Impact of Mineral Impurities in Solid Fuel Combustion

Impact of Mineral Impurities in Solid Fuel Combustion

This page intentionally left blank.

Impact of Mineral Impurities in Solid Fuel Combustion Edited by

R. P. Gupta CRC for Black Coal Utilisation University of Newcastle NSW, Australia

T. F. Wall CRC for Black Coal Utilisation University of Newcastle NSW, Australia

and

L. Baxter Sandia National Laboratories Livermore, California

KLUWER ACADEMIC PUBLISHERS NEW YORK, BOSTON, DORDRECHT, LONDON, MOSCOW

eBook ISBN: Print ISBN:

0-306-46920-0 0-306-46126-9

©2002 Kluwer Academic Publishers New York, Boston, Dordrecht, London, Moscow Print ©1999 Kluwer Academic / Plenum Publishers New York All rights reserved

No part of this eBook may be reproduced or transmitted in any form or by any means, electronic, mechanical, recording, or otherwise, without written consent from the Publisher

Created in the United States of America

Visit Kluwer Online at: and Kluwer's eBookstore at:

http://kluweronline.com http://ebooks.kluweronline.com

PREFACE This book contains papers presented at the Engineering Foundation Conference on mineral matter in fuels held on November 2-7, 1997 in Kona, Hawaii. The conference is one of a continuing series that was initiated by the CEGB Marchwood Engineering Laboratories in 1963. The conference was to be eventually organised by the Engineering Foundation as the need for multi-disciplinary work related to controlling ash effects in combustors became apparent. The conference covers both the science and the applications. The papers also present case histories, particularly for

current fuel technologies, developments in advanced technologies for power generation and mathematical modelling of these processes. Developments since 1963 have been slow, but steady, due to the complexity of the chemical and physical processes involved. However, the research presented here displays great improvement in our understanding of the mechanisms by which mineral matter will influence fuel use. Steve Benson from EERC presented a review and current status of issues related to ash deposition in coal combustion and gasification.

The application of new analytical tools, which have been detailed in the previous conferences, is presented. These include CCSEM, as well as new techniques for characterising sintering of ash, such as TMA, image analysis, X-ray diffraction crystallography

and thermal analysis. The new analytical techniques were extended to encompass widely differing fuels such as biomass. Ole H Larsen from ELSAM Denmark presented a review of these advanced techniques. Thermodynamic equilibrium calculations and knowledge of mineralogy has helped in understanding ash reaction mechanisms and evaporation of alkali species during combustion and gasification. Together with the analytical tools, and mathematical models based on these mechanisms, the leap to practical predictions is now possible. Several examples of this approach are detailed. For example, Jouni Pyykonen from VTT modelled ash deposition in a boiler using the mineral distribution from CCSEM and FLUENT, a computational fluid dynamics code. John Harb and his group from BYU extended the models to predict the changes in heat transfer in boilers with deposit growth. There has not been enough emphasis in the past on the estimates of thermal properties

of ash. A review article on thermal conductivity of ash like materials presented by Raj Gupta from CRC Black Coal is a significant step in the direction to rectify this problem. Thermal properties of deposits are shown to be strongly influenced by the deposit structure. A greater integration of data from advanced analytical techniques, mechanistic models and sophisticated CFD codes is expected in the future in order to predict the thermal performance of boilers. The conference is truly an international one, with participants representing all major laboratories working in this technical area. Many papers outline collaborative efforts, which have become a feature of modern research. This requires a special mention

V

vi

Preface

of the ELSAM project involving collaboration between Danish, US and Japanese researchers. The fuels of interest are somewhat continent-specific. In Europe, there is major interest in the use of biomass and its blending with other fuels. Larry Baxter from Sandia Laboratories reviewed the issues related to ash deposition in co-firing of biomass with coal. In Australia, the major interest is in coal. In Japan and the USA, all fuels are of interest. There were several studies on large power station boilers and pilot scale plants specially related to advanced technologies. The technologies considered include pulverised fuel, co-firing aid advanced power systems, such as PFBC and IGCC. FBC technology is highly suitable for low rank and low grade coals. The economic is pushing the usage of such fuels. A complete session was devoted to the ash problems and their solutions in FBC such as agglomeration and additives. There were a number of reviews and keynote papers in the field of FBC and other advanced technologies (Basu from BHEL India, Zhang from CRC Low Rank Coal and Ohman from ETC Sweden). The conference presents research in traditional atmospheric pressure systems but also considers newer systems where higher pressures are used and where the mineral matter in the fuel might be extracted as slag rather than the ash, which has been the traditional combustion product. Mr Sadayuki Shinozaki from CCUJ presented an extensive review of the advanced clean coal technologies. The Bryers Award for the best paper was presented to Lone Hansen of the Technical University of Denmark for her paper on “Ash Fusion Quantification by Use of

Thermal Analysis”, and Raj Gupta of The CRC for Black Coal Utilisation on his paper on “The Thermal Conductivity of Ash Deposits: Particulate and Slag Structures”. We also noted two excellent papers presented by Tim Heinzel of the University of Stuttgart and Bengt-Johan Skrifvars of Abo Akademi University, Finland.

It was particularly gratifying to have Richard Bryers at the conference to present the award named after him and also to give a keynote paper. We would like to thank a number of sponsors who provided support to allow the attendance of delegates who would otherwise not have been able to attend. The sponsors were: Cooperative Research Centre for Black Coal Utilisation, Australia; Cooperative Research Centre for Power Generation from Low-Rank Coal, Australia; Centre for Coal Utilisation, Japan; Idemitsu Kosan Co. Ltd., Japan; EBARA Corporation, Japan; IHI Co. Ltd., Japan; Kawasaki Heavy Industries, Japan; Electric Power Development Company, Japan and Japan Cement Association, Japan. We would also like to thank the Engineering Foundation for their assistance, particularly Richard Fein and Rosa Landinez who provided on-site support during the conference. We are grateful to all the session chairpersons for organising their sessions and reviewing the papers. In addition, Hongwei Wu and Chris Bailey, two PhD students from the CRC for Black Coal Utilisation, helped with arrangements. We would also like to thank Susan Safren of Plenum Publishing for her care, help and forbearance in the preparation of this book. R. P. Gupta

