Production of Ultra Clean Coal by Chemical Demineralization. Fuel 2001;80:2019-23. 11. Li C, Takanohashi T, Yoshida T. Effect of Acid Treatment on Thermal Extraction Yield in Ashless Coal Production. Fuel 2004;83:727-32. 12. Okuyama N, Komatsu N, Shigehisa T, Kaneko T, Tsuruya S. Hyper-coal Process to Produce the Ash-free Coal. Fuel Processing Technology 2004;85:947-67.
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7
Kinetic Study on the Lignite-CO2 Gasification in the Presence of K2CO3 V.C. Bungay1, B.H. Song1, S.D. Kim2, J.M. Sohn3, H.M. Shim4, Y.J. Kim4, G.T. Kim4, S.R. Park4, and Y.I. Lim5 1
Chemical Engineering Department, Kunsan National University, Gunsan, Jeonbuk 573-701, Korea: [email protected] 2 Department of Chemical & Biomolecular Engineering, Korea Advanced Institute of Science and Technology, Daejeon 305-701, Korea 3 Department of Mineral Resources & Energy Engineering, Chonbuk National University, Jeonju, Jeonbuk 561-756, Korea 4 SK energy Institute of Technology, Daejeon 305-712, Korea 5 Department of Chemical Engineering, Hankyong National University, Anseong, Gyonggi-do 456-749, Korea
Abstract The catalytic CO2 gasification of Mongolian lignite with K2CO3 has been performed in a thermogravimetric analyzer (TGA) and the kinetics of the reaction was studied. The gasification temperature was from 600°C to 900°C at ambient pressure. Catalyst was added to the lignite samples with 5-15 wt% loading by physical mixing or impregnation. The kinetic parameters like the rate enhancement factor and activation energy of the reaction have been evaluated using the gas-solid reaction models in the literature. It was observed at 600°C that the addition of K2CO3 increased the carbon conversion by 45% with respect to the uncatalyzed reaction. The enhancement in gasification rate at this temperature ranged between 10-25 times that of the uncatalyzed reaction. For coal gasification at higher temperatures, the enhancement in gasification rate was between 56 times that of the uncatalyzed reaction which implies that any further increase in the catalyst loading will not significantly affect the rate of the gasification reaction. On the other hand, the experimental results showed no significant difference in gasification rate whether catalyst was added by physical mixing or by impregnation. This study further confirms that K2CO3 exhibits good mobility which makes it an effective catalyst regardless of catalyst addition method. In addition, carbon conversion increases with increasing CO2 partial pressure as a result of the reduction of K2CO3 volatilization during gasification with N2-CO2 gas mixture. Most of gasification conversion behavior could be well predicted with the extended modified volumetric model.
1
1. Introduction Gasification of coal with CO2 and steam has been one of the main processes of producing inexpensive, clean and effective fuel. Specifically, gasification of low-rank coals draws much attention due to its higher reactivity relative to high-rank coals [1-2]. Alkali metal carbonates, either alone or in combination with other compounds, are known to enhance the reactivity of coal and other carbonaceous materials. The presence of these compounds greatly reduced the gasification temperature thus improving the process from the economic point of view. For coal gasification with CO2 [3,4] and steam [5-10], potassium carbonate (K2CO3) has been used and proven to improve the rate of coal conversion and significantly lower the gasification temperature. In most kinetic studies, thermogravimetric method is commonly employed being cost-effective, simple and provides rapid analysis of gasification reactions involving numerous parameters. Such method employs a thermogravimetric analyzer (TGA) which generates a curve that shows the variation of the sample weight with time and its derivative. This data is then converted to a rate-conversion curve and interpreted using various gas-solid reaction models. In this work, the kinetics of catalytic gasification of Mongolian lignite in the presence of K2CO3 was studied. The performance of the catalyst was reported at different temperatures, CO2 partial pressure, catalyst loading and catalyst addition method. Possibility of recycling the catalyst was also presented along with the evaluation of goodness of fit using various gas-solid reaction models. 2. Experimental section Mongolian lignite was used in this study with the proximate and ultimate analysis shown in Table 1. Coal samples were pulverized to particle sizes less than 0.250 mm. Potassium carbonate (99.5% purity) was used as catalyst and purchased from Samchun Pure Chemicals Co. Ltd. The catalyst was used as received and added to the coal sample by physical mixing and impregnation at 5%, 10% and 15% wt loading. Physical mixing was done by the addition of K2CO3 of particle size less than 0.250 mm to the coal sample and mechanically mixed in a sealed container. Impregnation was done by the addition of an aqueous solution containing the desired amount of K2CO3 to the coal sample and mixed thoroughly. Removal of the solvent was done in a rotary vacuum evaporator, air dried overnight and sieved to particle size less than 0.250 mm. Gasification experiments were carried out in a thermogravimetric apparatus (TGA Q50, TA Instruments) at temperatures ranging from 600°C to 900°C using CO2-N2 gas
2
mixtures. About 20 mg of the lignite sample was accurately weighed and filled in a platinum plan. Initially, the sample was heated isothermally at 110°C to remove the moisture and heated to the desired gasification temperature at ambient pressure to remove the volatile matter under N2 environment. The gas was then switched to CO2-N2 gas mixture at 0.60 atm CO2 partial pressure for the gasification of the fixed carbon. The residue was then treated with a mixture of air-N2 gas to obtain ash. This procedure was done to coal samples where temperature, catalyst loading and catalyst addition method were varied. To investigate the effect of CO2 partial pressure to the gasification reaction, CO2 partial pressure was varied from 0.20 atm to 0.80 atm from 600°C to 900°C with 5% wt loading of K2CO3. To evaluate the possibility of catalyst recycling, lignite with similar loading of K2CO3 was physically mixed and gasified at 700°C and 0.60 atm CO2 partial pressure. To maximize the use of the catalyst, recycling was done for the lignite-gasification with 0.60 atm CO2 partial pressure in the presence of 5% wt K2CO3 at 700°C. This gasification temperature was chosen due to the attainment of complete conversion for the given sample size at approximately 30 min of gasification time with CO2 and to investigate the catalytic effect rather than the thermal effect experienced at higher temperatures. In addition, the said temperature was far below the melting points of K2CO3 in N2 and CO2 environment which are 900°C and 905°C, respectively [11]. Ash and catalyst that remained from this process was weighed and lignite was added to obtain the same catalyst loading on the assumption that the amount of catalyst present was constant. Table 1. Proximate and ultimate analyses of Inner Mongolian lignite Proximate analysis (wt.% ar) Ultimate Analysis (wt.% daf) Moisture 23.34 C 65.33 Volatile matter 30.01 H 4.67 Fixed carbon 35.55 N 0.57 Ash 11.10 S 0.23 a O 18.10 a by difference Data of mass versus time from the gasification process was used to evaluate kinetic parameters using various gas-solid reaction models – homogeneous model (HM) [12], shrinking-core model (SCM) [1,13], random pore model (RPM) [14], modified volumetric model (MVM) [15] and the extended modified volumetric model (EMVM) [16]. From these parameters, the performance of the catalyst was evaluated based on its ability to lower the activation energy and to enhance the rate of the gasification process.
3
3. Results and Discussion 3.1. Uncatalyzed lignite-CO2 gasification To serve as a baseline data, uncatalyzed gasification of the lignite sample was carried out from 600°C to 900°C using CO2-N2 mixture at 0.60 atm CO2 partial pressure. As shown in Fig. 1a, about 90% conversion was attained at about 10 min for the given sample weight at 900°C. This shows that gasification was mainly influenced by temperature and such runs will serve as a basis for comparison for the catalyzed reactions. Using various gas-solid reaction models, the reaction rate constant was determined and from the Arrhenius plot shown in Fig. 1b, the activation energy and frequency factor of the uncatalyzed reaction was found to be 159.3 kJ-mol–1 and 2.5×108 hr–1, respectively.
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Figure 1. (a) Time-conversion and (b) Arrhenius plot for uncatalyzed lignite-CO2 gasification 3.2. Effect of temperature and catalyst loading The addition of K2CO3 dramatically increased carbon conversion as shown in Fig. 2. At 600°C, the presence of K2CO3 at 5% wt loading increased carbon conversion by 45% which signifies the positive effect of the catalyst. In addition, it is worth noting that at 800°C and 900°C, complete conversion was attained for the same gasification time and sample weight considered.
4
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1
Carbon conversion, X (-)
900°C 0.8 800°C 0.6 700°C 0.4 0.2 600°C 0 0
0.2
0.4
0.6
0.8
1
Dimensionless gasification time, t/t10 min (-)
Figure 2. Time-conversion plot for catalyzed lignite-CO2 gasification at various temperature and catalyst loading: (a) 5%, (b) 10% and (c) 15% wt loading K2CO3 To provide quantitative measure of the effect of the catalyst on the gasification reaction, Fig. 3 shows the gasification enhancement factor evaluated as the ratio of the reaction constant of the catalyzed reaction to that of the uncatalyzed reaction. From this plot, maximum enhancement in gasification rate was observed at 600°C which shows dominant catalytic effect rather than thermal effect. At temperatures higher than 600°C, enhancement in gasification rate was found to increase from 5-6 times that of the uncatalyzed reaction as catalyst loading was increased. As such, increase in the catalyst loading at these temperatures will not increase the rate of gasification reaction to a significant magnitude. One possible explanation for such saturation effect [3] in the reaction is the blocking of carbon pores by the catalyst therefore restricting the access of CO2.
5
Rate enhancement factor, kcat/kuncat (-)
27 15% Loading 10% Loading 5% Loading
24 21 18 15 12 9 6 3 600
700
800
900
Gasification temperature, T (oC)
Figure 3. Rate enhancement factor at different temperature and catalyst loading In terms of the ability of the catalyst to lower the activation energy of the gasification reaction, Fig. 4 shows the Arrhenius plots of the gasification reaction with and without the catalyst. From this plot, the activation energies of the gasification reactions with 5%, 10% and 15% wt loading of K2CO3 was found to be 132.90 kJ-mol-1, 124.88 kJ-mol-1 and 120.04 kJ-mol-1, respectively. 5 4 3
ln k
2 1 0 15% Loading 10% Loading 5% Loading Uncatalyzed
(1) (2) (3) 0.8
0.9
1
1.1
1.2
1000/T (K-1)
Figure 4. Arrhenius plots of gasification reactions 3.3. Effect of catalyst addition method In the previous runs, physical mixing was used to load the lignite samples with the catalyst due to its simplicity. To evaluate the effect of catalyst addition method, impregnation method was used to load 5% wt K2CO3 to the lignite sample. As shown in Fig. 5, reaction constants obtained from physically mixed and impregnated samples showed no significant difference with a square value of correlation index of 0.9970. Such observation is consistent with previous studies [9,17-19] wherein K2CO3 exhibited good mobility and proven to be effective whether loaded to coal via physical mixing or impregnation.
6
90
-1
kphysical mixing (hr )
80
5% loading
70
10% loading
60
15% loading
900°C
50 40 30
800°C
20 10
700°C
0
600°C
-10 -10 0
10 20 30 40 50 60 70 80 90
kimpregnation (hr-1)
Figure 5. Comparison of reaction constants for lignite sample physically mixed and impregnated with 5% wt loading of K2CO3 3.4. Effect of varying CO2 partial pressure For the variation of partial pressure of CO2 in a mixture of N2-CO2, Fig. 6 shows timeconversion plots for lignite gasification with 5% wt loading of K2CO3 at various temperatures. As expected, the subsequent increase in the concentration of the gasifying agent increased carbon conversion and hence the rate of gasification reaction. This positive effect in the gasification rate can be explained by the reduction of the rate of volatilization of K2CO3 in a N2-CO2 environment. Such observation was observed from a thermal stability study of K2CO3 near its melting point [11]. In this study, it was observed that the rate of volatilization of K2CO3 in pure N2 was nearly four times greater than the rate under pure CO2 environment. Therefore, the presence of increasing concentration of CO2 favored the backward reaction of the reversible decomposition of K2CO3. The activation energy of the reactions with varying CO2 partial pressure was found to be in the range of 132.7 ± 4.3 kJ-mol–1. Values obtained have relatively small deviation since activation energy is mainly a function of temperature and not concentration dependent.
a
b 0.80 atm CO2-0.20 atm N2 0.60 atm CO2-0.40 atm N2 0.40 atm CO2 - 0.60 atm N2 0.20 atm CO2 - 0.80 atm N2
0.8
Carbon Conversion, X (-)
Carbon Conversion, X (-)
1
0.6 0.4 0.2 0
1 0.80 atm CO2-0.20 atm N2 0.60 atm CO2-0.40 atm N2 0.40 atm CO2 - 0.60 atm N2 0.20 atm CO2 - 0.80 atm N2
0.8 0.6 0.4 0.2 0
0
0.2
0.4
0.6
0.8
1
Dimensionless gasification time, t/t120 min (-)
0.0
0.2
0.4
0.6
0.8
1.0
Dimensionless gasification time, t/t10 min (-)
7
d
Carbon Conversion, X (-)
1 0.8 0.6 0.4 0.80 atm CO2-0.20 atm N2 0.60 atm CO2-0.40 atm N2 0.40 atm CO2 - 0.60 atm N2 0.20 atm CO2 - 0.80 atm N2
0.2
1
Carbon Conversion, X (-)
c
0.8 0.6 0.4 0.80 atm CO2-0.20 atm N2 0.60 atm CO2-0.40 atm N2 0.40 atm CO2 - 0.60 atm N2 0.20 atm CO2 - 0.80 atm N2
0.2 0
0 0
0.2
0.4
0.6
0.8
0
1
0.2
0.4
0.6
0.8
1
Dimensional gasification time, t/t10 min (-)
Dimensionless gasification time, t/t10 min (-)
Figure 6. Time-conversion plots for lignite gasification at various CO2 partial pressures and gasification temperatures: (a) 600°C, (a) 700°C, (a) 800°C and (a) 900°C 3.5. Recycling of catalyst As shown in Fig. 7, the first re-use of the catalyst reduced the gasification rate by 60% and the succeeding recycle runs resulted to an average of 20% reduction in gasification rate. Decrease in rate may be attributed to the hindered diffusion of the catalyst to the coal surface as a result of the accumulation of ash particle and inefficient mixing of coal and catalyst. Another reason for deactivation of the catalyst is the possible reaction with compounds present in the ash [6,10,20]. Whether deactivation is physical or chemical in nature, recycle tests showed that K2CO3 present in the ash residue still possess some catalytic activity and can be used to some extent to enhance the rate of lignite-CO2 gasification. Catalyzed Reaction (5% K2CO3) 1st Recycle 2nd Recycle 3rd Recycle 4th Recycle 5th Recycle 6th Recycle Uncatalyzed
0.6
7 6
Catalyzed (5% K2CO3)
-1
0.8
Carbon Conversion, X (-)
b
1
Reaction Rate Constant (hr )
a
0.4
0.2
5 4 3 2
Uncatalyzed
1 0 0
0.2
0.4
0.6
0.8
Dimensionless gasification time, t/t30 min (-)
1
0 1
2
3
4
5
6
Reuse number
Figure 7. Plots of (a) carbon conversion and (b) reaction rate constants for gasification reactions with catalyst recycling 3.6. Kinetic modeling of catalytic lignite-CO2 gasification In the evaluation of kinetic parameters, several gas-solid reaction models were used. To evaluate which model best simulates the catalytic gasification reactions, goodness of fit was determined based on the square values of correlation index, R2. The validity of the models based on this parameter is shown in Fig. 8. From this figure, the extended 8
(EMVM) and modified (MVM) volumetric model and the random pore model was proven to correlate the gasification reactions at all temperatures. However, the simple models – shrinking core and homogeneous model – were observed to show goodness of fit at low temperatures. 600°C
0.98 0.96 0.94 0.92 0.90
EMVM
MVM
RPM
SCM
700°C
1.00
Square value of correlation 2 index, R
Square value of correlation 2 index, R
1.00
0.98 0.96 0.94 0.92 0.90
HM
EMVM
Gas-solid reaction models
800°C
0.98 0.96 0.94 0.92 0.90
EMVM
MVM
RPM
SCM
Gas-solid reaction models
RPM
SCM
HM
HM
900°C
1.00
Square value of correlation 2 index, R
Square value of correlation 2 index, R
1.00
MVM
Gas-solid reaction models
0.98 0.96 0.94 0.92 0.90
EMVM
MVM
RPM
SCM
HM
Gas-solid reaction models
Figure 8. Square value of correlation index (R2) for modeling catalytic of lignite-CO2 gasification using various gas-solid reaction models at different temperatures 4. Conclusions The catalytic activity of K2CO3 was proven to enhance gasification rate and lowering the gasification temperature to 800°C for the same given gasification time and sample weight. Increase in catalyst loading above 5% wt and addition of catalyst by physical mixing or impregnation showed no significant effect on enhancing gasification rate. Enhancement in gasification rate was observed to range from 10-25 times at 600°C while between 5-6 times at higher temperature. Increasing CO2 partial pressure results to increase in gasification rate due to reduced volatilization of the catalyst in CO2 environment. The extended modified volumetric model showed goodness of fit in all gasification reactions.
Acknowledgement This work was supported by Energy Efficiency and Resources R&D program (2009T100100675) under the Ministry of Knowledge Economy, Republic of Korea and SK Energy Co. Ltd.
9
References [1] Ye DP, Agnew JB, Zhang DK. Gasification of a South Australian low-rank coal with CO2 and steam: kinetics and reactivity studies. Fuel 1998;77:1209-19. [2] Beamish BB, Shaw KJ, Rodgers KA, Newman J. Thermogravimetric determination of the carbon dioxide reactivity of char from some New Zealand coals and its association with the inorganic geochemistry of the parent coal. Fuel Process Technol. 1998;53:243-53. [3] Li S, Cheng Y. Catalytic gasification of gas-coal char in CO2. Fuel 1995;74:456-8. [4] Sun Q, Li W, Chen H, Li B. The CO2-gasification and kinetics of Shenmu maceral chars with and without catalyst. Fuel 2004;83:1787-93. [5] Yeboah YD, Xu Y, Sheth A, Godavarty A, Agrawal PK. Catalytic gasification of coal using eutectic salts: identification of eutectics. Carbon 2003;41:203-14. [6] Wang J, Sakanishi K, Saito I, Takarada T, Morishita K. High-Yield Hydrogen Production by Steam Gasification of HyperCoal (Ash-Free Coal Extract) with Potassium Carbonate: Comparison with Raw Coal. Energy Fuels 2005;19:2114-20. [7] Wang J, Jiang M, Yao Y, Zhang Y, Cao J. Steam gasification of coal char catalyzed by K2CO3 for enhanced production of hydrogen without formation of methane. Fuel 2009;88:1572-9. [8] Wang J, Yao Y, Cao J, Jiang M. Enhanced catalysis of K2CO3 for steam gasification of coal char by using Ca(OH)2 in char preparation. Fuel 2010;89:310-7. [9] Sharma A, Takanohashi T, Saito I. Effect of catalyst addition on gasification reactivity of HyperCoal and coal with steam at 775-700 °C. Fuel 2008;87:2686-90. [10] Sharma A, Takanohashi T, Morishita K, Takarada T, Saito I. Low temperature catalytic steam gasification of HyperCoal to produce H2 and synthesis gas. Fuel. 2008;87:491-7. [11] Lehman RL, Gentry JS, Glumac NG. Thermal stability of potassium carbonate near its melting point. Thermochim Acta 1998;316:1-9. [12] Jaffri GR, Zhang JY. Catalytic gasification of Fujian anthracite in CO2 with black liquor by thermogravimetry. J Fuel Chem Technol 2007;35:129-35. [13] Wen CY. Noncatalytic heterogeneous solid fluid reaction models. Ind End Chem 1968;60:34-54. [14] Bhatia SK, Perlmutter DD. A random pore model for fluid-solid reactions. I. Isothermal, kinetic control. AIChE J 1980;26:335–379. [15] Kasaoka S, Sakata Y, Tong C. Kinetic evaluation of the reactivity of various coal chars for gasification with carbon dioxide in comparison with steam. Int Chem Eng 1985;25:160-75. [16] Wu Y, Wu S, Gao J. A study on the applicability of kinetic models for Shenfu coal char gasification with CO2 at elevated temperatures. Energies 2009;2:545-55. [17] Wen WY. Mechanisms of alkali metal catalysis in the gasification of coal, char, or graphite. Catal Rev 1980;22:1-28. [18] Wigmans T, Elfring R, Moulijn JA. On the mechanism of the potassium carbonate catalysed gasification of activated carbon: the influence of the catalyst concentration on the reactivity and selectivity at low steam pressures. Carbon 1983;21:1-12. [19] Spiro CL, McKee DW, Kosky PG, Lamby EJ. Observation of alkali catalyst particles during gasification of carbonaceous materials in CO2 and steam. Fuel 1984;63:686-91. [20] Formella K, Leonhardt P, Sulimma A, van Heek KH, Jüntgen H. Interaction of mineral matter in coal with potassium during gasification. Fuel 1986;65:1470-2.
10
Oviedo ICCS&T 2011. Extended Abstract
Prediction of Steam Reforming of the Simulated Coke Oven Gas with a Detailed Chemical Kinetic Model K. Norinaga1, R. Sato2, and J.-i. Hayashi1 1 Institute for Materials Chemistry and Engineering, Kyushu University, Kasuga, Fukuoka, 816-8580, Japan. 2 Division of Chemical Process Engineering, Graduate School of Engineering, Hokkaido University, Sapporo, 060-8628, Japan *Corresponding author: [email protected] (A/Prof. Norinaga) Abstract The detailed chemical kinetic modeling with a reaction mechanism consisting of thousands of elementary step like reactions has been successfully applied to predict combustion and pyrolysis characteristics of hydrocarbon fuels. This approach, however, has seldom been used for predicting steam reforming of aromatic hydrocarbons. In this study, the predictive capability of the existing detailed chemical kinetic model was critically evaluated for the thermal conversions of aromatic hydrocarbons in the presence of hydrogen and steam. Published experimental data of steam reforming of simulated coke oven gas (mixture of hydrogen, steam, and an aromatic hydrocarbon such as benzene, toluene, and naphthalene) in a tubular flow reactor (total pressure 160 kPa, 1073-1673 K, and residence time up to 2 s) were used for the kinetic model evaluations. Simulation using the kinetic model that consists of more than 200 chemical species and more than 2000 elementary step like reactions, and a plug flow reactor model precisely reproduced the experimental results for major products such as CO, CO2, CH4 as well as the conversions of source aromatic hydrocarbons. The coke yield is also fairly well predicted when assumed that computationally obtained total yields of polycyclic aromatic hydrocarbons and acetylene, both are the most potential coke precursors, correspond to the experimentally determined coke yields. Besides, influences of hydrogen partial pressure and residence time on the product distributions were also well predicted. 1. Introduction Volatile matter released at the primary stage of pyrolysis and gasification of solid fuels frequently contains tarry components, besides non-condensable gases such as H2, H2O, CO, CO2 and CH4. For using such product gases as synthesis gas for chemicals productions or fuel gas for gas engine or fuel cells, they should preferably be tar-free. This is commonly achieved by quenching to recover the condensable tar and further expensive gas treatment. An exhaustive reforming of tar into dry gas is thus highly desirable to reduce the cost for the separation. It can be found literatures dealing with thermal reforming of tar derived from pyrolysis of biomass1-4 and coal,5-7 as well as tar emitted from metallurgical coke ovens.8-15 Catalytic and non-catalytic approaches to reforming tar have been reported. Since deactivation of the catalyst by coking as well as sulfur/chlorine poisoning due to H2S/HCl contained in the product gas from solid fuels are likely to be unavoidable, catalytic reforming of the tar-containing gas from biomass and coal seems to be not at practical application phase still at fundamental study phase.16, 17 Non-catalytic methods are also being studied, aiming at a more robust reforming technology, and are already included some of the practically developing gasification processes such as the Carbo-V process developed by CHOHREN18 and a fluid bed gasifier coupled with high temperature thermal reformer developed by ENERKEM 19.
1
Oviedo ICCS&T 2011. Extended Abstract
An experimental study on the gas phase reforming characteristics of the model compounds of tar such as benzene, toluene, and naphthalene was reported.8 It should be important to interpret the observations in the reforming of the each aromatic hydrocarbons mechanistically toward deeper understanding of the chemistry and kinetics of the tar reforming. The detailed chemical kinetic approach to developing a reaction mechanism consisting of hundreds or thousands of elementary-like reaction steps is a promising method for elucidating an accurate description of the phenomena that occur in gas phase. To date, numerical simulations with detailed chemical kinetic models have been performed to predict the chemistry and kinetics of combustion,20 as well as the pyrolysis of hydrocarbons.21-25 However, few studies have reported on the steam reforming of aromatic hydrocarbons. In this study, numerical simulations for the steam reforming of aromatic hydrocarbons were conducted using a detailed chemical kinetic model. The kinetic model was evaluated critically by comparing the simulation results with the published experimental results by Jess 8 who made a systematic study on the kinetics of the thermal conversion of aromatic hydrocarbons, such as naphthalene, benzene, and toluene, in the presence of hydrogen and steam using a flow reactor. 2. Kinetic Model and Numerical Simulation A reaction mechanism for hydrocarbon combustion and polycyclic aromatic hydrocarbon growth developed by Richter and Howard26 was used as the primary basis for the kinetic model to simulate the steam reforming of aromatic hydrocarbons. The reaction mechanism consisted of 2216 reactions, including 257 chemical species from the smallest species of hydrogen radical to the largest molecule of coronene. This mechanism successfully predicted characteristics of the COG partial oxidation at a pilot scale test plant.15 The calculations were performed with the PLUG code in the DETCHEM program package (DETCHEMPLUG).27 Boundary conditions necessary for the calculations such as pressure, linear velocity, composition of the feed gas at the reactor inlet were determined from the experimental conditions reported.8 The reaction conditions used in the experiments are summarized in Table 1. Another required input such as a temperature profile along the reactor length was given as a polynomial function fit to the measured temperature profiles reported in the literature. Detailed description of the numerical simulation can also be found elsewhere.28 Table 1 Reaction Conditions for Thermal Conversions of Aromatic Hydrocarbons in a Tubular Flow Reactor a content, vol.% (rest nitrogen)
benzene toluene naphthale ne a
hydrocarbo n 0.6 0.6
hydrogen 40 40
steam 20 20
residence time at 1373 K, s 0.5 1
temperature at isothermal zone (TR), K 1073 - 1673 973 - 1673
0.5
0 – 48(48)
0-32(20)
0.3-2(0.5)
1073 - 1673
Standard conditions in brackets; total pressure 160 kPa
2
Oviedo ICCS&T 2011. Extended Abstract
3. Results and Discussion Figure 1 shows the wall temperature profile in which the quasi isothermal zone temperature (denoted as TR, reference temperature) is 1573 K used for the computations (upper) and computed mole fraction profiles of major components for the thermal conversion of naphthalene along the reactor length. Naphthalene starts to decompose and products such as CH4, CO, and CO2 start to form at the reactor length of 0.15 m at which temperature is around 1400 K. Naphthalene decomposes almost completely at around 0.3 m and little changes in any species’ concentrations occur beyond 0.4 m, most likely due to the temperature drop. Figure 2 compares the conversions of source aromatic hydrocarbons derived from the numerical simulation with those obtained by the experiments as a function of TR. The computationally predicted conversions were calculated based on the mole fraction of each feed hydrocarbon at the reactor outlet (length = 0.5 m). Excellent agreements are obtained for the conversions of these aromatic hydrocarbons. Benzene is a bit more refractory than naphthalene while toluene is most reactive. This trend is traced by the kinetic model perfectly. 1600
‐ , er 1200 u ta re 800 p 400 m et ‐, n o it ca rf el o m
0 0.6 H2
0.5 0.4 0.3 0.2
H2O
0.1 0.0 0.012
‐ 0.010 , n 0.008 o it ca 0.006 rf el 0.004 o 0.002 m 0.000 0.0
CH4
CO
naphthalene
CO2
0.1
0.2
0.3
0.4
0.5
distance from reactor inlet, m Figure 1. Temperature profile used for the numerical simulation (upper) and computationally obtained mole fraction profiles of major components (middle and lower) in thermal conversion of naphthalene in the presence of hydrogen and steam along the reactor flow direction. Velocity, temperature, and composition of the gas at reactor inlet are 0.086 m/s, 571 K, and naphthalene 0.5 vol%; H2 20 vol.%; H2O 48 vol% (rest nitrogen), respectively.
3
Oviedo ICCS&T 2011. Extended Abstract 100 80
toluene
, % 60 n io sr ev 40 n o c
naphthalene benzene
20 0 900
1100
1300
1500
1700
temperature (TR), K
Figure 2. Numerically (lines) and experimentally (symbols) derived conversions in thermal conversions of aromatic hydrocarbons (benzene, toluene, and naphthalene) in the presence of hydrogen and steam as a function of TR. Residence time 0.5 s (naphthalene and benzene) and 1.0 s (toluene) at TR = 1373 K; the experiments were conducted with a constant volume rate of the feed gas, therefore at higher and lower temperatures the residence time is slightly shorter and longer respectively.
The product distribution obtained at the steam reforming experiment of naphthalene, is shown in Figures 3. The computationally obtained product yields are also drawn for evaluating the kinetic model critically. For the naphthalene steam reforming, yields of CH4 and total yields of CO and CO2 are excellently reproduced by the present simulation. The yields of the other major products such as benzene and soot (plus condensed products) are also predicted well, although under-predictions are observed in benzene yields at around 1500 K and soot yields at temperature higher than 1500 K. Actually it is difficult to evaluate the soot yield based on the numerical simulation because the kinetic model considers chemical reactions and species in gas-phase. Here it is assumed that the total yield of thirty-three polycyclic aromatic hydrocarbons (PAHs), which are included in the kinetic model and are the most potential coke precursors, corresponds to the experimentally determined yields of coke. 100
‐% C , s ld iey t c u d o r p d n a e n el a h t h p a n la u id se r
50
80 60
40
naphthalene
30
40
20
20
10
0 30
0 20
20
CH4
15
CO+CO2
benzene
10 10
5
0 80 60
0 40
soot and condensed products
30
40
20
20
10
0 900
1100
1300
1500
1700
0 900
C2 hydrocarbons
1100
1300
1500
1700
temperature (TR), K Figure 3. Numerically (lines) and experimentally (symbols) derived product distributions in thermal conversion of naphthalene in the presence of hydrogen and steam as a function of TR.
4
Oviedo ICCS&T 2011. Extended Abstract
To understand why the kinetic model under-predict the yields of C2 hydrocarbons and organic cracking products, the yield of each component involved in the organic cracking products was examined. Figure 4 compares the yields of CH4, C2H2, C2H4, and C2H6 determined experimentally with those determined numerically for the benzene conversion. The yields of CH4 and C2H4 are excellently predicted and the C2H6 yield is fairly predicted, whereas the C2H2 yield is significantly over-predicted. This indicates that the gaps found in the yields of C2 hydrocarbons in the experiment of naphthalene steam reforming (Figure 3), and the gaps found in the yield of organic cracking products in the benzene steam reforming are attributed to the over-predicted C2H2 yield. C2H2 is one of the most potential coke precursors and easily form solid carbon deposited on the reactor wall29 and soot.30 In the flow reactor experiments, C2H2 should be consumed by these gas-solid reactions extensively and does not survive as much as predicted by the present numerical simulation with the chemical kinetic model which includes only homogeneous gas phase reactions. This augment is supported by the results demonstrated in Figure 5 which shows soot yields in the steam reforming of naphthalene (left) and benzene (right). Two types of predicted yields are drawn; the dashed line is obtained by assuming that soot yields correspond to the total yields of the PAHs while the solid line is obtained by assuming that the soot yields correspond to the C2H2 yields plus the total PAHs yields. The soot yields are significantly under-predicted when the total yields of PAHs account for the soot yields. But once the C2H2 yield is added to the total PAHs yields, the resulted predicted values are very similar with the experimental yields of soot. This result implies that the prediction of the soot yield is possible with the present kinetic model, which involves no carbon deposition chemistry though, when we assumed that the PAHs as well as C2H2 forms soot quickly and their yields correspond to the experimentally determined soot yields.
50 ‐% CH4 C ,s 40 tc 30 u d 20 ro p g 10 in kc 0 ar 20 c ci 15 C2H4 n ag r 10 o f o s 5 d le iy 0 900 1100
50
C2H2
40 30 20 10 0 5 4
C2H6
3 2 1 0 1300
1500
1700
900
1100
1300
1500
1700
temperature (TR), K Figure 4. Numerically (lines) and experimentally (symbols) derived yields of organic cracking products in thermal conversion of benzene in the presence of hydrogen and steam as a function of TR.
5
Oviedo ICCS&T 2011. Extended Abstract
30
80
naphthalene
benzene
% 60 ‐ C , ld ei 40 y t o o s 20 0 900
20
10
1100
1300
1500
0 1700 900
1100
1300
1500
1700
temperature (TR), K temperature (T R), K Figure 5. Numerically (lines) and experimentally (symbols) derived yields of soot in thermal conversions of naphthalene (right) and benzene (left) in the presence of hydrogen and steam as a function of TR. Dashed lines are obtained by assuming polycyclic aromatic hydrocarbons to be soot and condensed products, while solid lines are obtained by assuming acetylene besides polycyclic aromatic hydrocarbons to be soot and condensed products.
Figure 6 shows conversion and soot yield in thermal conversion of naphthalene at TR = 1373 K, and at varying residence time and hydrogen volume percent in the feed gas. Influence of hydrogen in inhibiting naphthalene conversion and soot formation is obvious and is well reproduced by the numerical simulation. The predictions are better when more than 12 % of hydrogen is included in the feed gas, while the model underpredicts naphthalene conversion and soot yield at the lower hydrogen volume percent. Ignorance of direct conversion of naphthalene into soot which is very likely favored at lower hydrogen partial pressure, would lead to the under-predictions of the naphthalene conversion as well as the soot formation, and thus contribute to the gaps between the experiments and the computations. Effect of temperature on the soot formation kinetics is examined in Figure 7. Experimental values for the soot yield at TR = 1473 K exhibit maxima at residence time around 0.8 s, and those at 1573 K also exhibit a maximum value at around 0.4 s. The decrease in the soot yield at longer residence time is attributable to the soot gasification by steam. These trends are well captured by the numerical simulation. Experimental values at 1673 K show no maximum, but the computations suggest that soot yield has a peak at around 0.1 s. These results imply that soot formation is inevitable and soot is a primary product regardless of the temperature in thermal conversion of naphthalene.
6
Oviedo ICCS&T 2011. Extended Abstract 100
f o n o is re v n o c
80
0 vol.% H2
% , e n el 60 a h t 40 h p a n
6 12
20
48
0 100
% ‐ C , d le iy t o o s
80
0 vol.% H2
60
6 40
12 48
20
0 0.0
0.2
0.4
0.6
0.8
1.0
residence time, s Figure 6. Influences of hydrogen and residence time on the conversion of naphthalene and the yields of soot. Reaction temperature 1373 K. Lines are numerically obtained, whereas symbols are experimental values. 100
1473 K
80 60 40 20 0 100
% ‐ C , ld ei y t o o s
1573 K
80 60 40 20 0 100
1673 K
80 60 40 20 0 0.0
0.5
1.0
1.5
2.0
residence time, s Figure 7. Yield of soot as a function of residence time and TR in thermal conversion of naphthalene: numerically predicted curves and experimental values (symbols).
4. Conclusions A detailed chemical kinetic modeling approach was first applied to simulate the thermal conversions of aromatic hydrocarbons including naphthalene, benzene and toluene, in the presence of hydrogen and steam. The critical evaluation of the kinetic model
7
Oviedo ICCS&T 2011. Extended Abstract
proposed by Richer and Howard26 was conducted by comparing the computationally obtained values with published experimental data of steam reforming of simulated coke oven gas in a tubular flow reactor.8 It was found that the numerical simulation using the kinetic model that consists of more than 200 chemical species and more than 2000 elementary step like reactions, and a plug flow reactor model precisely reproduced the experimental results for major products such as CO, CO2, CH4 as well as the conversions of source aromatic hydrocarbons. The soot yield is also fairly well predicted when assumed that the computationally obtained total yields of polycyclic aromatic hydrocarbons and acetylene, both are the most potential soot precursors, correspond to the experimentally determined soot yields. Besides, influences of hydrogen partial pressure and residence time on the product distributions were also well predicted. Acknowledgement. The authors are grateful to New Energy and Industrial Technology Development Organization (NEDO) for their financial supports to this study. References 1. Caballero, M. A.; Aznar, M. P.; Gil, J.; Martin, J. A.; Frances, E.; Corella, J. Ind. Eng.Chem. Res. 1997, 36 (12), 5227-5239. 2. Corella, J.; Toledo, J. M.; Aznar, M. P. Ind. Eng. Chem. Res. 2002, 41 (14), 33513356. 3. Corella, J.; Caballero, M. A.; Aznar, M. P.; Brage, C. Ind. Eng. Chem. Res. 2003, 42 (13), 3001-3011. 4. Hosokai, S.; Kishimoto, K.; Norinaga, K.; Li, C. Z.; Hayashi, J. I. Energy Fuels 2010, 24 (5), 2900-2909. 5. Hayashi, J. I.; Takahashi, H.; Iwatsuki, M.; Essaki, K.; Tsutsumi, A.; Chiba, T. Fuel 2000, 79 (3-4), 439-447. 6. Hayashi, J. I.; Iwatsuki, M.; Morishita, K.; Tsutsumi, A.; Li, C. Z.; Chiba, T. Fuel 2002, 81 (15), 1977-1987. 7. Matsuhara, T.; Hosokai, S.; Norinaga, K.; Matsuoka, K.; Li, C. Z.; Hayashi, J. I. Energy Fuels 24 (1), 76-83. 8. Jess, A. Fuel 1996, 75 (12), 1441-1448. 9. Miura, K.; Kawase, M.; Nakagawa, H.; Ashida, R.; Nakai, T.; Ishikawa, T. J. Chem. Eng. Jpn. 2003, 36 (7), 735-741. 10. Hashimoto, T.; Onozaki, M. J. Jpn. Inst. Energy 2006, 85 (5), 364-370. 11. Hongwei, C.; Yuwen, Z.; Xionggang, L.; Weizhong, D.; Qian, L. Energy Fuels 2009, 23 (1), 414-421. 12. Cheng, H.; Lu, X.; Liu, X.; Zhang, Y.; Ding, W. J. Natural Gas Chem. 2009, 18 (4), 467-473. 13. Cheng, H.; Lu, X.; Zhang, Y.; Ding, W. Energy Fuels 2009, 23 (6), 3119-3125. 14. Norinaga, K.; Hayashi, J. Energy Fuels 2010, 24, 165-172. 15. Norinaga, K.; Yatabe, H.; Matsuoka, M.; Hayashi, J. I. Ind. Eng. Chem. Res. 2010, 49 (21), 10565-10571. 16. Hepola, J.; Simell, P. Applied Cat. B-Environmental 1997, 14 (3-4), 305-321. 17. Tomishige, K.; Miyazawa, T.; Kimura, T.; Kunimori, K.; Koizumi, N.; Yamada, M. Applied Cat. B: Environmental 2005, 60 (3-4), 299-307. 18. Blades, T. Industrial Bioprocessing 2007, 29 (5), 9-10. 19. Chornet, E. In Converting non-homogeneous blomass residues into alcohols, 8th
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Oviedo ICCS&T 2011. Extended Abstract
World Congress of Chemical Engineering: Incorporating the 59th Canadian Chemical Engineering Conference and the 24th Interamerican Congress of Chemical Engineering, Montreal, QC, 2009; Montreal, QC, 2009. 20. Warnatz, J.; Maas, U.; Dibble, R. W., Combustion; 3rd Edition. Springer-Verlag: Heidelberg, New York, 2000. 21. Dean, A. M. J. Phys. Chem. 1990, 94 (4), 1432-1439. 22. Dagaut, P.; Cathonnet, M.; Boettner, J.-C. Int. J. Chem. Kinet. 1992, 24 (9), 813837. 23. Sheng, C. Y.; Dean, A. M. J. Phys. Chem. A 2004, 108 (17), 3772-3783. 24. Ziegler, I.; Fournet, R.; Marquaire, P. M. J. Anal. Applied Pyro. 2005, 73 (2), 212230. 25. Norinaga, K.; Deutschmann, O. Ind. Eng. Chem. Res. 2007, 46 (11), 3547-3557. 26. Richter, H.; Howard, J. B. Phys. Chem.Chem.Phys. 2002, 4 (11), 2038-2055. 27. Deutschmann, O.; Tischer, S.; Kleditzsch, S.; Janardhanan, V.; Correa, C.; Chatterjee, D.; Warnatz, J. DETCHEM V2.0, http://www.detchem.com. 28. Norinaga, K.; Janardhanan, V. M.; Deutschmann, O. Int. J. Chem. Kinet. 2008, 40 (4), 199-208. 29. Norinaga, K.; Hüttinger, K. J. Carbon 2003, 41 (8), 1509-1514. 30. Richter, H.; Howard, J. B. Prog. Energy Combust. Sci. 2000, 26 (4), 565-608.