T. F. Wall L. Baxter

CONTENTS

Keynote Papers Ash-Related Issues During Combustion and Gasification . . . . . . . . . . . . . . . . . . . Steven A. Benson and Everett A. Sondreal Mineral Characterization for Combustion: The Contribution from the Geological Sciences . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

1

23

Colin R. Ward The Development of Power Technologies for Low-Grade Coal . . . . . . . . . . . . . . K. Basu

33

Low-Rank Coal and Advanced Technologies for Power Generation . . . . . . . . . . . Dong-ke Zhang, Peter J. Jackson, and Hari B. Vuthaluru

45

The Thermal Conductivity of Ash Deposits: Particulate and Slag Structures . . . R. P. Gupta, T. F. Wall, and L. Baxter

65

The Development of Advanced Clean Coal Technology in Japan: Mineral Matter Issues . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Sadayuki Shinozaki

Factors Critically Affecting Fireside Deposits in Steam Generators Richard W. Bryers

..........

85

105

SECTION I Mineral Matter, Ash and Slag Characterisation Advanced Analytical Characterization of Coal Ashes—An Idemitsu Kosan—Elsam Cooperation Project . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Ole Hede Larsen, Flemming J. Frandsen, Lone A. Hansen, Signe Vargas, Kim Dam-Johansen, Karin Laursen, Takeo Yamada, and

133

Tsuyoshi Teramae vii

viii

Contents

A Novel Application of CCSEM for Studying Agglomeration in Fluidised Bed Combustion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Mika E. Virtanen, Ritva E. A. Heikkinen, H. Tapio Patrikainen, and

147

Risto S. Laitinen

Thermomechanical Analysis and Alternative Ash Fusibility Temperatures . . . . . S. K. Gupta, R. P. Gupta, G. W. Bryant, L. Juniper, and T. F. Wall

155

Ash Fusibility Detection Using Image Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . Klaus Hjuler

171

Ash Fusion Quantification by Means of Thermal Analysis . . . . . . . . . . . . . . . . . Lone A. Hansen, Flemming J. Frandsen, and Kim Dam-Johansen

181

Sticking Mechanisms in Hot-Gas Filter Ashes . . . . . . . . . . . . . . . . . . . . . . . . . . . John P. Hurley, Bruce A. Dockter, Troy A. Roling, and Jan W. Nowok

195

Classification System for Ash Deposits Based on SEM Analysis . . . . . . . . . . . . Karin Laursen and Flemming J. Frandsen

205

Determination of Amorphous Material in Peat Ash by X-Ray Diffraction . . . . Minna S. Tiainen, Juha S. Ryynänen, Juha T. Rantala, H. Tapio Patrikainen, and Risto S. Laitinen

217

System Accuracy for CCSEM Analysis of Minerals in Coal . . . . . . . . . . . . . . . . R. P. Gupta, L. Yan, E. M. Kennedy, T. F. Wall, M. Masson, and K. Kerrison

225

The Microstructure and Mineral Content of Pulverised Coal Chars . . . . . . . . . F. Wigley and J. Williamson

237

SECTION II The Use of Low-rank and Low-grade Coals and Cofiring

Fireside Considerations when Cofiring Biomass with Coal in PC Boilers . . . . . Allen L. Robinson, Larry L. Baxter, Gian Sclippa, Helle Junker, Karl E. Widell, Dave C. Dayton, Deirdre Belle-Oudry, Mark Freeman, Gary Walbert, and Philip Goldberg

247

Summary of Recent Results Obtained from Using the Controlled Fluidised Bed Agglomeration Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Marcus Öhman and Anders Nordin

Deposition and Corrosion in Straw- and Coal-straw Co-fired Utility Boilers: Danish Experiences . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Flemming J. Frandsen, Hanne P. Nielsen, Peter A. Jensen, Lone A. Hansen, Hans Livbjerg, Kim Dam-Johansen, Peter F. B. Hansen (1) and Karin H. Andersen (2), Henning S. Sørensen, Ole H. Larsen, Bo Sander, Niels Henriksen, and Peter Simonsen

259

271

Contents

ix

Development of Blast-Furnace Gas Firing Burner for Cofiring Boilers with Pulverized Coal . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Takashi Kiga, Takehiko Ito, Motoya Nakamura and Shinji Watanabe

285

Changes in Slagging Behaviour with Composition for Blended Coals . . . . . . . . Nicholas J. Manton, Jim Williamson, and Gerry S. Riley

297

Role of Inorganics During Fluidised-Bed Combustion of Low-Rank Coals . . . Hari Babu Vuthaluru and Dong-ke Zhang

309

Role of Additives in Controlling Agglomeration and Defluidization During Fluidised Bed Combustion of High-Sodium, High-Sulphur Low-Rank Coal . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Temi M. Linjewile and Alan R. Manzoori The Agglomeration in the Fluidized Bed Boiler During the Co-Combustion of Biomass with Peat . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Ritva E. A. Heikkinen, Mika E. Virtanen, H. Tapio Patrikainen, and Risto S. Laitinen Ash Fusion and Deposit Formation at Straw Fired Boilers . . . . . . . . . . . . . . . . .

319

333

341

Lone A. Hansen, Flemming J. Frandsen, Henning S. Sørensen, Per Rosenberg, Klaus Hjuler, and Kim Dam-Johansen SECTION III

Case Studies in Conventional and Advanced Plant

Influence of Metal Surface Temperature and Coal Quality on Ash Deposition in PC-Fired Boilers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Karin Laursen, Flemming J. Frandsen, and Ole Hede Larsen Full Scale Deposition Trials at 150MW E PF-Boiler CO-Firing Coal and Straw: Summary of Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

357

367

Karin H. Andersen, Flemming J. Frandsen, Peter F. B. Hansen, and Kim Dam-Johansen Slagging Tests on the Suitability of Alternative Coals in a 325 MW E PC Boiler . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Timm Heinzel, Jörg Maier, Hartmut Spliethoff, Klaus R. G. Hein. and Werner Cleve

Predicting Superheater Deposit Formation in Boilers Burning Biomasses . . . . . . Rainer Backman, Mikko Hupa, and Bengt-Johan Skrifvars

383

405

Deposit Formation in the Convective Path of a Danish 80 MW TH CFB-Boiler CO-Firing Straw and Coal for Power Generation . . . . . . . . . . Peter F. Binderup Hansen

417

Research on the Melting Points of Some Chinese Coal Ashes . . . . . . . . . . . . . .