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Oviedo ICCS&T 2011. Extended Abstract
Invention of Quantitative Method of Char and Soot to Clarify Soot Production and Reaction Behavior in Coal Gasification S. Umemoto*, S. Kajitani and S. Hara Energy Engineering Research Laboratory, Central Research Institute of Electric Power Industry (CRIEPI), 2-6-1 Nagasaka Yokosuka, Kanagawa 240-0196, Japan *Corresponding author: [email protected]
Abstract Coal gasification is one of the key technologies for utilization of coal. Coal gasification mainly consists of coal pyrolysis, char gasification and gas phase (including volatile matters) reactions. “Soot” is a fine solid carbon particle produced from volatile matter decomposition. Soot has lower gasification reactivity than char. However, there is no proper quantitative method of soot because char and soot are mixed together in the solid products from coal gasification, and it is difficult to identify soot in the mixture by the conventional methods. In this study, a novel quantitative method of char and soot utilizing a laser diffraction particle size analyzer was developed. Moreover, coal gasification experiments were performed using a pressurized drop tube furnace (PDTF). The yield of soot with low gasification reactivity did not decrease, while char was promptly consumed by CO2 gasification. Hence, the carbon conversion ratio increased initially, but was kept around 0.8 even at 1673 K.
1. Introduction Integrated coal gasification combined cycle (IGCC) plants have been developed worldwide to use coal more efficiently and cleanly. There are many types of gasifiers in the world. Most of them are oxygen-blown entrained flow type. In Japan, an air-blown entrained flow gasifier has been developed by CRIEPI and Mitsubishi Heavy Industries, Ltd [1]. 250 MW IGCC power plant using the air-blown gasifier was constructed and is in operation [2]. Furthermore, CRIEPI has proposed an “oxy-fuel IGCC system with CO2 recirculation for CO2 capture”, which includes an O2/CO2 blown gasifier [3]. In any type of a gasifier, coal gasification consists of coal pyrolysis, char gasification and gas phase (including volatile matters) reactions. Char gasification is believed to control the overall conversion rate. Therefore, a great number of papers concerning to char gasification have been published [4–6]. On the other hand, another type of solid carbon,
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1
Oviedo ICCS&T 2011. Extended Abstract
“soot” (coke) is produced from the volatile matter decomposition at high temperature above 1200 K. Besides, soot prepared from coal pyrolysis in inert gas has lower gasification reactivity than char [7, 8]. However, the amount of soot produced in coal gasification has never been quantified in any experiment, because char and soot are mixed together in the solid products from coal gasification and it is difficult to identify soot in the mixture by the conventional methods. In this study, a novel char and soot quantitative method was proposed. Moreover, coal gasification tests with CO2 utilizing a pressurized drop tube furnace (PDTF) were conducted to clarify the production and the gasification behavior of soot.
2. Experimental section Table 1 shows the properties of coals (NL coal from Australia, DT coal from China, MN coal from Indonesia and TN coal from Indonesia). Table 1 Ultimate and proximate analyses of coals used Ultimate analysis (wt%, d.a.f.)
Proximate analysis (wt%, db
C
H
N
O(diff.)
FC
VM
Ash
NL
82.5
5.1
1.4
11.0
57.4
28.7
13.8
DT
83.2
4.3
0.3
12.2
61.5
27.3
11.2
MN
81.1
5.7
2.0
11.2
51.0
40.7
8.4
TN
71.3
5.7
1.6
21.4
43.2
47.8
9.0
Coal
A drop tube furnace (DTF) shown in Figure 1 was used for coal pyrolysis tests and gasification tests. Coal pyrolysis tests were conducted using DTF without the sampling probe at ambient pressure and 1673 K to examine the char and soot quantitative method. Residence time was set at about 3 s. Coarse particles were trapped in the ash pot at bottom of furnace. On the other hand, fine particles were carried by gas and caught in the trapping bottle and the cartridge filter. The ash pot samples are almost char particles because char particles are courser than soot particles. The trapping bottle samples include both of char particles and soot particles. And the cartridge filter caught almost only soot particles. The ash pot samples were used as the char samples and the cartridge filter samples were used as the soot samples for the examination of the char and soot quantitative method.
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2
Oviedo ICCS&T 2011. Extended Abstract
Coal gasification tests with CO2 were also
N2
conducted using pressurized DTF with the
Particle feeder
sampling probe to clarify the production and reaction behavior of soot. The residence time
Reaction tube (I.D. 50 mm)
was controlled by traversing the sampling probe. The temperature was 1473 to 1673 K; the furnace pressure was 0.5 MPa; the partial pressure of CO2 was 0.05 MPa (N2 balance);
H2O
Heater Cartridge filter
N2 Air Ash pot
and total oxygen to carbon ratio (O/C
Sampling probe
[mol/mol]) was 1.4.
Trapping bottle Sampling filter
Figure 1 Schematic of the Drop Tube Furnace (DTF) 3. Results and Discussion 3.1 Char and soot quantitative method using a laser diffraction particle size analyzer Figure 2 shows the particle size distribution of the samples prepared by NL coal pyrolysis. The char sample, the soot sample and their mixtures are measured by a laser
Laser diffraction intensity [–]
cummulative passing [vol%]
Summation = index I
100 Soot 80
Char 20%
60 40
50% 40% 60%
Char
20 0 –1 10
80% 0
1
2
10 10 10 Particle size [μ m]
0.016 Char 80% 0.012 0.008
50% 0% (Soot)
0.004 0 0
3
10
Figure 2 Particle size distribution of
36
0.02
10 20 30 40 50 60 70 Detection element number [–]
Figure 3 Laser diffraction intensity used for determination of index I (NL coal)
char, soot and their mixture (NL coal) diffraction particle size analyzer (SALD-2200, Shimadzu). The distribution of the soot sample was completely different from the distribution of the char sample. However, the
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3
Oviedo ICCS&T 2011. Extended Abstract
distribution did not change proportionately with changes in the char mixing ratio. For example, the percentage of cumulative distribution under 10 μm reached 90% for the sample with 20% of the char mixing ratio, whereas reached only 30% for the sample with 40% of the char mixing ratio. Therefore, the raw data of the laser diffraction measurements were proposed to utilize for the soot quantitative method. Figure 3 shows the laser diffraction intensity data used for calculation of the particle size distribution shown in Figure 2. The detection elements which have the number under 36 did not detect any intensity for the soot sample (the char mixing ratio = 0%). The numbers of the detection elements mean the location of the element and the smaller number elements are located nearer to the center. In the laser diffraction particle size analyzer, the first peak of the laser diffraction intensity corresponds to the particle size. The peak for smaller particle shifts to the side where the detection element number is larger. The detection elements which have the number under 36 were not able to detect any Soot mixing ratio[%]
diffraction intensity for pure soot sample because the intensity too low. 100 80 60was 40 20 Thus, 0 the 0.5
index I, which was defined as the summation of normalized intensity of the detection index I [–]
NL coal 0.4 DTthe coal elements from number 0 to 36, was proposed to evaluate char and soot mixing ratio. MN coal Figure 4 shows the relationship between the 0.3 index I and char mixing ratio for the calibration curve 0.2 0.1 0 0
20 40 60 80 Char mixing ratio [%]
100
Figure 4 Calibration curve of index I
1
Carbon conversion, Yield [–]
Carbon conversion, Yield [–]
v.s. Char mixing ratio (1) 1473 K
0.8
Carbon conversion
0.6 0.4
Char
0.2 Soot 0 0
1
2 3 4 5 Residence time [s]
6
1
(2) 1673 K
0.8 Carbon conversion
0.6 0.4
Char Soot
0.2 0 0
1
2 3 4 5 Residence time [s]
6
Figure 5 Carbon conversion, char yield and soot yield obtained from CO2 gasification using PDTF (TN coal, initial CO2 0.05 MPa) (1) 1473 K (2) 1673 K
Submit before 31 May 2011 to [email protected]
4
Oviedo ICCS&T 2011. Extended Abstract
from NL coal, DT coal and MN coal. The index I increased monotonously with the char mixing ratio for any coals. Using the calibration curve drawn in Figure 4, the char and the soot mixing ratio in solid products of coal gasification can be quantified.
3.2 Results of the coal gasification test using DTF Figure 5 shows the carbon conversion, the char yield and the soot yield in TN coal gasification with CO2 at 1473 K and 1673 K. The carbon conversion increased as the temperature increased. However, the carbon conversion did not reach 1.0 and was still kept around 0.8 even at 1673 K and 3 seconds of residence time. The soot yield which was quantified by the proposed method was about 0.2 at any temperature. Furthermore, the soot yield did not decrease very much as residence time advance, while the char yield decreased by CO2 gasification. Consequently, the carbon conversion increased initially because of the char gasification but did not reach 1.0 because of the soot remaining.
4. Conclusions A novel char and soot quantitative method utilizing a laser diffraction particle size analyzer was developed. Coal gasification tests were performed using a pressurized drop tube furnace. The yield of soot with low reactivity did not decrease while char was promptly consumed by CO2 gasification. Hence, the carbon conversion ratio increased initially, but was kept around 0.8 even at 1673 K.
Acknowledgement. A part of the presented work was supported by New Energy and Industrial Technology Development Organization (NEDO) program “Innovative zero-emission coal gasification power generation project”, P08020.
References [1] Inumaru, J. A 2T/D pressurized two-stage entrained bed coal gasifier and test resutls. Proceedings of International Conference on Coal Science. 1989; 2: 297–300 [2] Watanabe T. Results and estimations of the 5,000 hour durability test at the Nakoso air blown IGCC plant (including other activities). presented at Gasification Technologies Conference 2010, Washington, D.C., 2010 [3] Oki Y, Inumaru J, Hara S, Kobayashi M, Watanabe H, Umemoto S, et al. Development of
Submit before 31 May 2011 to [email protected]
5
Oviedo ICCS&T 2011. Extended Abstract oxy-fuel IGCC system with CO2 recirculation for CO2 capture. Proc. Clearwater Coal Conference, 2011; 4: 1066–1073 [4] Miura K, Hashimoto K, Silveston P. Factors affecting the reactivity of coal chars during gasification, and indices representing reactivity. Fuel, 1989; 68: 1461–1475 [5] Kajitani S, Suzuki N, Ashizawa M, Hara S. CO2 gasification rate analysis of coal char in entrained flow coal gasifier. Fuel, 2006; 85: 163–169 [6] Roberts D, Harris D. Char gasification in mixtures of CO2 and H2O: Competition and inhibition. Fuel, 2007; 86: 2672–2678 [7] Miura K, Nakagawa H, Nakai S, Kajitani S. Analysis of gasification reaction of coke formed using a miniature tubing-bomb reactor and a pressurized drop tube furnace at high pressure and high temperature. Chemical Engineering Science, 2004; 59: 5261–5268. [8] Kajitani S, Nakagawa H, Miura K, Hara S. Study on coal pyrolysis property in pressurized entrained flow coal gasifier - Rapid pyrolysis property with pressurized drop tube furnace -. Proceedings of International Conference on Coal Science, Okinawa. 2005: 3D10
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6
NEW CANDLE PROTOTYPE FOR HOT GAS FILTRATION INDUSTRIAL APPLICATIONS M. Rodríguez-Galán1, M. Lupión1*, B. Alonso-Fariñas1, J. Martínez-Fernández2. 1
Department of Chemical and Environmental Engineering, ETS Ingenieros-University of Seville (Spain) 2 Departamento de Física de la Materia Condensada, Physics Faculty-University of Seville (Spain)
Abstract The improvement of gas cleaning technologies is crucial for the establishment of advanced clean power generation coal-based technologies such as Integrated Gasification Combined Cycle (IGCC) or Pressurized Fluidized Bed Combustion (PFBC) which need high performance of the syngas clean-up process. New materials and advanced operating strategies at higher temperatures that could give lower energy penalty are required to be developed. A large scale high temperature filtration pilot is currently in operation at the ETSI University of Seville, with financial support from the European Commission and the Spanish Ministry of the Environment. This pilot plant allows testing different filters and pulses cleaning strategies using real coal ash under an extensive range of operating conditions such as temperature and pressure. The aim of the on-going research is the evaluation of the alternatives for hot gas filtration technologies and the optimization of the operation and performance of the filtering elements. A new experimental campaign has been carried out to test a new type of silicon carbide candle. The prototype is fabricated from pyrolyzed wood and other materials as result of a novel environment friendly patented process (BioSiC®). This process provides good thermo-mechanical, chemical and structural stability in an extensive range of temperatures. During the testing campaign, main parameters for the characterization of the prototype have been studied such as filtration velocity, permeability, porosity, pressure drop across the filter, cleaning pulse interval, baseline pressure drop, filtration efficiency and durability of the filter. Optimal operating conditions and optimal pulse cleaning strategies for filter elements have been also determined. Additionally, a model to predict the behaviour of the elements under diverse operating conditions has been determined.
1
In general, the experimental results showed that the prototypes are suitable for industrial applications under the operating conditions indicated in this study, typical of those needed for hot gas cleaning of coal combustion and gasification flue gases. However, the analysis of the results shows possible improvements in the performance of the elements that should be faced in the next experimental campaign. This paper describes the main characteristics of the new material developed, and the results obtained from experimentation, as well as major conclusions extracted from the analysis. 1.
Introduction
Commercial ceramic filters type candles made of silicon carbide (SiC) are commonly used in industrial facilities for power generation such as coal IGCC Power Plants. Silicon carbide is a relatively new material in technological and industrial applications although it was discovered and manufactured a century ago. Nowadays it is used as a structural material in applications which require hardness, high temperature strength, high thermal conductivity, a low coefficient of thermal expansion and good wear and abrasion resistance [1]. Biomorphic silicon carbide (BioSiC®) is a SiC-based advanced ceramic material, fabricated through a cost effective and environmental friendly process that uses cellulose as its starting point. This material shows excellent thermo-mechanical performance, chemical and structural stability, in a wide interval of temperatures. It is possible to tailor the thermal and electrical properties of these materials to a wide range of values to satisfy device requirements [2]. This paper describes the main characteristics of the new material developed, and the results obtained from an extensive experimental campaign with the goal of characterizing the performance and behaviour of the prototypes developed. Optimal operating conditions and optimal pulse cleaning strategies for filtration applications at high temperature and pressure have been also determined
2.
Experimental section
A large scale high temperature filtration pilot is currently in operation at the ETSI University of Seville. This pilot plant allows testing different filters and pulses cleaning strategies using real coal ash under an extensive range of operating conditions such as 2 * Corresponding author: Tel.: +34954481181; Fax: +34 954461775; E-mail address: [email protected]
tempperature and d pressure. The aim of o the on-gooing researrch is the evaluation e o the of alternnatives for hot gas filtrration techn nologies andd the optimization of thhe operationn and perfoormance of the filteringg elements. The diagram off the pilot is i illustrated in Figuree 1. The faacility is described in detail elsew where [3-6]..
Figure 1: Baasic diagram m of the testt facility The objective of o the testinng campaign n presentedd in this papper is the evaluation e o the of s carbbide candle fabricated from pyroolyzed perfoormance off a new proototype of silicon woodd and otherr materials as result of o a novel environmen e nt friendly patented p prrocess (BioS SiC®). BioS SiC® is a ceellular silicoon carbide material m wiith hierarchical porositty obtained from reacttive infiltrattion of carrbon templaates throughh wood py yrolisis. In this t approaach, a woodd precursor is selectedd among thee many woood species available a coommerciallyy and convverted to carrbon by expposition to high h temperaatures (10000ºC) in an inert atmospphere. The resulting carbon tempplate, whichh retains thhe wood prrecursor's microstructu m ure, is then machined to t near net--shape and melt infiltraated with liiquid Si in vacuum at hightempperatures. Reaction R of Si S and C prroceeds by a solution-rreprecipitaioon mechanism to produuce SiC. Th he final matterial consissts of a SiC skeleton with w a microostructure cllosely resem mbling that of the woood precursorr, with poroosity often filled f residuual Si that can be remooved by acid leaching. Final propperties of thhe material, such as po orosity, poree size and ddistribution, can be taillored by selection of thhe wood precursor. 3
An appropriate selection of physical-chemical properties and geometries suitable for use as a material constituent of BioSiC® ceramic filters has been carried out. Table 1 shows the main characteristics of the BioSiC® material [7]. Table 1: Main characteristics of the BioSiC® material Mechanical Properties at room Temperature Density (g/cm3)
1.1 -2.6
Young modulus (GPa)
25-250
Compressive strength (MPa)
1500
Bending Strength (MPa)
430
Toughness (MPa·m1/2)
3
Dilatation coefficient (1/K)
3.5 · 10-6
Thermal Properties Anisotropic Thermal Conductivity
Variation up to 30%
Tailorable values in a wide range (W/mK)
50 - 120
Electrical Properties Anisotropic Electrical Resistivity
Ratio over 10
Tailorable values in a wide range (cm)
0.02-20
Adjustable dependence with temperature BioSiC® ceramic candles prototypes were designed and manufactured to be tested in the hot filtration pilot facility at ETSI of the University of Seville. Structural and geometric characteristics of prototypes were defined, as well as the conditioning of the filtration facility in order to accommodate the characteristics of the new prototypes. The dimensions of the prototypes are 18 mm outer diameter, 14 mm internal diameter, 225 mm total length with 10 mm neck for the fixation to the supporting plate which holds the elements. The number of filtering elements simultaneously tested was 36 with a total effective area of filtration of 0.43 m2.
4
Fig gure2: BioS SiC® prototyype for testiing in the filtration faciility 2.1
Characteriization testss
characterissation Folloowing the methodolog m gy proposedd in previouus testing campaigns, c tests were done in first placce with the aim of defi fining the test matrix annd the base case, that is, i the deterrmination of o the param meters that ccould influeence on thee performannce of the prototypes p such as thee maximum m filtration velocity, th he minimum m pressure drop acrosss the protottypes, or cleeaning pulse intervals [5][8]. 2.2
Operationnal tests
The aim of thee operationnal tests is the estimaation of thhe influencee of the crritical meters on th he performaance of the prototypes.. Critical paarameters sttudied connnected param to shhort term tesst are: (1) thhe pressure drop acrosss the filterinng element, (2) the cleaning pulsee interval defined d as the t frequen ncy of the cleaning c op peration or the period time betw ween conseccutive cleanning operatiions, (3) thee baseline pressure p drop, which is i the presssure drop im mmediately after cleanning and (4) filtration efficiency e of the elemeent. In relatiion to long g term, thhe durabilityy or deterrioration off the filtering elemennts is invesstigated.
3. Reesults and discussion d 3.1. C Characterization tests 5
Relevant parameters such as the value of the maximum filtration velocity rate compatible with the operation and the residual pressure drop of the “virgin” element among others were determined by means of characterization tests. In Figure 3, the evolution of the pressure drop with filtration velocity for virgin elements and clean gas is shown. It is observed a linear increase of the pressure drop with filtration velocity, as Darcy´s Law indicates [9] [10]. BioSiC Candles. CLEAN FILTER. (7 bar(g), 370 ºC) 2500
2000
Pressure drop (mmwc)
y = 24.555x 2 + 319.12x R² = 0.9981
1500
1000
500
0 2.50
3.00
3.50
4.00
4.50
5.00
Filtration velocity(cm/s)
Figure 3: Virgin element. Pressure drop vs. filtration velocity at 370 ºC The effect of the temperature has been also investigated. Figure 4 shows the evolution of the pressure drop across the prototypes at two levels of temperature, 235 and 370 ºC. It is clear than higher temperatures imply higher values of pressure drop. 2500
Pressure Drop (mm wc)
2000
1500 235 370 1000
500
0 2.50
3.00
3.50 Filtration velocity(cm/s)
4.00
4.50
Figure 4: Virgin element. Pressure drop vs. filtration velocity at two levels of temperature When loaded gas is injected, the operation becomes non-viable at gas velocities above 3cm/s as it is illustrated in Figure 5. During unstable operating conditions there is a rapid increase in the maximum pressure drop as well as a reduction in time between 6
pulses. On the contrary, Figure 6 shows an example of stable operating conditions, where the pressure drop after a cleaning pulse remains at similar levels. 2000
Pressure Drop (mm wc)
1800
1600
1400 Particle concentration: 13 g/Nm3 Filtration velocity: 3.3 cm/s Gas cleaning pressure: 16±0.2 barg
1200
1000
0
500
1000
1500
2000
2500
Time (s)
Figure 5: Unstable operation with gas velocity >3 cm/s 2200
Pressure Drop (mm wc)
2000 1800 1600 1400 Particle concentration: Particle concentration:1313 g/Nm3 g/Nm3 Filtration velocity: Filtration velocity:3.3 2.8cm/s cm/s Gas cleaning Gas cleaningpressure: pressure:16±0.2 16±0.2barg barg
1200 1000 0
500
1000
1500
2000
2500
3000
3500
4000
Time(s)
Figure 6: Stable operation with gas velocity <3 cm/s Table 2 summarises the experimental test matrix obtained as result of the analysis of the characterization tests. Table 2: Experimental test matrix Operational parameter
Level
Filtration velocity (cm/s)
2.7 – 3.0
Particle loading (mg/Nm3)
11,000-13,000-17,000
Pressure in the vessel (barg)
7
T (ºC)
235-550
∆Pmax (mmwc)
1,800 – 1,950
Pulse cleaning pressure (barg)
14 – 16 – 19
Duration of the pulse (ms)
700
7
The temperature range was selected from 235 ºC to 550 ºC in order to compare the prototypes tested with other commercial filtering elements already tested in the same facility. This way, 235ºC is the operating temperature of the dedusting system using Dia-Schumalith candles in ELCOGAS IGCC Power Station; 370 ºC is the maximum operating of 3M bag filters; and 550 ºC is the maximum temperature achievable by the filtration facility [5][6][8]. 3.2 Filtering element performance In order to determine the performance and behaviour of the filtering elements tested, the efficiency has been considered as one of the main parameters to be taken into account during the experimentation. To establish the efficiency, measurements of the particle concentration at the inlet and the outlet of the filtration vessel were made by using the isokinetic sampling system at high temperature and high pressure developed by the Chemical and Environmental Engineering Department [5]. In general, it is observed that the concentration of particles at the inlet of the vessel is in a range between 11,000 and 16,000 mg/Nm3 while the outlet is around 200 mg/Nm3 in all the cases. According to these data and applying the expression given above, the filtration performance is >98 %. Operational parameters and their influence on the performance of the filtering elements have been also identified and analyzed. In particular, the influence of filtration velocity, permeability, porosity, pressure drop across the filter, cleaning pulse interval, baseline pressure drop, filtration efficiency and durability of the filter has been investigated. Optimal operating conditions and optimal pulse cleaning strategies for filter elements have been determined. Figure 7 shows the effect of the cleaning pressure in the evolution of the pressure drop across the prototypes. The graphs show the evolution of the pressure drop for three different values of cleaning pressure: 14 barg, 16 barg and 19 barg.
8
Pressure Drop (mm wc)
2100
1900
1700
1500
Particle concentration: 16.7 g/Nm3 Filtration velocity: 2.7 cm/s Gas cleaning pressure: 14±0.2 barg
1300
1100 0
500
1000
1500
2000
2500
3000
Time (s)
Pressure Drop (mmwc)
2100
1900
1700
1500 Particle concentration: 16.7 g/Nm3 Filtration velocity: 2.7 cm/s Gas cleaning pressure: 16±0.2 barg
1300
1100 0
500
1000
1500
2000
2500
3000
3500
4000
Pressure Drop (mm wc)
Time(s)
Particle concentration: 16.7 g/Nm3 Filtration velocity: 2.7 cm/s Gas cleaning pressure: 19±0.2 barg
2100 1900 1700 1500 1300 1100 0
500
1000
1500
2000
2500
3000
3500
4000
4500
Time (s)
Figure 7: Effect of pressure pulse cleaning of the pressure drop across the prototypes. Cleaning pressure: 14, 16, 19 barg
It is shown that the pressure drop after cleaning is slightly lower at higher cleaning pressures, that is, the interval between cleaning cycles is higher at higher cleaning pressure. While the baseline pressure drop after 19 barg pulses is about 1300 mmwc, this value is around 1400 mmwc and 1500 mmwc when cleaning at 14 and 16 barg respectively. Therefore, it can be concluded that the cleaning pressure has a positive influence in the operation. It must be pointed out though that the cleaning operation at higher pressure results in a faster deterioration of the filtering elements, and higher
9
OPEX (increase of nitrogen consumption). In this sense, it s also noted that cleaning pressures above 19 barg showed no significant benefit. 3.3 Maximum Pressure Drop Tests at different values of maximum pressure drop were carried out. Figure 8 compares two values of maximum pressure drop, 1800 and 1950 mmwc; the rest of operational parameters remaining constant.
Pressure Drop (mm wc)
1900
1700
1500
Particle concentration: 13 g/Nm3 Filtration velocity: 2.9 cm/s Gas cleaning pressure: 19±0.2 barg Max. Pressure Drop: 1800/1730 mm wc
1300
1100 0
250
500
750
1000
1250
1500
1750
2000
2250
2500
2750
Time (s)
Pressure Drop (mm wc)
2500
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1000 Particle concentration: 13 g/Nm3 Filtration velocity: 2.9 cm/s Gas cleaning pressure: 19±0.2 barg Max. Pressure Drop: 1950/1860 mm wc
500
0 0
500
1000
1500
2000
2500
3000
3500
4000
Time (s)
Figure 8: Effect of the cleaning by Maximum Pressure Drop
It is illustrated that an increase of the maximum pressure drop implies more separation between cleaning cycles, and therefore better performance of the filtering elements. In this regards, an increase of 15% in the maximum pressure drop resulted in 7 times less frequency cleaning cycles. Nevertheless, higher values of maximum pressure drop result in a faster deterioration of the filtering elements. 3.4 Baseline pressure drop The influence of the operation time was observed by analyzing the evolution of the baseline. In this regard, Figure 9 shows how the pressure drop increases with time. This
10
is due to the gradual saturation of the filters, and also affects to the cleaning parameters, since higher cleaning pressures are required. 1900 1800
Pressure Drop (mm wc)
1700 1600 1500 BASELINE 1400 1300
Particle concentration: 11.7 g/Nm3 Filtration velocity: 3.0 cm/s Gas cleaning pressure: 16±0.2 barg
1200 1100 0
1000
2000
3000
4000
5000
6000
7000
8000
9000
10000 11000 12000 13000 14000 15000
Time (s)
Figure 9: Evolution of the baseline pressure drop with time
3.5. Modelling Experimental data obtained during the testing have been used for the developing of a model based on the Darcy’s law in order to predict the behaviour of the filtering element during the filtration process. General approaches pursued by most researchers on the flow of fluids through porous medium started with the Darcy’s law. As the Reynolds numbers for flow through the filter and cake is very small, the differential pressure drop can be described by the well known Darcy’s law: dP KU dz
(1)
Following the methodology proposed in [11], the total pressure drop, ∆P, comprises the pressure drop across the cake and that across the filter medium. Hence the total pressure drop can be expressed as:
P Pm Pc
(2)
Therefore the following relationship can be established by integration of eq. (1) [10-12]:
P K m z mU K c z c U
(3)
The solid balance in one single element shows the increase of the cake as a measurement of the thickness, and it is shown in the expression: zc
gUC t p (1 )
(4)
11
The term of the thickness is replaced in the general expression of the pressure drop through the cake, as follows:
Pc k c
g p (1 )
CU 2 t
(5)
The general expression of pressure drop in the filtration process can be written in a more simplified form using a coefficient b:
PT k m U bCU 2 t
(6)
The model is now being processed and will be available to be presented at the ICCST Conference.
4. Conclussions The analysis of experimental results, and based on the described in the preceding paragraphs, it can be drawn the following conclusions: In relation to the performance of filtration, the prototypes have an average yield around 98% and an outlet particle concentration around 200 mg/Nm3. This value of particle matter concentration is relatively high, and may be due to the fact that the prototypes are operating at high values of filtration velocity. It is expected that scale commercial filters with improved design to present better efficiency. With respect to the filtration rate, it has been observed that the real capabilities of the prototypes are generally higher than those specified for commercial filter elements. It is noted that the pressure drop across the prototypes elements is an order of magnitude bigger than across commercial filters. The dependence of cleaning performance, the residual pressure drop and the separation of the cleaning cycles with the next variables have been determined: -
Cleaning pressure
-
Maximum pressure drop
-
Pulse cleaning duration
-
Filtration velocity
-
Time
In this sense, the values of effective cleaning pressure have been set around 16 barg. It has been reached maximum pressure drop values up to 1950 mmwc without having seen any deterioration in the prototypes.
12
The study of the pulse cleaning duration concludes that the influence of the parameter is not significant if the pulse is short (500 – 800 ms). The influence of the filtration velocity on cleaning parameters has been investigated. An increase in gas filtration rate implies a reduction in cleaning performance and an increase in the pressure drop and in the fouling velocity. The analysis of the baseline pressure drop indicates that cleaning requirements increase as the experiment progress which implies a decrease in the separation of cleaning cycles. It has been found that, at high temperature (550 ºC), expansion, contraction or significant deformation of the material have been not observed.
Acknowledges This work is carried out with the financial support of the Spanish Innovation and Science Ministry and the European Commission (former ECSC Research program). References [1] Presas M, Pastor JY., Llorca J, Arellano López AR, Martínez Fernández J, Sepúlveda R. Microstructure and fracture properties of biomorphic SiC. Refractory metals & Hard Materials, July 2005. [2] Lupion M, Rodriguez-Galan M, Barbosa V, Cruz JL. Operating experiences of new filtering materials for syngas filtration at high temperature and high pressure.4th International Freiberg Conference on IGCC & XtL Technologies. Dresden (Germany). 2010. [3] Navarrete B, Lupión M, Gutiérrez F J, Cortés V J, Coca P, García Peña F. Improving the ELCOGAS IGCC dedusting system: facility plant erection and testing. International Freiberg Conference on IGCC & XtL Technologies 2005. [4] Lupión M, Navarrete B, Gutiérrez FJ, Cortés VJ. Design and operation experiences of a hot gas filtration test facility for IGCC power generation. Advanced Gas Cleaning Technology. Proceedings of the 6th International Symposium on Gas Cleaning at High Temperatures 2005. ISBN4 915245-61-6. [5] ECSC Project 7220-PR-141 Technological Improvement of Hot Gas Filtration for Onstream IGCC Plants in the European Union (GASFIL). Final Report 2007. [6] Lupión M, Gutiérrez FJ, Navarrete B, Cortés VJ. Assessment performance of hightemperature filtering elements. Fuel, 2010. Volume 89, Issue 4, Pages 848-854 13
[7] Martínez-Fernández J, Qispe J, Barbosa JV. Biomorphic Ceramics Materials for High Temperature and pressure industrial filtration Processes. 4th International Freiberg Conference on IGCC & XtL Technologies. Dresden (Germany). 2010. [8] Lupion M, Navarrete B, Alonso-Fariñas B, Rodriguez-Galan M. Hot gas filters for coal-based power generation systems: Operating Experiences. International Symposium on Gas Cleaning at High temperature (GCHT-8). Taiyuan, Shanxi (China). August 2010. [9] Seville, J P K; Clift R; Withers, C J; Keidel, W. Rigid Ceramic Media for filtering hot gases. Filtration & Separation, 1989. Pages 265-271. [10] Seville, J P K; Clift, R. Gas cleaning in demanding applications. 1997. J P K Seville (Ed), Blackie Academic & Professional, London, UK Seville. [11] Lupión M., Alonso-Fariñas B., Rodríguez-Galán M, Navarrete B. Modelling of the performance of filtering elements at high temperature and high pressure. 4th International Congress on Energy and Environmental Engineering and Management. Mérida (SPAIN). May 2011. ISBN13: 978-84-9978-014-6. [12] Duo, W; Seville, J P K; Kirkby, N F; Büchele, H; Cheung, C K. Patchy cleaning of rigid gas filters-II. Experiments and model validation.1997.Chemical Engineering Science, Volume 52, Issue 1, Pages 53-164.
14
FLUID DYNAMIC SIMULATION OF DRY FILTER FOR REMOVAL OF PARTICULATES FROM COAL AND BIOMASS GASIFICATION C. B. da PORCIÚNCULA1, N. R. MARCILIO1, M. GODINHO2 e A.R.SECCHI3 1
Dequi - Federal University of Rio Grande do Sul - Brazil e-mail: {cleiton,nilson}@enq.ufrgs.br 2 University of Caxias do Sul - RS e-mail:[email protected] 3 COPPE/PEQ – Federal University of Rio de Janeiro - RJ e-mail:[email protected]
ABSTRACT – The presence of particulates and other solid contaminants is very common in gasification and combustion processes. These by-products consist essentially of tar, fly ash and char, which may cause severe problems of corrosion and deposit in turbines, heat exchangers and other process equipments. The retention of such particulates in a bench filter comprised of a spheres glass bed has been studied in order to remove those particulates prior the process equipments cited above, as well as avoiding pollutant emissions to the atmosphere. In a first step, computation fluid dynamic simulations (CFD) were applied to this system in steady-state and transient conditions. The original filter design has two conical sections with the filtering bed between them, and its dimensions are to be adapted in a gasification and combustion pilot plant. Different modeling approaches were tested (Lagrangian, Eulerian, and Lagrangian-Eulerian), and the previous results show a pressure drop fourfold its initial value. The Darcy model for flow in porous media was employed in this modeling, and the permeability of the filter bed was estimated by semi-empirical correlations. Two different values of dimensions for the porous media were tested: 50 and 100 mm. Laboratory tests in a minor scale have been carried out in order to estimate the properties of the porous media and also evaluate the rise of pressure drop along the time. KEY WORDS: dry filter; particulate; fluid dynamic simulation, porous media.
1. Introduction One of the major drawbacks encountered in gasification and combustion processes is the presence of solid contaminants. Tar, ash, and other particulates cause severe problems in equipments upstream like turbines, combustion chambers, heat exchangers and boilers. The particulates greater than or equal to 5 μm diameter are most difficult to be removed, which can bring about problems concerning air pollution, incrustation and corrosion. The technology of dry filters is very promising in the sense of retaining such particulates. Yang and Zhou (2007), Hassler and Nussbaumer (1999) cite that sand filters might remove up to 99.9% of particulates of a process stream. Stanghelle et al. (2007) studied the granular filtration of biomass gasification by-products at temperatures around 550oC, and their results show high removal efficiencies combined with low pressure drop when the material to be filtered is fly ash, by utilizing a filter medium of aluminum oxide spheres. Neiva and Goldstein (2002) compared the pressure drop with classical literature models (Darcy, Karman-Kozeny and Happel laws) in a ceramic fiber bed, where the particulates have been arisen from char of a coal gasifier.
In terms of mathematical modeling and simulation of filtration processes, one may highlight the works of Deuschle et al. (2008), a computational fluid dynamics study of a diesel filter, Dittler and Kasper (1999), by simulation of a two-dimensional filter to prevent pressure drop among different regeneration efficiencies, Moghadasi e al. (2004) by means of a mixed filter media comprised of a compacted sand and glass sphere bed. The objective of this work is to provide a full CFD analysis of pressure drop in a pilot-plant filter still to be constructed. Initially, the simulations were carried out with estimated values of permeability from empirical correlations. Further on, other simulations were developed based on experimental values of permeability measured from a bench scale plant, and both results were compared. 2. Mathematical Modeling The governing equations are the Navier-Stokes (1) in a three-dimensional domain. The turbulence model k-ε was employed coupled with those equations (2 and 3) together with the continuity equation (4) as described below:
(
)
∂ ( ρu i ) ∂ ρu j u i ∂P ∂ + =− + ∂t ∂x j ∂xi ∂x j
⎛ ∂ui ⎜μ ⎜ ∂x j ⎝
⎡⎛ μ ∂ ( ρk ) + ∇ ⋅ (ρu i k ) = ∇ ⋅ ⎢⎜⎜ μ + T ∂t σk ⎢⎣⎝
⎞ ⎤ ⎟⎟∇k ⎥ + Pk − ρε ⎠ ⎥⎦
⎡⎛ μ ∂ (ρε ) + ∇ ⋅ (ρu i ε ) = ∇ ⋅ ⎢⎜⎜ μ + T σε ∂t ⎣⎢⎝
⎞ ⎟ + S ui ⎟ ⎠
⎞ ⎤ ε ⎟⎟∇ε ⎥ + (Cε 1 Pk − Cε 2 ρε ) ⎠ ⎦⎥ k
( )
∂ρ ∂ ρu j + =0 ∂t ∂x j
(1)
(2) (3)
(4)
Lagrangian transport particle implementation is written in terms of the second Newton law, as equation (5):
mp
dU p dt
=
(
CD ρ f Ap U f − U p U f − U p 2
)
+
πd p 2 (ρ p − ρ f )g 6
(5)
If considering an Eulerian-Lagrangian approach, one considers the diffusion term of the conservation equation per component:
(
)
∂Ci ∂ ρu j Ci ∂ = + ∂x j ∂x j ∂t
⎛ ∂Ci ⎜D ⎜ ∂x j ⎝
⎞ ⎟ ⎟ ⎠
(6)
Steady-state and transient simulations were performed for three cases: Lagrangiang track of particles, Eulerian-Lagrangian and Eulerian-Eulerian. The Darcy law was implemented in order to account the pressure drop along the filter bed:
U=
1 dV ΔP = μ Rt + R f A dt
(
)
(7)
Where the resistance of the cake, Rt , is a linear function of time:
Rt = αρUφt
(8)
3. Materials and Methods A bench scale filter was constructed in order to estimate resistance and permeability of the porous medium that could be applied to the pilot-plant filter. Simulations taking into account different bed widths (50 and 100 mm length) and also a mixed medium with properties of sands with different mean diameters (25 mm long each) were carried out by applying simulations in steady state and transient conditions to verify the behaviour of the pressure drop with time. The scheme of the filter is depicted in Figure 1 below:
Figure 1 - Scheme of the dry filter 4. Preliminary Results. The results of the steady-state simulations are shown in Table 1 below:
Table 1 - Results of steady state simulations Bed length (mm) 100 50 25 mm mixed – thicker sand 25 mm mixed – finer sand
Pressure drop (mmH2O) 93.2 83.1
Permeabilidade (m2) 4.4.10-9 4.4.10-9
0.4 0.4
Bed particle diameter (mm) 2 2
39.2
4.4.10-9
0.4
2
499.5
3.41.10-10
0.3
1
Porosity
From the Table 1 above, one can see that the larger values of pressure drop were obtained for the 100 mm bed and for the section of 25 mm of the mixed bed for finer sand. A similar test is described in the literature by Yang and Zhou (2007), where one might conclude that the finer are the particles, the larger is the pressure drop but the larger is the retention. Thus, there should be a trade-off between bed particle size and removal efficiency of the filter in order that the equipment may work out with reasonable pressure drops and high removal efficiency. For the transient regime, the results are illustrated in Figure 2 below:
Figure 2 - Pressure drop in transient simulation From Figure 2, one can see that there is no practical difference in any of the three mathematical approaches. All of them reproduce the same result for pressure drop in different times, so that we can conclude these three approaches are adequate. In Figure 3 it is possible to view an example of gas streamlines being distributed inside the filter chamber, for the geometry with 100 mm bed length.
Figure 3 - Gas streamlines in 100 mm bed The vortexes that are visualized in Figure 3 before the gas entering the bed is a picture of what might happen in practice: the deposit of particulates cake has a major probability to deposit on this areas. 5. Concluding remarks Based upon the results that were obtained by simulations, it is possible to predict a fourfold increase in pressure drop if considering a total cycle of operation about 10 hours. This result is the same one obtained for the three mathematical approaches that were tested.
Acknowledgments Authors are very grateful to RNC (Brazil Coal National http://www.ufrgs.br/rede_carvao) by the financial support for this research.
Network,
References
DEUSCHLE, T.; JANOSKE, U.; PIESCHE, M. A CFD model describing filtration, regeneration and deposit rearrangement effects in gas filter systems. Chemical Engineering Journal, v.135, p.49-55, 2008. DITTLER, A.; KASPER, G. Simulation of operational behaviour of patchily regenerated, rigid gas cleaning filter media. Chemical Engineering and Processing, v.38, p.321-327, 1999. HASLER P.; NUSSBAUMER TH. Gas Cleaning for IC Engine Applications from Fixed Bed Biomass Gasification. Biomass and Bioenergy, v.16, p.385-395, 1999. MOGHADASI, J.; STEINHAGEN, H.M.; JAMIALAHMADI M.; SHARIF, A. Theoretical and experimental study of particle movement and deposition in porous media during water injection. Journal of Petroleum Science & Engineering, v.43, p.163-181, 2004. NEIVA, A.C.B.; GOLDSTEIN, L. A procedure for calculating pressure drop during the build-up of dust filter cakes. Chemical Engineering and Processing, v.42, p.495-501, 2003. STANGHELLE, D.; SLUNGAARD, T.; SØNJU, O.K. Granular bed filtration of high temperature biomass gasification gas. Journal of Hazardous Materials, v.144, p.668-672, 2007. YANG G.; ZHOU J. Experimental Study on a New Dual-Layer Granular Bed Filter for Removing Particulates. Journal of China University of Mining and Technology, v.17, p. 201204, 2007.
Oviedo ICCS&T 2011. Extended Abstract
Capture of CO2 during low temperature biomass combustion in a fluidized bed using CaO. A new larger scale experimental facility J.R. Chamberlain, C. Perez Ros Gas Natural Fenosa, Avenida San Luis 77, 28033, Madrid, Spain Abstract
This paper outlines a new experimental test facility of 300kWt being commissioned in the grounds of Gas Natural Fenosa’s La Robla coal-fired power plant in the Leon region, Northwest of Spain, with the goal to advance the demonstration of the capture of CO2 with CaO in a circulating fluidized bed (CFB) combustor-carbonator reactor, where the combustion of biomass with air occurs simultaneously with the carbonation of CaO, thereby capturing the CO2 released from the combustion process. This process intends to exploit the high reactivity of most natural biomasses permitting the possibility of combustion at low temperatures (around 700ºC) and the capability of CaO to absorb CO2 at these temperatures. This is a niche application for the carbonate looping cycles, which is currently being developed for other post-combustion and pre-combustion processes. Previous results obtained in a 30kWt test facility made up of two interconnected CFB reactors (combustor-carbonator and combustor-calciner) located at the facilities of the Spanish Institute of Coal (INCAR-CSIC) in Oviedo, Spain, have demonstrated the experimental feasibility of in situ CO2 capture during a “lowtemperature” when operating at around 700ºC that maximizes both combustion and CO2 capture efficiencies in circulating fluidized beds fed with a continuous supply of CaO. CO2 capture efficiencies of over 80% have been obtained, remarkably close to those allowed by the equilibrium and the combustion mass balances, when an adequate stock of active CaO in the combustor-carbonator reactor, combined with an intense circulation of solids between this carbonator and calciner have been achieved. These positive results have justified the construction of the new larger scale test facility in La Robla, some 10 times larger, which is described in this abstract. Over the next year this facility should permit the validation of the smaller scale experimental work and provide crucial experimental results from longer duration experiments to further verify the process and provide data for models that need to be developed for the next scale up of the concept, both as a standalone process and as a possible co-combustion concept when integrated
1
Oviedo ICCS&T 2011. Extended Abstract
with an existing thermal power plant.
1. Introduction
In order to keep global warming below 2ºC, it is foreseen by the International Energy Agency (IEA) that Carbon Capture and Storage (CCS) must provide 20% of the global CO2 cuts required by 2050; the costs of doing so without CCS will be over 70% higher [1]. The integration of CCS with biomass will lead to CO2 “negative emissions” in the generation of electricity or in other energy products, which is a very attractive concept. This was initially recognized by Ishitani and Johansson [2] and the positive implication of negative emissions in energy technologies for long term climate change mitigation has been highlighted in many recent scenario exercises [3, 4].
The capture of CO2 in the niche application of the carbonate looping cycle of this work with CaO in a circulating fluidized bed (CFB) combustor-carbonator reactor, where the combustion of biomass with air and the carbonation of CaO takes place simultaneously has been presented before [5, 6]. Results were obtained in a small scale experimental 30kWt located at the facilities of the Spanish Institute of Coal (INCAR-CSIC) in Oviedo, Spain and many tests have been carried out at different carbonation reaction temperatures, solids inventories, superficial gas velocities, solids circulation flow rates and concentrations of CO2 generated by biomass combustion. Three different biomasses were tested: saw-dust, crushed olive pits and wood pellets and experiments up to 14 hours at steady state conditions were achieved. An example of previously published results [6] is shown in Figure 1.
2
Oviedo ICCS&T 2011. Extended Abstract
100
25
(a)
O2 analyzer
20
Capture Efficiency (%)
Concentration (% vol.)
CO2 O2 probe 15
10
5
80
60
40
20
(b)
Experimental Equilibrium
0 19:10
19:20
19:30
19:40
19:50
time (h:min)
20:00
20:10
0 19:10
19:20
19:30
19:40
19:50
20:00
20:10
time (h:min)
Figure 1. Example of experimental results in the combustor-carbonator reactor in a typical experiment. (a) Combustor-carbonator exit gas concentrations of CO2 and O2 measured by the on-line gas analyzer and O2 zirconia probe (b) Experimental capture efficiency and maximum capture efficiency allowed by equilibrium.
In this example, the average CO2 concentration at the exit of the combustor-carbonator reactor was 3.1 vol%. The average oxygen concentration at the exit of the combustorcarbonator was 7.6 vol.%. From a combustion mass balance, the CO2 produced by biomass combustion was estimated to be around 13.6 vol.% , therefore, the average CO2 capture efficiency was 77%. The average temperature in the combustor-carbonator was 690 ºC during the period, which according to the equilibrium of CO2 on CaO allows for a 2.4 vol% of CO2 (maximum efficiency allowed by the equilibrium of 81%, remarkably close to the experimental value).
These positive results have led to a decision by Gas Natural Fenosa to scale-up the process and design and construct a new larger test facility in La Robla.
2. Scale-Up of Test Facilities
As a result of these positive results from the 30kWt test facility, Gas Natural Fenosa took the decision to construct the new larger test facility, some 10 times larger at a nominal capacity of 300kWt, to advance and further demonstrate this niche concept.
3
Oviedo ICCS&T 2011. Extended Abstract
The smaller 30kWt test facility located in Oviedo consists of two interconnected CFB reactors: a 6.5m high carbonator and a 6m high air-fired calciner (it was assumed that the rate of sorbent reaction is similar that when operating the calciner in oxy-fired mode). Both reactors had an internal diameter of 0.1 m.
The new test facility located in the grounds of the Gas Natural Fenosa’s La Robla coal fired power plant also has a central part consisting of two interconnected CFB reactors. However, in this case the carbonator is a cylindrical reactor of some 12 metre high with a diameter of 600mm. In order to control the temperature of the biomass combustion in the range of 650-700ºC a heat exchanger has been installed in the carbonator that employs thermal oil as the cooling medium. The calciner has the same dimension as the carbonator, 12m high with a diameter of 600mm. This reactor is ceramic lined as it operates in the temperature range of 850-900ºC. As with the smaller experimental plant, the calciner is air fired. It is again assumed that the rate of sorbent reaction will be similar to that when operating the calciner in oxy-fired mode and the novelty of this concept that has to be demonstrated is the process of combustion and capture in the carbonator.
Figure 2. Images of the 300kWt experimental plant in La Robla power plant actually in the stages of commissioning.
As one of the objectives of this larger test facility is to undertake experiments of longer duration the power plant includes automated systems for biomass handling and injection via screw-feeds. Biomass storage for one week experimentation is contemplated, sufficient to provide both the nominal 1.7T/day of biomass to the carbonator and 2,5T/day of biomass to the calciner. Limestone storage and injection is also contemplated, capable of injecting over 0,5T/day for make-up if required so depending 4
Oviedo ICCS&T 2011. Extended Abstract
on sorbent deactivation and breakdown. There is obviously the capability to receive further supplies of biomass and limestone during plant operation for longer experimental runs. Air is preheated and supplied to independently to both the carbonator and calciner reactors in a controlled manner through variable speed forced draft fans. As biomass and air are injected at the bottom carbonator, which will contain a bed of predominantly CaO and if the temperature of combustion can be controlled to temperatures approaching 700ºC, via fuel supply and coolant flow control, combustion of the biomass and capture of CO2 through the reaction with the CaO to produce CaCO3 should occur simultaneously. The mixture of gases and solids that leave the top of the carbonator pass through two cyclones situated in series in order to separate the solids from the combustion gas without CO2. The solid separated in the first cyclone are then introduced into the calciner via a loop seal in order to permit the regeneration of CaO from CaCO3, thus separating the CO2. Biomass is added to the calciner as a fuel source to obtain the required temperature, above 850ºC. In a similar manner to the carbonator, in the calciner the solids, in this instance mainly CaO and a gas enriched in CO2, leave the top of the calciner and pass through two cyclones situated in series in order to separate the solids from the CO2 enriched gas. The solids separated in the first cyclone are returned to the carbonator via loop seal in order to close the loop of the process.
The experimental plant is fully automated and contemplates extensive monitoring equipment throughout the loops to permit the process to be followed at each step and thus characterised. Additionally, some of the energy input will be recovered from the hot gases leaving the reactors in heat exchangers that serve to preheat the air.
3. Summary
Carbonate looping is one of the emerging second generation CO2 capture technologies considered to be of promise as it employs a low cost and readily available sorbent and due to the process temperatures, it should be possible to recover and use much of the heat input required, reducing the final energy penalty. The niche concept being investigated in this work, where the combustion of biomass with air and the carbonation of CaO take place simultaneously is considered to be an attractive option for this technology, promoting a concept of negative emissions for the biomass consumed. However, to date much of the experimental work undertaken has been at a very small 5
Oviedo ICCS&T 2011. Extended Abstract
scale of various kWt. The new larger experimental plant being commissioned in the grounds of La Robla power plant of Gas Natural Fenosa is one of the required next steps to further develop and validate the concept. Over the next year, this larger experimental plant should deliver experimental results at a larger scale from experiments of longer durations to confirm and validate the previous smaller scale experimental work. Also, and more importantly, these results should determine the sorbent performance and make-up requirements, a key issue, contribute to both the technical and economical evaluation of the concept as well as generating data for the future scale up of this option to sizes in the order of several megawatts.
4. Acknowledgments
This work has been carried out thanks to the financial support from Spanish Centre for the Development of Industrial Technology (CDTI) under the auspices of the projects CENITCO2 and MENOSCO2. The participations of the Spanish Institute of Coal (INCAR-CSIC) in Oviedo and the Centre of Research for Energy Resources and Consumption (CIRCE) in Zaragoza in the project are also gratefully acknowledged.
5. References [1] International Energy Agency (IEA), World Energy Outlook, 2009 [2] Ishitani, H., Johansson, T. B. Energy supply mitigation options. In: R. T. Watson, M. C. Zinoyowera and R. H. Moss, editors. Climate Change 1995: Impacts, Adaptations, and Mitigation of Climate Change: Scientific-Technical Analyses, Cambridge, UK: Cambridge University Press; 1996, p. [3] Rhodes, J. S., Keith, D. W. Biomass with capture: negative emissions within social and environmental constraints: an editorial comment. Climatic Change 2008; 87: 321-8 [4] Obersteiner, M., Azar, C., Kauppi, P., Mollersten, K., Moreira, J., Nilsson, S., et al. Managing climate risk. Science 2001; 294: 786[5] Abanades Garcia, J. C., Alonso, M., Rodriguez, N. Experimental validation of in situ CO2 capture with CaO during the low temperature combustion of biomass in a fluidized bed reactor. Int. J. Green. Gas. Cont. 2010; d.o.i 10.1016/j.ijggc 2010.01.006: [6] Alonso, M., Rodriguez, N., Gonzalez, B., Arias, B., Abanades Garcia, J. C., Capture of CO2 during low temperature biomass combustion in a fluidized bed using CaO. Process description, experimental results and economics. GCGT-10, Energy Procedía 6
Oviedo ICCS&T 2011. Extended Abstract
2010;
7
Oviedo ICCS&T 2011. Extended Abstract
Measurement of Gasification Rate of Coal Char under High Pressure and High Temperature using A Mini Directly-heated Reactor K. Miura*, M. Makino, E. Sasaoka, S. Imai, R. Ashida Department of Chemical Engineering, Kyoto University, Kyoto 615-8510, Japan [email protected] Abstract A mini directly heated reactor (mini-DHR) was constructed to measure the gasification rate handily under high CO2 pressure of ~2 MPa in the presence of other gases, such as CO and H2, at T = ~ 1200°C. The mini-DHR was made of U-shaped SUS or Pt tubing of 3 mm I.D. The reactor itself was used as a heating element. An electric current of 75 – 150 A and a few volts were introduced to the reactor to heat up the reactor up to 900 to 1200°C. About 1 mg of char was placed in a platinum mesh basket of 1.0 mm I.D. and 10 mm high. The basket with the char sample was placed just above a thermocouple in the reactor. The conversion of char, X, was estimated by weighing the remaining char sample. The X vs. t relationships obtained under various conditions were analyzed to formulate a gasification rate equation in the presence of both CO2 and CCO for a char prepared from an Australian brown coal. 1. Introduction Enhancement of gasification reactivity of coal chars is very effective in increasing coal gasification efficiency.
Gasification reactivity of coal chars is believed to be
controlled mainly by catalytic effect of inherent minerals [1-4].
Then addition of
catalyst has been performed to further increase the gasification reactivity of coal char. A more cost-effective method, however, has been desired to enhance the gasification rate. We have recently proposed an upgrading method of low rank coal which consists of treatment of coal in non-polar solvent, such as 1-methylnaphthalene, at temperatures below 350°C [5]. The products obtained from the treatment are solvent-soluble fraction (extract) and insoluble fraction which we call “upgraded coal”. It was found that the gasification reactivity of the upgraded coal char was much larger than that of the raw coal for all the three coals tested. The CO2 gasification rate at 900°C of the upgraded coal char prepared from an Australian brown coal, Loy Yang coal, was surprisingly larger than the gasification rates of any other coal chars reported in the literature. Thus,
Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
it was found that the proposed upgrading method of low rank coal can be one of the ways of enhancing the gasification reactivity of coal char without using catalyst. In Japan oxygen blown gasification with recycled CO2 has been proposed to facilitate the CO2 separation and hence to increase the gasification efficiency under the NEDO “Innovative Zero-emission Coal Gasification Power Generation Project”. To realize the gasification concept practically, it is essential to increase the CO2 gasification reactivity of coal chars under high CO2 pressure of ~2 MPa at T = ~ 1200°C. This requests the development of methods to increase the gasification rate and to measure the gasification rate under such extreme conditions. In this work a mini direct heating reactor was constructed to measure the gasification rate under the extreme conditions. Then the char conversion vs. time relationships obtained under various conditions were analyzed to formulate a gasification rate equation in the presence of both CO2 and CO for a char prepared from an Australian brown coal. 2. Experimental 2.1 Samples. An Australian brown coal, Loy Yang coal (LY), was used as a low rank coal in this study. The detailed upgrading procedure has been described in a previous paper [5]. The coal was treated in 1-methylnaphthalene (1-MN) at 350°C under 2.2 MPa for 3 h and then separated into 1-MN-soluble fraction (extract) and insoluble fraction (residue) which we call “upgraded coal (UC)”. The elemental compositions of the LY coal and the upgraded coal (LY/UC) are given in Table 1. LY and LY/UC were carbonized for 30 min at 900°C in an inert atmosphere to prepare their chars. The char particles ranging from 75 to 150 μm in diameter were served to the gasification experiment. Table 1. Analyses of Loy Yang brown coal (LY) and the upgraded coal prepared from LY (LY/UC). Sample
Ultimate analysis [wt%, d.a.f.]
Ash
C
[wt%, d.b.]
H
N
O+S(diff.)
LY
66.7
4.7
0.9
27.7
1.5
LY/UC
77.4
4.0
1.0
17.6
2.9
2.2 Gasification Experiment.
Figure 1 shows the set up used for the gasification
experiment. The detail of the custom made reactor, mini direct heating reactor (miniDHR), is shown in Figure 2. The reactor was made of U-shaped SUS or platinum tubing of 3 mm I.D. The reactor itself was used as a heating element. An electric current of 75
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Oviedo ICCS&T 2011. Extended Abstract V5
reactor V1
MFC1
V3
regulator
V6
V12 NV3
MFC3
V9
P
BV1
He
P NV1 regulator
MFC2 V4
He or CO2 Reaction gas CO
V10
V11
V13 BV2
NV2
V7 V8
Fig. 1. Experimental set up for gasification at high temperature and high pressure.
Quartz wool
35 mm
Char in mesh basket (1 mg)
90 mm
ID: 3 mm SUS or Pt tube
Thermo-couples
Gas flow 600 cm3/min(NTP)
Fig. 2. Detail of the custom made mini directly heated reactor (mini- DHR). – 150 A and a few volts was introduced to the reactor to heat up the reactor up to 900 to 1200 °C. About 1 mg of char was placed in a platinum mesh basket of 1.0 mm I.D. and 10 mm high. The basket with the char sample and wrapped by quarts wool was placed just above a thermocouple in the reactor. The sample was heated up to a desired temperature in a He stream. Figure 3 shows a typical heating profile. It is shown that the temperature can be controlled very accurately. After reaching the gasification temperature the gas stream was switched to the stream containing CO2 or CO2/CO mixture to start the gasification. After the elapse of a predetermined reaction time the gas stream was switched back to the He stream, and cooled down to room temperature in the He stream. The the weight of basket containing the remaining char was measured by a micro balance to estimate the conversion of char, X. By repeating this procedure the accurate relationships between X vs. reaction time,t, could be obtained.To examine
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Oviedo ICCS&T 2011. Extended Abstract
1200
temperature [ºC]
1000 800
CO2 & He CO
600 400
He
30 s
200 0
3～5 V 60～130 A
0
30
60
90
120
150
180
210
time [s]
Fig. 3. A typical heating profile of the mini-DDHR. the validity of the mini-DHR, the X vs. t relationships obtained under an atmospheric pressure was compared with those obtained by a sensitive thermobalance (Shimadzu, TGH-50). The X vs. t relationships below atmospheric pressure were obtained by use of the thermobalance. 3. Results and Discussion 3.1 Examination of the validity of mini-DHR.
Figure 4 compares the X vs. t
relationships measured by using the mini-DHR (keys) and those obtained by the thermobalance (broken lines) at 900 °C under 0.1 MPa of CO2 pressure (pCO2) for the LY and LY/UC chars. Good agreements between the mini-DHR and TG experiments show the validity and accuracy of the gasification rate measurement using the mini-DHR. The gasification rates 1.0 LY LY/UC TG
X[-]
0.8 0.6 0.4 0.2
( = 900 ℃, pCO2 = 0.1 MPa 0 0
100
200
300
400 t [s]
500
600
700
Fig. 4. Comparison of the X vs. t relationships measured by using the mini-DHR (keys) and those obtained by the thermo-balance (broken lines). 3.2 Formulation of gasification rate in the presence of CO2 and CO.
The gasification
rate of char, -rc, was formulated by the by combining the Langumuir-Hinshelwood type equation representing the gasification of each active site and the Random-Pore model [6]
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Oviedo ICCS&T 2011. Extended Abstract
describing the active site number as a function of X as follows: − rC =
k1 pCO2 f 0 dX / dt = 1 − ψ ln ( 1 − X) 1− X 1 + (k1' k 2 ) pCO + (k1 k 2 ) pCO2
(1)
where ψ is the structural parameter of the char, f0 is the initial value of the number of active sites, and the rate constants, k1, k1’, k2, are the rate constants of the following elementary reactions: k1
CO2 + Cf → ← CO + C(O) k 1’
(2)
k
C(O) →2 Cf + CO
(3)
The structural parameter ψ was found to be set equal to 0 from the accurate X vs. t relationships obtained by use of the thermo-balance at atmospheric pressure of CO2. The X vs. t relationships measured under a wide a range of CO2 pressures and at four different temperatures are shown in Figures 5 and 6.
The -rc values at X = 0.3 were
estimated from these data and used to estimate the rate parameters k1f0 and k1/k2. 1
1
0.8
pCO2 = 2.0 MPa
1100℃
1.0 MPa
0.6
X[-]
X[-]
0.8
0.6 MPa
0.6 1000℃ 0.4
0.4 0.2 MPa
0.2 0
T = 1200℃
0
100
200
300
0.2
400
500
0
0
20
40
60
t [s]
t [s]
(mini-DHR T = 900ºC、pt = pCO2 + pHe = 2.0 MPa)
(mini-DHR pCO2 = 1.0 MPa, pHe = 1.0 MPa)
Fig. 5. X vs. T relationships in CO2 (1)
Fig. 6. X vs. T relationships in CO2 (2)
The X vs. T relationships in CO2/CO gas mixtures were measured systematically under several selected conditions.
Figure 7 shows
typical X vs. T relationships
obtained. Using all of the X vs. T relationships obtained, -rc values at X = 0.3 were estimated and plotted against pCO2 at several different values as shown in Figure 8. Figure 9 shows final Arrhenius plots for estimating k1f0, k1/k2, and k1’/k2. The Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
2.5
-rC (X = 0.3) ×103 [s-1]
pCO = 0
0.8
X[-]
T = 900℃ pCO2 = 0.8 MPa
0.1 MPa
0.6 0.4
1.0 MPa
0.5 MPa
X = 0.3 0.2
pCO = 0 0.1 MPa 0.5 MPa 1.0 MPa
2.0
PCO = 0
0.04 MPa
1.5 0.1 MPa
1.0 0.5
0.5 MPa 1.0 MPa
0
0 0
2000
4000
0
6000
0.5
1.0 pCO [MPa]
1.5
2.0
2
t [s]
(T = 900℃)
Fig. 7. X vs. T relationships in CO2 /CO.
Fig. 8. -rc values in CO2 /CO.
k1f0 [MPa-1 s-1], k1’/k2, k1/k2 [MPa-1]
1000
k1/k2
100
10
k1’/k2 1
k1f0
0.1
0.01 6
7
8
9
104/T [K-1]
Fig. 9. Arrhenius plots for estimating k1f0, k1/k2, and k1’/k2. Table 2. Estimated values for k1f0, k1/k2, and k1’/k2. Ai k1f0 k1’/k2 k1/k2
1.6×107 MPa-1 s-1 0.16
MPa-1
6.2×10-4 MPa-1
Ei [kJ mol-1] This work Kajitani7) 2.0×102
2.22×102
-39
-23
-91
-48.1
estimated values are shown in Table 2. The activation energies estimated by Kajitani [7] are also shown for comparison purpose. The activation energies of k1f0, are so close to each other. Finally the experimentally obtained X vs. T relationships were compared with the X vs. T relationships calculated by the estimated rate equation in Figure 10.
Good
agreement was obtained between the X vs. T relationships, showing the validity of the rate equation estimated.
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Oviedo ICCS&T 2011. Extended Abstract
1 T = 900℃ ψ=0
calc
0.8
pCO = 0.6 MPa pCO = 0.1 MPa
pCO = 1.5 MPa
X[-]
2
2
0.6 0.4 0.2
pCO = 0.8 MPa pCO = 0.5 MPa 2
0 0
300
600
900
t [s]
1200
1500
Fig. 10. Comparison of experimentally obtained X vs. T relationships and calculated X vs. T relationships. 4. Conclusions A mini direct heating reactor (mini-DHR) was successfully constructed to measure the CO2 gasification rate of coal chars at high temperature of up to 1200°C and high pressure up to 2 MPa. The validity and accuracy of the gasification rate measurement by the mini-DHR were well clarified, indicating that the mini-DHR can be a handy apparatus for the gasification measurement under extreme gasification conditions. The gasification rate equation was successfully formulated by combining the Langmuir-Hinshelwood model and the Random-Pore mode. Acknowledgement. This work was commissioned by the Central Research Institute of Electric Power Industry (CRIEPI) under the NEDO “Innovative Zero-emission Coal Gasification Power Generation Project”. References [1] Walker, P. L.; Rusinko, F.; Austin, L.G. In Gas Reactions of Carbon, Advances in Catalysis, Vol.11, Academic Press, New York, 1959, p.133. [2] Essenhigh, R. H. In Chemistry of Coal Utilization 2nd Supplementary Volume, John Wiley & Sons, 1981, p.1153. [3] Van Heek, K. H.; Muehlen, H. J. Fuel 1985, 64; 1405-1414 [4] Miura, K.; Hashimoto, K.; Silveston, P. L. Fuel 1989; 68; 1461-1475. [5] Miura, K.; Ashida, R.; Umemoto, S.; Sakajo, A.; Saito, K.; Kato, K. Proceedings of the 25th Pittsburgh Coal Conference, Pittsburgh, 2008. [6] Bhatia, S.K; Perlmutter, D.D. AIChEJ 1980, 26,379-386. [7] Kajitani, S.; PhD thesis (Kyoto University, 2007).
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Oviedo ICCS&T 2011. Extended Abstract
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Oviedo ICCS&T 2011. Extended Abstract
IMPLEMENTATION OF COAL GASIFICATION IN A FLUIDIZED BED FIRING SYSTEM FOR BRICK TUNNEL KILN.
F. Chejne1, C. londono1, C. Gómez1, J. Espinosa1, F. Mondragon2, J.J Fernandez2, Erika Arenas3 L. C Cuartas4 1
2
Universidad Nacional de Colombia, Facultad de Minas. Grupo de Termodinámica Aplicada y Energías Alternativas, TAYEA. Medellín, Colombia. [email protected]
Universidad de Antioquia, Grupo Química de recursos Energéticos y medio Ambiente, Instituto de química y medio Ambiente, Medellín, Colombia. [email protected] 3
Grupo de Energía y Termodinámica. Instituto de Materiales y Medio Ambiente. Escuela de ingeniería. Universidad Pontificia Bolivariana, Medellín- Colombia, [email protected] 4
Ladrillera San Cristóbal - Las Playas, Medellín, Colombia. Conm. 427 01 45,ladrillera [email protected]
Abstract We present the results of the design, installation and commissioning of a fluidized bed gasifier coupled to a tunnel kiln for firing bricks. The main objective of the project was put gasification technology of coal in fluidized bed developed by the Universidad Nacional de Colombia-Medellín, Universidad Pontificia Bolivariana, Universidad de Antioquia and Ladrillera San Cristobal for firing brick in a tunnel kiln. The idea of burning the bricks in the tunnel kiln with coal gas by replacing the system that uses pulverized coal, is that significantly reduces specific fuel consumption and produces building blocks for a more cleanly, cheaply and with less waste, further emerges as a need to search for better alternatives to burning coal because it is an abundant resource in Colombia.
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1
Oviedo ICCS&T 2011. Extended Abstract
1. Introduction. Coal is one of the non-renewable energy still available and disseminated worldwide. However, improper use can be one of the most dangerous pollutants to the environment.
As the century progresses, there is a growing need for energy due to global economic growth. It is projected that fossil fuels will remain the main energy sources in the world in this century, and the coal should increase its participation in power generation, using the large reserves, estimated at 987.066 billion tonnes in the world with an estimated availability of this fuel from 216 to 500 years at current consumption rates. However, environmental concerns have grown about the use of coal with respect to greenhouse gas emissions of pollutants and particulate material available, therefore, there is a clear global need to develop environmentally friendly technologies for the management coal.
The most appropriate system to make the coal more competitive with other fossil energy resources to meet environmental requirements, is the gasification, which is obtained with a combustible gas that can be cleaned of contaminants to be used then processes combined cycle generating electricity with high efficiencies and a significant reduction of pollutant gases such as CO2, SOx, NOx and particulate matter. Have identified several of them, such as Integrated Gasification Combined Cycle (IGCC), Pressurized Fluidized Bed Combustor (PFBC) and Combustion in Pulverized Coal Injection (PCI) as the most viable alternatives to the use of clean coal IGCC to be the most efficient. In the development of these technologies employ high operating pressures, for example, 15-25 atmospheres for IGCC, PFBC atmospheres for 10-15, and less than 5 atmospheres for PCI. One of the most efficient technologies for coal gasification are fluidized bed reactors.This type of equipment are widely used in the chemical industry, energy, environmental and oil, due to good mixing of solids and high transfer rates and reaction that provides the fluid-solid contact made. Coal gasification in a fluidized bed has advantages over other technologies to allow the use of coal rubble, which is lower cost granular carbon, there is a reduction of NOx by working at low temperatures, the order of 850 ° to 900oC and allows recovery of SO2 in-situ.
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2
Oviedo ICCS&T 2011. Extended Abstract
For proper and optimal use of coal, leading to the award of new sub-products that maximize the use and greater value to this mineral, this project was developed as a major objective. Coal is an abundant resource in Colombia and as such should take advantage of the occasion, looking to get the best possible use. Should investigate ways to add value to this raw material displaying different aspects of traditional combustion, as demonstrated in the previous investigation DESIGN, INSTALLATION AND COMMISSIONING OF A fluidized bed gasifier For DRYING of BRICK , "held by the same research group, in which it was found that gasification is a more efficient alternative.
The idea of burning the bricks in the tunnel kiln with syngas by replacing the system that uses pulverized coal, significantly reduce the specific fuel consumption and produces a brick clean, cheaply and with less waste. The objetive industry is the brick firing clean, more efficient and environmentally friendly.
Currently the company Ladrillera San Cristobal has replaced a Hoffmann kiln for firing bricks by for tunnel kiln. Which allows significantly reducing specific fuel consumption and achieving better production quality, lower cost and less waste? The project aims to put the technology of coal gasification in fluidized bed developed by the Universidad Nacional de Colombia-Medellín, Universidad Pontificia Bolivariana, Universidad de Antioquia and San Cristobal Ladrillera in the process for firing brick in a tunnel kiln.
Some of the advantages that brings the implementation of this technology are: The brick comes out spotless avoiding a washing process and economy of the builders, fuel economy with the system currently used, facilitates the reuse of hot gases, the combustion gas produced by gasification is cleaner as there is no particulate matter (fly ash) in suspension, being a more efficient, reduce the emission of CO2, CO and other polluting gases in the atmosphere such as NOx and SOx, the fluidized bed gasifier can gasify particles, which is achieved use of energy from solid waste, increased business productivity and modernizing and simplifying the process of brick production.
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Oviedo ICCS&T 2011. Extended Abstract
2. Experimental section
Based on the proposed initial system design, it had to make modifications in accordance with recommendations of the employees of the company. One change relates to the arrangement of the burners in the tunnel kiln, which was initially contemplated that were arranged laterally and eventually he settled in the oven roof. Figure 1 shows a diagram of the plant.
Figure 1. Schematic illustration of the plant with the gasifier. The Figure 1 shows a schematic of the final gasification plant designed for Ladrillera San Cristobal. It can detail the complete set of gasifier, hopper, cyclone and exchangers.
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Oviedo ICCS&T 2011. Extended Abstract
Figure 2. Scheme of the gasification plant
Figure 2 shows the reactor, the high efficiency cyclone and high flow heat exchanger to generate steam entered the cyclone. In Figure 3, on the other hand, presents the current pattern of distribution of gas in the oven where it has a dual system for the supply of gas or coal. Besides the main components, the coal gasification system to operate and carry to the kiln gases to be burned there, require additional peripheral equipment, which must be selected in accordance with the requirements of each, among them are : blower, water pump, coal feeder, etc.
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Oviedo ICCS&T 2011. Extended Abstract
Figure 3. Displaying the gasification system with air lines and synthesis gas.
The minimum fluidization velocities for these two particle sizes and were to be 0.3 m / s for sample 1 and 0.21, for the other. According to this information was determined to work with a particle size of 0.8 to 1.3 mm to ensure a fluidized bed gasification tests. For tests initially hot burner was installed at the outlet of the heat exchanger as shown in Figure 4 in order to ensure the necessary conditions do not significantly affect the process of firing in the oven.
Figure 4. Preliminary tests burner installed.
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Oviedo ICCS&T 2011. Extended Abstract
For the first hot test starts with a Wood bed as shown in Figure 5, was later introduced to diesel to start the fire. It was done to the first gasification tests subsequently given the problem of formation of tars from the start come on only with coal.
Figure 5. Wood bed on.
Once the system has a temperature above 300 ° C in the freeboard the door is closed, parallel turns on the air supply and start the process posteriori power supply gradually to 720 kg / h. Before starting the gas supply to furnace gases are lit in the fireplace as shown in Figure 6.
Figure 6. Turning on the fireplace. Submit before January 15th to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
Once it reaches a temperature in the freeboard of about 700 ° C starts the water supply to the cyclone, which in turn goes to the reactor as gasification agent, the time when the temperature is around 800 ° C starts gas supply to the furnace.
For the first tests it lit the flame at the outlet of heat exchanger as shown in Figure 7, later made a rough cut of gas supply to the furnace for testing without affecting production processes up and got a flame in furnace as shown in Figure 7
Figure 7. Flame obtained by burning the synthesis gas burner outlet provisional in the oven.
During the process of setting up interim assembly was required given the complexity of the cooking process and to not change the heating curve of the oven without having the certainty of producing synthesis gas with the necessary conditions for the process 3. Results and Discussion In order to evaluate the energy and environmental performance of tunnel kiln using the synthesis gas in burning, is scheduled measurement of variables such as flow coal in the gasifier and tunnel kiln, ceramic work flow, flow tunnel furnace and air gasifier, temperatures, flow of gases in the chimney of the furnace, gas concentrations and emissions of particulate matter and acid mists tunnel kiln, among others.
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8
Oviedo ICCS&T 2011. Extended Abstract
It was noted that the unburned carbon is deposited on trucks that carry loads of bricks in the tunnel kiln fell more than 80% from a pickup truck waste by 8.181 kg to 1.766 kg on average. This means that the quality of the bricks is improved by 80% for not having sulfur coal residue on the surface, which generates the yellow stains on the bricks called called "efluorecencia. " During the experimental trials were conducted sampling for gas composition analysis via gas chromatography. The results are presented in Table 1.
Tabla 1. Composition of gas generated
% MASS GAS
Experimental date
DRY BASIS Simulated data
Flow Coal, Kg/h ~700
720
Flow air kg/h Máx 1900
1525
Flow water kg/h <120 GAS
130
% DRY
% DRY
BASIS
MASS BASE
% DRY MASS BASE
MOLAR CO
20
27
33,4
H2
7
0,5
1,3
CH4
2
0,5
0,0
CO2
17
27
12,3
N2
54
54
49,3
PCS COLD, MJ/Nm3 seco
3,8
6,0
3,2
5,0
PCS COLD, MJ/kg
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9
Oviedo ICCS&T 2011. Extended Abstract
seco
Dry syngas flow generated
Max 2705
2450
Max 2290
1435
summary kg/h Dry syngas flow generated summary, Nm3/h
According to gas chromatographic analysis (Table 1), we conclude that it is generating a fuel gas that contains high concentrations of CO. With these features, the calorific value of gas is 3.8 MJ/Nm3 depending on the percentage of H2 present in the syngas, which places it in the range from gas, able to meet the energy needs of the oven in San Ladrillera Cristobal SA
Based on nitrogen balance was calculated synthesis gas flow maximum that could be causing the gasifier (see Table 1). It gets the maximum value, since it corresponds to the maximum value that can supply air blower, however, the flow of air entering the gasifier is controlled according to desired conditions and therefore, under certain operating conditions is necessary to reduce the air supply.
The discrepancy between the simulated and the experimental ones is due mainly to nonguaranteed flows of air and steam for experimental tests similar to the simulated values and the type of coal is probably not the same and therefore has a gasification kinetics than that used in the model. During the simulation we used water vapor and not liquid water, as we introduce the gasifier. However, it is considered acceptable theoretical and experimental results.
4. Conclusions. It was possible to design, installation and commissioning of a coal gasifier fluidized bed capacity of 720 Kg per hour coupled to the oven tunnel Ladrillera San Cristobal, which can meet between 80% and 100% of energy needs. It was an innovative development of an assembly that involves the gasification process with the cooking, which is a delicate process that was put into operation successfully. Submit before January 15th to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
Importantly, the union of the three universities together with a private company with strong support from Colciencias be able to develop innovations in the near future will open new doors for the implementation of such technologies. With this type of project the San Cristobal Ladrillera put at the forefront of innovation and development of clean technologies is an approach that can easily emulate the other with less risk.
I raised from the beginning was to replace a burner that gives 60% of energy from the oven once the oven is working on an ongoing basis may be substituted for 80% of needs, which defined a capacity of 720 Kg / h to meet the needs of brick production of 300 tonnes per day. Initially contemplated the use of synthesis gas by the side of the oven, but burners were installed at the top by a decision of the company and to avoid distortions in the oven that was initially designed to be operated with fuel entering from the top.
We installed a control panel that allows you to operate the air blower, the regulation of food, water supply pumps, also were installed in U gauges and temperature gauges to ensure the conditions of fluidization and gasification. However, the system requires more control to ensure adequate fuel supply to the oven and automatic safety system that currently is done manually. These control systems require more investment not covered by the project beyond the scope of the project budget, hence the importance of a second phase to achieve funding from the company and government support to complete the requirements of control .
The experiments performed consisted of the implementation of the system, checking the conditions of fluidization, visual determination of the presence of combustible gases in the presence of air and visual assurance on the burners in the furnace. However, it is important to quantify the quality of gas at stable operating conditions and this requires the final coupling of the power system, and supply to all burners.