433

Shen Xianglin, Chen Ying, and Liu Haibin

x

Contents

Computer Controlled Scanning Electron Microscopy (CCSEM) Analysis of Straw Ash . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Henning Sund Sørensen

Prognosis of Slagging and Fouling Properties of Coals Based on Widely Available Data and Results of Additional Measurements . . . . . . . . . . . . . . Alexander N. Alekhnovich, Natalja V. Artemjeva, Vladimir V. Bogomolov, Vyacheslav I. Shchelokov, and Vasilij G. Petukhov The Slagging Behaviour of Coal Blends in the Pilot-Scale Combustion Test Facility . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Alexander N. Alekhnovich, Vladimir V. Bogomolov, Natalja V. Artemjeva, and Vladimir E. Gladkov

441

455

471

SECTION IV Studies at Rig Scale (Including Corrosion) In Situ Measurements of the Thermal Conductivity of Ash Deposits Formed in a Pilot-Scale Combustor . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Allen L. Robinson, Steven G. Buckley, Gian Sclippa, and Larry L. Baxter

485

Low Corrosivity of Coal Chlorine . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Elliott P. Doane and Murray F. Abbott

497

Laboratory Studies on the Influence of Gaseous HCl on Superheater Corrosion Keijo Salmenoja, Mikko Hupa, and Rainer Backman

513

The Role of Alkali Sulfates and Chlorides in Post Cyclone Deposits from Circulating Fluidized Bed Boilers Firing Biomass and Coal . . . . . . . . . . . . Bengt-Johan Skrifvars, Tor Laurén, Rainer Backman, and Mikko Hupa

Fly Ash Deposition onto the Convective Heat Exchangers during Combustion of Willow in a Circulating Fluidized Bed Boiler . . . . . . . . . . . . . . . . . . . . . . Terttaliisa Lind, Esko I. Kauppinen, George Sfiris, Kristina Nilsson, and Willy Maenhaut

525

541

Ash Behaviour in Biomass Fluidised-Bed Gasification . . . . . . . . . . . . . . . . . . . . . Antero Moilanen, Esa Kurkela, and Jaana Laatikainen-Luntama

555

Bench-Scale Biomass/Coal Cofiring Studies . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Deirdre Belle-Oudry and David C. Dayton

569

Iron in Coal and Slagging: The Significance of the High Temperature Behaviour of Siderite Grains During Combustion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Gary Bryant, Christopher Bailey, Hongwei Wu, Angus McLennan, Brian Stanmore, and Terry Wall

581

Contents

xi

SECTION V Develpments in Advanced Coal Technologies

Fractionated Heavy Metal Separation in Biomass Combustion Plants—Possibilities, Technological Approach, Experiences . . . . . . . . . . . . . Ingwald Obernberger and Friedrich Biedermann

595

Distributions of Major and Trace Elements in Entrained Slagging Coal Gasification Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Kenichi Fujii, Masamitsu Suda, Tadayoshi Muramatsu, and Masahiro Hara

609

Behavior of Inorganic Materials During Pulverized Coal Combustion . . . . . . . . Tsuyoshi Teramae, Toru Yamashita, and Takashi Ando Energy Production from Contaminated Biomass: Progress of On-Going Collaboration Projects . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Alexandre Grebenkov, Anatoli Iakoushev, Larry Baxter, Dave Allen, Helle Junker, and Jørn Roed

621

635

Triboelectrostatic Coal Cleaning: Mineral Matter Rejection In-Line Between

Pulverizers and Burners at a Utility

...............................

645

John M. Stencel, John L. Schaefer, Heng Ban, TianXiang Li, and James K. Neathery Development of Advanced PFBC . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Makoto Takai and Masahiro Hara

653

Operating Experiences of 71 MW PFBC Demonstration Plant . . . . . . . . . . . . . . . Hideki Goto and Syoichi Okutani

663

Development of an Innovative Fluidized Bed Cement Kiln System . . . . . . . . . . . Sadayuki Shinozaki, Isao Hashimoto, Katsuji Mukai, and Kunio Yoshida

675

The Influence of Pressure on the Behaviour of Fuel Carbonates . . . . . . . . . . . . . Arvo Ots, Tõnu Pihu, and Aleksander Hlebnikov

685

SECTION VI Modeling of Ash Behaviour and Ash Deposition

Modeling of Ash Deposit Growth and Sintering in PC-Fired Boilers . . . . . . . . . Huafeng Wang and John N. Harb Predicting Ash Behavior in Conventional Power Systems: Putting Models to Work . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Christopher J. Zygarlicke

Thermodynamic Modelling of the System to Characterise Coal Ash Slags . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Evgueni Jak, Sergei Degterov, Arthur D. Pelton, Jim Happ, and Peter C. Hayes

697

709

723

xii

Development of a Prediction Scheme for Pulverised Coal-Fired Boiler Slagging . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Jouni Pyykönen, Jorma Jokiniemi, and Tommy Jacobson Modelling the Initial Structure of Ash Deposits and Structure Changes Due to Sintering . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Hamid R. Rezaei, Rajender P. Gupta, Terry F. Wall, S. Miyamae, and K. Makino

Index . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

Contents

735

753

767

ASH-RELATED ISSUES DURING COMBUSTION AND GASIFICATION

Steven A. Benson and Everett A. Sondreal Energy & Environmental Research Center University of North Dakota PO Box 9018, Grand Forks, North Dakota 58202-9018 USA

phone [701] 777-5177, e-mail [email protected]

1. INTRODUCTION The effects of ash on the performance of combustion and gasification systems depends on the inorganic composition of the fuel and on operating conditions. Ash is known to be a major problem that results in decreased efficiency, unscheduled outages,

equipment failures, and high cleaning costs. The detrimental effects of fuel-associated inorganic components on combustor or gasifier process performance include slag-tapping problems, fireside ash deposition, corrosion and erosion of system parts, production of fine particulates that are difficult to collect, blinding of filtering media, formation of hazardous air pollutants, and production of precursors to the formation of secondary fine particulate. The amount of literature on ash-related issues is immense. Overviews of ashrelated issues and compilations of work by many investigators can be found by referring to the work of Baxter and DeSollar [1996], Couch [1994], Williamson and Wigley [1994],

Benson and others [1993b], Benson [1992], Bryers and Vorres [1990], Raask [1985, 1988], and Watt [1969]. Variability in chemical and physical properties is a major problem in burning and gasifying fuels in an environmentally acceptable manner. For coal, most utilization problems are related to the complex associations of the inorganic components. The association and abundance of major, minor, and trace elements in coal is dependent upon coal rank and depositional environment. The abundance and association of minerals in coal have been reviewed [Benson and others, 1993b; Raask, 1988]. The sulfur oxide emissions in power plants are derived from organic sulfur and mineral forms such as pyrite, gypsum, barite, and others. Trace elements that can cause pollution and operational problems are found associated with coal and waste-derived fuels. Many considered to be air toxics— Hg, Cd, Pb, As, and others—are associated with sulfides. The association, fate, and Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

1

2

S. A. Benson and E. A. Sondreal

behavior of air toxic metals have been reviewed and published in a Special Issue of Fuel

Processing Technology [Benson and others, 1994]. Additionally, the primary precursors to the formation of respirable particles (particles less than 2.5 micrometers) are partly derived from sulfur dioxide and nitrogen oxide emissions produced as a result of combustion [U.S. Environmental Protection Agency, 1996]. The prediction of ash behavior relative to the performance of conversion, hotgas cleanup, and air pollution control systems based on conventional ASTM (American Society for Testing and Materials) methods of analysis is severely limited because of the inadequacy of such methods to determine the distributed characteristics of the inorganic components as they exist in the coal. Computer-controlled scanning electron microscopy (CCSEM) [Jones and others, 1992] is used to determine the size, composition, and abundance of minerals in coals. In lower-ranked coals, chemical fractionation is used to determine the abundance of organically associated cations that are important to predict ash deposition behavior [Benson and Holm, 1985]. The effects of these minerals on system performance and emissions can be predicted with a much higher degree of certainty using CCSEM and other advanced methods of analysis [Benson and others, 1993a]. This paper, based mainly on work conducted over the last three decades at the Energy & Environmental Research Center (EERC), focuses on the influence of ash con-

stituents on the fireside performance of conversion systems. The assessment of fuel con-

stituents will primarily involve coal; however, the effects of biomass and wastes will also be contrasted and compared.

2. RESULTS AND DISCUSSION 2.1. Fuel Quality Currently, a wide range of materials is combusted or gasified for production of chemical feedstocks and energy, destruction of hazardous materials, and reduction in the volume of waste. Table 1 provides a summary of the bulk chemical composition

of selected fuels and waste materials that are currently being fired. The components of the fuel that have been considered most significant with respect to the fireside performance of fuels in combustion and gasification systems include alkali and alkalineearth elements and iron, which when combined with silicates can produce wall slagging and high-temperature convective pass fouling problems. At lower temperatures, the alkali and alkaline-earth elements combine with sulfur, halogen species, and sometimes carbonates, producing bonded deposits on heat-transfer surfaces. The alkali and alkaline-

earth elements can also contribute to the blinding of hot-gas filters and corrosion of system parts.

The abundance and mode or form of occurrence of major, minor, and trace inorganic components have an influence on their transformations upon conversion and ultimate fate in the conversion and environmental control systems upon gas cooling. Figure 1 illustrates the modes of occurrence of inorganic components in coals. Table 2 lists the modes of occurrence of vapor, minor, and trace elements. The modes of occurrence of the major elements in fuels are known and can be determined by a variety of methods, including CCSEM [Kong and others, 1996] and chemical fractionation [Benson and Holm, 1995]. The modes of occurrence for trace elements are not as well known because of the uncertainty in determining the association. Finkelman [1994] reviewed the modes

Ash-Related Issues During Combustion and Gasification

3

4

S. A. Benson and E. A. Sondreal

of occurrence of potentially hazardous elements in coal and levels of confidence in the

modes. The abundance in the various modes changes dramatically with coal range and fuel type. Biomass typically contains high levels of organically associated calcium, potas-

sium, and phosphorus in addition to silicates and aluminosilicate contaminants. When combined in the conversion process, low-melting-point silicates, sulfates, phosphates, and carbonates produce high- and low-temperature deposits.

The most significant problem in the effective utilization of the fuels is the high variability in the inorganic composition of the fuels. For example, coals vary considerably from mine to mine and within seams. In order to effectively use coals with high variability, planning and assessing the effects of fuel handling, preparation, and distribution within the coal-fired power plant are necessary.

2.2. Transformations/Growth/Sintering The inorganic constituents in fuels and waste undergo complex chemical and physical transformations during the conversion process and gas cooling. The types of transformations depend on the modes of occurrence of the inorganic material in the fuel or waste and on conversion conditions. The transformations include processes illustrated in Fig. 2. The inorganic components that undergo volatilization during conversion usually condense within the conversion and environmental control system. The components that have the potential to remain in the vapor state through the air pollution control system include various forms of vapor-phase Hg and Se and HF, and HCl. Most of the volatilized components will condense within the overall system. The condensation and/or reaction of the volatilized species usually produce liquid phases that result in the formation of sticky particles and deposits. Condensed phases that are known to cause significant problems upon condensation include sulfides, sulfates, phosphates, chlorides, fluorides, carbonates, and some oxides. Figure 3 illustrates the distribution of trace

Ash-Related Issues During Combustion and Gasification

5

elements as a function of particle size produced from gasification of a bituminous coal [Benson and others, 1994a]. Some trace elements are known to accumulate in various

areas within the conversion system. Examples of these phases include calcium arsenate, selenium compounds, barium sulfate, strontium sulfate, lead sulfate/sulfides, and antimony compounds [Benson and others, 1994a].

6

S. A. Benson and E. A. Sondreal

The nonvolatile components such as silicates, aluminosilicates, and iron-rich phases that pass through a liquid state are responsible for producing the bulk of the ash deposits. Significant experimental and modeling efforts have been conducted using laboratory-,

pilot-, and full-scale combustion testing to elucidate the mechanisms of formation of intermediate ash species. The intermediate species are in the form of gases, liquids, and solids. Computer models [Erickson and others, 1993; Wilemski, 1992] are able to predict the size, composition, and abundance of solid and liquid particles and the abundance and composition of gas-phase components upon conversion and gas cooling. Knowledge of the chemical and physical properties of the intermediate materials is a prerequisite to predicting transport and deposit growth.