Acknowledgement. . The authors express their sincere thanks for financial support, the Departamento Administrativo de Ciencia Tecnología e Innovación COLCIENCIAS, the Servicio
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Oviedo ICCS&T 2011. Extended Abstract
Nacional de Aprendizaje SENA, INCARBO and Ladrillera San Cristóbal to contribute to the development of research in our country. The brick is highlighted in particular Hildebrando Alvarez for their hard work in the process of construction and commissioning.
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12
Catalytic steam gasification of large coal particles S. Nel1, H.W.J.P Neomagus1*, J.R. Bunt1,2, R.C. Everson1 1
School for Chemical and Minerals Engineering, North - West University, Potchefstroom Campus, South Africa 2
Sasol Technology (PTY) Ltd, Box 1, Sasolburg, 1947, South Africa
*Corresponding author: Tel. + 27 18 2991991, Fax: + 27 18 2991535, E-mail: [email protected] Abstract The catalytic gasification of coal has been studied extensively in order to develop more efficient gasification processes.
Advantages of catalytic gasification include: increased
reaction rates, decreased operating temperature and increased process throughput. Previous studies have mainly focused on the catalytic gasification of small coal particles and pulverised coal. This study involves the development of a suitable catalyst impregnation method for large coal particles, as well as the investigation of the catalytic affect on steam gasification reactivity. The excess solution method was used to impregnate large coal particles (5 mm, 10 mm, 20 mm and 30 mm) with the selected catalyst, K2CO3. Results from XRF analysis indicated that catalyst loadings up to 0.83 wt.% (coal basis) were obtained for large coal particles, using the excess solution impregnation method. The potassium-impregnated large coal particles were used for low temperature steam gasification experiments, in order to determine the affect of catalyst addition on the coal reactivity. Results obtained for the reactivity of the parent coal were compared to that of the impregnated coal, which indicated that the addition of K2CO3 increased the reaction rate of large coal particles. It was also found that the addition of K2CO3 to the coal decreased the activation energy. Keywords: catalytic steam gasification; large coal particles; potassium carbonate; catalyst impregnation.
1. Introduction Gasification processes are receiving ever-increasing attention in terms of a cleaner and more energy efficient coal conversion technology. Various methods are being studied in order to develop more efficient and environmentally friendly coal conversion technologies [1]. One such method is catalytic gasification, which is considered to be an excellent optimisation method for gasification processes.
The catalysis of coal gasification processes has the
purpose of increasing coal reactivity towards steam, achieving higher reaction rates at reduced temperatures, and reducing reactor energy demand [2]. Extensive research regarding the catalysis of coal gasification has two main purposes. Firstly, to understand and identify the gasification kinetics of coals which contain catalytically active compounds, and secondly, to develop and investigate the catalysis of gasification processes [3-6]. Previous studies conducted on the topic of catalytic gasification have mainly focused on the impregnation and gasification of small coal particles and pulverised coal. This study will focus on the catalytic steam gasification of large coal particles (5 mm, 10 mm, 20 mm and 30 mm), in order to investigate the catalytic influence in terms of reactivity. The addition of a catalyst to large coal particles, with a suitable catalyst impregnation method, is envisaged to increase the steam-gasification reaction rate, and thereby increase the overall efficiency of the gasification process. Therefore, this study involves the development of a suitable catalyst impregnation method for large coal particles, as well as the investigation of the catalytic affect on steam gasification reactivity.
2. Experimental section
2.1 Impregnation The catalyst selected for this study is potassium carbonate (K2CO3). Various studies have shown that K2CO3 is one of the most effective catalysts for steam gasification of coal or char [7]. The following factors contribute to the exceptional effectiveness of K2CO3 as catalyst [7,8]: K2CO3 has the ability to significantly increase the gasification reaction rates; K2CO3 is mobile under gasification operating conditions, which leads to good dispersion throughout the coal sample; the catalytic effectiveness of K2CO3 is not influenced by coal rank. The impregnation method used for large particle impregnation is the excess solution method, where the coal particles are completely submerged in the catalyst impregnation solution for a
certain period of time [9]. Impregnation experiments were conducted for 21 days, at constant temperature (23 °C), and at a solution concentration of 0.5 M K2CO3. XRF analysis was used to determine the catalyst loading achieved during impregnation, for the various particle sizes.
2.2 Reactivity experiments The potassium-impregnated large coal particles were used to determine the effect of catalyst addition on the coal reactivity during low temperature steam gasification experiments in a large particle, home-built TGA. The main component of the TGA is the vertical tube reactor. During experimentation, the coal particle is placed in a quartz holder, which is positioned on a balance, and inserted at the bottom of the vertical tube reactor. The feed gas, nitrogen and steam, is introduced from the top of the vertical tube reactor. Reactivity experiments were carried out isothermally, at 800 °C, 825 °C, 850 °C and 875 °C, with a steam concentration of 90 mol%.
3. Results and discussion
3.1 Coal The coal used in this study was a washed medium rank-C bituminous Highveld coal from South Africa, having a density of 1 400 kg/m3. Results for the proximate analysis of the coal are given in Table 1: Table 1: Proximate analysis Characterisation analysis Proximate analysis Inherent moisture content Ash content Volatile matter Fixed carbon (by difference)
wt.%
% % % %
Coal sample (air-dried basis)
Coal sample (dry basis)
5.0 12.6 27.0 55.4
13.2 28.5 58.3
As seen from Table 1, the ash content of the coal is 12.6 wt.%. According to Pinheiro [10], this is a relatively low ash content compared to other South African coals. The low ash content of this coal makes it especially suitable for use in catalytic gasification studies. The main factor responsible for increased gasification rates is the interaction of the catalyst cation with the char. However, according to Lang [11], the catalyst cation reacts with the char, as well as with the mineral matter in the ash. Suzuki et al. [7] observed that large amounts of
ash decrease the catalytic effect of a catalyst.
Therefore, in order to minimise interactions
between the catalyst cation and the mineral matter, it is advantageous to use coal with a low ash content for catalytic gasification. According to Suzuki [7], coals with ash content values of 7.2-14.7 wt. % (dry basis) gave optimum results when used in catalytic gasification studies.
3.2 Impregnation Results from the XRF analysis are given in Table 2, and indicate the total amount of potassium (K) specie (wt.%) present in the coal after impregnation. Catalyst loading results, as presented in Table 2, are given as a weight percentage on a coal basis. Table 2: XRF results Particle size
Loading of K specie (wt.%) (coal basis)
5 mm
0.83
10 mm
0.76
20 mm
0.76
30 mm
0.68
The XRF results shown in Table 2 indicate the maximum catalyst loading achievable for the various particle sizes, with the excess solution impregnation method. Results show that the 5 mm particle size obtained the largest catalyst loading. From Table 2 it can also be seen that the potassium-loading decreases with increasing particle size. The catalyst loading obtained for large particle impregnation is significantly lower when compared to results where powdered coal was used during impregnation. Sharma et al. [12] used -75 μm coal particles and achieved a catalyst loading of 6 wt.% when impregnated with K2CO3, while Yeboah et al. [13] achieved a catalyst loading of 10 wt.% with K2CO3 and a coal particle size of -250 μm. The use of powdered coal and an impregnation method such as mechanical mixing allows the addition of a desired amount of catalyst, which means that the catalyst loading can be controlled. However, for large particle impregnation, the catalyst loading is largely dependent on the degree of dispersion of the impregnation solution throughout the particle. Consequently, the catalyst loading achievable through large particle impregnation depends on the chemical and physical structure of the coal particles.
3.3 Reactivity experiments Steam gasification experiments were conducted in order to determine the catalytic effect of K2CO3 on the reaction rate. The reactivity obtained for the parent coal was compared to that of the impregnated coal, to establish the effect of catalyst addition in terms of reactivity. Results obtained for the steam gasification experiments, for the 10 mm coal particles, are shown in Figures 1-4. Reactivity results are shown for experiments conducted at 800 °C, 825 °C, 850 °C and 875 °C, and are plotted as conversion (X) vs. time (hr). Results obtained for impregnated coal are termed “Cat”, while results obtained for the raw coal (uncatalysed coal) are termed “NonCat”.
Figure 1: Reactivity results at 800 °C
Figure 2: Reactivity results at 825 °C
Figure 3: Reactivity results at 850 °C
Figure 4: Reactivity results at 875 °C As can be seen from the results shown in Figures 1-4, the addition of K2CO3 as catalyst significantly increases the reaction rate when compared to the raw coal. As a result of the increased reaction rates, the experimental runs conducted with the impregnated coal reached complete conversion at a faster rate than the raw coal. Sharma et al. [12] also observed a significant increase in reaction rate at 775 °C, with the addition of K2CO3 at a catalyst loadings of 3 wt.%. Similar results were also obtained by Liu and Zhu [8], for the catalytic gasification of char with K2CO3. Sharma et al. [12] and Liu and Zhu [8] used fine powdered coal with the lowest catalyst loadings being 3 wt.% and 4 wt.%, respectively. However, the results shown in Figures 1-4 indicate that a significant increase in reaction rate can be observed at catalyst loadings as low as 0.76 wt.%. It can also be observed from the figures given above that the addition of K2CO3 increases the initial reaction rate of the reactivity experiments. The slopes of the initial reaction rate (up to 40% conversion), as seen in Figures 1-4, are used to determine the initial reactivity, R0, of the different experiments. The activation energy is then calculated from the slope obtained from an Arrhenius plot, as given in Figure 5.
Figure 5: Arrhenius plot The activation energy, Ea (kJ/mol), was calculated from the slopes obtained by the Arrhenius plot, as seen in Figure 5. Table 3 provides the values for the different activation energies Table 3: Values for activation energy Ea (kJ/mol) Raw coal sample
162
Coal sample with K2CO3
141
As seen from Table 3, the addition of K2CO3 decreases the activation energy by 13%. These values are in accordance with results obtained by Lee and Kim [6], whom also observed a 13% decrease in activation energy with the addition of K2CO3. The activation energy for the raw coal compares well with results obtained in a study conducted by Schmal et al [14]., where an activation energy of 165.3 kJ/mol was determined for a southern hemisphere subbituminous coal, at temperatures between 800 °C and 1000 °C. According to literature, the activation energy of coal catalysed by K2CO3, ranges between 142 and 310 kJ/mol, and similar results were obtained for this study [15].
4. Conclusions 1. Results obtained from this study indicate that the excess solution method is suitable for obtaining sufficient catalyst loadings for large coal particles. 2. Reactivity experiments conducted with 10 mm coal particles indicated that the addition of K2CO3 significantly increased the reaction rate compared to raw coal particles. It was also found that catalyst loadings as low as 0.76 wt.% had a noticeable effect on the reaction rate. 3. Results for the activation energies indicated that the addition of K2CO3 to the coal particles decreased the activation energy by 13%.
References [1] Zhu W, Song W, Lin W. Catalytic gasification of char from co-pyrolysis of coal and
biomass. Fuel processing technology 2008:89:890-896. [2] Jaffri G, Zhang J. Catalytic gasification characteristics of mixed black liquor and calcium catalyst in mixing (air/steam) atmosphere. Journal of Fuel Chemistry and Technology 2008:36(4):406-414. [3] Nishiyama Y. Catalytic gasification of coals-features and possibilities. Fuel Processing Technology 1991:29:31-42. [4] Wang J, Jiang M, Yao Y, Zhang Y, Cao J. Steam gasification of coal char catalyzed by K2CO3 for enhanced production of hydrogen without formation of methane. Fuel 2009:88:1572-1579. [5] Zhang Y,Hara S, Kajitani S, Ashizawa M. Modeling of catalytic gasification kinetics of coal char and carbon. Fuel 2009:doi:10.1016/j.fuel.2009.06.004. [6] Lee WJ, Kim SD. Catalytic activity of alkali and transition metal salt mixtures for steam char gasification. Fuel 1995:74(9):1387-1393. [7] Suzuki T, Mishima M, Kitaguchi J, Itoh M, Watanabe Y. The catalytic steam gasification of one australian and three japanese coals using potassium and sodium carbonates. Fuel Processing Technology 1984:8:205-212. [8] Liu Z, Zhu H. Steam gasification of coal char using alkali and alkaline-earth metal catalysts. Fuel 1986:65:1334-1338. [9] Dawes SB, Mazumber P. 2006. Method for preparing catalysts. Us patent: 7,097,880.
[10] Pinheiro, HJ. 1999. A techno-economic and historical review of the South African coal industry in the 19th and 20th centuries and analyses of coal product samples of South African collieries 1998-1999 (in bulletin 13. SABS:Pretoria. 97p.). [11] Lang, RJ. Anion effects in alkali-catayzed steam gasification. Fuel 1986:65:1324-1330. [12] Sharma, A, Takanohashi T, Morishita K, Takarada T, Saito I. Low temperature catalytic steam gasification of Hypercoal to produce H2 and synthesis gas. Fuel 2008:87:491-497. [13] Yeboah, YD, Xu Y, Sheth A, Godavarty A, Agrawal PK. Catalytic gasification of coal using eutectic salts: identification of eutectics. Carbon 2003:41:203-214. [14] Schmal M, Monteiro JLF, Castellan JL. Kinetics of coal gasification. Ind. Eng. Chem. Process. Des. Dev. 1982:21:256-266. [15] Yuh SJ, Wolf EE. K2CO3-catalysed steam gasification of supercritical extracted chars. Fuel 1983:62:738-741.
Oviedo ICCS&T 2011. Extended Abstract
Mobility of hazardous elements of Coal Cleaning Residues
Luis F. O. Silvaa; Frans Waandersd; Marcos L. S. Oliveiraa; Kátia da Boita a
Catarinense Institut of Environmental Research and Human Development – IPADHC,
Brazil. b
School of Chemical and Minerals Engineering North West University (Potchefstroom
campus) Potchefstroom 2531, South Africa * Corresponding author. E-mail address: [email protected] (L.F.O. Silva)
Abstract The geochemical characteristics of coal cleaning residues (CCR) occurring in the Santa Catarina, State, Brazil, were investigated. Around 3.5 million ton/year of coal waste are being dumped where coal beneficiation by froth flotation results in large amounts of CCR composed of coaly and mineral matter, the latter characterised by the occurrence of sulphide minerals and a broad array of leachable elements. The total and leachable contents of more than 60 elements were analysed. Atmospheric exposure promotes sulphide oxidation which releases substantial sulphate loads as well as Ca2+, K+, Mg2+, Cl- and Al3+. The metals with the most severe discharges were Zn, Cu, Mn, Co, Ni and Cd. Most trace pollutants in the CCR displayed a marked pH-dependent solubility, being immobile in near-neutral samples. The results highlight the complex interactions among mineral matter solubility, pH and the leaching of potentially hazardous elements.
1.
Introduction
Environmental pollution caused by industrial effluents rich in acid, metals and sulphate, such as resulting from run-off from coal cleaning plants before the coal is being utilized in thermoelectric power generation, are major contributors to the mineralisation of receiving waters and may prove toxic to both fauna and flora due to unacceptably high concentrations of, amongst others heavy metals. This is a universal concern and in two of the BRICS (Brazil-Russia-China-India-South-Africa) countries this phenomenon has been studied intensively [1-4]. Figure 1 (a) depicts an acidic pond created by coal discard washing water in South Africa and Figure 1 (b) shows the same problem occurring in Brazil.
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Oviedo ICCS&T 2011. Extended Abstract
Figure 1. (a) Coal cleaning residue pond in the South Africa situation and (b) a coal cleaning rejects pond close to a coal washing plant in Brazil. In this paper attention will be given to the Santa Catarina State, Brazil, coal mining region where coal has been used for power generation for nearly 80 years [5,6]. The characteristics of Brazilian coals render them difficult to wash and more aggressive coal cleaning technologies, such as gravimetric jigging, are implemented and the CCR are dumped in mined out areas or in flat areas near coal washing plants and the waste heaps are composed basically of mineral and residual coaly matter [7]. The problem is further aggravated by the fact that the waste dumps are instable and can collapse, that spontaneous combustion occurs with subsequent deleterious atmospheric emissions and acid leachate (AMD) forms, affecting the local environment negatively [7]. A number of measures to reduce the environmental impact of mining and washing activities are being applied, however, these measures have proven to be insufficient to prevent damage caused and furthermore there are no standardised methods for reducing the AMD potential in the area [7, 8]. Leaching of CCR dumps (percolation and runoff leachate) severely impact the soil, surface water and groundwater resources. Acid and sulphate-rich solutions are produced from the pyrite present in the waste dumps after the coal has been washed, leaving the typical dark red water. The following reactions are responsible for the pyrite oxidation and resultant pollution where oxidation of pyrite is kinetically enhanced [4] by the presence of anaerobic microorganisms: [9]: Æ 4FeSO4 + 4H2SO4 4FeS2 + 7O2 + 4H2O 4FeSO4 + 2H2SO4 + O2 Æ 2Fe2(SO4)3 + 2H2O 2Fe2(SO4)3 + 12H2O Æ 4Fe(OH)3 + 6H2SO4 _______________________________________________ 4FeS2 + 15O2 + 14H2O Æ 4Fe(OH)3 + 8H2SO4 The main aims of this study were to determine the geochemical and mineralogical Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
characteristics of coal cleaning residues and assess the leaching potential of these wastes for various elements.
2.
Experimental
A total of 18 fresh samples were collected immediately after the coal cleaning process plant. The samples were dried for 16 hours in a furnace at a temperature of 40ºC, homogenised and sieved through a -450μm mesh. Subsamples were ground to pass through a -5μm mesh for further analysis. The mineralogical study was carried out with a Bruker XRD-diffractometer. An ESEM coupled with EDX and a TEM was used for chemical analyses of individual particles. Prevention of possible mineralogical changes in individual solvents were done by preparing suspensions of 10 ml of each of the solvents mixed with 0.5 g of dried and
sieved CCR. The suspension was stirred for 1 min and then pipette onto carbon films
supported by Cu-grids and left to evaporate before TEM analyses. For chemical and leaching analyses, all samples were acid digested (first HNO3 hot extract followed by HF:HNO3:HClO4 acid) to retain volatile elements in the dissolute coal. The resulting solution was then analysed for major and selected trace elements by ICP-AES and ICP-MS, with analytical errors <3% for most of the elements and around 10% for Cd, Mo and P. Hg-analyses were done directly on solid samples using a LECO AMA 254 gold amalgam atomic absorption spectrometer. In order to study the leaching of elements, the compliance single batch leaching test EN 12457-2 [10] was applied and performed at a liquid to solid ratio of 10 L/kg and 24 h of agitation time. Deionised water was used as leachant and analyses were performed in duplicate.
3.
Results and Discussion
Mineralogy The mineralogy of the CCR was quite diverse. With some exceptions, the minerals detected in the CCR were those typically found in most coals and CCR [11-14]. The prominent minerals detected by means of XRD analyses in most of the samples taken, include: quartz (SiO2), kaolinite (Al2Si2O5(OH)4), microcline (KAlSi3O8),
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Oviedo ICCS&T 2011. Extended Abstract
muscovite ((Ba,K)Al2(Si3Al)O10(OH)2) and pyrite (FeS2). The
minerals
occurring
in
(Na0.5Al6(Si,Al)8O20(OH)10.H2O),
lesser
amounts
diopside
included:
(Ca,Mg(SiO3)2,
chlorite illite
K1.5Al4(Si6.5Al1.5)O20(OH)4), albite (NaAlSiO8), K-feldspar (KAlSi3O8), melilite (CaAl12MgSi3O14), mullite (Al6Si2O13), olivine (MgFeSiO4), talc (Mg3Si4O14(OH)2), zircon (ZrSiO4), galena (PbS), marcasite (FeS2), pyrrhotite (Fe(1−x)S), sphalerite (ZnS), aragonite
(CaCO3),
(Ca,Mg(CO3)2),
ankerite
oligonite
((Fe,Ca,Mg)CO3), (Fe(Mn,Zn)(CO3),
calcite siderite
(CaCO3),
dolomite
(FeCO3),
brushite
(CaPO3(OH).2H2O), monazite ((Ce,La,Th,Nd,Y)PO4), anhydrite (CaSO4), alunogen (Al2(SO4)3.17H2O),
barite
(BaSO4),
butlerite
(Fe(OH)SO4.2H2O),
calcantite
(CuSO4·5H2O), epsomite (MgSO4).7H2O), ferrohexahydrite (FeSO4.6H2O), hexahydrite (MgSO4.6H2O), gypsum (Ca(SO4)·2H2O), jarosite (KFe3+3(SO4)2(OH)6), melanterite (FeSO4.7H2O),
natrojarosite
(NaFe3(SO4)2(OH)6),
schwertmannite
(Fe3+16O16(OH)12(SO4)2), rozenite (FeSO4.4H2O), brucite (Mg(OH)3), hematite (Fe2O3), goethite (Fe(OH)3), gibbsite (Al(OH)3) and rutile (TiO2). Accessory species were also observed by X-ray diffraction and by microbeam (SEM/EDX, TEM/EDX; Figure 2) analyses and the use of water during coal mining in conjunction with atmospheric exposure promotes sulphide oxidation [15, 16]. Pyrite is known to react with water and dissolved oxygen to form sulphates and iron oxyhydroxides [9]. The oxidation of pyrite may release the trace pollutants such as As, Hg, Se or Pb to the environment [17]. The newly formed secondary minerals from AMD may play an important role for attenuating trace metals [18-20]. Jarosite and schwertmannite are environmentally relevant because Pb, As and Cr may be assimilated within their structures [21, 22]. Gypsum, in addition to jarosite and schwertmannite were the most prominent sulphate phases in the cleaned coal rejects, where their formation requires wet, oxidizing and acidic conditions.
Figure 2. Photomicrographs showing pyrite framboids (left), the most common form of syngenetic pyrite in coal and organic-rich shales, gypsum (SEM image, middle) and barite (SEM image right) [7]
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Oviedo ICCS&T 2011. Extended Abstract
Leaching of potentially hazardous elements The total soluble fraction of a residue is an important consideration for evaluating potential environmental impacts as CCR are not highly soluble in water and rarely the fractions of CCR surpasses 2 wt%. The major constituents solubilised from CCR are readily leachable salts from Ca2+ and SO42- ions, followed by K, Mg, Cl−, Fe and Al. The remaining elements, including most trace metals, were leached in much lower levels, but few of them are still of concern given their toxicity threshold [7]. According to the pH of leachates, samples can be classified in two main groups: (1) acidic samples, with pH values in the 3.8 to 4.5 range, and (2) neutral samples, with pH values ranging from 6.3 to 7.0. The low pH in group 1 is consistent with the occurrence of the slightly soluble jarosite, which was absent in group 2 samples. The non-metals were leached in the following order: SO42− > Se > NO3− with S being the most abundant and mobile constituent in the leachates. The order in which the alkalis were leached was in decreasing order: Na > K > Li > Rb > Cs with Na and K in concentrations of up to 200 mg/kg and the leachable concentrations of Li, Rb and Cs only up to 7 mg/kg. The alkaline earths leached in the following order: Ca > Mg > Sr > Ba > Be with Ca the main cationic species most easily released due to its high mobility and occurrence in various species. Leachable levels of Mg and Sr were 1 order of magnitude lower than those of Ca, while extractable Sr showed a wider range of variation with pH. Ba was highly insoluble, accounted for the <0.2% Ba leaching. Be was immobile under neutral conditions and extractable yields of this element and reached 20% of the total content at low pH [7]. The leaching of the metalloids followed the order: Si > B > As > Sb = Ge with the leachable contents of Si ≈ 35 mg/kg. B and As were leached in levels <0.8 mg/kg, entailing extractable proportions of < 1% of the total content. As, Sb and Ge were not critical elements due to their low abundance and leachable contents not mobile in CCR. The order for the transition metals: Fe > Mn > Zn > Cu > Co > Ni > V > Cd > Cr > Mo > Ti > W > Zr > Hf showed a well-defined pH-dependent leaching behaviour for most of these elements and the leachable contents were generally higher under acidic conditions, with very few exceptions. The leaching behaviour of Fe may be primarily linked to the occurrence of leachable Fe sulphates (e.g. jarosite), but in samples containing pyrite sludge, other phases also controlled Fe solubility [23]. Mn showed an order-of-magnitude variation in its total and leachable contents, suggesting a Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
considerable mobility with respect to the other transition metals and thus a persistent pollutant in neutralized AMD [24]. Besides Mn, Zn, Cu, Co and Ni were the main heavy metals released from the CCR, with a mean leachable contents of 14 mg/kg for Zn and 2 mg/kg for the remaining elements. The total contents of V and Cr were extracted to < 1%, revealing a low mobility whatever the pH. Zr and Hf were highly immobile in CCR and these elements were leached below the detection limit, regardless of the pH [7] The remaining metals leached according to the sequence: Al > Pb > Sn > Ga > Tl > Bi. Al leached at very low rates with respect to the high total contents with the only remarkable leachable contents measured at acidic pH. Pb is one of the most abundant toxic metals in coal [17] with a total content in the CCR reaching 100 mg/kg, but this metal was highly immobile with a leachable contents not exceeding 0.3 mg/kg regardless of pH and it can be assumed that the CCR will not result in Pb contamination. Sn, Ga and Tl were leached in very low levels in the most acidic samples (up to 0.4 mg/kg), while being immobile in the remaining samples and the total leachable contents of Bi in CCR was close to the detection limit [7]. Lastly the sequence for the Rare Earth elements (REE) and other metals was: Ce > Nd > U > Th > Y > La > Sc > Gd > Sm > Dy > Pr > Er = Yb > Eu > Ho > Tb > Tm = Lu. These elements were mostly associated with clay and detrital phosphate minerals and the acidity of the coal-forming environment exerted an influence on their concentration. In general, the total contents of these elements displayed a narrow variation among all the samples with Ce the REE most prominently leached (3 mg/kg), followed by Nd and La. The remaining REE’s were leached in concentrations <0.5 mg/kg and U, Y, Sc and Th were also released in low levels (<1 mg/kg). Significant percentages of around 2–6% Gd, Pr, Eu, Dy, Y and La were extracted in the most acidic CCR samples, while extractable yields decreased to negligible proportions at pH over 6 [7].
4.
Conclusions
In the present investigation the CCR were found to consist primarily of quartz, kaolinite, gypsum and pyrite, followed by feldspars and jarosite, the main weathering product. Pseudomorphic jarosite provided evidence of pyrite undergoing different oxidation rates, releasing high sulphate loads and a number of hosted trace metals. The major constituents solubilised from the CCR were found to be Ca and SO42−, largely released regardless of the pH, followed by K, Mg, Cl− and Al. The trace metals with the most severe discharges in leachates were Zn, Cu, Mn, Co, Ni and Cd. Very few samples Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
revealed potentially deleterious releases surpassing the limit values, which suggest that this issue could be solved if properly addressed. However, the acidity of leachates may not be easy to surmount as long as this characteristic is closely linked to the CCR origin. Acknowledgements One of the authors (FW) is greatly indebted to the main author (LFOS) who provided the information needed for this presentation. The NWU and THRIP are thanked for financial support to attend the conference.
References 1.
Silva , L. F. O., Moreno, T.,&Querol, X. (2009a).An introductory TEM study of Fenanominerals within coal fly ash. Science of the Total Environment, 407, 4972–4974.
2.
Silva, L. F. O., Oliveira, M. L. S., da Boit, K. M., & Finkelman, R. B. (2009b). Characterization of Santa Catarina (Brazil) coal with respect to human health and environmental concerns. Environmental Geochemistry and Health, 31, 475–485.
3.
Silva, L. F. O., & Oliveria, M. L. S. (2010). A preliminary study of coal mining drainage and environmental health in the Santa Catarina region, Brazil. Environmental Geochemistry and Health. doi:10.1007/ s10653-010-9322-x.
4.
Maree J.A., Treatment of industrial effluents for neutralization and sulphate removal, PhD unpublished PhD, North-West University 2006,
5.
Pires, M., & Querol, X. (2004). Characterization of Candiota (South Brazil) coal and combustion byproduct. International Journal of Coal Geology, 60, 57–72.
6.
Kalkreuth, W., Holz, M., Kern, M., Machado, G., Mexias, A., Silva, M. B., et al. (2006). Petrology and chemistry of Permian coals from the Paraná Basin: 1. Santa Terezinha, LeãoButiá and Candiota Coalfields, Rio Grande do Sul, Brazil. International Journal of Coal Geology, 68, 79–116.
7.
Silva, L.F.O, Izquierdo, M., Querol, X., Finkelman, R.B., Oliveira, M.L.S., Wollenschlager, M., Towler, M., Pérez-López, R., Macias, F., Leaching of potential hazardous elements of coal cleaning rejects, Environ Monit Assess (2010) 175:109-126.
8.
Akcil, A., & Koldas, S. (2006). Acid mine drainage (AMD): Causes, treatment and case studies. Journal of Cleaner Production, 14, 1139–1145.
9.
Barnes, H. L., and Romberger, S. B., 1968. Chemical aspects of acid mine drainage. Journal WPCF, 40(3), 371 – 384.
10. European Committee for Standardisation (2002). Characterisation of waste—leaching— compliance test for leaching of granular waste materials and sludges— Part 2: One stage batch test at a liquid to solid ratio of 10 L/kg for materials with particle size below 4 mm.
Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract EN 12457-2:2002.
11. Sakurovs, R., French, D., & Grigore, M. (2007). Quantification of mineral matter in commercial cokes and their parent coals. International Journal of Coal Geology, 72, 81–88.
12. Huang, X., & Finkelman, R. B. (2008). Understanding the chemical properties of macerals and minerals in coal and its potential application for occupational lung disease prevention. Journal of Toxicology and Environmental Health. Part B, 11, 45–67.
13. López, I. C., & Ward, C. R. (2008). Coposition and mode of occurrence of minerals matter in some Colombian Coals. International Journal of Coal Geology, 73, 3–18.
14. Silva, L. F. O., & Oliveria, M. L. S. (2010). A preliminary study of coal mining drainage and environmental health in the Santa Catarina region, Brazil. Environmental Geochemistry and Health. doi:10.1007/ s10653-010-9322-x.
15. Devasahayam, S. (2006). Chemistry of acid production in black coal mine washery wastes. International Journal of Mineral Processing, 79, 1–8.
16. Weber, P. A., Skinner, W. M., Hughes, J. B., Lindsay, P., & Moore, T. A. (2006). Source of Ni in coal mine acid rock drainage, West Coast, New Zealand. International Journal of Coal Geology, 67, 214–220.
17. Finkelman, R. B. (1994). Modes of occurrence of potentially hazardous elements in coal: Levels of confidence. Fuel Processing Technology, 39, 21.
18. Bigham, J. M., Carlson, L., & Murad, E. (1994). Schwertmannite, a new iron oxyhydroxysulphate from Pyhasalmi, Finland, and other localities. Mineralogical Magazine, 58, 641–648.
19. Webster, J. G., Swedlund, P. J., Webster, K. S. (1998). Trace metal adsorption onto AMD Fe(III) oxyhydroxysulphate. Environmental Science & Technology, 32, 361–1368.
20. McCarty, D. K., Moore, J. N., & Marcus, W. A. (1998). Mineralogy and trace element association in an acid mine drainage iron oxide precipitate; comparison of selective extractions. Applied Geochemistry, 13, 165–176.
21. Simona, R., Andreas, B., & Stefan, P. (2004). Formation and stability of schwertmannite in acidic mining lakes. Geochimica Et Cosmochimica Acta, 68, 1185–1197.
22. Stoffregen, R. E., Alpers, C. N., & Jambor, J. L. (2000). Alunite–jarosite crystallography, thermodynamics, and geochronology. Sulfate minerals: Crystallography, geochemistry, and environmental significance. Reviews in Mineralogy 40, 453–479 (Mineralogical Society of America).
23. Rigol, A., Mateu, J., Gonzalez-Nunez, R., Rauret, G., & Vidal, M. (2009). pH Stat vs. single extraction tests to evaluate heavy metals and arsenic leachability in environmental samples. Analytica Chimica Acta, 632, 69–79.
24. Kimball, B. A., Callender, E., & Axtmann, E. V. (1995). Effects of colloids on metal
Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract transport in a river receiving acid mine drainage, upper Arkansas River, Colorado, U.S.A. Applied Geochemistry, 10, 285–306.
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Oviedo ICCS&T 2011. Extended Abstract
Modeling of coal slag viscosity: Effect of the volume fraction of solid particles Arne Bronsch, Patrick J. Masset Freiberg University of Mining and Technology, Centre for Innovation Competence Virtuhcon, Reiche Zeche, Fuchsmühlenweg 9, D-09596 Freiberg, Germany, E-Mail corresponding author: [email protected]
Abstract This work deals with the modeling of the viscosity for several slag systems. The viscosity of an artificial composition (SiO2 and CaO) and two ashes from German lignites (silica-rich and silica-free) were measured with a high temperature viscometer up to 1700 °C under Argon. Several models from literature were used to calculate the viscosities without taking into account the influence of solid particles in the slag. Large discrepancies were observed. Therefore, the volume fraction of solid particles was included in the new calculations using the Einstein-Roscoe-Model. Using the Einstein-Roscoe-Model a rather good agreement was observed between experimental data and the calculated values. This allows predicting the viscosity for multiphase slag systems.
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Oviedo ICCS&T 2011. Extended Abstract
1
Introduction
The viscosity of molten slag has been widely investigated since five decades. Systems from different origins like artificial laboratory prepared glasses [1] and ashes [2], coal fly ash [3], magmatic systems [4], [5] or gasifier slags [6] were analysed and viscosity models were developed empirically and/or further improved by other authors. A summary of empirically developed viscosity models is given in [7]. The developed empirical models are based on the principle of the Arrhenius equation. The viscosity is a function of the temperature and the mineral composition (measured at room temperature) of the ash. Several simplifications are commonly used: -
No change of the slag composition during thermal treatment.
-
No influence of the atmosphere on the slag.
-
Effect of solid particles on the slag viscosity is neglected.
For technical applications with small or slow changes in feedstock’s composition, e.g. coal-burning power plants or blast furnaces, these models are sufficient. But for newer applications, the prediction of slag viscosity is important to maintain operation conditions. For example the entrained flow gasifier, which is used to generate syngas for utilization of coal for gasoline production, uses the benefice of a solid slag shield to prevent high temperature corrosion of the inner walls of the reaction chamber. This slag shield consists of a solid and a liquid slag layer. For optimum conditions, the viscosity of the liquid slag layer should range from 12 Pa s to 25 Pa s [8]. If this condition is not fulfilled, the reaction chamber of the gasifier may be damaged [9]. Other fields of viscosity knowledge are material sciences and basic research, respectively. There the Newtonian or non-Newtonian behavior of melts is used to describe the glass transition in macroscopic, mechanical terms [10]. The understanding of the rheological behavior of molten magmatic systems is critical for geophysical problems like mechanisms of magma generation or seismic behavior [11]. Especially for gasifier conditions in reducing atmospheres and at high temperatures, the calculation of slag viscosity fails. Oxides like FeOx are reduced and may form liquid iron and gaseous O2. The liquid iron could influence the viscosity of the slag in several ways, the slag is thinned or new phases with other slag compounds are created. Another factor is the influence of solid particles on the viscosity of slags. Multicomponent slags
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Oviedo ICCS&T 2011. Extended Abstract
form a mixture of solid particles and liquid phases when the liquidus temperature is reached, so that the melt is a mixture of solid and liquid fractions (mushy). The Einstein-Roscoe-Model was used to describe the influence of solid particles on the mixture viscosity. This model depends on the work of Einstein which deals with the estimation of the molecule sizes of a solved material in a non-dissociated dilution [12]. Roscoe extended this work to solid spheres and developed Eq. 1.
η = η 0 (1 − af )− n Where
is the dynamic viscosity with particle influence in Pa s,
viscosity in Pa s, shape and
(1)
and
is the solid-free
are factors depending on the particle size and the particle
being the particle volume fraction in the liquid slag. For spherical particles and
of a uniform size Einstein suggested volume fraction is limited to
to be 1.35 and 2.5, respectively [13]. The
= 0.74 in case of a constant factor a = 1.35. These
values were successfully used for the modeling of the viscosity of glass melts of SiO2Na2O compositions mixed with B2O3 particles [14]. For an examination of the influence of spinel particles (MgAl2O4) to the viscosity of a 20Al2O3-28CaO-10MgO-42SiO2 (in mass-%) mixture with increasing particle fraction (0 to 22 mass-%) at a constant temperature (1646 K/1373 °C) the absolute viscosity increases from 2.2 to 6.3 Pa s at a shear rate of 1 s-1 and a particle size range of 0.1 to 0.22 mm. The effect of increasing particle size for a constant temperature, shear rate and mass fraction of solid particles was investigated, too. It was concluded that the viscosity was increasing (3.5 to 10 Pa s) with increasing particle size (range 0.1 to 0.21 mm and 0.44 to 0.99 mm). For solid loadings over 8 mass-% the slag showed an apparent Bingham behavior. The factors
and
with values of 4.24, 3.29 and 3.56 for
were described as particle-size dependent, and values of 1.28, 2.36 and 2.24 for
when
fine, medium and coarse particles are added. It should be mentioned that a fixed value of 2.5 for
2 2.1
shows good fits to experimental data when only
was varied [15].
Experimental section Sample classification
Three samples were prepared: artificial slag (Slag T) and natural based ashes of German lignites, named Slags A and B (see Table 1). The composition analysis was realized
3
Eliminado: Table 1
Oviedo ICCS&T 2011. Extended Abstract
using X-ray fluorescence method (XRF). The XRF-analysis device is a BRUKER AXS Tiger (wavelength dispersive). X-ray diffraction analysis (XRD) was done on a BRUKER AXS D8 Discover with cobalt tube, equipped with Goebel mirrors, a XYZ Eliminado: Table 2
table and a Våntec-1 detector (see Table 2. Table 1: Composition of the slag in mass-% (trace elements not mentioned). Sample Slag A Slag B Slag T
Al2O3 1.0 2.1 -
CaO 39.2 28.3 31.4
Fe2O3 11.8 12.7 -
K2O 1.6 -
MgO 23.1 2.4 -
Na2O 5.0 -
SiO2 29.6 68.6
SO3 16.8 24.2 -
Total 98.5 99.3 100 Eliminado: -
Table 2: XRD phase analysis on ash (in mass-%). Phase Akermanite Anhydrite Braunmillerite Calcite Diopside Gehlenite Hematite Larnite Lime Maghemite Magnetite Mayenite Periclase Quartz Sodium sulfate Srebrodolskite Srebrodolskite+ Braunmillerite Wollastonite Total
Formula Ca2MgSi2O7 CaSO4 Ca2FeAlO5 CaCO3 CaMgSi2O6 Ca2Al2SiO7 α-Fe2O3 Ca2SiO4 CaO γ-Fe2O3 Fe3O4 Ca12Al14O33 MgO SiO2 Na2SO4 Ca2Fe2O5 Ca2Fe2O5+Ca2FeAlO5 CaSiO3
Ash A 19.4 4.2 22.7 22.7 9.7 21.3 100
Ash B 1.7 40.2 4.4 2.6 0.2 0.7 3.9 0.4 2.4 2.0 0.6 2.0 25.9 9.3 3.7 100
Eliminado:
The ashes from German lignites were prepared by thermal treatment under air at 815 °C for 24 hours. To avoid any exchange of volatiles between the two coals, the samples were prepared separately. After cooling and checking of possible coal/coke odds (which would cause a second ashing) the ashes were weighted and filled into glass flasks. Slag T is a mixture of CaO (Sigma Aldrich, 99.9 % purity) and SiO2 (Sigma Aldrich, 99.8 %) powders fired for 6 hours at 900 °C before mixing. Slag A is silica-free with high amounts of CaO and MgO with a high melting point. Slag B offers a medium content of SiO2 and CaO. Artificial Slag T is developed to form a mixture of solid and liquid phases below a temperature of 1600 °C, based on the CaOSiO2 phase diagram with eutectic temperature close to 1437 °C [16].