Ash-Related Issues During Combustion and Gasification

7

Ash particle transport, deposit growth, and sintering mechanisms depend upon the velocity, temperature, and composition of the gas phase; the temperature, size, and composition of the particles; the composition and temperature of the deposition surface; and the physical characteristics of the particles and deposition surface. Ash transport mechanisms include both small-particle (less than 5 and large-particle (greater than Small-particle transport is dominated by thermophoresis, electrophoresis, and diffusion. Large-particle transport mechanisms are dominated by inertial impaction. These mechanisms have been described in detail by Rosner [1986]. Deposit growth and sintering are dominated by the abundance and viscosity of the liquid-phase components [Raask, 1988]. The initial sticking of particles for ash deposit growth is not completely quantified. Viscosity of the surface of the particles has been found useful in predicting the stickiness of ash. Sticking coefficients for deposit growth are defined as the ratio of ash sticking rate to ash firing rate. The critical value for sticking is in the range of to poise [Wilemski and others, 1992]. Strength development in materials is the result of a combination of sintering processes. According to Watt [1969], there are four primary mechanisms: vapor transport, surface diffusion, volume diffusion, and viscous or plastic flow. These processes are all distinct, but likely occur simultaneously. The mechanism of viscous flow sintering [Frenke, 1945] appears to be

the dominant mechanism in creating the high-temperature silicate-based melts in ash deposits. The other mechanisms likely play a more significant, if not dominant, role at lower temperatures. The process involved in the viscous flow sintering of particles is illustrated in Fig. 4. On the basis of Raask [1988], the onset of sintering occurs at to poise, moderate sintering at to rapid sintering at to formation of nonflowing slag at to slow movement of slag at to and good slag flow at 250 poise. Currently, methods used to predict slag flow behavior are based on empirical relationships derived from bulk chemical analysis of the slag and viscosities measured as a function of temperature with a rotating-bob viscometer. The computational methods

8

S. A. Benson and E. A. Sondreal

have been reviewed and revised by a number of investigators, including Watt [1969], Schobert and others [1985], Kalmanovitch and Frank [1990], and Folkedahl [1997]. Many of the methods are limited in their range of application for predicting the viscosity of slags. Only a limited amount of work has been conducted in developing relationships for various chemical compositions at high viscosities above poise, which is the upper limit for the rotating-bob viscometer. Research is being conducted at the EERC to measure high viscosities at low temperatures using a heated-stage microscope. The method uses the Frenke [1945] relationship to calculate the viscosity of the melt based on the neck growth between particles. The technique has the potential to measure viscosity up to poise. Tests are currently being conducted on particles derived from glasses and homogeneous slags of known composition. Figure 5 illustrates the sintering of sodalite spheres. The physical parameter controlling slag flow is viscosity. For example, the maximum viscosity at which a slag can be tapped from a cyclone-fired boiler is 250 poise. The temperature at a viscosity of 250 poise is called T250. Generally, for a coal to be appropriate for cyclone firing, it must have a slag T250 of 2,600°F or lower. In addition, the slag must be of sufficient thickness such that it will retain the burning coal particles, thereby transferring the coal ash to the slag and completing the combustion in the cyclone. In order to maintain slag flow, a boiler has to be operated at a temperature higher than T250. Figure 6 illustrates the measured viscosity temperature curves for a Powder River Basin (PRB) coal slag under various atmospheres [Folkedahl, 1997]. Some PRB coals

exhibit extremely low viscosities, making it difficult to maintain a sufficiently thick slag layer. These low-viscosity slags also freeze rapidly, as shown in Fig. 6. Atmosphere affects the viscosity of the slag. The most pronounced effect on viscosity resulted from changing the test atmosphere from air or reducing gas to a simulated combustion gas (calculated based on coal composition and excess air), which caused an extremely high freezing temperature.

Ash-Related Issues During Combustion and Gasification

9

As shown in Figure 6, another factor that influences the flow of slag is freezing, or crystallization. The temperature at which a sudden transition in viscosity occurs is called the temperature of critical viscosity Crystallization causes a rapid increase in viscosity on cooling below this temperature. Crystallization can cause sluggish slag flow and

must be considered in slags that exhibit such behavior. In estimating when a slag is completely molten, the following consideration should be used: if

then slag removal temperature =

if

then slag removal temperature =

2.3. Combustion 2.3.1. Pulverized Fuel. Extensive work has been carried out at the EERC for the past 30 years on ash fouling and slagging related to coal firing. Much of the early work was conducted on lower-rank western U.S. coals. More recently, bituminous and other fuels such as waste and biomass fuels have been tested. This work has been conducted using bench-, pilot-, and full-scale testing. To date in the ash-fouling unit (750,000 Btu/hr combustor), the EERC has conducted over 725 pulverized coal-fired tests and over 200 water-slurried fuel tests. Numerous runs have been made using a bench-scale drop-tube furnace system equipped with fly ash and ash deposition probes. The general mechanisms of ash deposit formation are illustrated in Fig. 7. Field testing has been conducted at more than 20 plants over the past 10 years in efforts to determine the mechanisms of ash formation and deposition aimed at developing ways to predict and mitigate ash-related problems. The key to better understanding ash behavior problems at the EERC over the last decade has been in developing advanced methods for sampling and aerodynamically classifying ash particulates by size, measuring deposit growth on cooled probes in pilot-plant

and full-scale systems, and analyzing inorganic species in coal and coal ash by a combination of 1) chemical fractionation for determining exchangeable cations and soluble