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Oviedo ICCS&T 2011. Extended Abstract
2.2
Viscosity measurements
The measurements were realized with different equipment to due the different melting temperature range. All used viscometer devices and the investigated samples are described in Table 3, ¡Error! No se encuentra el origen de la referencia. and ¡Error! No se encuentra el origen de la referencia..
Eliminado: Table 3 Eliminado: Fig. 1 Eliminado: Fig. 2
Table 3: Devices and measurement conditions for investigated samples. Sample
Device
Heating system
Measuring principle
Slag A
Ravenfield HT Viscometer
Graphite heaters
Couette
Slag B
Anton Paar Physica MCR 301
Slag T
Bähr Thermoanalyse GmbH
Induction
Fig. 1: Detail of measuring chamber of viscometer by Bähr Thermoanalyse GmbH.
3
Crucible/Bob Material
Atmosphere
High purity Boron nitride
Nitrogen
Molybdenum
Argon
PtRh30
Air
Searle
Fig. 2: Schematic view of (a) Ravenfield HT Viscometer and (b) Anton Paar Physica MCR 301.
Results and discussion
3.1 Using
Thermodynamic calculations the
software
FactSage® 6.2
by
Thermfact
and
GTT-Technologies,
thermodynamic calculations were performed to determine the fractions of solid and liquid phases as a function of the temperature and initial composition of the ash. The Eliminado: Table 1
starting compositions of the ash are given in Table 1. To estimate the volume fraction, the density
of solid and liquid phases were required. 5
Oviedo ICCS&T 2011. Extended Abstract
The volume fraction f =
of the solid species was calculated using Eq. (2)
Vsolid = Vsolid + Vliquid
∑m
∑m solid ,i
solid ,i
ρ solid ,i
(2)
ρ solid ,i + ∑ mliquid ,i ρ liquid ,i
where V is the volume, m the mass and ρ the density of a solid or liquid component i designed by indices solid or liquid , respectively. The density
for liquid and solid phases was calculated by neglecting the temperature
influence using Eq. (3) ρ phase = ∑ ρi ⋅ xi
(3)
where ρ phase is the density of the solid or the liquid phase summed by the accordant density ρ i of constituents i and the molar fraction xi . The densities of species ρ i in the researched slags were calculated by the work of Lange and Carmichael [17], Lee and Gaskell [18] and Swamy and Dubrovinsky [19]. The volume fractions of solid and liquid slag components for Slag T are given in Fig. 3. The solid phase is consisting of SiO2 (cristobalite) in a temperature range between 1437 °C (eutectic) and 1600 °C (liquidus). The liquid slag fraction is a mixture of CaOSiO2. A slow volume increase of solid particles is determined below liquidus
Volume fraction
temperature up to the eutectic.
1.0 0.8 0.6 0.4 0.2 0.0 1400
1450 Liquid Slag
1500
1550 T in °C Solid particles
1600 Eutectic
1650
1700
Liquidus
Fig. 3: Slag T - Calculated volume fractions of solid and liquid phases.
The volume fraction results of Slag A are given in Fig. 4. A low silica content and high amounts of calcia and magnesia lead to a high liquidus temperature (ca. 2280 °C). Between 1340 °C to 2280 °C several solid phases are occurring, especially magnetite and spinel. Below liquidus temperature there is a strong increase in solid particle volume. Below 1800 °C until the solidification temperature the volume fraction vary
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Oviedo ICCS&T 2011. Extended Abstract
slowly. Calculations by FactSage® 6.2 show liquid slag up to a temperature below
Volume fraction
1100 °C with a mass fraction of < 0.1 %. 1.0 0.8 0.6 0.4 0.2 0.0 1100 1200 1300 1400 1500 1600 1700 1800 1900 2000 2100 2200 2300 T in °C Liquid Slag
Solid particles
Solidification
Liquidus
Fig. 4: Slag A - Calculated volume fractions of solid and liquid phases.
In Fig. 5 calculated volume fractions of solid and liquid phases of Slag B are represented. The temperature range where solid and liquid phases simultaneously exist next to each other is very narrow, ca. 1240 °C to 1320 °C. The formed solid phases in
Volume fraction
the described temperature range are wollastonite, perovskite and pyroxenes.
1.0 0.8 0.6 0.4 0.2 0.0 1100
1150 Liquid Slag
1200
1250 T in °C Solid particles
1300 Solidification
1350
1400
Liquidus
Fig. 5: Slag B - Calculated volume fractions of solid and liquid phases.
3.2
Viscosity modeling
The modeling of viscosity with particle influence was performed by three steps: 1. Calculate viscosity without particle influence by a viscosity model. 2. Define the solid volume fraction by thermodynamic calculations with FactSage® 6.2 software and referenced partial molar densities. 3. Insert the particle volume fraction by Eq. 1 to the results of step 1. For the first step, 9 viscosity models were used which give different results for one sample. In the third step the particle influence was calculated with constant factors and
.
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Oviedo ICCS&T 2011. Extended Abstract
The Slag T viscosity predictions are shown in Fig. 6 for selected models. The
Eliminado: Fig. 6
temperature range where solid and liquid phases exist simultaneously is between the continuous and dashed line and was computed by FactSage® 6.2. Predictions with deviations more than 1 magnitude where excluded from discussions. Closest results of predictions and measurements where observed with the models of Watt-Fereday, Bomkamp and Streeter. Viscosity in Pa s
1000 100 10 1 0.1 1400
1500
T in °C
1600
1700
Slag T Eutectic Solidus Bomkamp Bottinga-Weill Lakatos Sage-McIlroy Streeter Watt-Fereday Fig. 6: Results of viscosity measurements and predictions with particle influence on Slag T.
In case of Slag A an interesting effect is observed. All models underestimate the measured viscosity by two or more magnitudes, Fig. 7. As explained in the introduction
Eliminado: Fig. 7
all models were developed empirically by examinations of silica-rich slags. Slag A is
Viscosity in Pa s
silica-free, during the calculation the absence of silica results in errors on the models. 100 1 0.01 0.0001 1200 1300 1400 1500 1600 1700 1800 1900 2000 2100 2200 2300 T in °C Slag A Bomkamp Lakatos
Solid fraction = 1 Bottinga-Weill S²
Solid fraction = 0 Kalmanovitch-Frank Urbain
Fig. 7: Results of viscosity measurements and predictions without particle influence on Slag A.
The introduction of the particle influence by the Einstein-Roscoe-Model influenced the viscosity prediction (Fig. 8). In comparison to predictions without particle influence, the viscosity values were raised. Fittings of measurements and predictions with a deviation factor of 3 were reached in a small temperature range from 1400 to 1600 °C. The 8
Eliminado: Fig. 8
Oviedo ICCS&T 2011. Extended Abstract
alternating viscosity behavior observed on higher temperatures levels were less predicted by the model of Bottinga-Weill. There is a small viscosity increase with increasing temperature, explained by the used constants tabulated in [20]. The BottingaWeill-Model was developed with magmatic melts. This kind of melts content a high number of components which are similar to slag compositions. In case of the Bomkamp-Model the strong increasing viscosity with increasing temperature was alleviated, an almost linear behavior is calculated. Viscosity in Pa s
1000 100 10 1 0.1 0.01 1200 1300 1400 1500 1600 1700 1800 1900 2000 2100 2200 2300 T in °C Slag A Solid fraction = 1 Solid fraction = 0 Bomkamp Bottinga-Weill Kalmanovitch-Frank Lakatos S² Urbain
Fig. 8: Results of viscosity measurements and predictions with particle influence on Slag A.
Slag B viscosity predictions are given in Fig. 9 by several models. Models like WattFereday, Bomkamp and Lakatos show deviations by a factor of 2 above a temperature of 1315 °C. All selected viscosity predictions increase rapidly below 1350 °C and reach similar values to measured viscosity maximum at ca. 1270 °C. A temperature shift of
Viscosity in Pa s
75 K is observed.
100
1
0.01 1100 Slag B Bomkamp S²
1200
1300
T in °C
1400
Solid fraction = 1 Bottinga-Weill Urbain
1500
1600
Solid fraction = 0 Lakatos Watt-Fereday
Fig. 9: Results of viscosity measurements and predictions with particle influence on Slag B.
9
Eliminado: Fig. 9
Oviedo ICCS&T 2011. Extended Abstract
4
Conclusion
In this work measured and predicted viscosities of three different slag systems were compared. The viscosity predictions were improved by the introduction of the EinsteinRoscoe-Model for describing the influence of solid particles to the viscosity. An important value on the Einstein-Roscoe-Model is the volume fraction of solid particles. The mass fraction of solid particles was calculated by FactSage® 6.2 software. Density parameters for occurring species were taken from literature to calculate the volume fraction finally. Thermodynamic calculations allow to estimate the composition of a slag for several temperatures and was used as input into viscosity models to predict the flow behavior of a slag.
Acknowledgement This research has been founded by the Federal Ministry of Education and Research of Germany in the framework of Virtuhcon (project number 03z2FN12) Eliminado:
References 1. Richet, P. Viscosity and configurational entropy of silicate melts. Geochimica et Cosmochimica Acta 1984, 48, 471–483. 2. Urbain, G.; Cambier, F.; Deletter, M.; Anseau, M.R. Viscosity of silicate melts. Transactions & Journal of the British Ceramic Society 1981, 80, 139–141. 3. Moriyama, R.; Takeda, S.; Onozaki, M.; Katayama, Y.; Shiota, K.; Fukuda, T.; Sugihara, H.; Tani, Y. Large-scale synthesis of artificial zeolite from coal fly ash with a small charge of alkaline solution. Fuel 2005, 84, 1455–1461. 4. Shaw, H.R. Viscosities of magmatic silicate liquids; an empirical method of prediction. Am J Sci 1972, 272, 870–893. 5. Ledzki, A.; Migas, P.; Stachura, R.; Klimczyk, A.; Bernasowski, M. Dynamic viscosity of blast furnace primary and final slag with titanium and alkali admixtures. Archives of Metallurgy and Materials 2009, 54, 499–509. 6. Hurst, H.J.; Novak, F.; Patterson, J.H. Viscosity measurements and empirical predictions for some model gasifier slags. Fuel 1999, 78, 439–444. 7. Vargas, S.; Frandsen, F.J.; Dam-Johansen, K. Rheological properties of hightemperature melts of coal ashes and other silicates. Prog. Energy Combust. Sci. 2001, 27, 237–429. 8. Innes, K. Coal Quality Handbook for IGCC; Cooperative Research Centre for Black Coal Utilisation: Newcastle, 1999. 9. Buhre, B.J.P.; Browning, G.J.; Gupta, R.P.; Wall, T.F. Measurement of the Viscosity of Coal-Derived Slag Using Thermomechanical Analysis. Energy & Fuels 2005, 19, 1078–1083. 10. Webb, S.L.; Dingwell, D.B. The onset of non-Newtonian rheology of silicate melts. Physics and Chemistry of Minerals 1990, 17, 125–132.
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Salto de página
Oviedo ICCS&T 2011. Extended Abstract
11. Shaw, H.R. Rheology of Basalt in the Melting Range. Journal of Petrology 1969, 10, 510–535. 12. Einstein, A. Eine neue Bestimmung der Moleküldimensionen. Annalen der Physik 1906, 324, 289–306. 13. Roscoe, R. The viscosity of suspensions of rigid spheres. British Journal of Applied Physics 1952, 3, 267–269. 14. Reddy, R.G.; Weizenbach, R.N., Eds. The Paul E. Queneau Int. Symp. Extractive Metallurgy of Copper, Nickel and Cobalt, 1993. 15. Wright, S.; Zhang, L.; Sun, S.; Jahanshahi, S. Viscosity of a CaO-MgO-Al2O3-SiO2 melt containing spinel particles at 1646K. Metallurgical and Materials Transactions B 2000, 31, 97–104. 16. Verein Deutscher Eisenhüttenleute, (.; Allibert, M.; Gaye, H.; Janke, D.; Keene, B.J.; Kirner, D.; Kowalski, M.; Lehmann, J.; Mills, K.C.; Neuschütz, D.; et al. Slag Atlas, 2nd ed; Verlag Stahleisen GmbH: Düsseldorf, 1995. 17. Lange, R.A.; Carmichael, I.S.E. Densities of Na2O-K2O-MgO-MgO-FeO-Fe2O3Al2O3-TiO2-SiO2 liquids: New measurements and derived partial molar properties. Geochimica et Cosmochimica Acta 1987, 51, 2931–2946. 18. Lee, Y.; Gaskell, D. The densities and structures of melts in the system CaO-“FeO”SiO2. Metallurgical and Materials Transactions B 1974, 5, 853–860. 19. Swamy, V.; Dubrovinsky, L.S. Thermodynamic data for the phases in the CaSiO3 system. Geochimica et Cosmochimica Acta 1997, 61, 1181–1191. 20. Bottinga, Y.; Weill, D.F. The viscosity of magmatic silicate liquids; a model calculation. Am J Sci 1972, 272, 438–475.
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Oviedo ICCS&T 2011. Extended Abstract
Effect of ash components on devolatilisation behaviour of coal and biomass – product yields, properties and heat requirement D. Reichel, M. Klinger, S. Krzack, B. Meyer TU Bergakademie Freiberg, Department of Energy Process Engineering and Chemical Engineering, Fuchsmühlenweg 9 – Haus 1, 09596 Freiberg, Germany; [email protected] Abstract The influence of inorganic constituents on the pyrolysis behaviour of different biomass materials and brown coals was studied in a fixed bed reactor within a temperature range of 250 und 700 °C. The tests were carried out with untreated and demineralized samples. Inorganic elements dominant in the used brown coals are calcium, silicon, iron, magnesium and for some samples also sodium and alumina. Biomass inorganic constituents mainly involve potassium, silicon and calcium. Nearly total demineralization was accomplished for brown coals via HCl and HF treatment. To prevent exceeding structure determination for biomass materials only HCl was used to remove elements like potassium, magnesium, calcium and phosphorus. Yields, product compositions and HHV were determined for each pyrolysis temperature to create mass and energy balances. Demineralization causes an increase in total liquid yields, while gas yield and char yield (only slightly) decrease. Biomass materials show a stronger effect and main decomposition stage is shifted to higher temperatures. Gas composition is also affected by acid treatment, whereas differences occur between the various fuels. Furthermore, the pyrolysis process becomes more endothermic for the demineralized samples.
1. Introduction Pyrolysis plays an important role within thermochemical conversion processes for the production of heat, power and chemicals. Brown coals and different biomass materials like wood or straw are available for that purpose especially in Germany. Due to the complex macromolecular structure of these fuels a multitude of reactions take place during thermal decomposition and the occurring mechanism are actually not known in detail. But not only fuel structure and their relating composition and properties influence the pyrolysis behaviour. The generated product yields and their properties are also affected by process parameters (e.g. temperature, pressure, heating rate) and the amount
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Oviedo ICCS&T 2011. Extended Abstract
and type of embedding of the present inorganic elements. Some research works deal with the catalytic effect of alkali and earth alkali metals (Na, K, Ca) or transition metals (Ni, Co, Fe) on brown coal pyrolysis [1][2][3][4]. However, much more research effort was spent on influence of various inorganic species on pyrolysis behaviour of different kinds of biomass [5][6][7][8] and the biopolymers cellulose, hemicellulose (xylan) and lignin [5][9][10][11] for upgrading of the produced liquid and gas phase.
2. Experimental Section
2.1. Sample characterization Three German brown coals (Lusatia, Rhineland, Central Germany) as well as three biomass samples (spruce wood chips incl. bark, wheat straw pellets, maize silage pellets) were used in a raw and a demineralized state for the investigations into pyrolysis behaviour in a fixed bed reactor. Ultimate and proximate analyses of raw samples are given in Table 1. Materials were chosen due to their different ash content (1.45 wt.% for spruce wood up to 17.67 wt.% for Central Germany brown coal) and composition (see Table 2) as well as various contents of biopolymers in case of biomass. Table 1 Results of ultimate, proximate analysis and heating value determination of raw samples Proximate analysis in wt.-% M (r) A (d) VM (d) FC (d)
Ultimate analysis in wt.-% (daf) C
H
N
S
O
in kJ/kg (d) HHV
LHV
1
Brown coal Rhineland RBC Lusatia LBC Central Germany CGBC Biomass2 Wheat straw WS
51.12 5.47 50.70 54.00 11.99 52.08
43.83 35.93
69.04 69.54
5.01 5.41
0.79 0.64
0.05 25.11 1.40 23.01
25,267 24,603
24,233 23,565
50.60 17.67 51.78
30.55
72.18
5.81
0.64
2.45 18.93
24,289
23,245
6.81
18.67
49.62
6.06
0.69
0.14 43.49
17,939
16,724
Maize silage MS 13.07 4.16 80.90 14.93 48.32 6.38 1.12 0.11 44.07 Spruce wood 12.53 1.45 78.54 20.01 52.06 5.92 0.25 0.11 41.66 incl. bark SWC 1 determination of ash content at 450 °C; 2 determination of ash content at 550 °C
18,559
17,225
20,154
18,881
8.16
73.17
With regard to potential catalytic activity of inorganic elements main differences mainly occur in sodium, potassium, calcium and iron but also in silicon content. Please note that determination of ash is done under oxidizing conditions according to DIN 51719 for coal and DIN CEN/TS 14775 for biomass. Due to that fact the inorganic elements are usually present in the highest oxidation state, ash contains a lot of oxygen. Under inert respectively reducing conditions during pyrolysis other mineral phases are present. For further considerations and balancing of the pyrolysis process the sum of
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Oviedo ICCS&T 2011. Extended Abstract
inorganic elements (XRF analysis) exclusive total sulphur is used instead of ash content. Thus the oxygen content (calculated as difference) increases. Table 2 Inorganic constituents present in brown coal and biomass samples (XRF analysis) Data given in S C Na Mg Al Si P Cl K Ca wt.% dry feed (total)
Fe
sum traces
Brown coal Rhineland RBC b.d.l. 0.26 0.40 Lusatia LBC 0.05 b.d.l. 0.25 Central Germany b.d.l. b.d.l. 0.17 CGBC Biomass Wheat straw WS 0.41 b.d.l. 0.23 Maize silage MS 0.01 0.02 0.12 Spruce wood 0.11 0.00 0.04 incl. bark SWC
0.05 0.09
0.12 b.d.l. 1.70 b.d.l.
0.64 2.64
0.03 0.02
0.03 0.01
1.21 1.83
0.43 0.90
0.04 0.07
0.96
1.91
0.00
4.00
0.01
0.05
2.99
0.93
0.11
0.02 0.08
1.46 0.81
0.03 0.22
0.20 0.18
0.16 0.03
1.75 0.70
0.38 0.23
0.01 0.09
0.01 0.02
0.01
0.04
0.02
0.11
0.00
0.13
0.41
0.01
0.04
b.d.l. below detection limit
2.2. Demineralization procedure The brown coal samples were demineralized according to the procedure of Samaras et al. [12] using diluted hydrochloric acid (1:1) for 1 h first. To reach nearly complete demineralisation (removal of SiO2, Al2O3) diluted hydrofluoric acid (8%) was added to the sample for 1 h. At last another washing step with HCl for 1 h was enclosed to remove residues of HF. Demineralisation of the biomass samples was done according to Vamvuka et al. [13] with diluted HCl (1:1) for 1 h. Washing with distilled water was accomplished after the acid treatment until the chlorine concentration in the remove liquid can be neglected. Total demineralization can´t be reached for biomass because treatment with HF will lead to strong changes in biomass structure.
2.3. Pyrolysis equipment The pyrolysis reactor set-up consists of a fixed bed reactor (di=20 mm, l=335mm), a vertically movable tube furnace (max 1.15 kW), two cooling traps designed like washing flasks (1st ice water cooled, 2nd cooled by saturated NaCl solution), a mass flow controller to set the purge gas stream, a gas collection system composed of six gas sampling tubes (each with a volume of 1 l) and two filters. It is possible to take a gas sample of each gas sampling tube by using gas sampling bags. The second cooling trap is filled with a defined amount of tetrahydrofuran acting as an absorbent to realize a nearly complete precipitation of condensable products.
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Oviedo ICCS&T 2011. Extended Abstract
2.4. Experimental procedure Pyrolysis investigations were carried out with raw and demineralized brown coal and biomass samples within a temperature range between 250 and 700 °C. The samples were grinded to a particle size < 2 mm and dried at 105 °C until weight constancy. The fixed bed reactor was prepared with 20 to 30 g of dried sample to reach nearly the same sample volume for each experiment. The furnace preheating temperature was set to a value of 200 K above desired pyrolysis temperature to achieve a high heating rate (up to 120 K/min). After flushing the reactor with Argon to remove air from the system the furnace was pulled over the reactor and the sample was heated up. The furnace was removed slightly before the sample thermocouple inside the reactor reached the desired temperature. The whole system is flushed with a constant Argon stream of 50 ml/min (STP) during the whole experiment. After cool down all the parts of the system were detached, weighed and char as well as liquid product were recovered for further analyses.
2.5. Product analysis The composition of pyrolysis gas was analysed via gas chromatography (Micro GC Agilent 3000A, 3 columns: mole sieve, PLOTU, alumina). For char characterization ultimate, proximate and XRF analysis as well as heating value determination was done according to German standardization. The removed tar from the cold traps was characterized via ultimate analysis and heating value determination. The water content of the generated tars/oil was determined by Karl Fischer titration. The titrant used was HYDRANAL Composite 5K together with HYDRANAL KetoSolver as solvent.
3. Results and Discussion 3.1. Level of demineralization Results of XRF analysis of raw and demineralized samples were used to calculate level of demineralization obtained by acid treatment (see Figure 1). Carbonate, sodium, magnesium and for the brown coal samples also phosphorus were completely removed. Differences in level of demineralization for the used fuels are the result of various types of embedding of inorganic elements (bound to organic matter or as mineral grains like carbonates, sulphates, silicates and so on) and their variable resistance against acid treatment. Compared to brown coals (>91.6 %) the biomass samples reach a significantly lower level of demineralization (WS: 34.1 %; MS: 47.7; SWC: 84.3) because no
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Oviedo ICCS&T 2011. Extended Abstract
treatment with HF was done to remove the silicon and aluminium containing species. The relatively high value for wood is related to the fact that calcium carbonate is the dominant inorganic species, while less silicon is present. However, the elements supposed to have a catalytic activity (Na, K, Mg, Ca) are removed to a satisfying amount.
Level of demineralisation in %
80
80
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20 RBC
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MS
SWC
0
0 Al
Si
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Fe traces total
Mg
P
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Ca
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Fe traces total
Figure 1 Demineralisation levels obtained by acid treatment (HCl, HF) of brown coal and biomass
3.2. Product yields In general, comparison of pyrolysis behaviour from brown coals and biomass shows significant differences, due to the fact of varying compositions (see Table 1). While brown coals possess a very continuous decrease in char yield according to rising temperatures (s. Figure 2 a), biomass shows a main decomposition stage within a narrow temperature range, starting at 200 to 250 °C and ending at 350 to 400 °C with a mass loss up to about 70 wt.% (daf) (s. Figure 2 b). Wheat straw starts pyrolysis first followed by maize silage and spruce wood. The influence of demineralization on pyrolysis product yields is exemplary shown in Figure 2 for Rhenish brown coal and wheat straw, since similar trends occur within the material groups. Brown coals possess only a slight decrease in char yield by demineralization, which is not steady within the investigated temperature range (Figure 2 a). Less char amounts after acid treatment are also reported from [13], but no effect during TG experiments was found by [2]. On contrary, liquid yields increase to the disadvantage of gas, whereat reaction water shows an increase (Figure 2 c) and in case of tar/oil no similar behaviour can be observed (RBC ↓, LBC and CGBC ↑). Increase in tar/oil yield for CGBC and LBC corresponds to [4]. The biomass materials show a stronger influence caused by acid treatment. For the herbaceous ones (WS, MS), containing high amounts of hemicellulose, pectin and cellulose, the main decomposition
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Oviedo ICCS&T 2011. Extended Abstract
stage is shifted to higher temperatures (strong influence for WS, see Figure 2 b). Due to high amounts of potassium in raw biomass, cellulose pyrolysis is shifted to lower temperatures as already reported by [5][7]. Finally, less char remains for demineralized samples (700 °C: WS -17.7; MS -10.4; SWC -10.2 %), since potassium is supposed to catalyse polymerization reactions [11]. Production of total liquid amount is significantly rose by acid treatment (700 °C: WS +24.9; MS +10.2; SWC +9.3 %), while gas yields decrease about 11 % for SWC and up to 30.0 % for WS at 700 °C (s. Figure 2 b and d). The obtained trends correspond to the work of [7] and partly to that of [6][8]. The strong increase of liquids is caused by an enhancement of tar/oil, whereas reaction water production is nearly unaffected except for SWC which shows a slight rise up to 500 °C. a)
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Figure 2 Pyrolysis product yields for Rhenish brown coal (RBC, left column) and wheat straw (WS, right column). Solid lines for raw samples, dashed lines for demineralized samples. a), b) char and gas yields. c), d) liquid product yields (tar/oil, reaction water and total liquid).
3.3. Gas yields and composition Main gas components being released during pyrolysis of brown coal and biomass are CO2, CO, H2 and CH4. Furthermore, saturated and unsaturated hydrocarbons (C2 to C4) as well as H2S and COS, the latter especially in the case of brown coals, are produced. In general, demineralization leads to lower CO2 and H2 yields during pyrolysis, while
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Oviedo ICCS&T 2011. Extended Abstract
CH4 and CO are increased for brown coals (see Figure 3). The different biomass fuels show various trends with subject to temperature. Acid treatment causes a slight shift to higher temperatures for the start of CO2 production in brown coal (except for LBC) and biomass pyrolysis, whereas the trend according to temperature is similar for raw and
yield in l(STP)/kg Feed (daf)
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demineralized samples, but as already depicted at lower yields.
0 250
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Figure 3 Released cumulative yields of main gas components during brown coal pyrolysis. a) CO2, b) CO, c) H2, d) CH4. Solid lines for raw samples, dashed lines for demineralized samples.
Especially in case of wheat straw pyrolysis the effect on CO2 yield caused by
demineralization is very strong and the released amount of the raw sample is bisected. On contrary, CO production increases for demineralized brown coals, whereas start temperature is not affected. With subject to increasing temperatures CO is released in two phases, which can be seen especially for LBC and RBC (Figure 3 b). The temperature for the first plateau is shifted to about 100 K higher values by acid treatment. The increase in CO and decrease in CO2 yields during brown coal pyrolysis corresponds to results of Zou et al. [1], who found that presence of calcium (present in high amounts in the used brown coals) leads to lower levels in CO and aliphatic hydrocarbons. The latter fact is also closed to the results of this study, where a rise of saturated hydrocarbons (C2–C4) was obtained. On the contrary, Yang et al. [3] confirm only the behaviour of CO2 by adding Ca and Fe to untreated lignite.
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Oviedo ICCS&T 2011. Extended Abstract
CO release during biomass pyrolysis is shifted to higher temperatures. Demineralized SWC as well as WS produce a similar amount of CO which is below the value for raw samples up to 600 °C and rises further for increasing temperatures while the raw feeds show a plateau phase. These characteristics are also reported by di Blasi et al. [6]. The obtained CO yield for demineralized maize silage was on contrary below the raw sample. Methane yields are also affected by demineralization (Figure 3 d), but the trends vary for brown coal and biomass. CH4 production starts in significant amounts at about 300 °C for raw and demineralized brown coals whereupon the trends with subject to temperature are very similar. But, in case of demineralized brown coals higher CH4 yields are produced, which is contrary to the results of [3] and to the behaviour of biomass samples. The demineralized biomass materials produce a significantly smaller amount of CH4 above 400 °C for SWC and above 500 °C for WS. Differences range between 12 (SWC) and 20 l(STP)/kg Feed (daf) for WS. Indeed, the demineralized maize silage delivers a higher methane yield than the raw sample. In dependence on pyrolysis temperature the methane yield shows a maximum at around 600 °C for WS and MS. Yields
of
saturated
gaseous
hydrocarbons
(C2H6, C3H8, C4H10)
decrease
for
demineralized biomass and behave the other way around for brown coal. In case of unsaturated hydrocarbons (C2H4, C3H6, C3H4) yields decline for CGBC and LBC, but rise for the other fuels. This may be attributed to missing alkali metals present in the raw biomass samples and RBC.
3.3. Energy distribution Based on the determined yields and higher heating values of pyrolysis products the energy distribution each product as well as the difference between chemically bounded energy within products and feed ΔEchem can be calculated. Results are shown exemplarily for pyrolysis at 400 °C in Table 3. For liquids an increase in chemically bounded energy occurs from raw to demineralized samples, but no general trend could be observed for the other products. This is a result of different levels of influence by acid treatment on product yields and composition between and within the different material groups. The general trend which arises for the effect of demineralization on ΔEchem is that occurring reactions are going to be more endothermic than for untreated samples within the investigated temperature range. At 400 °C the values change from partly strongly exothermic (e.g. LBC, WS) to slightly endothermic for all samples except CGBC. Due to
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Oviedo ICCS&T 2011. Extended Abstract
the fact, that the creation of a complete energy balance for the pyrolysis process including chemically bound and physically enthalpies is actually not possible because of missing data for e.g. heat capacities, enthalpy of evaporation of tar and so on. It should be noted that the given values for ΔEchem can only be considered as trends. Table 3 Energy distribution (absolute values) to pyrolysis products at 400 °C, ΔEchem as well as trends in heat requirement (basis of calculation: yields YP,i (daf) and HHVP,i (daf)) product yield*product HHV trend difference char liquid, total tar/oil reaction water gas products-feed in kJ/kg (daf) in % raw 21,180 4,117 4,064 53 268 -370 -1.5 exo RBC de 21,480 4,355 4,259 96 318 368 1.5 endo LBC
raw de
19,409 19,276
4,854 4,864
4,754 4,685
100 178
595 490
-1,019 46
-4.1 0.2
exo endo
CGBC
raw de
21,372 20,776
4,079 5,606
4,007 5,560
72 46
267 307
-437 -159
-1.7 -0.6
exo exo
raw de raw
9,453 8,116 8,356
7,796 12,511 10,104
7,377 11,944 9,534
419 566 571
668 387 495
-877 1,758 -45
-4.7 9.8 -0.3
exo endo exo
de raw de
8,843 11,801 11,195
11,338 7,971 9,116
10,807 7,527 8,616
531 444 499
343 326 339
495 -223 339
2.6 -1.2 1.8
endo exo endo
WS MS SWC
4. Conclusions Amount and composition of inorganic species have a significant impact on the thermal degradation characteristics of brown coals as well as biomass materials. Due to the fact of various macromolecular structure, varying ash content, and dominant inorganic species, the results differ between brown coals and biomass fuels and furthermore within the two material groups. The higher tar/oil yields produced during pyrolysis of demineralized biomass, to the disadvantage of gas (mainly decrease of CO2, CO, H2), indicate that inorganic matter (K, Ca) may act as catalysts for the breakdown of oxygen containing functional groups via decarboxylation or decarbonylation. Thermal decomposition of acid treated biomass produces higher amounts of light oxygen bearing species e.g. phenol, organic acids or methanol (see [5]) leading to an increase in tar/oil oxygen content. The influence of mineral matter on degradation characteristics for brown coal results mainly from high amounts of calcium, iron and sodium (only RBC). The main effect is an increasing reaction water production caused by demineralization. This leads to higher amounts of total liquids to the disadvantage of gas and char. Strong influence was also observed in gas composition. The decrease in CO2 and the increase in CO by acid treatment may be related to missing calcium, acting as catalyst for the
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Oviedo ICCS&T 2011. Extended Abstract
formation of acidic oxygen bearing species (see [1]). This is confirmed by the observed higher oxygen content in tar/oil of acid treated brown coals. The obtained increase in endothermic reactions after demineralization is associated with rising liquid yields. This indicates the effect of inorganic constituents, present in brown coal and biomass, on the occurring decomposition mechanisms.
Acknowledgement. The investigations were financially supported by the German Federal Ministry of Agriculture, Food and Consumer Protection (biomass pyrolysis) as well as from the German Federal Ministry of Research, Vattenfall, RWE, Mibrag and Romonta within the research project “German centre for energy raw materials”. References [1] Zou X, Yao J, Yang X, Song W, Lin W. Catalytic effect of metal chlorides on the pyrolysis of lignite. Energ Fuel 2007;21:619-24 [2] Liu Q, Hu H, Zhou Q, Zhu S, Chen, G. Effect of inorganic matter on reactivity and kinetics of coal pyrolysis. Fuel 2004;83:713–8 [3] Yang JB, Cai, NS. A TG-FTIR study on catalytic pyrolysis of coal. J Fuel Chem Tech 2006;34:650–4 [4] Tyler RJ, Schafer HNS. Flash pyrolysis of coals: influence of cations on devolatilization behaviour of brown coals. Fuel 1980;59:487–94 [5] Jensen A, Dam-Johansen K. TG-FTIR study of potassium chloride on wheat straw pyrolysis. Energ Fuel 1998;12:929–38 [6] Di Blasi C, Branca C, D´Errico G. Degradation characteristics of straw and washed straw. Thermochim Acta 2000;364:133–42 [7] Raveendran K, Ganesh A, Khilar KC. Influence of mineral matter on biomass pyrolysis characteristics. Fuel 1995;74:1812–22 [8] Wang Z, Wang F, Cao J, Wang J. Pyrolysis of pine wood in a slowly heating fixed-bed reactor: Potassium carbonate versus calcium hydroxide as a catalyst. Fuel Process Technol 2010;91:942–50 [9] Patwardhan PR, Satrio JA, Brown RC, Shanks BH. Influence of inorganic salts on primary pyrolysis products of cellulose. Bioresource Technol 2010;101:4646–55 [10] Varhegyi G, Antal Jr MJ. Simultaneous thermogravimetric-mass spectroscopic studies of the thermal decomposition of biopolymers. 1. Avicel cellulose in the presence and absence of catalysts. Energ Fuel 1998;2:267–72 [11] Nowakowski DJ, Jones JM. Uncatalysed and potassium-catalysed pyrolysis of cell-wall constituents of biomass and their model compounds. J Anal Appl Pyrol 2008;83:12–25 [12] Samaras P, Diamadopoulos, E, Sakellaropoulos GP. Acid treatment of lignite and its effects on activation. Carbon 1994;32:771–6
[13] Vamvuka D, Troulinos S, Kastanaki E. The effect of mineral matter on the physical and chemical activation of low rank coal and biomass materials. Fuel 2006;85:1763–71
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Oviedo ICCS&T 2011. Extended Abstract
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Oviedo ICCS&T 2011. Extended Abstract
La Pereda CO2: a 1.7 MW pilot to test postcombustion CO2 capture with CaO A. Sánchez-Biezma1, J Paniagua1, L. Diaz2, E. de Zarraga2, J. López3, J Alvarez3, B. Arias* 4, M. Alonso4, J.C. Abanades4 1 Endesa Generación, Riberal del Loira 60, 28042 Madrid (Spain) 2. Hunosa, Avenida de Galicia 44, 33005 Oviedo (Spain) 3.Foster Wheeler Energía S.L.U, Gabriel García Márquez, 2, 28232 Madrid (Spain) 4. Spanish Research Council, CSIC-INCAR* (Spain) * [email protected] Abstract Postcombustion CO2 capture using CaO as a regenerable solid sorbent is a new and rapidly development technology. Previous studies have shown that the technology has the potential to achieve a substantial reduction in capture cost and energy penalties respect to stand-alone oxy-fired systems. However, for these benefits to be realized, it is essential to demonstrate the operation under realistic conditions in terms of solid circulation flows between reactors and purge rates of deactivated sorbent and fuel ashes. This communication describes the 1.7 MWt pilot test facility built in La Pereda power plant (Spain) to test the concept. The experimental work-plan includes the operation of the CFBC calciner in oxy-combustion mode under different O2/CO2 ratios, with O2 and CO2 supplied from liquefied tanks. The experiments with different limestones, coals with different compositions (including high S and ash content) and a range of operating conditions in stationary state and in dynamic conditions, will supply the necessary data for model validation purposes and scaling up. If results are successful at this 1.7 MWt scale, and the enabling oxyfired CFB technology develops at the expected pace, this technology could be ready for demonstration at large scale before 2020. 1. Introduction CO2 Capture and Storage (CCS) is recognised as a critical technology for combating climate change within a portfolio of technologies, including greater energy efficiency and renewable energy. In parallel to the commercial deployment of the most mature CO2 capture technologies, it is necessary to develop a new generation of technologies able to reduce the energy penalty and cost associated to the capture process. Between the new emerging technologies, there is a category which includes new processes that use new materials in reactors or systems already commercially established on a large scale. Calcium looping fits in this category of technologies, as it relies in the utilization of Submit before May 15th to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
circulating fluidized bed reactors operating with circulating materials (mainly CaO particles and their derived products and coal ashes) and key operating conditions close to those present in large scale CFB boilers. Calcium looping uses the reversible reaction of CO2 with CaO to separate it from a flue gas stream in a cyclic process. The best CaO precursor for large scale CO2 capture in power plants are natural limestones and dolomites due to its low cost. The CO2 absorption capacity of these CaO natural sorbents is known to decay rapidly with the number of carbonation calcination cycles [1] and a large body of literature is now available trying to design methods to overcome it [2]. However, it is now well known [3] that there is a residual activity of the CaO particles that allows for the operation of the capture loop by using sufficiently large solid circulation rates. The objective of this work is to describe the design of a 1.7 MWth pilot plant aimed at the capture of 70-95% of the CO2 contained in a side stream of the flue gases from the existing CFB power plant. 2. Description of the process In a postcombustion Ca-looping capture process, CO2 is captured from the combustion flue gas of a power plant by using CaO as sorbent in a circulating fluidized bed carbonator operating between 600-700ºC. The stream of partially carbonated solids leaving the carbonator is directed to a second circulating fluidized bed where solids are calcined to regenerate the sorbent (CaO) and to release the CO2 captured in the second CFB. In this reactor, coal is burned under oxy-fuel conditions at temperatures above 900 ºC to calcine the CaCO3 formed in the carbonator and to produce a highly concentrated stream of CO2. Therefore, exhaust gases from the calciner contains the CO2 captured from the flue gases of the power plant and the CO2 resulting form the oxy-fired combustion of coal in the calciner. The option of calcining the sorbent using an oxyfired circulating fluidized bed, first proposed by Shimizu et al [4], is one of the strengths of this process and may speed up the development of calcium looping technology as oxy-combustion CFB technology is already a more mature technology in a nearcommercial stage [5]. One of the main features of this process is the possibility of producing additionally power from the different high temperature sources involved in the system. Thus, calcium looping technology shows a low energy penalty compared with other CO2 capture technologies [6, 7].In fact, the calciner can be viewed as a new oxy-fired fluidized bed which increases the total installed power.
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Oviedo ICCS&T 2011. Extended Abstract
Some experimental works for testing calcium looping in small fluidized bed pilot plants with different configurations at small scale (10s KW) have been published in recent years demonstrating the concept and showing that CO2 can be effectively removed using CaO under realistic operation conditions in a CFB carbonator reactor [8, 9, 10, 11]. The experimental information obtained in these small facilities has shown that the inventory of active solids in the carbonator bed has been showed as one of the most important variables [12, 13, 14]. The quantitative information and the modelling work available from laboratory studies on the carbonator reactor have been used to design the 1.7 MWt pilot plant. However, the importance of the oxy-fuel CFB calciner, as well as the need to obtain in the pilot plant certain parameters (bed inventories, and solid circulation rates) at specific conditions, has required the application of all necessary know how on CFB technology achieved by Foster Wheeler during decades of experience in the design and construction of CFB boilers. The main features of the pilot plant are described in the next section. 3. Pilot plant design An agreement between ENDESA (a major European utility), Foster Wheeler (a leading world manufacturer of fluidized bed combustion technology), HUNOSA (the biggest coal mining company in Spain and owner of a CFB power plant) and CSIC (Spanish Research Council) was signed in 2009 to progress in the industrial application of Calooping. This consortium was substantially reinforced by leading R&D partners in the field in Europe (IFK of the U. Stuttgart in Germany, Lappeenranta University in Finland, Imperial College in the UK) and in Canada (University of Ottawa), through the project “CaOling” (Development of postcombustion CO2 capture with CaO in a large testing facility, www.caoling.eu), funded by the European Union 7th Framework Programme-FP7 (Dec 2009-Nov 2012). The pilot plant is being built and integrated with “La Pereda” power plant which consists in a circulating fluidized boiler with an installed capacity of 50 MWe. The plant is located in Asturias (North of Spain) and is owned by HUNOSA. Following the objective of reproduce operation conditions representative to those encountered on industrial CFB boilers, a minimum size of the facility was established according to standard design criteria of this type of reactors and the pilot plant was dimensioned to treat a maximum flow of flue gas corresponding to the generation of 1.7 MWth in “La Pereda” power plant
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Oviedo ICCS&T 2011. Extended Abstract
(about 2400kg/h). The most important part of the CO2 capture pilot plant consists in two interconnected circulating fluidized bed (see Figure 1). Typical operation temperatures are around 650 ºC in the carbonator and 920 ºC in the calciner. Each reactor is equipped with a high efficiency cyclone (cut size 5 μm) and a loop seal which design allows controlling the solid flow between reactors. The pilot plant was designed to work in an operating window where process variables can be varied as indicated in Table 1. The variables chosen to define the operation window were the flow of fresh limestone, the heat demand in the calciner and the solid circulation. Once defined the gas flow to be treated in the pilot plant and the desired capture efficiency in the carbonator, the rest of the variables involved in the process can be calculated by solving the respective mass and energy balances. Regarding the dimensions of the reactors, the diameter was fixed to achieve gas velocities similar to those encountered in CFB boilers (3-5 m/s). To ensure enough gas residence time and inventory of solids, the risers were designed with a height of 15 m. The maximum flow of limestone was fixed to a maximum value of 1 kgCaCO3/kg coal fed in the calciner as these scenarios of high purge of solids can acceptable if it is used a feedstock in the cement industry. On the other hand, the minimum make up flow was determined by fixing the maximum heat demand in the calciner to a value of 1.7 MW. Table 1. Main inputs to the pilot plant. Flue gas flow to carbonator (kg/h) Maximum coal flow to calciner (kg/h) Maximum fresh limestone flow (kg/h) Oxygen flow to calciner (kg/h) CO2 flow to calciner (kg/h) Air flow to calciner (kg/h)
680-2400 325 300 300-600 700-2250 600-2500
Circulation rates expected inside the operation window can vary between 5-10 kg/m2s, which are in the typical range of CFB reactors. These solid circulations between reactors allows for effective CO2 capture with a conservative value of calcium conversion to CaCO3 in the carbonator. Other important aspect of the design was the heat dissipation in the carbonator, which is the addition of two terms: the flow of reacting CO2 (exothermic carbonation reaction) and the hot solids circulating from the calciner. It was calculated that the heat dissipation requirements in the carbonator can vary between 0.1 and 0.7 MWth. A water cooling systems using double-pipe cooling retractable tubes was Submit before May 15th to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
designed to control the total heat extracted in the reactor. This system gives a great flexibility and permits to operate the plant in a wide range of conditions. A stream of flue gas from “La Pereda” will be taken after the electrostatic precipitator and will be sent to the CO2 capture pilot plant. A fan with an average gas flow of 1400 kg/h will be used to increase the pressure of the flue gas before entering the carbonator. The gas leaving the carbonator at high temperature (approximately 650 ºC) will be cooled down and returned to main flue gas stream of “La Pereda” power plant before the electrostatic precipitator. No final purification of CO2 is planned in this pilot. Therefore the gas stream from the calciner will be mixed with the decarbonated flue gas leaving the carbonator before being cooled down. The coal feeding system was dimensioned to introduce a maximum flow rate of 325 kg/h. The system can use different types of coals or pretreated solid fuels. The fuel is discharged in a feed hopper using a bigbag handling system. From the feed hopper, coal is pneumatically transported to an intermediate bin. From this bin, coal is discharged to a common hopper where limestone and coal is mixed. A rotary feeder isolates the mixture feeding from furnace overpressure. Finally a screw feeder drives the solids into the calciner bed. As was indicated before, coal can be burned in the calciner using an oxy-fired or air-fired mode. O2 and CO2 is supplied from tanks of liquefied gases using several atmospheric vaporizers and an electrical acconditioner. O2 and CO2 flows are controlled and blended in a mixer skid, enabling to modifify O2 and CO2 concentration. The temperature of this flow is increased in a gas heater. An additional steam line joins to the flow previous to the calciner inlet to simulate different compositions of the gas, increasing the operational flexibility of the plant. In order to work under air-firing mode, a fan will be installed to supply the air needed for the combustion of coal. To work inside the limits of the operating window, the limestone feeding system was dimensioned to give a maximum mass rate of 300 kg/h. Limestone reception is discharged directly from trucks to a feed silo. From this silo, limestone is pneumatically transported to an intermediate bin before being mixed with coal. To remove the solids from the system, two pneumatically slide gates are located under the reactors. The calciner has an automatic purge removal system which consists in a water cooled screw conveyor. This screw conveyor discharges the cold ashes to a chain conveyor that leaves the material to a container. The carbonator material purge can only be made manually Submit before May 15th to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
and after stop the plant. The instrumentation of the pilot plant was defined attending to control requirements and to obtain information to analyze the experimental results (inventory of solids, solid samples, gas composition, temperature…). There will different ports along the risers, stand pipes and loop seals to measure different parameters inside the reactors and to collect solid samples. Gas composition will be analyzed at the exit of each reactor and along the risers.
Figure 1. Photograph of the 1.7 MWth pilot plant The first experimental campaign will start in the second half of 2011, and will run for 12 months. The experimental plan will explore the impact of the main operation conditions on CO2 capture efficiency and other pollutant emissions. Also, due to the nature of calcium looping technology, which involves a stream of sorbent circulating between two interconnected CFB reactors, the controllability and stability of the interconnected solid circulation system is a key aspect that has to be analysed and evaluated in this experimental facility during the testing program. From the information and experience obtained during the operation of the 1.7 MWth pilot plant, a conceptual design of a preindustrial plant of 20-30 MWth will be carried out. The construction of this medium size plant will be the next step in the scaling up route, if the results obtained in the 1MWth pilot plant are successful and the validity of calcium looping under realistic CFB
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Oviedo ICCS&T 2011. Extended Abstract
conditions is demonstrated.
4. Conclusions Calcium looping is a promising technology for post-combustion CO2 capture. In order to scale calcium looping at large scale, a 1.7 MWth pilot plant has been design and is being constructed close to a 50 MWe CFB power plant in Asturias (Spain). This facility is expected to enter into full operation in the first half of 2011. The aim of this plant will be to advance in the experimental validation of this technology and to speed up the development of the calcium looping towards commercial scale. It has been possible to design this pilot using design information, equipment and material that are standard in the existing large scale CFBC power plants. This highlights the possibility for a rapid scaling up of calcium looping technology if the testing program on this pilot confirms the good results obtained so far in small lab scale units.
Acknowledgements This work is being funded by the European Commission 7th Framework Programme under the CaOling project
References [1] Curran G.P., Fink C.E., Gorin E. CO2 acceptor gasification process. Studies of acceptor properties. Adv. Chem. Ser. 1967; 69: 141-161. [2] Blamey J., Anthony E.J., Wang J., Fennel P.S. The calcium looping cycle for largescale CO2 capture. Progress in Energy and Combustion Science 2010; 36: 260-279. [3] Grasa G.S., Abanades J.C. CO2 capture capacity of CaO in long series of carbonation/calcination cycles. Ind. Eng. Chem. Res. 2006; 45: 8846-8851. [4] Shimizu T., Hirama T., Hosoda H., Kitani K., Inagaki M., Tejima K. A twin-bed reactor for removal of CO2 from combustion processes. Trans. IChemE. 1999; 77 (part A): 62-68. [5] Myöhänen K., Hyppänen T., Pikkarainen T., Eriksson T., Hotta A. Near Zero CO2 emissions in coal firing with oxyfuel CFB boiler. Chem. Eng. Technol. 2009; 3: 355363. [6] Romeo L.M., Abanades J.C., Escosa J.M., Paño J., Jiménez A., Sánchez-Biezma A., Ballesteros J.C. Oxyfuel carbonation/calcination cycle for low cost CO2 capture in existing power plants. Energy Conversion Management 2008; 49: 2809-2814. Submit before May 15th to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
[7] Ströle J., Galloy A., Epple B. Feasibility study on carbonate looping process for postcombustion combustion CO2 capture from coal-fired power plants. Energy Procedia 2009; 1: 1313-1320. [8] Abanades J. C., Rubin E. S., Anthony E. J. Sorbent cost and performance in CO2 capture systems. Ind. Eng. Chem. Res. 2004; 43 (13): 3462-3466. [9] Lu D.Y., Hughes R.W., Anthony E.J. Ca-based sorbent looping combustion for CO2 capture in pilot-scale dual fluidized beds. Fuel Processing Technology 2008; 89: 13861395. [10] Charitos A., Hawthorne C., Bidwe A.R., Sivalingam S., Schuster A., Spliethoff H., Scheffknecht G. Parametric investigation of the calcium looping process for CO2 capture in a 10kWth dual fluidized bed. International Journal of Greenhouse Gas Control 2010; 4: 776-784. [11] Alonso M., Rodríguez N., González B., Grasa G., Murillo R., Abanades J.C. Carbon dioxide capture from combustion flue gases with a calcium oxide chemical looping. Experimental results and process development. Int. Journal of Greenhouse Gas Control 2010; 4: 167-176. [12] Hawthorne C., Charitos A., Perez-Pulido C.A., Bing Z., Scheffkenecht G. Design of a dual fluidized bed system for the post-combustion removal of CO2 using CaO. Part I. CFB carbonator model. Proceedings of the 9th International Conference on Circulating fluidized Beds, Hamburg, Germany 2008, 759-764. [13] Alonso M., Rodríguez N., Grasa G., Abanades J.C. Modelling of a fludized bed carbonator reactor to capture CO2 from a combustion flue gas. Chemical Engineering Science 2009; 64: 883-891. [14] Rodríguez N., Alonso M., Abanades J.C. Experimental investigation of a circulating fluidized bed reactor to capture CO2 with CaO. AIChE Journal 2010b; 57: 1356-1366.
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Oviedo ICCS&T 2011. Extended Abstract
Fluidized Bed Desulfurization using Lime Obtained after Slow Calcination of Limestone Particles
F. Scala1, R. Chirone1, P. Meloni2, G. Carcangiu3, M. Manca4, G. Mulas4, A. Mulas4 1
Istituto di Ricerche sulla Combustione – CNR, P.le Tecchio 80, 80125 – Napoli (Italy). E-mail: [email protected]; 2 Università di Cagliari, Dipartimento di Ingegneria Chimica e Materiali, Piazza d’Armi 1, 09127 – Cagliari (Italy); 3 Istituto di Geoingegneria e Tecnologie Ambientali – CNR, Piazza d’Armi 1, 09127 – Cagliari (Italy); 4 Calcidrata S.p.A – Via Valsugana 6, 09123 – Cagliari (Italy). Abstract In this work we have tested the fluidized bed desulfurization performance of lime particles obtained by means of a proprietary limestone slow calcination pre-treatment technique. This performance was compared to that of the parent untreated limestone particles. The occurrence of particle fragmentation and attrition during the fluidized bed operation was also investigated with a specific test protocol for both raw limestone and pre-treated lime sorbent. Two particle size ranges were tested under typical fluidized bed coal combustion conditions. The experiments were complemented by porosimetric and morphological (SEM) analyses of the sorbent. Results showed that limestone pretreatment was able to preserve the high mechanical strength of the parent particles as opposed to the fast in situ calcination typically active in fluidized beds. In addition, a high calcium reactivity and final conversion were observed for the pre-treated lime particles, leading to a SO2 capture capacity per unit mass of sorbent much higher than that obtained with the untreated limestone. Simple economic evaluations suggest that the use of the pre-treated lime in place of limestone can involve significant economies for fluidized bed coal combustor operators.
1. Introduction Removal of sulfur oxides generated during fluidized bed (FB) combustion is typically accomplished by in situ injection of limestone or dolomite [1]. At atmospheric pressure, sorbent particle sulfation proceeds according to the following reactions: CaCO 3( s ) ⇔ CaO ( s ) + CO 2( g )
(1)
CaO( s ) + SO2( g ) + 1 2 O2( g ) ⇔ CaSO4 ( s)
(2)
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Oviedo ICCS&T 2011. Extended Abstract
i.e., the sorbent first calcines (reaction 1) to yield porous calcium oxide which, in turn, is able to remove SO2 (reaction 2) producing calcium sulfate. Particle sulfation most typically conforms to a core-shell sulfation pattern: a sharp reaction front establishes in the sorbent particles between the porous unreacted CaO core and the compact CaSO4 outer shell [2-6]. Extensive sulfation of the core is prevented by the onset of a strong diffusional resistance to SO2 migration through the shell. Calcium conversion seldom exceeds 30-40%, so that over-stoichiometric sorbent feeding is required, resulting in the increased production of solid waste. The particle size distribution (PSD) of the bed material is an important factor in FB combustors, as it affects fluid-dynamics, heat transfer and pollutants formation. When the bed material contains an SO2 sorbent, like limestone or dolomite, its PSD also affects the desulfurization efficiency in the boiler. If the sorbent particles are too fine, they rapidly escape as fly ash, and calcium conversion decreases because of the insufficient residence time. On the contrary, if the sorbent particles are too coarse, conversion decreases because of limited penetration of sulfur into the particle. The PSD of the sorbent establishing at steady state in the boiler is the result of the interplay of a number of processes. In particular, attrition and fragmentation phenomena can substantially affect the sorbent PSD and, in turn, the performance of the desulfurization process [7-9]. Attrition and fragmentation of limestone during FB combustion have been thoroughly characterized over the last decade [9-15]. Key phenomenological features and mechanistic pathways of sorbent attrition in FB combustors have been disclosed with the aid of a comprehensive test protocol consisting of different and mutually complementary test procedures [9,11,13,14]. In particular, sorbent attrition phenomena have been classified into: i) primary fragmentation, which occurs immediately after the injection of sorbent particles in the hot bed as a consequence of thermal stresses and internal overpressures due to CO2 emission; ii) attrition by abrasion, related to the occurrence of surface wear as the FB emulsion phase is sheared by the passage of bubbles, generating mostly fine/quickly elutriable fragments; iii) secondary fragmentation, a result of highvelocity impacts of sorbent particles against targets (bed material, reactor walls/internals), occurring mostly in the jetting region of the FB and in the exit region of the riser and the cyclone of circulating FB reactors. The critical influence of the progress of calcination and sulfation on attrition of limestone in FB combustors has long been recognized [7,9,13]. These reactions bring about relevant modifications of the mechanical and morphological properties of the Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
sorbent particles which significantly affect the mechanisms and extent of particle breakage. For example, the progress of sulfation decreases the attrition rate with respect to that of the native porous lime due to the formation of a more compact and tougher sulfate shell at the periphery of the particle [9]. The aim of this work was to explore the possibility to enhance the sulfur capture capacity of the sorbent during FB combustion, and at the same time to increase its mechanical resistance with respect to attrition and fragmentation phenomena. A proprietary limestone slow calcination pre-treatment technique was developed by Calcidrata S.p.A., which appears to be promising for the production of a relatively cheap desulfurization sorbent with suitable characteristics. An Italian limestone was pre-treated with this technique and then tested in a lab-scale FB reactor for its SO2 capture performance and attrition behavior. The results were compared with those obtained by using the same untreated raw limestone under the same operating conditions.
2. Experimental section Apparatus. The sorbent sulfation experiments were carried out in a stainless steel
atmospheric bubbling fluidized bed reactor 40mm ID and 1m high. The fluidization column was heated by two 2.2kW electric furnaces. The temperature of the bed, measured by means of a chromel-alumel thermocouple, was kept constant by a PID controller. The distributor of fluidization gas was a perforated plate with 55 holes 0.5mm in diameter disposed in a triangular pitch. Batches of material could be fed to the reactor via a hopper connected sideways to the upper part of the freeboard. The latter was equipped with a three-way valve. By operating this valve it was possible to convey flue gases alternately to two removable filters of sintered brass (100% filtration efficiency for > 10μm-particles). This device allowed time-resolved collection of elutriated fines. The fluidizing gas flow, composed by a mixture of air and N2-SO2, was measured by means of two high precision mass flowmeters which were specifically calibrated for each gas used. Analysis of CO2 and SO2 concentrations in the flue gas was accomplished by means of two NDIR analyzers on line. Further details can be found in [9]. Procedures. The reactor was charged with a bed made of sand (150g), and then heated to
the temperature of 850°C prior to each experiment. The fluidizing gas superficial velocity was 0.75m/s. Experiments were carried out by feeding a 20g sorbent batch in the bed while keeping a flow of dry air containing sulfur dioxide (1800ppm) and 8.5% by volume of oxygen. Under these conditions calcination and sulfation occurred at the Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
same time. Calcium conversion degree during sulfation was calculated as a function of time by working out the SO2 concentration at the exhaust. SO2 oxidation to SO3 inside the reactor was accounted for following the procedure detailed by Scala et al. [9]. Rates of fines generation by attrition of bed material were determined by measuring the amount of fines carried over by the fluidizing gas and elutriated from the reactor. The assumption underlying this procedure was that the residence time of elutriable fines in the reactor could be neglected and that elutriation rate could be assumed equal to the rate of fines generation by attrition at any time during limestone conversion. Elutriated fines were collected by means of the two-exit head by letting the flue gas flow alternately through sequences of filters (one was in use while the previous one was replaced) for definite periods of time. In order to prevent hydration and/or recarbonation of the collected material, each filter was readily put in a drier after being used where it was cooled down before it was weighed. The difference between the weights of the filters before and after operation, divided by the time interval during which the filter was in operation, gave the average fines generation rate relative to that interval. Attrition of sand could be neglected [9]. Particle size distribution of bed sorbent at the end of the run was determined by retrieving the bed material from the reactor and subjecting it to particle size analysis. Retrieval of sorbent particles after sulfation could be easily accomplished by discharging the bed from the reactor and sieving the sorbent out of the sand. This operation was carried out gently in order to avoid further attrition of particles, but rapidly because of the propensity of calcined sorbent to absorb moisture when in contact with ambient air. The sorbent was eventually characterized from the standpoint of particle size distribution by sieving. Materials. The bed material consisted of mixtures of sorbent and sand. Sand belonged to
the nominal size range 900-1000µm. Minimum fluidizing velocity was 0.4m/s. The sorbent used in this work was a high-calcium Italian limestone (Boundstone in Dunham’s classification) quarried from Mesozoic carbonate succession (Est Sardinia) and commonly named “Biancone di Orosei”. Chemical analysis of the raw limestone was carried out by a Philips PW 1400 XRF spectrometer, operating with a Rh tube at 30kV and 60mA, and gave a CaCO3 content of 98,83% by weight.
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Oviedo ICCS&T 2011. Extended Abstract
A
B
Figure 1 A) Slaking kinetics of pre-treated CaO (62°C at 25’); B) SEM micrograph
(Secondary Electrons) of the surface of a pre-treated CaO particle.
The sorbent was used either as received or after a proprietary pre-treatment process. This pre-treatment process consisted in the slow calcination of limestone at mild temperature and under controlled gaseous environment. After this pre-treatment, the sorbent consisted mostly of CaO (95.0%). Batches of both as received and pre-treated sorbent were sieved in the two nominal size ranges: 200-300 and 400-600μm. Microstructural-morphological SEM investigation of the CaO manufactured by Calcidrata S.p.A. was performed, on conductive samples, by a Zeiss Leo 50 XP apparatus operating with 20kV of accelerating voltage and an electron source of LaB6. Pore size distribution and total porosity of CaO were obtained by means MIP (mercury intrusion porosimetry) techniques using
a Micromeritics Autopore IV porosimeter
operating at 2000bar. Slaking kinetics test of CaO was carried out in wet condition in a special dewar device (Fapa instrument) according to UNI EN 459-2:2002 standard.
3. Results and Discussion Pre-treated lime characterization. Quicklime manufactured by Calcidrata S.p.A. by
means of the slow calcination treatment is very reactive as results from the slaking kinetics test (Fig. 1A). In fact, after addition of water, the temperature of the lime sample rises up to 60°C in few tens of seconds. SEM investigations (Fig. 1B) indicate that the microstructure of the pre-treated lime is composed of equigranular CaO crystallites, of pseudocylindrical shape, with dimensions in the range 2÷3μm. MIP porosity is around 50%, with a unimodal dimensional distribution of pores. Pore size radius classes are mostly concentrated in the interval 0.4÷0.9μm. The calculated specific surface area is over 15m2/g.
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Oviedo ICCS&T 2011. Extended Abstract
0.5
0.5
A
B
d = 400 - 600 μm d = 200 - 300 μm
0.4
0.43
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0.3
XCa, -
XCa, -
0.3
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0.0 0
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400
Time, min
Time, min
Figure 2 Calcium conversion vs. time for sorbent particles of different size sulfated
batchwise in FB at 850°C and 1800 ppm SO2. A) Raw limestone; B) Pre-treated CaO. Desulfurization performance. Figure 2 reports the degree of calcium conversion XCa as a
function of time during the FB sulfation of batches of either sorbent. Two particle sizes have been tested. For the raw limestone calcium conversion at the end of the test was relatively low, especially for the larger particle size. The pre-calcined sorbent performed better with a 34-43% calcium conversion after about 300-350min. The increase of calcium conversion was particularly evident for the 400-600μm particles. In order to better compare the performance of the two sorbents, the sulfur capture data have been worked out to obtain the SO2 capture capacity of the sorbents. This quantity represents the grams of SO2 captured per gram of sorbent, and is a more practical way to rank the sorbent performance. Figure 3 shows the SO2 capture capacity as a function of time for the same experiments reported in Fig. 2. Comparison of the results for the two sorbents highlights the much better performance of the pre-calcined lime. 0.5
A
SO2 capture capacity, g(SO2)/g(sorbent)
SO2 capture capacity, g(SO2)/g(sorbent)
0.5
d = 400 - 600 μm d = 200 - 300 μm
0.4
0.3
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0.19 0.1
0.08
B
0.46
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0.37 0.3
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0.1
d = 400 - 600 μm d = 200 - 300 μm
0.0
0.0 0
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0
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200
300
400
Time, min
Figure 3 SO2 capture capacity vs. time for sorbent particles of different size sulfated
batchwise in FB at 850°C and 1800 ppm SO2. A) Raw limestone; B) Pre-treated CaO.
Submit before 31 May 2011 to [email protected]
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1.0
d = 400 - 600 μm d = 200 - 300 μm 0.8
0.6
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Cumulative particle undersize distribution, -
Cumulative particle undersize distribution, -
Oviedo ICCS&T 2011. Extended Abstract
1.0
d = 400 - 600 μm d = 200 - 300 μm 0.8
0.6
0.4
0.2
B 0.0 0
Particle size, μm
100
200
300
400
500
Particle size, μm
Figure 4 Cumulative particle size distribution of sorbent of different size sulfated
batchwise in FB at 850°C and 1800 ppm SO2. A) Raw limestone; B) Pre-treated CaO. Under similar operating conditions, 2-4 times less sorbent is needed to obtain the same SO2 capture performance. This result relies on the combination of two effects: the better calcium exploitation of the pre-calcined sorbent (Fig. 2), and the lower molecular weight of CaO with respect to CaCO3. This last effects determines a larger number of moles of Ca available for reaction with SO2 in the pre-calcined lime per unit mass of sorbent. Attrition behavior. Figure 4 reports the cumulative particle size distribution of the
sorbent samples discharged from the bed after FB sulfation, for both particle sizes tested. The raw limestone (Fig.4A) exhibits a moderate fragmentation for the 200-300μm particles, and a significant fragmentation for the 400-600μm particles. Limestone fragmentation is mostly caused by the rapid calcination of the particles upon feeding in the hot bed, which generates significant overpressures inside the particles during CO2 release [9]. Larger particles are related to higher internal overpressures, which determine a higher degree of fragmentation. The population of particles below the initial particle size range (fragments) accounts for 4 and 23% of the sample mass for the 200-300 and 400-600μm particles, respectively. The pre-calcined sorbent exhibits a different behaviour (Fig.4B). Very limited fragmentation is evident at the end of the run. For this sorbent the fragments account for less than 1% of the sample mass for both particle size ranges. This result is certainly caused by the absence of significant calcination in the fluidized bed for the pre-calcined lime, but also indicates that the slow calcination/treatment is effective in strengthening the sorbent structure. Figure 5 shows the fines elutriation rate measured during FB sulfation tests with the two sorbents. The elutriation rate shows an initial high peak due to particle rounding off. Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
0.30
0.30
A
B 0.25
Elutriation rate, g/min
Elutriation rate, g/min
0.25
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d = 400 - 600 μm d = 200 - 300 μm
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0.05
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Figure 5 Fines elutriation rate vs. time during batchwise sulfation of sorbent of different
size in FB at 850°C and 1800 ppm SO2. A) Raw limestone; B) Pre-treated CaO. As sulfation proceeds, the fines elutriation rate decreases until a steady value is reached. This decay occurs over a time scale comparable to that over which calcium is sulfated, and should be related to the progress of reaction through the formation of a sulfate layer, harder than the oxide, on the particle surface [9,11]. A comparison between the two sorbents shows the raw limestone generates more fines especially during the first 20min. The total quantity of elutriated fines during the raw limestone experiments was 0.92 and 0.82g for the 200-300 and 400-600μm particles, respectively. For the pre-calcined lime the total quantity of elutriated fines was approximately halved, namely 0.49 and 0.34g for the 200-300 and 400-600μm particles, respectively. This result further confirms that the sorbent pre-treatment is able to give a large mechanical resistance to the particles.
4. Economic analysis
A simple case study was carried out on the basis of the typical operating conditions of an Italian full-scale FB unit burning international low-sulfur coals and using a limestone similar to that used in this work. The analysis of the solid residues coming out from the plant (coal ash + spent sorbent) gave an average content of unused CaCO3 and CaO of ∼7 and ∼20%, respectively. Taking into account that the total quantity of solid residues
is of the order of 160,000 tons per year, and that the raw limestone has an average cost of 30€ per ton, the following “penalties” associated with the inefficient use of traditional limestone in the plant can be estimated (yearly costs): a) the cost for unused CaCO3 (336,000€); b) the cost for unused CaO (1,714,560€); c) the cost associated to the consumption of coal that is burned to generate the necessary heat to produce unused Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
CaO by calcination (521,600€) – this cost was calculated assuming an average cost of 100 € per ton of coal; d) the cost for CO2 emissions deriving from limestone calcination in the FB and from the fraction of coal burning for unused CaO production (606,239€) – this cost was calculated assuming an average cost of 15 € per ton of produced CO2; e) the cost associated to landfilling of unused CaCO3 and CaO and of ash from the fraction of coal burning for unused CaO production (1,958,085€) – this cost was calculated assuming an average landfilling cost of 45 € per ton of residue. Summing up these costs, the total penalty associated to the inefficient use of limestone for this plant is 5,136,484€. From this simple estimation, it is clear that considerable economies can be obtained by the plant operator if these penalties can be minimized. With this respect, the use of the pre-treated sorbent developed by Calcidrata S.p.A. would be associated with a number of advantages. First, since the sorbent is precalcined, no unused CaCO3 would be present in the process, no coal would be necessary for CaO production, and no CO2 would be produced by sorbent calcination. This means that costs a), c) and d) would become zero. Second, the fraction of unused CaO would decrease because of the better calcium conversion and lower sorbent attrition in the FB. As a consequence, also the cost associated to landfilling of unused CaO would decrease. Thus, on the basis of these simple considerations the pre-treated sorbent appears to be promising for its use in FB coal-burning plants. It will be possible to carry out the detailed calculation of these economies once the cost per ton of the pre-treated sorbent will be available. Finally, a further advantage for the sorbent supplier should be mentioned. Since the FB operator requires a well defined particle size distribution of the sorbent, after milling a significant fraction of undersize particles are typically produced. While this residue has practically no market for limestone, it is highly valuable for lime and can be used to produce slaked lime for the building market. On the other side, it must be highlighted that since lime is hygroscopic and irritant, the storage and conveying devices should be sealed in order to avoid contact with moisture and with plant operators. This would of course add an initial cost to upgrade the storage and conveying devices of the plant.
Acknowledgement
The experimental support of G. Somma and A. Coppola is gratefully acknowledged.
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Oviedo ICCS&T 2011. Extended Abstract
References [1] Anthony EJ, Granatstein DL. Sulfation phenomena in fluidized bed combustion systems. Prog Energy Combust Sci 2001;27:215-36. [2] Dam-Johansen K, Østergaard K. High-temperature reaction between sulphur dioxide and limestone-I. Comparison of limestones in two laboratory reactors and a pilot plant. Chem Eng Sci 1991;46:827-37. [3] Dam-Johansen K, Østergaard K. High-temperature reaction between sulphur dioxide and limestone-II. An improved experimental basis for a mathematical model. Chem Eng Sci 1991;46:839-45. [4] Montagnaro F, Salatino P, Scala F. The influence of sorbent properties and reaction temperature on sorbent attrition, sulfur uptake, and particle sulfation pattern during fluidized-bed desulfurization. Combust Sci Technol 2002;174:151-69. [5] Duo W, Laursen K, Lim J, Grace JR. Crystallization and fracture: product layer diffusion in sulfation of calcined limestone. Ind Eng Chem Res 2004;43:5653-62. [6] Montagnaro F, Salatino P, Scala F. The influence of temperature on limestone sulfation and attrition under fluidized bed combustion conditions. Exp Therm Fluid Sci 2010;34:352-8. [7] Chandran RR, Duqum JN. Attrition characteristics relevant for fluidized bed combustion. In: Grace JR, Shemilt LW, Bergougnou MA, editors. Fluidization VI, New York: Engineering Foundation; 1989, pp 571–80. [8] Couturier MF, Karidio I, Steward FR. Study on the rate of breakage of various Canadian limestones in a circulating transport reactor. In: Avidan AA, editor. Circulating Fluidized Bed Technology IV, New York: AIChE; 1993, pp 672–8. [9] Scala F, Cammarota A, Chirone R, Salatino P. Comminution of limestone during batch fluidized-bed calcination and sulfation. AIChE J 1997;43:363–73. [10] Di Benedetto A, Salatino P. Modelling attrition of limestone during calcination and sulfation in a fluidized bed reactor. Powder Technol 1998;95:119–28. [11] Scala F, Salatino P, Boerefijn R, Ghadiri M. Attrition of sorbents during fluidized bed calcination and sulphation. Powder Technol 2000;107:153–67. [12] Chen Z, Lim CJ, Grace JR. Study of limestone particle impact attrition. Chem Eng Sci 2007;62:867–77. [13] Scala F, Montagnaro F, Salatino P. Attrition of limestone by impact loading in fluidized beds. Energy Fuels 2007;21:2566–72. [14] Scala F, Montagnaro F, Salatino P. Sulphation of limestones in a fluidized bed combustor: the relationship between particle attrition and microstructure. Can J Chem Eng 2008;86:347–55. [15] Yao X, Zhang H, Yang H, Liu Q, Wang J, Yue G. An experimental study on the primary fragmentation and attrition of limestones in a fluidized bed. Fuel Process Technol 2010;91:1119–24.
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Oviedo ICCS&T 2011. Extended Abstract
Synthetic gas bench study of CO2 capture from PCC power plants
E. Ruiz, J. M. Sánchez, M. Maroño, J. Otero CIEMAT, Avda. Complutense, 22, 28040, Madrid (SPAIN), [email protected] Abstract
Much attention is currently paid to the global warming effect due to CO2 emissions. Physical adsorption of CO2 using regenerable sorbents is a promising approach for CO2 capture from flue gases. The Combustion and Gasification Division of CIEMAT recently participated, under the supervision of Tecnicas Reunidas, in the CENIT CO2 Project, funded by CDTI, and led by ENDESA Generación, with the relevant collaboration of UNIÓN FENOSA, 12 industrial partners and 16 research institutions. The role of CIEMAT was to study the potential for CO2 capture of different sorbents, at a representative scale and under realistic conditions resembling those of flue gases after the desulphurisation tower of conventional pulverized coal combustion (PCC) plants. Firstly, comparative adsorption-desorption studies for CO2/N2 mixtures were performed for different sorbents: one alumina-based, one zeolitic-type and one activated carbon. The zeolitic type sorbent was identified as most promising and was subjected to a subsequent in-depth study of its stability, tolerance to poisoning and durability over cycles. Short term tests were performed to assess the effect of competitive adsorption e.g. H2O, and deactivation due to SO2 and NOx. Long term tests were conducted to study the effect of the number of adsorption-desorption cycles. CO2 adsorption breakthrough curves (exit CO2 concentration about 1 %) of the fresh zeolitic material under different gas compositions were obtained. The addition of water had a detrimental effect on CO2 adsorption, shifting CO2 breakthrough curve to shorter times, probably due to competitive adsorption of H2O. The simultaneous presence of H2O and SO2 has a synergic negative effect on CO2 adsorption. However, when H2O, SO2 and NO coexist in the testing gas, the synergic negative effect on CO2 adsorption is slightly attenuated. The adsorbent was also studied on consecutive adsorption-desorption cycles under the different gas compositions. In all cases, there was a certain loss in CO2 adsorption capacity on passing from the first to the second cycle, being almost constant in successive cycles.
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Oviedo ICCS&T 2011. Extended Abstract
1. Introduction
Recently, much attention is being paid to the global warming effect caused by excessive CO2 emissions. Carbon dioxide is contained, among others, in exhaust gas from combustion of fossil fuels. In post-combustion capture, the CO2 in the exhaust gas is captured using chemical or physical solvents (e.g. amine scrubbing). This is a fairly mature technology for CO2 separation at low temperature and in gases with low CO2 content. However, several drawbacks are often reported in literature such as large scale equipment, chemicals handling, and reduction of thermal efficiency [1-2]. Therefore, new approaches to capture CO2 are now being tackled, e.g. physical adsorption using regenerables adsorbents. Target materials should exhibit high adsorption capacity, selectivity and adsorption rate for CO2 capture. The process of capturing CO2 through physical adsorption methods consists of a first stage of CO2 separation, and a subsequent stage of regeneration in which CO2 is concentrated.
Commercially available devices based on physical adsorption processes operate in adsorption-regeneration cycles. At the adsorption stage, the gas passes through the adsorbent where CO2 is retained, ideally in a selective way. Once the bed of adsorbent reaches CO2 saturation, the gas to treat is forced to flow through another bed of adsorption while the saturated one is regenerated.
Adsorption-desorption cycle time is of the order of minutes and depends on the degree of purification required. The separation of the adsorbent and captured gas is a key stage in the process of adsorption. There are several methods to make the process of adsorption-desorption: -
Temperature Swing Adsorption (TSA): adsorption and desorption of CO2 occurs
at different temperatures. -
Pressure Swing Adsorption (PSA): adsorption and desorption of CO2 occurs at
different pressure. When vacuum is required for regeneration then it is called "Vacuum Swing Adsorption (VSA)".
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Oviedo ICCS&T 2011. Extended Abstract
The use of regenerable adsorbent solids could be a potentially valid alternative to absorbing liquids for capture of CO2 in combustion gases from power plants, although the separation process should be cost-effective, what happens when a highly efficient CO2 adsorbent material is used. It is necessary that the solid adsorbent has a high CO2 adsorption capacity, more than about 90 mg/g of adsorbent, high selectivity towards capturing this component in the gas mixture and the difference of temperatures between the process of adsorption and desorption of CO2 should be the smallest possible. In addition, since the exit of the exhaust gas temperature is around 120 - 150 ° C, the adsorbent material should be able to operate at temperatures relatively high to avoid the stages of cooling and condensation of corrosive compounds.
A number of solids capable of capturing CO2 are being studied, and, according to literature, they can be broadly divided into alumina-based materials, zeolite-type materials, mesoporous materials and carbon materials with different degrees of modifications to enhance CO2 adsorption features. Certain type of zeolite is used as adsorbent of CO2 in industrial applications [3] and it has been considered as a potential candidate for CO2 adsorption in postcombustion capture. However, it is necessary to evaluate its performance under realistic operating conditions and at a scale representative of the new process application.
2. Experimental section
Adsorption and desorption tests were performed in an existing bench-scale plant which was subsequently adapted for postcombustion capture of CO2 (Figure 1). The plant can treat up to 20 Nm3/h of both simulated and actual combustion off-gases. Process temperature can be up to 500ºC at about atmospheric pressure. The plant has been described in detail elsewhere [4].
The analytical monitoring of the adsorption-desorption processes was carried out by means of an FTIR analyzer.
Temperature and pressure drop along the bed and gas constituent flow rates were recorded during the adsorption/desorption tests.
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Oviedo ICCS&T 2011. Extended Abstract
Adsorption/desorption cycles were carried out in an atmospheric pressure fixed bed reactor.
The tested adsorbents were a commercial activated carbon and a synthetic activated alumina impregnated with Na2CO3 (10% weight), both of them in the form of cylindrical pellets, as well as a commercial zeolite-based material simulating one prepared from ashes of thermal power stations, in the form of spherical pellets.
Figure 1. Diagram of the bench-scale plant modified to study CO2 capture from combustion off-gases
Conditions of the adsorption-desorption comparative tests are listed in Table 1. Operating conditions utilized for the zeolite stability study are listed in Table 2. Desorption tests were carried out by flowing N2 at 110 º C and at atmospheric pressure. Table 1. Adsorption-desorption comparative tests. Operating conditions Adsorbent Adsorption gas (v/v) Adsorption T. (ºC) Desorption gas (v/v) Desorption T. (ºC) Space velocity (Nl h-1 g-1) Number of cycles
Activated alumina 15 % CO2 in N2 47 N2 47 0,88 3
Activated carbon 15 % CO2 in N2 47 N2 110 0,88 1
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Zeolite 15 % CO2 in N2 47 N2 110/47 0,88 5
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Oviedo ICCS&T 2011. Extended Abstract
Table 2. Zeolitic adsorbent stability study. Operating conditions Adsorption composition (%) 15 CO2, N2 balance 15 CO2, 10.7 H2O, N2 balance 15 CO2, 10.7 H2O, 5 O2, 0.0174 SO2, N2 balance 15 CO2, 10.7 H2O, 5 O2, 0.0174 SO2, 0.0495 NO, N2 balance
Temperature (˚C) 47 47 47 47
3. Results and Discussion
In order to compare the performance of the different materials studied, the CO2 breakthrough adsorption curves of the fresh adsorbents were obtained.