10

S. A. Benson and E. A. Sondreal

minerals, 2) CCSEM for mineral grains, and 3) scanning electron microscopy point count (SEMPC) for analyzing phases in fly ash, deposits, and slags. Bulk composition and crystalline phases are determined by x-ray fluorescence (XRF) and x-ray diffraction (XRD), respectively. Ash transformation research at the EERC has, over time, delineated the hightemperature ash chemistry and particle-size distributions and mass transport mechanisms that control the rate of deposition and deposit strength development during combustion [Benson and others, 1996; McCollor and others, 1996; Zygarlicke and others, 1992]. The analysis of mineral grains and organically associated cations in low-rank coal can be used to predict patterns of fusion, vaporization, heterogeneous and homogeneous condensation, and coalescence or fragmentation. The products of these transformations are a wide combination of fused mineral fragments, inorganic vapor species, and condensed surface coatings and fine particulates. Ash deposition depends on the transport of these products to the cooler heat-transfer surface by a combination of vapor and smallparticle diffusion, thermophoresis, electrophoresis, and large-particle inertial impaction. Cohesion in deposit is provided by van der Waals, surface tension, electrostatic, and interlocking particle effects. Deposit growth and strength development depend most critically on the formation of low-viscosity melt phases at a microscopic level, which can be predicted by viscosity calculations based on the distributed CCSEM and SEMPC analysis

of precursors and products. Project Sodium, an industry-sponsored study [Benson and others, 1988] provided a detailed understanding of sodium-related ash-sintering mechanisms that are critical in high-temperature superheat sections of pf- and cyclone-fired boilers. The partitioning of ash species as a function of state and particle size was shown to be determined by their associations in the coal. Organically associated sodium was confirmed to be the most significant predictive factor of ash sintering and deposit growth. The sodium in hightemperature deposits was found to be concentrated in the amorphous melt phase and not

Ash-Related Issues During Combustion and Gasification

11

in the crystalline phases. Most of the sodium-containing liquid phase identified in the high-temperature deposits (above l,900°F or 1,038 °C) was derived from aluminosilicates, but mixed alkali and alkaline-earth sulfates were found to be important at lower temperatures. Figure 8 illustrates the relationship between the surface tension and viscosity

and the relative deposition potential of the coal. At low surface tension-to-viscosity ratios, the deposition potential was high as determined in a pilot-scale combustion system. This relationship is consistent with the Frenke model, in which the sintering rate is related to surface tension/viscosity. Project Calcium, a multiclient program for high-calcium U.S. low-rank coals [Hurley and others, 1992], characterized several types of ash fouling deposits of operational significance below 1,900°F (1,038°C), where the thermodynamic equilibrium favors the formation of sulfate phases, including calcium sulfate crystals that impart strength to deposits. Deposit formation was studied both in utility boilers using cooled probes and in laboratory sintering tests. Entrained ash was aerodynamically sized, collected, and characterized to determine composition and size relationships. Deposits were collected from the same location as the entrained ash. Massive deposits forming on the upstream side of the leading tubes in the temperature zone below 1,900°F were rapidly sulfated and hardened unless removed at once by sootblowing. On the tubes behind the lead tubes, enamel-like deposits were formed by upstream deposition of small calciumenriched ash particles smaller than 3 the enamel hardened almost immediately and was most effectively removed by thermal shock using water blowers. A common type of powdery deposit formed on the downstream side of tubes from ash smaller than was easily removed by suitably angled sootblower nozzles; however, this deposit type had the greatest potential for reducing heat transfer if left unremoved because of its large surface coverage. Numerous field tests were conducted as part of Project Calcium to collect and characterize entrained ash and deposits. Figure 9 shows the partitioning of

12

S. A. Benson and E. A. Sondreal

Ash-Related Issues During Combustion and Gasification

13

the major inorganic ash components in the various size fractions of ash and the composition of the deposit formed on the upstream side of a reheater tube in a tangentially fired combustor. A predictive PC model of low-temperature deposition [LEADER] was developed that estimates deposition rates, deposit strength, and heat exchange loss on the basis of the coal analyses, including the ultimate analysis, the XRF ash analysis, and CCSEM coal mineral analysis. 2.3.2. Fluidized-Bed Combustion Systems. Agglomeration and deposition in fluidizedbed combustors (FBCs) [Hajicek and others, 1985; Henderson and others, 1995; Benson and others, 1982; Goblirsch and others, 1982, 1983] are determined by the same type of sulfate-based fouling that occurs in the lower-temperature zones of conventional boilers, except that the fluidized-bed material introduces another reactant and an entirely different fluid dynamic regime. Agglomeration of bed particles by sticky surface coatings

causes progressive accretion that leads to pressure and temperature fluctuations and ultimately the forced shutdown of the FBC. A high concentration of sodium in the coal ash is once again an important predictive factor. The EERC has investigated this problem in pilot-scale bubbling and circulating fluidized beds, in utility atmospheric fluidized-bed combustors (AFBCs) burning U.S. lignite and subbituminous coal and operated with sand and limestone beds, and in the Tidd clean coal demonstration of pressurized fluidized-bed combustion (PFBC). The mechanisms of ash formation in FBC systems are illustrated in Fig. 10. Much work has also been conducted on inorganic transformations and agglomerate formation during firing of Australian low-rank coals [Manzoori, 1990; Manzoori and others, 1992; Manzoori, 1992; Manzoori and Argarwal, 1993, 1994]. The primary contributor to increasing the sintering propensity of the bed particles was NaCl. More recently, Mann [1997] reviewed the literature related to the fate of alkali in FBC systems. Mann also examined sorbents that would remove alkali from PFBC flue gas to levels below that specified by turbine manufacturers. Methods have been established for determining the rate of agglomeration and deposition based on

14

S. A. Benson and E. A. Sondreal

sieve analysis and SEM image analysis on the spent-bed material. The link between coal ash properties and coating/agglomerate formation is determined by the chemical fractionation and the CCSEM and SEMPC methods discussed above. The results of an ongoing investigation are being used to optimize the selection and replacement rate of bed material in relation to sulfur control requirements and the effect of coal ash constituents on agglomeration. Hot-gas filter ash characterization studies are being performed by the EERC in partnership with EPRI, the U.S. Department of Energy (DOE), and a consortium of companies to develop predictive models of ash-related problems in hot-gas filters for PFBC and integrated gasification combined-cycle (IGCC) systems. This program has been successful in identifying the causes of ash bridging between candle filters and related filter breakage that occurred during certain operating conditions at the Tidd PFBC Clean Coal demonstration project [Hurley and others, 1996]. Analysis of composition and size distributions down to 0.1 in diameter using the EERC’s SEMPC method established that ash bridging was caused by the smaller particles derived from the coal, cemented by a sticky potassium-enriched surface layer, rather than by larger particles derived from the dolomite sulfur sorbent. Thermochemical equilibrium modeling indicated that about 2% of the coal ash would exist in a liquid phase of potassium carbonate and sulfate at the conditions on the filter. The laboratory tensile strength of a simulated ash filter cake was shown to increase with temperature. Aged deposits from the Tidd filter were found to be

more sulfated than other filter cake material, which could account for the hardening and expansion of deposits that led to filter breakage.