The adsorption process was characterised in terms of net CO2 capture capacity and useful adsorption time. They are defined as:
- Net CO2 capture capacity: amount of CO2 adsorbed until the CO2 exit concentration reaches 1% v/v. - Useful adsorption time: time during which the CO2 exit concentration is less than 1% v/v.
The experimental results obtained in dry gas adsorption tests for activated alumina, activated carbon and zelolite are shown in Table 3, 4 and 5 respectively.
A comparison between CO2 breakthrough adsorption curves corresponding to activated alumina, activated carbon and zeolite is depicted in Figure 2.
Table 3. Experimental results. Activated Alumina Test Net CO2 capture capacity (mg CO2/gadsorbent) Useful adsorption time (sec) 1st cycle 22.4 330 rd 3 cycle 3.6 54 Table 4. Experimental results. Activated carbon Test Net CO2 capture capacity (mg CO2/gadsorbent) Useful adsorption time (sec) st 1 cycle 4.2 78
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Oviedo ICCS&T 2011. Extended Abstract
Table 5. Experimental results. Zeolite Test 1st cycle 2nd cycle 3rd cycle 4th cycle 5th cycle
Net CO2 capture capacity (mg CO2/gadsorbent) 134.2 85.9 82.4 80.3 77.5
Useful adsorption time (sec) 1716 1188 1140 1110 1068
1.0
ZEOLITE fresh zeolite
Exiting CO2 (%)
0.8
after reg. 110ºC N 2 after reg. 47ºC N2
0.6
after reg. 110ºC N2
after reg. 47ºC N2 0.4
ALUMINA fresh
0.2
after reg. 47ºC N2
ACTIVATED CARBON fresh
0.0
0
200
400
600
800 1000 1200 1400 1600 1800 2000 2200
Time (sec)
Figure 2. Comparative between CO2 breakthrough curves of the different adsorbents Analyzing the experimental results shown in Table 3 and Figure 2, respectively, it can be seen that there is a significant adsorption capacity loss in the case of activated alumina on passing from the first adsorption cycle to the following one. In fact, Table 3 shows a sharp decline in both the net CO2 capture capacity and the useful time of adsorption. Figure 2 shows a shift in the breakthrough curve to lower adsorption times, in line with the trend observed in the adsorption parameters.
The activated carbon exhibits a low adsorption capacity, even in the case of using fresh adsorbent, as seems to be evident from the experimental results shown in Table 4 and Figure 2 respectively. For this reason, only looking at performance in the first cycle of adsorption, it was ruled out both the subsequent regeneration run and to conduct successive cycles of adsorption-desorption.
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Oviedo ICCS&T 2011. Extended Abstract
For the zeolite adsorbent, there was a considerable reduction in CO2 adsorption capacity on passing from the first to the second cycle, remaining, however, almost constant in successive cycles. The capacity loss in the second adsorption cycle was probably due to an inefficient regeneration at the temperature level imposed by the postcombustion CO2 capture process, as it is clear from the experimental results shown in Table 5 and Figure 2, respectively. In fact, Table 5 shows a sharp decline in both the net CO2 capture capacity and the useful adsorption time on passing from the first cycle to the following one and a more gradual reduction in successive cycles. Figure 2 shows a shift in the breakthrough curve to lower adsorption times, in line with the trend observed in the adsorption parameters.
In the case of the zeolitic adsorbent the value of the regeneration temperature (47 ° C and 110 ° C) was alternated in the successive adsorption-desorption cycles in order to analyse the influence of the regeneration temperature in the subsequent behaviour of the adsorbent. The analysis of the adsorption parameters in Table 5, and the trend of the breakthrough curves in Figure 2, follows that regeneration at 110 ºC is slightly more effective than at 47 ° C. However, it is necessary to analyse if the energy penalty associated with this increase in the desorption temperature is compensated by the improvement in the effectiveness of regeneration.
On comparing the CO2 breakthrough curves of activated alumina, activated carbon, and zeolite in the presence of dry gas (Figure 2), it can be derived that the behaviour of the latter is far superior to those of the activated alumina and activated carbon. Therefore, this material was selected at the most promising at this point and was subsequently subjected to an in-depth stability study under more realistic conditions.
CO2 breakthrough (CO2 exit concentration about 1 %) adsorption curves of the fresh zeolitic material under the different gas composition are shown in Figure 3. Addition of water has a detrimental effect on CO2 adsorption, shifting CO2 breakthrough curve to shorter times, probably due to competitive adsorption of H2O. The simultaneous presence of H2O and SO2 has a synergic negative effect on CO2 adsorption. It seems that both H2O and SO2 compete with CO2 for adsorption sites on the zeolite surface. However, when H2O, SO2 and NO coexist in the testing gas, the synergic negative effect on CO2 adsorption is slightly attenuated. Submit before 31 May 2011 to [email protected]
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Oviedo ICCS&T 2011. Extended Abstract
1.0 FRESH ZEOLITE
Exiting CO2 (%)
0.8
CO2+N2 CO2+H2O+N2
0.6
CO2+H2O+O2+SO2+N2
0.4
CO2+H2O+O2+SO2+NO+N2
0.2 27mg/g 54mg/g 70 mg/g
134 mg/g
0.0 0
250
500
750
1000
1250
1500
1750
Time (sec) Figure 3. CO2 breakthrough adsorption curves of the fresh zeolite adsorbent The adsorbent was also studied on consecutive adsorption-desorption cycles under the different gas compositions shown in Table 2. In all cases, there was a considerable reduction in CO2 adsorption capacity on passing from the first to the second cycle, being almost constant in successive cycles, probably due to an inefficient regeneration at the temperature level imposed by the postcombustion CO2 capture process. As an example, the long term performance of the zeolitic material in the presence of all typical constituents of the combustion flue gas is shown in Figure 4.
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Oviedo ICCS&T 2011. Extended Abstract
1.0
Exiting CO2 (%)
0.8
4º
ZEOLITA ENSAYO MULTI-CICLO MULTY-CYCLE TEST O+O CO2CO +H2 2+SO 2+NO+N 2 +H O+O +SO +NO+N 2
2
2
2
2
0.6 28 mg/g
6º
2º
0.4
1º 0.2 0.0 0
50 100 150 200 250 300 350 400 450 500 550
Time (sec) Figure 4. CO2 breakthrough adsorption curves of the zeolite adsorbent on consecutive adsorption-desorption cycle.
4. Conclusions
On the one hand, activated alumina rendered a poor behaviour of adsorption, with low adsorption capacities (including for the fresh sample) and low effectiveness of regeneration under the operating conditions imposed by the potential application of the process (at the exit of the desulfurization tower), because the temperatures are lower than the required for an effective regeneration of the adsorbent. For this reason, it was discarded as promising adsorbent.
The activated carbon was also discarded as promising adsorbent due to the observed low adsorption capacity even for the first cycle with fresh adsorbent.
The adsorption/desorption behaviour of the zeolite is far superior to those of the activated alumina and activated carbon. For this reason, the zeolite was selected as the most promising adsorbent for further deepen in the study of the CO2 adsorption/desorption processes under different operating conditions and over successive cycles. However, there are some limitations for the application of the tested
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Oviedo ICCS&T 2011. Extended Abstract
zeolite to postcombustion CO2 capture, due to a lack of selectivity and regenerability under conditions representative of flue gases from atmospheric pulverized coal boilers. Moreover, the tolerance and durability studies make evident that the CO2 adsorption capacity of the zeolite decreased on passing from the first to the second cycle, the presence of water diminished CO2 adsorption capacity by competitive adsorption and the simultaneous presence of H2O and SO2 has a synergic negative effect on CO2 adsorption capacity, although this synergic negative effect on CO2 adsorption is slightly attenuated under the coexistence of H2O, SO2, O2 and NO Acknowledgement.
Financial support from Centro para el Desarrollo Tecnológico Industrial of the Ministerio de Ciencia e Innovación of Spain (Project CENIT CO2) is acknowledged. The authors thank project partners for supplying the different adsorbents studied.
References [1] Bredesen R, Jordal K, Bolland O. High-temperature membranes in power generation with CO2 capture. Chem Eng Proc 2004;43:1129–58. [2] Simmonds M, Hurst P, Wilkinson MB, Watt C, Roberts CA. A study of very large scale post combustion CO2 capture at a refining & petrochemical complex. In: Gale J, Kaya Y, editors. Proceedings of the 6th International Conference on Greenhouse Gas Control Technologies, London: Elsevier; 2003, p. 39–44. [3] Zhao Z, Cui X, Ma J, Li R. Adsorption of carbon dioxide on alkali-modified zeolite 13X adsorbents. International Journal of Greenhouse Gas Control 2007;1:355–9. [4] Ruiz E, Sánchez JM, Maroño M, Otero J. CO2 capture from flue gases by physical adsorption. In: Abstracts of International Symposium about Capture and Storage of CO2, Seville, 2008.
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10
The influence of particle size on the steam gasification of coal G. Hennie Coetzee, Hein W.J.P Neomagus*, Raymond C. Everson School of Chemical and Minerals Engineering, Energy Systems, North-West University, Potchefstroom Campus, Private Bag X6001, Noordbrug 2520, South Africa *corresponding author: Tel. + 27 18 2991991, Fax: + 27 18 2991535, E-mail: [email protected]
Abstract The conversion of coal with steam into synthesis gas is an important intermediate reaction for e.g. the production of synthetic fuels. Due to the importance of this reaction, numerous reaction kinetic studies have been carried out on steam gasification, mainly on powdered coal samples. However, very few studies have addressed the steam gasification of large particles (> 1 mm) although several combustion and gasification technologies are based on the conversion of large particles. In this study, the effect of particle size (5 mm – 30 mm) on the steam gasification reaction rate is studied.
For this purpose, a medium rank, bituminous coal was investigated and particles with a density between 1400 and 1500 kg/m3 were selected using a mercury submersion technique. The results from this method were also compared with the results obtained from helium picnometry and mercury intrusion, so as to obtain more information about the porous structure of the coal.
The reactivity experiments were carried out in a large particle TGA, constructed in-house. The TGA is operated in such a way that the single coal particles are converted in the chemical reaction controlled regime. Steam and nitrogen were fed to the reactor, with a feed composition of 90 mol% steam. The data obtained from the reactivity experiments were modelled by several particle models, including (modifications of) the random pore and shrinking core model. Keywords: steam gasification; large coal particles; reaction kinetics; mercury submersion density analysis.
1. Introduction Coal is not only one of the most important sources of energy, but also the fastest growing energy source in the world[1]. According to WCI[1], South Africa is the fifth largest producer of coal and has the seventh largest coal reserve. South Africa’s abundant coal reserves allow it to meet the increasing energy demand. However, the coal reserves will eventually deplete, and therefore the efficient utilization of coal is crucial.
Comentario [A1]: Capacity increase?
South Africa’s unique synfuels and petrochemical industry is largely driven by SASOL, where coal is used as a feedstock to the gasification process. The coal is converted to synthesis gas, which is then converted to fuel using the Fisher-Tropsch process. The global demand for syngas is increasing at a rate of around 10% per year, which results in intensive research regarding the optimization and improvement of the gasification process[2].
Comentario [A2]: No problems here with quoting Sasol?
Due to the large size distribution of coal particles fed to the gasifier, it is important to extend the knowledge of particle size influence on the steam gasification reaction rate. Previous studies conducted on the steam gasification of coal were done on small particles or pulverised coal (-1 mm) [3-6]. Therefore, this study investigates the influence of large coal particles (5-30 mm) on steam gasification reactivity.
2. Experimental The coal used for this study originates from a Highveld coalfield, number 4 seam, and is mostly used for synthetic fuel production[7]. The reactivity experiments were done on single large coal particles, which were hand selected on a dimension basis. The length, width and height of the particles are used to categorise the size of the coal particle. The hand selected coal particles are then further sorted according to bulk density using the mercury submersion analysis. 2.1 Mercury submersion The mercury submersion analysis determines the bulk density of a coal particle using mercury as the submersion fluid. The coal particle is submerged and the buoyancy force of the coal particle is measured using a laboratory scale. The experimental setup is shown in Figure 1.
Comentario [A3]: Not always, inform that you use single particles for 20 and 40 and multiple (how many?) for 5 and 10
Figure 1: Mercury submersion experimental setup.
Single coal particles are used when determining the bulk density of the 20 and 30mm particles, while an average bulk density is calculated for the 5 and 10 mm particles.
2.2 Reactivity experiments The reactivity experiments were carried out in a large particle TGA, constructed in-house. The TGA consists of a vertical tube furnace which has an inner pipe diameter of 53mm. The reactants, steam and nitrogen, are fed from the top of the furnace as shown in Figure 2. Steam is generated using heating coils and water is fed to the heating coils using a variable speed peristaltic pump. The sample is situated inside a sample basket, which is located on top of a laboratory mass balance. The reactivity experimental setup is shown in Figure 2.
Figure 2: Reactivity experimental setup.
The samples were charred at the reactivity temperature in an inert nitrogen atmosphere. The reactivity experiments were carried out isothermally at 775 ˚C, 800 ˚C, 825 ˚C, 850 ˚C and 900 ˚C, in a 90 mol% steam concentration atmosphere. Single coal particles were used for the reactivity experiments of the 20 mm and 30 mm particles, while coal samples with several particles were used for the reactivity experiments of the 5 mm and 10 mm.
3. Results and discussion 3.1 Mercury submersion The mercury submersion method, which is a non-destructive method, was developed to accurately determine the bulk density of a single large coal particle. Since this method is non-destructive, the density of a single large particle can be measured, after which the same particle can be used for reactivity experiments. 20 mm (70 particles) and 30 mm (80 particles) single coal particles were hand selected and the bulk density of each coal particle was determined using the mercury submersion method. The results are shown in Figure 3 and 4.
Figure 3: Density analysis of 20 mm particles.
Figure 4: Density analysis of 30 mm particles.
From Figure 3 and 4 it is seen that the mean bulk density is between 1400-1500 kg/m3 and for this reason a density cut between 1400-1500 kg/m3 was used for the reactivity experiments. A density cut is used to reduce the deviation in ash percentage[8,9] and maceral composition[8,10,11] for the single coal particle which will increase the homogeneity of coal samples used. This will result in smaller deviations obtained in the reactivity experiments. 3.2 Coal characterisation
The Proximate, Ultimate, Calorific value and XRF analysis were done on a representative coal sample. The analyses were outsourced to Advanced Coal Technology Laboratories (Pty) Ltd. The summarised results for these analyses are shown in Table 1: Table 1: Coal characterisation results and comparison with literature.
Procedure Proximate analysis Moisture content (%) Ash content (%) Volatile matter content (%) Fixed carbon content (%) Fuel ratio (%) Ultimate anlysis H/C ratio O/C ratio Other Calorific value (MJ/kg) Alkali index
Result
Hoffman (2009)
Pinheiro (1999)
7.0 18.4 23.4 51.2 2.2
6.5 20.5 22.3 50.7 2.3
5.6 28.3 22.0 44.1 2.0
0.062 0.267
0.071 0.308
0.053 0.175
22.59 5.9
22.28 5.8
19.69 3.3
The results obtained for the coal characterisation was compared to Hoffman[12] and Pinheiro[13], which studied coal from the same region and seam (Highveld seam 4). It was found that the characterisation results compares well to what was found by Oberholzer[12]. However, the results obtained by Pinheiro do not compare well and the difference may be as a result of the different areas of the coal field mined over the past 10 years.
The mercury intrusion and CO2-BET analysis were done on a density cut (1400 – 1500 kg/m3) of the 5, 10, 20 and 30 mm particles. The analysis was done to determine the effect of particle size on the physical structure of the chars. The different coal samples was charred at 900 ˚C for 1h and prepared for mercury intrusion and CO2-BET. The short coal method was used for CO2-BET analysis at 273.15 K (which includes BET, Langmuir, DubininRadushkevich and Horvart-Kawazoe models). The results obtained for mercury intrusion and CO2-BET (BET model results) are shown in Table 2. Table 2: Mercury intrusion and CO2-BET results.
CO2-BET Particle size mm 30 20 10 5
Surface area (m2/g) 185 188 183 168
Mercury intrusion Particle size mm 30 20 10 5
Porosity (%) 16 16 16 23
From the results presented in Table 2, it is seen that there is no considerable physical structural change between the 10, 20 and 30 mm particles. However, the results obtained suggest that there are structural changes between the 5 mm particles and the other three particle sizes.
The deviation can be attributed to the sample selection method used to
determine the bulk density of the 5 mm coal particles.
3.3 Reactivity experiments
The reactivity experiments were done in order to determine the influence of coal particle size on the reactivity of steam gasification of coal. 5, 10, 20 and 30 mm coal particles were gasified at 775, 800, 825, 850 and 900 ˚C, using a steam concentration of 80 mol%. The coal samples for the 5 mm and 10 mm particles were weighed to ensure sufficient sample weight, in order to overcome large particle TGA limitations. Adequate repeatability of the 5 mm and 10 mm TGA runs were obtained after 2 to 3 runs. However, up to 4 runs were required for the 20 mm and 30 mm particles to obtain repeatable results. Repeatability tests for the steam gasification of the 30 mm single coal particles, at 900 ˚C, are shown in Figure 5.
Figure 5: Results to determine the repeatability of reactivity experiments at 900 ˚C.
From Figure 5 it is seen that 3 runs were used to validate repeatability of the results. The average conversion (up to 90%) for the 3 runs where determined and is also shown in Figure 5. The same was done for all the particle sizes at the various gasification temperatures, and the average conversion is used from this point on to compare results. Figure 6 shows the results for the average conversion of 30 mm coal particles for 775, 800, 850 and 900 ˚C.
Figure 6: Average conversion for the 30 mm coal particles.
The average conversion results are shown in Figure 6 for the 30 mm particles at the various gasification temperatures. The 95% confidence errors obtained from the multiple runs are also included in Figure 6 and range from 0.01 for the 850 ˚C runs to 0.1 for the 825 ˚C runs. From Figure 6 it can be seen that an increase in temperature results in an increase in reaction rate. For the increase in temperature (from 775 to 800 ˚C and 800 to 825 ˚C) a large decrease in reaction time is seen. However, the decrease in reaction time for an increase in reaction temperature (from 825 to 850 ˚C and 850 to 900 ˚C) is much smaller than what is observed for the lower temperatures. The same trend is observed for the 5, 10 and 20 mm reactivity experiments.
To determine the influence of particle size on the reactivity of steam gasification of coal, the average conversion of four particle sizes at a set temperature is compared. Figure 7 illustrates the average conversion (up to 90%) of the 5, 10, 20, and 30mm particles at 900 ˚C.
Figure 7: Comparison of average conversion of all the particle sizes at 900 ˚C.
From Figure 7 it is seen that the particle size does have an influence on the reactivity of steam gasification of coal at 900 ˚C. There is no distinct difference in the average conversion (up to 90%) of the 20 and 30 mm particle size, which can be attributed to the fracturing of large coal particles at high temperatures. A study conducted by Bunt & Waanders14on the thermal breakage of coal, concluded that particles larger than 25 mm fracture to form multiple smaller particles. However, thermal breakage of coal particles larger than 25 mm is not observed at lower gasification temperatures (<900 ˚C). From Figure 7 it is noted that the particle sizes varying between 5 mm and 20 mm do have an influence on the reactivity of steam gasification. As the particle size decreases, the time required for 90% conversion decreases. However, this trend is not observed for studies done on smaller coal particles (<2mm) [15,16].
4. Conclusion •
Results obtained from CO2-BET and mercury intrusion analyses, indicate that for a particle size range between 10 and 30 mm, no structural changes are observed. However, the deviation in structural parameters for the 5 mm particles can be attributed to the sample selection method.
•
At 900 ˚C, the 30 mm particles break into smaller particles, which produce reactivity results similar to that obtained for the 20 mm particles. However, for the lower temperatures the 30 mm particles do not fracture and a noticeable difference is observed between the reactivity results obtained for the 20 and 30 mm particles.
•
As the particle size is reduced, the reaction time required for 90% conversion is reduced (at constant temperature).
References
[1] WCI. (2010). World Coal Institute [Web:] http://www.worldcoal.org [Date accessed: 25 October 2010]. [2] VAN DYK, J.C., KEYSER, M.J. and COERTZER, M. 2005. Syngas production from South African coal sources using Sasol-Lurgi gasifiers. International Journal of Coal Geology, 65, 243-253, 2006. [3] SCHMAL, M. MONTEIRO, J.L.F. AND TOSCANI, H. (1983). Gasification of high ash content coals with steam in a semibatch fluidised bed reactor. Industrial and Engineering Chemistry Process Design and Development. 22:563 [4] EVERSON, R.C., NEOMAGUS, H.W.J.P., KASAINI, H. AND NJAPHA, D. (2005). Reaction kinetics of pulverised coal-chars derived from inertinite-rich coal discards: Gasification with carbon dioxide and steam. Fuel 85:1076. [5] ZHANG, L., HUANG, J., FANG, Y. AND WANG, Y. (2006). Gasification reactivity and kinetics of typical anthracite chars with steam and CO2. Energy & Fuels 20(3):1201. [6] ROBERTS, D.G. AND HARRIS, D.J. (2007). Char gasification in mixtures of CO2 and H2O: Competition and inhibition. Fuel 86(17-18):2672. [7] JEFFREY, L. S. 2004. Characterization of the coal resources of South Africa. In: SAIMM Proceedings of Sustainability of coal, 2004. The South African Institute of Mining and Metallurgy. [8] WANG, W., QIN, Y., WEI, C., Li, Z., GUO, Y. AND ZHU, Y. (2006). Partitioning of elements and macerals during preparation of Antaibao coal. International Journal of Coal Geology 68:223-232. [9] AKTAS, Z., KARACAN, F., AND OLCAY, A. (1998). Centrifugal float – sink separation of fine Turkish coals in dense media. Fuel Processing Technology 55:235-250. [10] VAN NIEKERK, D. AND MATHEWS, J.P. (2010). Molecular representation of Permian-aged vitrinite-rich and inertinite-rich South African coals. Fuel 89:73-82. [11] MAROTO-VALER, M.M., TAULBEE, D.N, ANDRESEN, J.M., HOWER, J.C. AND SNAPE, C.E. (1998). Quantitative 13C NMR study of structural variations within the vitrinite and inertinite maceral groups for a semifusinite-rich bituminous coal. Fuel 77:805813. [12] HOFFMAN, J. W. 2009. The reaction kinetics involved in steam gasification of large coal particles- the influence of temperature. Potchefstroom, B.Ing.
[13] PINHEIRO, H.J. 1999. A techno-economic and historical review of the South African coal industry in the 19th and 20th centuries AND analyses of coal product samples of South African collieries 1998-1999. (In Bulletin 113. SABS: Pretoria. 97p.) [14] BUNT, J. R. AND WAANDERS, F., B. 2009. Pipe reactor gasification studies of a South African bituminous coal blend. Part 1 – Carbon and volatile matter behavior as function of feed coal particle size reduction. [15] HANSON, S., PATRICK, J.W, AND WALKER, A. (2002). The effect of coal particle size on pyrolysis and steam gasification. Fuel 81:531-537. [16] YE, D.P., AGNEW, J.B, AND ZHANG, D.K. (1998). Gasification of a South Australian low-rank coal with carbon dioxide and steam: kinetics and reactivity studies. Fuel 77:12091219.
Effect of iron and calcium catalysts on pyrolysis and steam gasification of wood
K. MURAKAMI, M. SATO, T. KATO, and K. SUGAWARA Faculty of Engineering and Resource Science, Akita University 010-8502, Tegatagakuencho 1-1, Akita city, Akita, JAPAN
Abstract The influence of iron and calcium catalysts on the gas evolution behavior in the pyrolysis and steam gasification of Japanese cedar (Sugi) up to 900°C was examined in this study. The pyrolysis of Sugi without catalyst produced 20 wt% of char and 32 wt% of tar. In the pyrolysis of the Sugi impregnated with iron up to 10 wt%, the char yield increased to 35 wt% and the tar yield decreased to a few percent. Both CO and CO2 yields were enhanced by the presence of the impregnated iron.
The structure of
char produced by pyrolysis of Sugi without catalyst was amorphous, whereas the turbostratic carbon and Fe3C was formed by the loading of iron.
In the steam
gasification of Sugi without catalyst, hydrogen began to evolve from about 600°C, and the maximum rate of hydrogen evolution was reached at 900°C.
By the addition of
iron catalyst, the hydrogen evolution peak shifted to the lower temperature and the amount of hydrogen evolution increased. In the temperature range between 500 and 800°C, the hydrogen evolution profiles were similar to the CO2 evolution profiles. The comparison between the results of steam gasification and pyrolysis suggests that the iron catalysts enhance the uptake of carbon into the char during pyrolysis and the conversion of the carbon into CO2 to form hydrogen during steam gasification.
It was
also shown that the calcium catalysts had the effect of the promotion of hydrogen evolution at the lower temperature.
1.
Introduction Biomass fuels such as wood-based materials, agricultural residues, forestry waste
and so on have lately attracted considerable attention as future resources of energy and chemical materials for the solution of the global warming issue, because the biomass is a carbon neutral.
In order to use the waste wood effectively, many studies such as the
power generation by the direct combustion, the thermal conversion (e.g. gasification, liquefaction, pyrolysis, etc), and the biochemical conversion (e.g. ethanol fermentation) have been carried out. Among them, the steam gasification of wood is considered to be one of the promising thermochemical conversion processes and the hydrogen produced during steam gasification can be used as clean fuel.
In the gasification
process, however, a large amount of tar is also evolved, resulting in the blockage and fouling of process equipments such as fuel lines, engines and turbines. Previously, we studied the influence of ion-exchanged catalysts on the pyrolysis and gasification of Australian brown coal and reported that the gas yields increased and the tar yields decreased by loading the metallic catalysts such as iron and nickel [1,2], indicating that these metallic catalysts promote to decompose the tar into low-molecular weight compounds.
The catalytic gasification is expected to suppress tar evolution and
produce hydrogen rich gas.
In this study, the cheap catalysts, iron and calcium, were
loaded on the Japanese cedar (Sugi) by the impregnation method and the catalyst-impregnated Sugi samples were pyrolyzed under argon flow and gasified under steam/argon flow.
The purpose of this study is to clarify the influence of the
impregnated catalysts on the gas and tar yields during pyrolysis and steam gasification.
2. Experimental 2.1
Samples
Japanese cedar (hereafter referred to as Sugi) used in this study was ground between 125 - 250 μm.
The analyses of Sugi are as follows: C 46.8 wt% (dry), H 5.8
wt% (dry), N 0.1 wt% (dry), O 47.0 wt% (diff), and Ash 0.3 wt% (dry). of iron and calcium by impregnation was performed as follows.
The loading
A weighed amount of
Sugi (10 g) was immersed in 200 ml of aqueous solution containing iron (II) nitrate or calcium hydroxide so that the amount of metal loading was to be 1 - 10 wt%.
Then the
water was evaporated at 60°C under vacuum using a rotary evaporator.
The
impregnated samples are referred to as the metal loading in wt% and the metal species, e.g. 1 wt%Fe. 2.2
Pyrolysis and steam gasification A quartz boat containing 0.5 g of the sample was placed in a horizontal fixed-bed
type reactor (inner diameter 24.5 mm, length 50 mm).
At first, the sample was heated
from room temperature to 100°C under the gas flow of argon (60 ml-STP/min).
After
heating at 100°C for 1 h, the sample was heated at the heating rate of 10°C/min to 900°C under the mixed gas flow of argon (60 ml-STP/min) and steam (60 ml-STP/min), and kept at the temperature for 1 h. The gaseous products such as H2, CO, CO2, and CH4 were analyzed at every 5 min by a gas chromatograph (Yanaco G 2800, TCD) attached to the reactor. The pyrolysis experiments of the samples were also performed using the same reactor and the same conditions as the steam gasification experiments except for the gas flow of argon (60 ml-STP/min).
3. 3.1
Results and Discussion Pyrolysis of iron-impregnated Sugi Fig. 1 shows the product distribution of the iron-impregnated Sugi samples after
pyrolysis up to 900°C. As shown in this figure, char yield increased from 20 wt% (daf) for the Sugi without iron to 35 wt% (daf) for the 10 wt%Fe.
On the contrary, the
tar yield drastically decreased from about 32 wt% (daf) to a few percent by the presence of impregnated iron up to 10 wt%.
Fig. 2 shows the carbon balance of the
iron-impregnated Sugi samples after pyrolysis up to 900°C. component “Others” contains tar as well as acids and alcohols.
In this figure, the The carbon remaining
in the char for the 10 wt%Fe was about two times larger than that for the Sugi. the amounts of CO and CO2 slightly increased by loading iron catalyst.
Also,
From the
XRD measurement (Rigaku RAD-C), the structure of char produced by pyrolysis of the Sugi was amorphous, while the turbostratic carbon and Fe3C were formed by the loading of iron.
These results suggest that the iron promoted to decompose the volatile
matter including tar, so that a part of decomposed components was evolved as the low-molecular weight compounds and the other was incorporated to form iron carbides. 3.2
Steam gasification of iron-impregnated Sugi and calcium-impregnated Sugi Fig. 3 shows the hydrogen and CO2 evolution profiles during the steam gasification
of the iron-impregnated Sugi.
For the Sugi without iron, the hydrogen started to
produce from about 600°C, and the maximum rate of hydrogen evolution was reached at 900°C. With increasing the amount of iron, this hydrogen evolution peak shifted to lower temperature gradually.
As for the CO2 evolution, two peaks (first peak between
300 and 650°C, second one above 650°C) appeared, and both peaks increased with increasing the amount of iron.
Interestingly, the evolution behaviors of second CO2
peak were similar to those of hydrogen evolution. Fig. 4 shows the carbon balance after steam gasification of the iron-impregnated Sugi.
From this figure, the decreases
in char and CO yields and the significant increase in CO2 yield were observed when the amount of iron increased.
From the above results, the iron is considered to play roles
of promoting the uptake of carbon into the char during pyrolysis (Fig. 2) and the reaction of carbon with water to produce hydrogen and CO2 during steam gasification (Fig. 3).
The
hydrogen
evolution
profiles
during
calcium-impregnated Sugi are shown in Fig. 5.
steam
gasification
of
the
By loading calcium on the Sugi, the
hydrogen evolution rate accelerated between 500 and 800°C.
Unlike the case of iron,
however, the peak temperature was almost the same independent of the amount of calcium. Fig. 6 shows the hydrogen yields from the steam gasification of various samples. In the case of iron-impregnated Sugi, the amount of evolved hydrogen significantly increased with the amount of iron (from about 80 mmol/g-Carbon for the Sugi to 130 mmol/g-C for the 10 wt%Fe).
As for the calcium-impregnated Sugi, the promotion of
hydrogen formation by the presence of calcium catalyst was small.
Currently, the
steam gasification behaviors of calcium-impregnated Sugi is studied in detail.
4.
Conclusions The pyrolysis and steam gasification of iron- and calcium-impregnated Japanese
cedar (Sugi) were carried out in this study. In the steam gasification of Sugi, the presence of iron and calcium shifts the hydrogen evolution peak to the lower temperature and increases the hydrogen yields. In particular, the iron catalyst affects the amount of hydrogen evolution during steam gasification of wood significantly.
References [1] Murakami K, Shirato H, Ozaki J, Nishiyama Y. Effects of metal ions on the thermal decomposition of brown coal. Fuel Proc Technol 1996; 46, 183-94.
[2] Murakami K, Shirai M,
Arai M. Pyrolysis behavior of nickel-loaded Loy Yang brown coals: Influence of calcium additive. Energy Fuels 2002; 16, 752-5.
[3] Murakami K, Fuda K, Sugai M. Pyrolysis of wood
impregnated with iron: Influences of impregnated iron on the product distribution and the char structure. Journal of MMIJ 2008; 124, 143-7 (in Japanese).
100
Yield / wt% (daf)
80 60 40
100
Others CH4
80
CO2 CO
Yield / mol%
Others CH4 CO2 CO Tar Char
60
Char
40
20
20
0
Sugi
1 wt%Fe
5 wt%Fe
0
10 wt%Fe
Fig. 1 Product distribution of iron-impregnated Sugi pyrolyzed at 900°C.
Sugi
1 wt%Fe
5 wt%Fe
10 wt%Fe
Fig. 2 Carbon balance of iron-impregnated Sugi pyrolyzed at 900°C.
(a) H2
5
Sugi 1 wt%Fe 5 wt%Fe 10 wt%Fe
-1
Rate / mmol･g-Carbon ･min
-1
6
4 3 2 1 0 100
300 500 700 Temperature / ℃
900
20 40 Time / min
60
(b) CO2
100
Sugi 1 wt%Fe 5 wt%Fe 10 wt%Fe
4 3 2 1
60 40
0
20
100
300 500 700 Temperature / ℃
900
20 40 Time / min
60
0
Fig. 3 (a)Hydrogen and (b)CO evolution profiles during steam 2
gasification of iron-impregnated Sugi.
-1
4 3 2 1 0 300 500 700 Temperature / ℃
900
20 40 Time / min
60
Fig. 5 Hydrogen evolution profiles during steam gasification of calcium-impregnated Sugi.
Amount of hydrogen / mmol g-Carbon-1
-1
Sugi 1 wt%Ca 5 wt%Ca 10 wt%Ca
5
100
Sugi
1 wt%Fe
5 wt%Fe
10 wt%Fe
Fig. 4 Carbon balance of iron-impregnated Sugi gasified with steam up to 900°C.
6 Rate / mmol･g-Carbon ･min
Others CH4 CO2 CO Char
80 Yield / mol%
5
-1
Rate / mmol･g-Carbon ･min
-1
6
150
100
50
Fe Ca
0 0
2
4 6 8 10 Loading / wt% Fig. 6 Amount of hydrogen evolution during steam gasification.
International Conference on Coal Science & Technology, Oviedo, October 9−13th, 2011.
THERMODYNAMIC EFFICIENCY ANALYSIS OF GASIFICATION OF HIGH ASH COAL AND BIOMASS Rodolfo Rodrigues, Nilson R. Marcilio, Jorge O. Trierweiler Department of Chemical Engineering. Federal University of Rio Grande do Sul (UFRGS). Rua Eng. Luis Englert, s/n – Campus Central. 90040-040. Porto Alegre, Brazil. Phone: +55 51 3308 3956. E-mail: [email protected], [email protected], [email protected] Marcelo Godinho Department of Chemical Engineering. University of Caxias do Sul (UCS). Rua Francisco Getúlio Vargas, 1130. 95070-560. Caxias do Sul, Brazil. Phone: +55 54 3218 2100. E-mail: [email protected] Abstract: Since the coal is one of major source of energy it is predicted that will continue to play an important role in the world energy demands. Although as a fossil resource, i.e. a non-renewable resource, that should take into account atmosphere emissions. Recently, the gasification has proven to be very promising to enable the conversion of carbonbased fuels for applications in synthesis and co-generation. In this sense, the joint processing of coal and renewable carbon-based fuel (biomass) allows to use coal in a cleaner way. This work evaluates the gasification potential of coal and biomass available in Brazil. For that a thermodynamic approach is used. High ash coal and biomasses of bigger energetic potential for co-generation are evaluated. Predictions evaluate the process performance through the estimation of efficiency by changing operational conditions, concerning the gasifying agent (air and steam). The preliminary results show the range of 70 to 80% efficiency by using only air as gasifying agent with up to 65% of stoichiometric air demand. On the other hand, using air and steam those values can reach 85% to close to 100% with smaller than 40% of stoichiometric air demand. Keywords: gasification, high ash coal, biomass, cold gas efficiency, equilibrium model.
1. INTRODUCTION Since the coal is one of major source of energy it is predicted that will continue to play an important role in the world energy demands. Although as a fossil resource, i.e. a non-renewable resource, that should take into account atmosphere emissions. The gasification is a technology for thermal processing of coal. Recently, this technology has proven to be very promising to enable the conversion of carbon-based fuels into products for applications in synthesis and co-generation. In this sense, the joint processing of coal and renewable carbon-based fuel such as biomass allows to use coal in a cleaner way. This technology called cogasification leads to make up deficiencies of one kind of fuel by a combination with other fuel (synergy). This work evaluates the gasification potential of coal and biomass available in Brazil. For that a thermodynamic approach is used through thermodynamic equilibrium model. 2. EVALUATION METHODOLOGY Brazilian coal reserves are about 7 billion tons (0.8% of world total). The coal is subbituminous and high ash content (nearly 50%) [1]. On the other hand, the main Brazilian biomasses of bigger energetic potential to co-generation [2] are evaluated. According to processing characteristics, the biomasses are here classified in three groups: agricultural, forestry, and industrial residues. For this study, rice husk [3] and coconut residue [4] are considered in the agricultural group. In the forestry group is considered sawmill wood waste [5] and in the industrial group are considered sugar cane bagasse [6], sugar cane straw [6] and footwear leather waste [7]. The proximate and ultimate analyses for six biomasses and high ash coal are presented in Tab. 1. 1
International Conference on Coal Science & Technology, Oviedo, October 9−13th, 2011.
Table 1. Characterization of the fuels considered in this study. Coal [1]
Rice husk [3] Proximate analysis (% wt, db) Moisture (wb) 11.7 12.00 Volatile matter 18.7 67.80 Fixed carbon 25.1 13.60 Ash 56.2 18.60 Ultimate analysis (% wt, db) C 31.6 38.30 H 2.1 4.00 N 0.7 0.50 O 8.3 38.60 S 1.1 – Cl <0.007 – 15,491 HHV (kJ/kg, db) 11,900
Coconut residue [4]
Sawmill wood waste [5]
Sugar cane bagasse [6]
Sugar cane straw [6]
Footwear leather waste [7]
83.74 70.61 19.14 10.25
12.93 86.48 12.93 0.59
50.20 79.90 18.00 2.20
29.40 83.30 12.80 3.90
14.10 77.30 16.90 5.80
48.23 5.23 2.98 33.19 0.12 – 22,807
50.91 6.13 0.23 42.14 – – 20,100
44.60 46.20 50.72 5.80 6.20 8.76 0.60 0.50 12.78 44.50 43.00 25.40 0.10 0.10 1.88 0.02 0.10 0.46 18,100 17,400 18,448 wt = weight, db = dry base, wb = wet base.
This study applies a thermodynamic equilibrium model for multiphase to evaluate the coal and biomass gasification. The proposed model finds the final composition minimizing the total Gibbs energy of an ideal mixture [8]. The model assumes one solid-phase consisting of solid carbon and one gas-phase consisting of 70 species (combinations of C, H, O, N, S, and Cl elements). The simulations have been carried out through Python codes using Cantera software libraries [9] and thermodynamic properties of Burcat's database [10]. Python and Cantera are free to use software and available for Windows (Microsoft Corporation) and Linux (The Linux Foundation) platforms. Predictions evaluate the process performance through the estimation of efficiency by changing operational conditions, concerning the amount of gasifying agent per amount of feed fuel: air at 25°C and 1 atm, and steam at 200°C and 1 atm. The air amount is expressed by “equivalence ratio” () and the steam amount is expressed by “steam-to-carbon ratio” (stm). Equivalence ratio represents the ratio among oxygen fed per oxygen required to complete combustion (stoichiometric oxygen demand). The cold gas efficiency is used to measure the process performance. 3. RESULTS AND DISCUSSION Sensitivity analyses for biomass and coal gasification are done to a range of operational process conditions. The next figures show the results apart for the six biomasses and high ash coal. The evaluation of cold gas efficiency for several operational conditions identifies the operating ranges to formation of fuel gas with highest heating value to each fuel. Equilibrium temperature is the attainable temperature for each operating point in adiabatic conditions. Fig. 1 illustrates the evaluating parameters for agricultural residues. Rice husk (Fig. 1a-b) already attains higher than 75% efficiency by using only air as gasifying agent in 0.35 < < 0.45 up to 80% in ≈ 0.4. The joint use of steam and air should raise efficiency to 90% in stm > 0.75 and lower air demand ( < 0.3). While coconut residue (Fig. 1c-d) reaches lower efficiency and temperature values due to high moisture content (83.74%) that suggests further studies to evaluate a previous drying stage. The sawmill wood waste (Fig. 2) has the particularity that the addiction of a steam fraction reveals no improvement in outcomes. So 80% efficiency might already be attained with only air as gasifying agent, equivalence ratio close to 0.35.
2
International Conference on Coal Science & Technology, Oviedo, October 9−13th, 2011.
(a) Cold gas efficiency (%) to rice husk.
(b) Equilibrium temperature (ºC) to rice husk.
(c) Cold gas efficiency (%) to coconut residue.
(d) Equilibrium temperature (ºC) to coconut residue.
Figure 1. Cold gas efficiency (%) and equilibrium temperature (ºC) to agricultural residues for a range of operational condition (air and steam demand).
(a) Cold gas efficiency (%) to sawmill wood residue.
(b) Equilibrium temperature (ºC) to sawmill wood residue.
Figure 2. Cold gas efficiency (%) and equilibrium temperature (ºC) to forest residues for a range of operational condition (air and steam demand).
3
International Conference on Coal Science & Technology, Oviedo, October 9−13th, 2011.
The analysis of Fig. 3 shows a region of < 0.4 where there are not gasification reactions for the coal's case. This operational range represents 0% efficiency (Fig. 3a) and temperature lower than 25°C (Fig. 3b). From equivalence ratio about 0.65 it can achieve efficiency up to 70% using only air as gasifying agent. Hence, the progressive increment of steam-to-carbon ratio can increase efficiency to 95% with stm around 0.75 and < 0.5.
(a) Cold gas efficiency (%) to coal.
(b) Equilibrium temperature (ºC) to coal.
Figure 3. Cold gas efficiency (%) and equilibrium temperature (ºC) to high ash coal for a range of operational condition (air and steam demand).
Fig. 4 shows the evaluate parameters for industrial residues. As well as coconut residues, sugar cane bagasse has high moisture content (50.2%) that make difficult to reach higher efficiency values adding a steam fraction (Fig. 4a-b). Sugar cane straw (Fig. 4c-d) attains over 80% efficiency from ≈ 0.4 equivalence ratio, and 90% efficiency could be attained with steam-to-carbon ratio higher than 0.5. Footwear leather waste (Fig. 4e-f) can attain about 70% maximum efficiency just for using air as gasifying agent in range of 0.5 to 0.55. Whereas the application of steam can increase efficiency to 90% from 1.0 steam-to-carbon ratio.
(a) Cold gas efficiency (%) to sugar cane bagasse.
(b) Equilibrium temperature (ºC) to sugar cane bagasse.
4
International Conference on Coal Science & Technology, Oviedo, October 9−13th, 2011.
(c) Cold gas efficiency (%) to sugar cane straw.
(d) Equilibrium temperature (ºC) to sugar cane straw.
(e) Cold gas efficiency (%) to footwear leather waste.
(f) Equilibrium temperature (ºC) to footwear leather waste.
Figure 4. Cold gas efficiency (%) and equilibrium temperature (ºC) to industrial residues for a range of operational condition (air and steam demand).
For all cases presented, the figures show maximum equilibrium temperatures at = 1 and stm = 0. This temperature is known as adiabatic flame temperature. The addition of steam causes an increasing of efficiency by the formation of more H2. This addition also decreases the adiabatic flame temperature as shown in = 1. The region of = 1 also illustrates the conversion of all useful gas (fuel gas) that corresponds to cold gas efficiency equal to zero. Failure to account for an ash fraction in the proposed model would have great effect on the simulated values, especially for coal that has high ash content (56.2%). This would allow evaluating the effects of co-gasification of higher ash fuels (coal) together with lower ash fuels (biomass). 4. CONCLUSIONS This study presents a general analysis of gasification of biomass and high ash coal through a thermodynamic equilibrium model. Predictions evaluate the process performance against operational conditions. The preliminary results show the range of 70 to 80% efficiency by using only air as gasifying agent with up to 65% of stoichiometric air demand. On the other hand, using air and steam those values can reach 85% to close to 100% with smaller than 40% of stoichiometric air demand. Next steps for this study include the model validation by literature data and the 5
International Conference on Coal Science & Technology, Oviedo, October 9−13th, 2011.
evaluation of gasification of coal-biomass blends. At last, a complete evaluation of the best cogasification conditions will be available to blends processing. ACKNOWLEDGMENT The study received financial support from the Brazilian Research Foundation (CNPq) under project 551386/2010-0 and also from the Brazil Coal National Network (http://www.ufrgs.br/rede_carvao). REFERENCES [1] Kalkreuth, W., Holz, M., Kern, M., Machado, G., Mexias, A., Silva, M., Willett, J., Finkelman, R., Burger, H., 2006. Petrology and chemistry of Permian coals from the Paraná basin: 1. Santa Terezinha, Leão-Butiá and Candiota coalfields, Rio Grande do Sul, Brazil. Int. J. Coal. Geol. 68 (1–2), 79–116. DOI: 10.1016/j.coal.2005.10.006 [2] ANEEL, 2008. Atlas electrical energy of Brazil. Tech. Rep., National Agency of Electrical Energy (ANEEL), Brasília, Brazil. [3] Hoffmann, R., 1999. A method to evaluate the regionalised energy generation, with power plants lower than 1 MW, by the residual biomass administration—The rice husk’s case. Ph.D. thesis, Department of Mechanical Engineering. Federal University of Rio Grande do Sul, Porto Alegre. HDL: 10183/11967 [4] Andrade, A. M., Passos, P. R. A., Marques, L. G. C., Oliveira, L. B., Vidaurre, G. B., Rocha, J. D. S., 2004. Pyrolysis of coconut residues (Cocos nucifera L.) and analysis of charcoal. R. Árvore. 28 (5), 707–714. DOI: 10.1590/S0100-67622004000500010 [5] Wander, P. R., 2001. Utilization of wood wastes as alternative of renewable energy for the sustainable development of northeast region of the state of Rio Grande do Sul, Brazil. Ph.D. thesis, Department of Mechanical Engineering. Federal University of Rio Grande do Sul, Porto Alegre. HDL: 10183/2303 [6] Seabra, J. E. A., 2008. Technical-economic evaluation of options for whole use of sugar cane biomass in Brazil. Ph.D. thesis, Faculty of Mechanical Engineering. State University of Campinas, Campinas. URL: http://libdigi.unicamp.br/document/?code=vtls000446190 [7] Godinho, M., 2006. Combined gasification-combustion of solid wastes of leather and shoe industry. Ph.D. thesis, School of Engineering. Federal University of Rio Grande do Sul, Porto Alegre. HDL: 10183/8958 [8] Rodrigues, R., Marcilio, N. R., Trierweiler, J. O., Godinho, M., Pereira, A. M. S., 2010. Cogasification of footwear leather waste and high ash coal: A thermodynamic analysis. In 27th Int. Pittsburgh Coal Conf. Istanbul, Turkey. [9] Goodwin, D. G., 2009. Cantera: An object-oriented software toolkit for chemical kinetics, thermodynamics, and transport processes. Caltech, Pasadena. URL: http://cantera.googlecode.com [10] Burcat, A., Ruscic, B., 2005. Third millennium ideal gas and condensed phase thermochemical database for combustion with updates from active thermochemical tables. Argonne National Laboratory, Illinois, Report TAE 960 Table 4. URL: ftp://ftp.technion.ac.il/pub/supported/aetdd/thermodynamics
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Oviedo ICCS&T 2011. Extended Abstract
How closely do low volatile bituminous coals prepared by hydrous pyrolysis of woody biomass and low-rank coals correspond to prime coking coals? Sylvia Kokonya1, Miguel Castro Diaz1, Clement Uguna1, Colin Snape1 and Andrew D. Carr2 1
Department of Chemical & Environmental Engineering, Faculty of Engineering, University Park, Nottingham NG7 2RD, UK 2 Advanced Geochemical Systems, Burton-on-the-Wolds, Leicestershire, UK [email protected]
Abstract It is established that hydrous pyrolysis of woody biomass can produce coal-like materials and that this treatment of low-rank coals can generate coals in the bituminous coal rank range.
Conditions where artificially matured coals containing 20-25%
volatile matter with mean vitrinite reflectances in the range of 1.2-1.4 can be generated have been identified that, in principle, should have compositions and properties corresponding quite closely to those of prime coking coals. Temperatures above 380oC are required for short reaction times (ca. 1 day) meaning that either superheated steam or supercritical water has to be used, depending upon the water pressure.
The
properties of an artificially matured coal have been compared to those of prime coking coals where fluidity development has been investigated using a combination of high temperature 1H NMR and rheometry. Although the high temperature hydrothermal treatment increased the fluid material (5%) in the artificially matured coal, its rheological behaviour was not improved and was quite different to that of prime coking coals. Further investigation will be required to determine whether the high temperature hydrothermal treatment of other low-rank coals and using higher pressures could be a potential route to produce alternatives to prime coking coals.
1
Oviedo ICCS&T 2011. Extended Abstract
1. Introduction Coalification of coal is generally considered to be dependent of geological time, temperature and pressure. Some authors have also suggested that water is an important factor [1-3]. Indeed, these authors found that more bitumen was generated under low pressure hydrous pyrolysis when compared to anhydrous or non-hydrous (absence of water) pyrolysis of coal and petroleum source rocks. Therefore, the role of water cannot be ignored since the presence of water has been found to increase the yield of bitumen or fluid materials during pyrolysis experiments. As the presence of water can enhance the amount of bitumen or fluid material under low pressure hydrous pyrolysis conditions, water have also been found to cause a reduction in bitumen and liquid product yields under high pressure conditions. Some authors [2, 4-7] observed that gas, bitumen and liquid product yields decreased at 500 bar liquid water pressure and higher when compared to low pressure during hydrous pyrolysis of petroleum source rocks. As water can have both beneficiary and negative effects depending on the pressure applied during the hydrous pyrolysis experiment as suggested by the above authors, it is important to conduct hydrous pyrolysis experiments under low pressure conditions where higher conversion to liquid materials from coals or biomass is required. Therefore, the aim of this work is to investigate whether hydrous pyrolysis of a low-rank coal can generate a coal in the bituminous coal rank range.
2. Experimental section A low rank UK coal (coal A) of Westphalian age with initial vitrinite reflectance of 0.75% Ro was used. Two good coking coals, namely an Australian coal (coal B) and a North America coal (coal C) were used for comparison. Coal B has a volatile matter content of 25.2 wt% daf and vitrinite reflectance of 1.13% Ro. Coal C has a volatile matter content of 24.0 wt% daf and a vitrinite reflectance of 1.22% Ro. Pyrolysis experiments were conducted on 5g of coal A (1-2 mm particle size) at 350oC (temperature accuracy +/– 1oC) for 24 hours under hydrous pyrolysis conditions. The pyrolysis equipment (Figure 1) comprised of a Parr 4740 series stainless steel pressure vessel (25 ml cylindrical) rated to 550 bar at 350oC and connected to a pressure gauge rated to 690 bar. Heat was applied by means of a fluidised sand bath which was controlled by a temperature gauge connected to it. Temperature was monitored by means of an additional K-type thermocouple connected externally to a computer which 2
Oviedo ICCS&T 2011. Extended Abstract
recorded the temperature every 10 seconds. The hydrous experiment was conducted at 190 bar. The un-extracted coal was first weighed and transferred to the vessel, after which 15 ml of water was added before the pressure vessel was assembled for the experiments. The reaction vessel was flushed with nitrogen gas to replace air in the reactor head space, after which 2 bar pressure of nitrogen was pumped into the pressure vessel to produce an inert atmosphere during the experiments. The sand bath, which was connected to a compressed air source, was pre-heated to the required temperature and left to equilibrate. Then, the pressure vessel was lowered onto the sand bath and the experiment left to run with a constant air flow through the sand bath for 24 hours, after which the sand bath was switched off and left to cool to ambient temperature before product recovery.
The total internal volume of the empty pressure vessel and its
associated pipe work was estimated to be about 34 ml by pressurising with nitrogen gas to 2 bar and measuring the volume of gas released.
Figure 1. Schematic diagram of pyrolysis equipment. For 1H NMR and rheometry analyses, the hydrous pyrolysis sample was dried in a vacuum oven at 40°C for 4 hours to remove the moisture. A Doty 200 MHz 1H NMR probe was used in conjunction with a Bruker MSL300 instrument to determine fluidity development with temperature. A flow of 25 dm3/min dry nitrogen was used to transfer heat to the samples and to remove the volatiles that escaped from the container. Below the sample region, a flow of 60 dm3/min of dry air prevented the temperature rising 3
Oviedo ICCS&T 2011. Extended Abstract
above 50°C to protect the electrical components. In addition, air was blown at 20 dm3/min into the region between the top bell Dewar enclosing the sample region and the outer side of the probe to prevent the temperature from exceeding 110°C. The sample temperature was monitored using a thermocouple in direct contact with the sample container. The solid echo pulse sequence (90o-τ-90o) was used to acquire the data. A pulse length of 3.50 μs was maintained throughout the test. Approximately 100 mg of sample was packed lightly into a boron nitride container, and 100 scans were accumulated using a recycle delay of 0.3 seconds. The samples were heated from room temperature to the final temperature at approximately 3°C/min. The spectra obtained were deconvoluted into Gaussian and Lorentzian distribution functions, which enabled the calculation of the fraction of total hydrogen that is mobile and its mobility (T2L). The higher the concentration and mobility of the fluid phase the higher the fluidity, and thus, fluidity depends on both concentration and mobility of the fluid or mobile phase. Rheological measurements were performed in a Rheometrics RDA-III high-torque controlled-strain rheometer. The amount of sample for analysis was 1.5 g and the samples were compacted under 5 tonnes of pressure in a 25 mm die to form disks with thickness of approximately 2.6 mm. The tests involved placing the sample disk between two 25 mm parallel plates which had serrated surfaces to reduce slippage. The sample disk was heated quickly to 330°C and heated to 520°C at a rate of 3°C/min. The furnace surrounding the sample was purged with a constant flow of nitrogen to transfer heat to the sample and remove volatiles.
The sample temperature was monitored using a
thermocouple inside the furnace.
A continuous sinusoidally varying strain with
amplitude of 0.1 % and frequency of 1 Hz (6.28 rad/s) was applied to the sample from the bottom plate throughout the heating period. The stress response on the top plate was measured to obtain the complex viscosity (η*), phase angle (δ) and plate gap (ΔL) as a function of temperature.
3. Results and Discussion The hydrous pyrolysis of coal A produced a carbonaceous material with a vitrinite reflectance of 1.35% Ro. Figure 2 shows the high-temperature rheometry results for the coal before and after the hydrous pyrolysis treatment.
4
Oviedo ICCS&T 2011. Extended Abstract
Coal A (after hydrous pyrolysis)
Coal A (before hydrous pyrolysis) 6
10
1.1
90.0
1.0
75.0
0.9
60.0
6
10
1.2
60.0
η* ( ) [Pa-s]
η* ( ) [Pa-s]
45.0
δ( ) [°]
105
ΔL ( ) [mm]
45.0
δ( ) [°]
ΔL ( ) [mm]
0.8
75.0
1.1
1.0 105
90.0
0.9
4
10 350.0
400.0
450.0
500.0
0.7
30.0
0.6
15.0
0.5 550.0
0.0
Temp [°C]
30.0 0.8
4
10 350.0
400.0
450.0
500.0
0.7 550.0
15.0
0.0
Temp [°C]
Figure 2. Complex viscosity (η*), phase angle (δ) and plate gap (ΔL) as a function of temperature for coal A before and after hydrous pyrolysis.
The complex viscosity in coal A before the hydrous treatment does not change significantly with temperature. The phase angle increases gradually with temperature but does not cross the threshold of 45° that is characteristic of liquid-like viscoelastic materials. The increase in plate gap can be attributable to the thermal expansion of the rheometer plates rather than expansion of the sample since the increase in plate gap is less than 0.3 mm throughout the test. After hydrous pyrolysis, the material does not develop any fluid material and the viscoelastic characteristic remain constant with temperature.
Therefore, any liquid material promoted by the hydrous pyrolysis
treatment is not able to change the viscoelastic properties of the coal.
Figure 3 shows the rheometry results for the good coking coals B and C, which have similar vitrinite contents (1.1% - 1.2%) as coal A after the hydrous treatment. These results show that significant amount of liquid material are required to cause viscoelastic changes and expansion/collapse of the sample mass that are characteristic of good coking coals. This might be achieved in low-rank coking coals by either changing the hydrous pyrolysis conditions to maximise the liquid product and/or selecting those coals that already develop fluidity during pyrolysis.
5
Oviedo ICCS&T 2011. Extended Abstract
Coal B
Coal C
7
10
0.85
7
90.0
0.8
10
75.0
6
10
0.55 3
10 350.0
375.0
400.0
425.0
450.0
475.0
0.5 500.0
75.0
0.0
60.0
η* ( ) [Pa-s]
-0.5
45.0
-1.0
30.0
-1.5
15.0
δ( ) [°]
5
10
30.0 0.6
104
0.5
ΔL ( ) [mm]
45.0
δ( ) [°]
0.65
ΔL ( ) [mm]
η* ( ) [Pa-s]
60.0
105
90.0
106
0.75
0.7
1.0
4
10 15.0
0.0
103 350.0
375.0
400.0
Temp [°C]
425.0
450.0
475.0
-2.0 500.0
0.0
Temp [°C]
Figure 3. Complex viscosity (η*), phase angle (δ) and plate gap (ΔL) as a function of temperature for coals B and C.
Figure 4 shows the percentage of fluid hydrogen as a function of temperature in coal A before and after hydrous pyrolysis. The hydrous treatment seems to increase 5% the amount of fluid material in the coal between 400-450°C.
Figure 4. Percentage of fluid hydrogen as a function of temperature in coal A before and after hydrous pyrolysis.
However, Figure 5 shows that the mobility of the fluid phase in coal A before hydrous treatment increases from 375°C to 440°C and reaches a maximum in T2L of 140μs, whereas coal A after hydrous treatment does not have a maximum in mobility and this is fairly constant with temperature (around 85μs). These results indicate that although the
6
Oviedo ICCS&T 2011. Extended Abstract
hydrous pyrolysis treatment of the coal causes a small increase in fluid material and an increase in vitrinite reflectance, these changes will not be sufficient to produce a good coking coal.
Figure 5. Mobility of the fluid hydrogen as a function of temperature in coal A before and after hydrous pyrolysis.
4. Conclusions The high temperature hydrous pyrolysis of a low-rank coal increased the amount of fluid material by approximately 5% as the vitrinite reflectance was raised to 1.3. However, the total amount of liquid phase (18% fluid material) was not able to affect the viscoelastic properties of the coal, and hence, its rheological behaviour was completely different to that of good coking coals. These results indicate that low-rank coals that do not produce fluid material during pyrolysis cannot be converted into good coking coals by this high temperature hydrothermal treatment. Still, further investigation is required to ascertain whether other low-rank coals that develop fluid material during hydrous pyrolysis can be converted into prime coking coals. In addition, the pressure of the water phase in hydrous pyrolysis may also aid the retention of fluid material in the coal.
Acknowledgement The authors than the Research Fund for Coal & Steel for financial support (DENSICHARGE project, Contract No. RFCR-CT-2010-00007).
7
Oviedo ICCS&T 2011. Extended Abstract
References [1] Behar F, Lewan MD, Lorant F, Vandenbroucke M. Comparison of artificial maturation of lignite in hydrous and nonhydrous conditions. Org Geochem 2003;34:575–600. [2] Carr AD, Snape CE, Meredith W, Uguna C, Scotchman IC, Davis RC. The effect of water pressure on hydrocarbon generation reactions: some inferences from laboratory experiments. Petroleum Geoscience 2009;15:17–26. [3] Lewan MD. Experiments on the role of water in petroleum formation. Geochimica et Cosmochimica Acta 1997;61:3691–3723. [4] Price LC, Wenger LM. The influence of pressure on petroleum generation and
maturation as suggested by aqueous pyrolysis. Org Geochem 1992;19:141–159. [5] Landais P, Michels R, Elie M. Are time and temperature the only constraints to the simulation of organic matter maturation? Org Geochem 1994;22:617–630. [6] Michels R, Landais P, Torkelson BE, Philp RP. Effects of effluents and water
pressure on oil generation during confined pyrolysis and high-pressure hydrous pyrolysis. Geochimica et Cosmochimica Acta 1995a;59:1589–1604. [7] Michels R, Landais P, Torkelson BE, Philp RP. Influence of pressure and the presence of water on the evolution of the residual kerogen during confined, hydrous and high-pressure hydrous pyrolysis of Woodford shale. Energy & Fuels 1995b;9:204–215.
8
Biomass Characterisation and link with Char Morphology and Ashing Behaviour Cheng Heng PANG1, Tao WU2, Edward LESTER1,a 1
School of Chemical and Environmental Engineering, University of Nottingham, University Park,
NG7 2RD Nottingham, United Kingdom 2
Division of Engineering, University of Nottingham, 315100 Ningbo, China
a
[email protected]
+44 115 9514974
Abstract
The advantages of blending biomass with coal for power generation include the reduction of CO2 emissions and the dependency on non-renewable fossil fuels. However, there is a need to understand the role of biomass during direct combustion and co-firing, particularly in terms of the effect of biomass on milling, combustion performance and ash slagging and fouling behaviour. This work aims to relate properties of biomass in its original form to its char morphology and ashing behaviour at high temperatures. The lignocellulosic composition of raw biomass was determined using standard biological assays. Char morphology was studied using oil immersion microscopy. Elemental composition of ash produced at 650°C was determined using the SEM/EDAS, while the ash characteristics were studied using a newly developed advanced ash fusion test. Preliminary results have shown good correlations between lignocellulosic composition and char morphology as well as ash behaviour at elevated temperatures.
1. Introduction
Biomass is significantly different from coal in terms of its high moisture and volatile contents, coupled with the comparatively low amount of fixed carbon and ash content. Whilst this creates a negative effect on the calorific value of biomass, the thermal conversions of both coal and biomass, however, follow a similar order of reactions [1]:
(1) Pyrolysis i.e. devolatilisation and volatile release where materials often soften and swell while ejecting gaseous products of moisture and hydrocarbons, resulting in the formation of char (2) Conversion of the intermediate carbon rich char particles in the presence of an oxidant gas, i.e. combustion (3) Deposition or collection of mineral residues along with any unburnt carbon particles
Generally, certain degree of overlapping occurs between step (1) and (2) [2] but the process by which fuels are converted to heat are principally the same [1].
The thermal conversion of biomass into energy is said to be more dependant on the pyrolysis step due to the presence of high moisture and volatile matters. The energy generated during pyrolysis is directly linked to the characteristics of raw biomass [3, 4] while the intermediate char particles determine the fate of carbon in ash as they combust to further provide heat or gasify to produce syngas [5, 6, 7, 8, 9]. Therefore, it is more important to study biomass in its original form and intermediate char structures, as well as the relationship between the two.
Biomass has three major organic components of cellulose, hemicellulose and lignin. These lignocellulosic components constitute plant cell walls, which are the main remains after biomass has been dried. Cellulose is the most abundant organic compound on earth and is the principal component of plant cell walls [10]. Lignin is the second most abundant organic substance in plants after cellulose [10].
Whilst electricity generation from oil and natural gas generates little or no ash respectively [11], coal combustion can produce significant amount of ash that inevitably leads to ash related problems. Ash slagging and fouling have always been a major factor with regards to boiler design and operation particularly as power stations now generally buy coals from all over the world.
A new challenge for boiler operators has been the introduction of biomass which has become increasingly popular in recent years, as a carbon neutral energy source. The low melting temperature of biomass ash, however, is problematic and has limited the utilization of biomass in both direct combustion and co-firing. This phenomenon is mainly due to the high alkaline content of biomass ash, particularly that of herbaceous biomasses [12]. Potassium is the main source of alkali in most biomass fuels [13, 14] located in the inherent mineral matter and is the main cause of ash deposition and corrosion. This is in contrast to coal, where sodium is the dominant and most problematic alkali metal [13, 14].
This work aims to investigate the relationship between the lignocellulosic components of biomass and the intermediate char morphology, as well as its ash characteristics using the advanced ash fusion test developed at The University of Nottingham.
2. Materials and Methods
Nine different biomasses were used for this study (Table 1). These biomasses cover a range of biomass types including energy crops, agricultural wastes and industrial wastes. A standard UK high volatile bituminous coal, Daw Mill was also used in this work.
Table 1. Types of samples used and their associated lignocellulosic composition. Starch together with the lignin, cellulose and hemicellulose reported add up to 100wt%.
Fuel
Type
Cellulose Content (wt%)
Hemicellulose Content (wt%)
Lignin Content (wt%)
Corn Stover
Agricultural Waste
24.1
30.4
10.4
Wheat Shorts
Agricultural Waste
22.2
0.0
10.6
Miscanthus
Energy Crop
57.9
30.8
11.4
Wheat
Industrial Waste
9.6
30.6
7.7
Sunflower Seed
Industrial Waste
48.4
40.4
11.2
Rapeseed
Industrial Waste
51.5
38.3
10.2
Olive Residue
Industrial Waste
70.5
19.8
9.7
Distillers Dried Grain (DDG)
Industrial Waste
1.5
76.9
12.0
Distillers Dried Grain with Solubles (DDGS)
Industrial Waste
7.9
71.3
8.3
Daw Mill
low rank high volatile bituminous
NA
NA
NA
The lignocellulosic compositions of the biomass were determined using standard biological assays. The biomass samples were chosen to show a great degree of variability (Table 1).
2.1. Char Preparation and Analysis
Raw biomass samples were riffled and sieved to produce 300-600 micron size range. A fixed bed furnace was used to pyrolyse the biomass in a high heating rate environment. Ceramic crucibles were filled with 10 g of raw biomass and then covered with a ceramic lid (to allow pyrolysis but with reduced air ingression in order to minimise combustion) before being placed directly into fixed bed furnace preheated to 1000°C. Samples were left in the furnace for 3 minutes to allow pyrolysis to be completed, after which the crucibles were removed and placed in a dessicator to avoid the reintroduction of moisture. It is understood that this method is not a direct simulation of a fluidised bed combustion system but nevertheless this process represents a high ramp rate, high temperature environment where the relative changes in each biomass type can be compared.
Prepared char samples were embedded in epoxy liquid resin to prepare polished blocks for optical characterisation. Polishing was carried out using aqueous Pedemat Struers Rotopol polisher with increasing alumina grades to produce a scratch free surface. Afterwards, a Zeiss Leitz Ortholux II POL-BK microscope with oil-immersion objective
creating a total of 320x magnification was used to analyse char morphology. Composite images (3090 x 3900 pixels) from mosaics of 15 x 15 were obtained from the digital camera Zeiss AxioCam attached to the microscope and operated using the KS400 V3.1 software. Manual point counting was also carried out (in a total of 500 points) to determine the different char morphotypes.
2.2. Ash Preparation and Analysis
The ash behaviour of each sample was investigated using the advanced ash fusion test developed at the University of Nottingham. The samples were ashed in a laboratory muffle furnace at 650°C following a standard protocol before being pelletized. Ash pellets of 1 cm diameter weighing 0.500g were prepared. These pellets were then heated up to 1520°C in an SDAF2000d Ash Fusion Analyser. The ash fusion system used a Viewse VC-523D high resolution black and white CCD camera to capture images of the pellets at a rate of 1 frame per degree increase whilst heating. These images were then analysed using Matlab (using a specifically developed code) to produce a profile of height/initial height versus temperature. These profiles were found to be unique to each sample.
Particulate ash samples, produced at 650°C, were also analysed using the SEM/EDAS in order to determine residual mineral compositions.
3. Results and Discussions
3.1. Char Analysis
The different char morphotypes were studied using polished sectioned blocks under the oil immersion microscope. This method is particularly useful in quantifying the different characteristics of biomass chars, such as area, length, width, diameter, wall thickness and macro porous calculation, observed across the cross section. Figure 1 shows the typical char morphotypes produced from biomass. Although various classification systems exist
for the coal chars [15, 16], these classes tend to relate to the macerals of coal and do not fully describe the different biomass char morphotypes [1]. The major difference is the presence of cellular pores in biomass chars which are similar to its original structure. Therefore, apart from the common criterion of thick walls, thin walls and solid, cellular and porous structures should be included to accommodate the unique characteristics of biomass.
(a)
(b)
(c)
(d)
(d)
Figure 1. The different biomass char morphotypes obtained from automated image analysis. (a) thin walled cell structure; (b) thin walled porous structure; (c) thick walled cell structure; (d) thick walled porous structure; (e) solid.
Table 2 details the char morphology data of the different biomass types studied. It is believed that the nature of this difference is linked to the lignocellulosic compositions of biomass. This is because each plant cell type is assigned to a specific function and associated with different amounts of lignin, cellulose and hemicellulose. Hence, the characteristics of cells are apparent e.g. xylem cells of the vascular tissue responsible for the transport of water are commonly heavily lignified to provide the necessary strength to withstand the negative pressure [17], therefore the thicker and stronger cell walls (primary and secondary walls) would produce thick cellular char structure (figure 1c) similar to its original form. In contrast, ‘weaker’ parenchyma cells of the ground tissue equipped with only primary walls would easily swell, rupturing the internal structure during the volatile release to produce thin porous chars (Figure 1b). Table 2. Percentage of different char morphotypes obtained from manual point counting Thin Walled Cell Structure 26.0
Thin Walled Porous Structure 14.8
Thick Walled Cell Structure 26.8
Thick Walled Porous Structure 30.4
Miscanthus Olive Residue
58.4
7.2
34.4
0.0
0.0
47.2
0.8
52.0
0.0
0.0
Rapeseed
1.6
8.8
64.0
24.0
1.6
Shorts
Solid 2.0
Corn Sunflower Seed
39.2
34.4
16.8
8.0
1.6
13.6
4.8
75.2
6.4
0.0
Wheat
20.0
40.8
14.4
20.0
4.8
DDG
6.4
11.2
30.4
50.4
1.6
DDGS
6.4
7.2
29.6
55.2
1.6
Figure 2. The linear relationship between percentage of cell structure and amount of cellulose and lignin in wt%, R2=0.9287
As expected, Figure 2 shows a linear increase in the percentage of chars with cellular structure as the amount of cellulose and lignin increases. Lignin is a characteristic of secondary thickening commonly found in cell walls of cells with specific function to provide strength, such as sclerenchyma and xylem. In such cases, lignin provides the required compressive strength [10]. Cellulose on the hand, through the winding of microfibrils [10], provides the necessary tensile strength, similar to that of steel [17]. With the adequate compressive and tensile strength to prevent the expansion of cells (and the compression of adjacent cells), coupled with the relatively low reactivity of the cellulose and lignin, the cellular structure is retained during volatile and moisture release. The lack of such strengthening factors makes the cell walls relatively easier to rupture
during pyrolysis, thereby breaking the boundaries between adjoining cells to produce larger pores. Figure 3 illustrates this effect.
Figure 3. The linear relationship between percentage of porous structure and amount of cellulose and lignin in wt%, R2=0.9325
It is believed that the lignocellulosic compositions of raw biomass could potentially be used to predict char morphology and hence, char reactivity. More work is needed on this.
3.2. Ash Analysis
The mineral compositions of each sample (biomass and coal) are summarized in Table 3.
Table 3. Mineral composition of samples tested in molar percentage Corn Olive Wheat stover DDG DDGS miscanthus residue rapeseed shorts Mol Mol Mol Mol Elem Mol % Mol % Mol % % % % %
Sunflower seed Mol %
wheat Mol %
coal Mol %
Na2O
0.65
6.44
9.46
0.52
0.42
0.00
0.64
0.63
0.00
1.55
MgO
12.83
27.18
21.12
4.03
6.56
17.58
31.74
21.18
16.37
4.03
Al2O3
0.00
0.00
0.00
1.26
4.77
0.00
1.20
1.16
0.00
20.42
SiO2
13.19
2.96
4.89
63.62
25.02
0.00
3.72
4.23
19.17
54.68
P 2O 5
12.89
28.23
23.97
2.45
1.39
20.36
29.24
3.14
24.29
0.35
SO3
6.70
1.47
6.14
3.45
2.19
10.65
0.66
10.04
2.20
6.54
K2O
24.33
31.58
28.93
8.39
5.42
19.85
29.60
30.42
29.68
2.13
CaO
14.01
1.73
5.14
15.88
52.38
31.56
2.63
28.56
7.79
5.99
TiO2
0.00
0.12
0.00
0.00
0.31
0.00
0.19
0.26
0.00
1.40
Fe2O3
15.39
0.28
0.35
0.40
1.56
0.00
0.38
0.37
0.49
2.91
An example of an ash profile is given in Figure 4. The shrinking temperature i.e. the temperature at which the horizontal section of each profile ends is tabulated in Table 4. This temperature signifies the first melting event whereby the ash pellet starts to shrink due to sintering.
Figure 4 – a typical ash profile showing the four phases of transformation during heating. Table 4.Shrinking temperatures of samples obtained from the advanced ash fusion test
Fuel Shrinking Temperature (°C)
Corn stover
DDG
DDGS
720
590
620
Miscanthus
Olive Residue
790
1110
Rapeseed
Wheat Shorts
Sunflower seed
Wheat
Daw Mill Coal
790
690
780
650
920
Figure 5 shows the relationship between the ash elemental compositions and the ash behaviour. This graph is produced using data from the biomass and coal samples tested. Group 1 denotes the amount of potassium and sodium according to the periodic table classification, while Group 2 denotes the amount of calcium and magnesium in the ash samples. As the ratio of Group 1/Group 2 increases, the inverse of the shrinking temperature increases linearly. This trend shows the lowering of shrinking temperature and hence melting temperature with the increase in Group 1/Group 2 ratio. This is because potassium is the main source of alkali in most biomass fuels [13, 14] located in the inherent mineral matter and it is therefore the main cause of ash deposition and corrosion. This is in contrast to coal, where sodium is the dominant and most problematic alkali metal [13, 14].
Figure 5. The linear relationship between ash behaviour (1/shrinking temperature) and ash composition (Group 1 elements/Group 2 elements), R2=0.9053
Figure 6 shows the link between lignocellulosic compositions of biomass and its respective ash elemental composition. This graph is produced using the different biomass samples tested. As the amount of cellulose (in different samples) increases, the Group1/Group 2 ratio decreases linearly. This is in accordance to the different minerals associated with different cells and their functions. Cellulose being the principal component of cell wall [10], is a good indicator of the number of cells present in the samples.
In addition, since cellulose is more abundant in secondary walls than primary walls [10], an increase in the amount of cellulose could potentially indicate an increase in the number of secondary walls within the samples. These walls are commonly associated with xylem cells which are normally dead at maturity i.e. the loss of cytoplasm and all of its contents [17], therefore the need for macronutrients such as potassium (required for the osmotic potential of cells and as enzyme cofactor) is reduced, hence the decrease in the Group 1/Group 2 ratio.
Figure 6. The relationship between ash composition (Group 1/Group 2 according to periodic table) and lignocellulosic composition (amount of cellulose in weight percentage). R2=0.9514
These two relationships could potentially be combined to link the increase in shrinking temperature of ash with cellulose content. This would relate biomass in its original form (lignocellulosic compositions) to its ash behaviour (shrinking temperature) and therefore be a potential means of predicting ash slagging/fouling propensity using only the characteristics of the biomass fuel in its original form.
4. Conclusions 1) A set of biomass samples were tested for their respective lignocellulosic composition, char morphotypes, ash elemental composition and ash behaviour at elevated temperatures. 2) An increasing linear relationship was established between the percentage of chars with cellular structure and the amount of cellulose and lignin. A decreasing trend was found for the porous structure. 3) An exponential increase in chars with thick cellular structure was observed with the increase in cellulose and hemicelluloses. A decreasing linear relationship was obtained between the percentage of thin porous chars and the amount of cellulose and hemicellulose. 4) A linear relationship between 1/shrinking temperature and group 1/group 2 ratio was found, potentially linking the ash behaviour to its elemental composition 5) A linear relationship between Group 1/ Group 2 ratio and the amount of cellulose in biomass was found, potentially relating the lignocellulosic composition to the ash elemental composition 6) A linear relationship between 1/shrinking temperature and the amount of cellulose in biomass was found, potentially linking biomass in its original form to its subsequent ash characteristics at higher temperatures 7) The lignocellulosic compositions of raw biomass could potentially be used to predict the char morphology and ash behavior of corresponding biomass. References
[1]Avila Claudio, Cheng Heng Pang, Tao Wu, Edward Lester, Morphology and Reactivity Characteristics of Char Biomass Particles, Bioresource Technology 102, 2011, 5237-5243 [2]Alonso, M.J.G., et al., Physicochemical transformations of coal particles during pyrolysis and combustion. Fuel 80 (13),2001, 1857–1870. [3]Lu, Q., et al., Overview of Fuel Properties of Biomass Fast Pyrolysis Oils, Energy. Convers. Manage. 50, 2009, 1376-1383. [4]Di Blasi, C., Influences of Physical Properties on Biomass Devolatilization Characteristics, Fuel 76, 1997, 957-964. [5]Asadullah, M., et al., Effects of Biomass Char Structure on Its Gasification Reactivity, Bioresource Technology 101, 2010, 7935-7943. [6]Cloke, M., et al., Char Characterization and Its Application for a Coal Burnout Model, Fuel 82, 2003, 1989-2000. [7]Wornat, M., Hurt, R., Yang, N., Headley, T., Structural and compositional transformations of biomass chars during combustion. Combust. Flame 100, 1995, 131– 143. [8]Ulloa, C., Borrego, A.G., Helle, S., Gordon, A.L., García, X., Char characterization and DTF assays as tools to predict burnout of coal blends in power plants. Fuel 84, 2005, 247–257. [9]Senneca, O., Kinetics of pyrolysis and gasification of three biomass fuels, Fuel Proc. Technol. 88,2007, 87–97. [10]Evert, R.F., Esau’s Plant Anatomy, John Wiley and Sons Inc, 2006 [11]Regan, J.W., Impact of Coal Characteristics on Boiler Design, Coal Technology ’82 5th International Coal Utilization Exhibition and Conference, 1982 [12]Fernandez Llorente, M.J., Carrasco Garcia, J.E., Comparing Methods For Predicting The Sintering Of Biomass ash in Combustion, Fuel 84, 2005, 1893-1900. [13]Baxter Larry, L. et al, The Behavior of inorganic material in biomass-fired power boilers: field and laboratory experiences, Fuel Processing Technology 54, 1998, 47-78. [14]Sjaak, V. L., Koppejan, J., The Handbook of Biomass Combustion and Co-firing, Earthscan, 2008.
[15]Kwiecinska, B., Petersen, H.I., Graphite, Semi-graphite, natural coke and natural char classification-ICCP System, Int, J. Coal Geol. 57-2, 2004, 99-116. [16]Lester, E., et al., The Procedure Used to Develop a Coal Char ClassificationCommission III Combustion Working Group of the International Committee for
Coal
and Organic Petrology, Int. J. Coal Geol. 81, 2010, 333-342. [17]Taiz, L., Zeiger, E., Plant Physiology Fourth Edition, Sinauer Associates, 2006