2.4. Gasification 2.4.1. Fluidized Bed—Transport Reactor. The transport reactor demonstration unit (TRDU) (shown in Fig. 11) at the EERC is a 100-kg/hr pressurized circulating fluidizedbed gasifier that models the larger gasifier of similar design being tested by DOE at Wilsonville, Alabama, as a facility for evaluating hot-gas cleaning systems [Mann and others, 1995; Swanson and Ness, 1997]. The EERC aims to demonstrate acceptable filter performance at the pilot scale before long-term tests are performed at Wilsonville. The TRDU has also proven to be a valuable tool for evaluating ash deposition under the conditions of fluid dynamics and ash chemistry found in a circulating bed gasifier [Benson and Sondreal, 1996]. Deposits impeding operation of the unit formed in bends near the top of the gasifier loop, at the top of the riser, and at the cyclonic gas-bed disengager when a U.S. subbituminous coal was gasified that contained very finely divided clay and quartz in association with high levels of organically bound calcium. Analysis of these deposits by SEMPC found them to be enriched in low-melting-point calcium magnesium aluminosilicates derived from the coal ash and sulfided dolomite. A similar chemical composition was observed in the finer fraction of the circulating bed and in bridging deposits on the candle filters installed during the test. The bulk chemical composition of size-fractionated ash and the filter-bridging deposits are illustrated in Fig. 12. The results illustrated the important linkage between the operability of an advanced power system and the chemistry and morphology of the coal ash.

2.4.2. Entrained Flow Gasification. Coal ash behavior in reducing environments (CABRE) [Benson and others, 1994b] is a phased investigation being performed by the EERC under the support of different industrial consortia to develop analytical methods to characterize and classify reduced phase species found in species that condense to form

Ash-Related Issues During Combustion and Gasification

15

low melting eutectics in gasification systems and to understand the high-temperature ash transformation mechanisms, develop predictive methodologies, and identify the properties of volatile sulfide species that condense to form low melting eutectics in gasification and hot-gas cleaning systems. Device-specific results are proprietary to the sponsors, but some general observations can be made. Corrosive alkali–iron sulfide eutectics are chemically stable in a gasification environment only below about 750 °C. CaS is stable at higher temperatures up to about 900 °C and may form CaS–FeS solid solutions at lower temperatures where FeS is stable. Eutectics formed from FeS, and are believed to be principal causes of ash sintering at temperatures below 700 °C. Gas transport processes that move volatile species from higher-temperature to lower-temperature zones serve to reduce the viscosity of eutectic melts and intensify sintering below 700 °C. Sulfide forming on the surface of an aluminosilicate particle is illustrated in Fig. 13.

16

S. A. Benson and E. A. Sondreal

2.5. Fuel Quality Planning Predictive indices of ash fouling and slagging in conventional boilers developed for bituminous coals have been shown by experience to be largely unsuitable for low-rank coals. The EERC has recently compiled a series of indices for low-rank coals into a computer software package designated as the Predictive Coal Quality Effects Screening Tool or PCQUEST [Zygarlicke and others, 1996]. PCQUEST, as illustrated in Fig. 14, is

based primarily on the EERC’s ash transformation and transport research and related computer modeling, and the models used rely on detailed mineralogical and chemical fractionation analyses in addition to the conventional bulk measurements of coal moisture, principal elements, and ash oxides. Eight indices are calculated by PCQUEST to

Ash-Related Issues During Combustion and Gasification

17

predict furnace wall slagging, high-temperature fouling, low-temperature fouling, slag tapping, opacity, tube erosion, coal grindability, and sootblower requirements. Utilities and coal companies have used the indices to map coal quality at the mine and follow performance at the power plant, and results have recently been validated [Zygarlicke and

McCollor, 1997]. Dynamic models to predict ash formation, transport, deposit growth, and removability have been developed [Erickson and others, 1995; Allan and others, 1996]. These models uses an ash transformation model to predict the particle-size composition distribution [Erickson and others, 1993; Wilemski and others, 1992] of the ash, which becomes input to the ash transport portion of the model. The ash transport model uses smallparticle (diffusion and thermophoresis) and large-particle (inertial impaction) mechanisms. Once the particles are transported to the surface, the deposited particles increase in temperature because of the insulating properties of the ash. The high temperatures increase the level of liquid in the deposit, resulting in increases in the ash-sticking coefficients. Deposit strength develops through the formation of low-viscosity liquid phases. The overall process is illustrated in Fig. 15 for prediction of fouling deposit formation.

3. SUMMARY AND CONCLUSIONS Improving the operability and performance of combustion/gasification systems and associated environmental systems hinges on our understanding the transformation processes involved in the formation of inorganic gas, liquid, and solid species that impede

heat transfer and gas flow, and promote degradation of materials. Understanding ash behavior allows for proper selection of fuels, operating conditions, and system modification and design. Variability in chemical and physical properties is a major problem in the burning and gasifying of fuels in an environmentally acceptable manner.

The prediction of ash behavior relative to the performance of conversion, hot-gas

18

S. A. Benson and E. A. Sondreal

cleanup, and air pollution control systems on the basis of conventional ASTM methods

of analysis is severely limited because of the inadequacy of such methods to determine the distributed characteristics of the inorganic components as they occur in the coal. CCSEM can be used to determine the size, composition, and abundance of minerals in coals as a more valid basis for prediction. Significant progress has been made in understanding ash formation mechanisms and fate in conversion and environmental control systems. Assessing and predicting ash

behavior in pf-fired boilers have made significant advances over the years. The behavior of ash in FBC systems is becoming better understood. The reducing environment present in gasification systems results in the formation of a wide range of species that are not

well understood with regard to their effects on performance. Challenges remain in developing accurate methods for measuring the composition of highly variable fuels and feedstocks, assessing and monitoring the variability of fuels, developing methods to predict the impact of ash, meeting the environmental requirements related to ash components, and applying this information to a wide range of fuels and wastes.

4. REFERENCES Allan, S.E., Erickson, T.A., and McCollor, D.P. (1996). “Modeling of Ash Deposition in the Convective Pass of a Coal-Fired Boiler.” L. Baxter and R. DeSollar (Eds.), Applications of Advanced Technology to AshRelated Problems in Boilers, New York: Plenum Press, pp. 451–470. Baxter, L., and DeSollar, R. (Eds.) (1996). Applications of Advanced Technology to Ash-Related Problems in Boilers. New York: Plenum Press.

Benson, S.A. (Ed.) (1992). Inorganic Transformations and Ash Deposition During Combustion. New York: American Society of Mechanical Engineers for the Engineering Foundation.

Benson, S.A., and Holm, P.L. (1985). “Comparison of Inorganic Constituents in Three Low-Rank Coals.” Ind. Eng. Chem. Prod. Res. Dev., 24, 145. Benson, S.A., and Sondreal, E.A. (1996). “Impact of Low-Rank Coal Properties on Advanced Power Systems.” Proceedings of the Pittsburgh Coal Conference, September 1996, pp. 484–498.

Ash-Related Issues During Combustion and Gasification

19

Benson, S.A., Karner, F.R., Goblirsch, G.M., and Brekke, D.W. (1982). “Bed Agglomerates Formed by Atmospheric Fluidized-Bed Combustion of a North Dakota Lignite.” Prepr. Pap.—Am. Chem. Soc., Div. Fuel Chem., 27(1), 174–181. Benson, S.A., Fegley, M.M., Hurley, J.P., Jones, M.L., Kalmanovitch, D.P., Miller, B.G., Miller, S.F., Steadman, E.N., Schobert, H.H., Weber, B.J., Weinmann, J.R., and Zobeck, B.J. (1988). “A Detailed Evaluation of

Sodium Effects in Low-Rank Coal Combustion Systems.” Project Sodium Final Technical Report, July. Benson, S.A., Hurley, J.P., Zygarlicke, C.J., Steadman, E.N., and Erickson, T.A. (1993a). “Predicting Ash Behavior in Utility Boilers.” Energy & Fuels, 7(6), 746–754. Benson, S.A., Jones, M.L., and Harb, J.N. (1993b). “Ash Formation and Deposition.” D.L. Smoot (Ed.), Fundamentals of Coal Combustion for Clean and Efficient Use. New York: Elsevier, Chapter 4, pp. 299–373. Benson, S.A., Erickson, T.A., Zygarlicke, C.J. (1994a). “Transformations of Trace Metals in Coal Gasification.”

Presented at the Joint AFRC/JFRC Pacific Rim International Conference on Environmental Control of Combustion Processes, Maui, HI, October 16–20, 1994. Benson, S.A., Erickson, T.A., Brekke, D.W., Folkedahl, B.C., Tibbetts, J.E., and Nowok, J.W. (1994b). “Coal

Ash Behavior in Reducing Environments.” DOE/METC-94, Vol. 1, pp. 322–331. Benson, S.A., Steadman, E.N., Mehta, A.K., and Schmidt, C.E. (Eds.) (1994). Trace Element Transformations in Coal-Fired Power Plants, Special Issue of Fuel Processing Technology, 39 (1–3). Benson, S.A., Steadman, E.N., Zygarlicke, C.J., and Erickson, T.A. (1996). “Ash Formation, Deposition, and Erosion in Conventional Boilers,” L. Baxter and R. DeSollar (Eds.), Applications of Advanced Technology to Ash-Related Problems in Boilers. New York: Plenum Press, pp. 1–15. Benson, S.A., Pavlish, J.H., Zygarlicke, C.J. (1998). “Trace Elements in Low-Rank Coals.” Proceedings of the Fifteenth Annual Pittsburgh Coal Conference, September 14–18, 1998. ISBN 1-890977-15-2. Brooker, D.D., and Oh, M.S. (1995). “Iron Sulfide Deposition During Coal Gasification.” S.A. Benson (Ed.), Special Issue of Fuel Processing Technology, Ash Chemistry in Fossil Fuel Processes, 44(1–3), 181–190. Bryers, R.W. (1995). “Utilization of Petroleum Coke and Petroleum Coke/Coal Blends as a Means of Raising Steam.” Fuel Processing Technology, 44, 121–144. Bryers, R.W., and Vorres, K.S. (1990). Proceedings of the Engineering Foundation Conference on Mineral Matter

and Ash Deposition from Coal. Santa Barbara, CA: United Engineering Trustees Inc. Couch, G. (1994). “Understanding Slagging and Fouling During Combustion.” J. Williamson and F. Wigley (Eds.), The Impact of Ash Deposition on Coal Fired Plants: Proceedings of the Engineering Foundation Conference. London: Taylor & Francis.

Erickson, T.A., Allan, S.E., McCollor, D.P., Hurley, J.P., Srinivasachar, S., Kang, S.G., Baker, J.E., Morgan, M.E., Johnson, S.A., and Borio, R. (1995). “Modeling Fouling and Slagging in Coal-fired Utility Boilers.” Special Issue of Fuel Processing Technology, Ash Chemistry in Fossil Fuel Processes, 44(1–3),

155–171. Erickson, T.A., O’Leary, E.M., Folkedahl, B.C., Ramanathan, M., Zygarlicke, C.J., Steadman, E.N., Hurley, J.P., and Benson, S.A. (1993). “Coal Ash Behavior and Management Tools,” J. Williamson and F. Wigley (Eds.), The Impact of Ash Deposition in Coal Fired Plants, London: Taylor & Francis, pp. 271–282.

Finkelman, R.B. (1994). “Modes of Occurrence of Potentially Hazardous Elements in Coal, Levels of Confidence.” Fuel Processing Technology, 39(1–3), 21–34. Folkedahl, B.C. (1997). “A Study of the Viscosity of Coal Ash and Slag,” Ph.D. Thesis, The Pennsylvania State

University. Frenke, J. (1945). “Viscous Flow of Crystalline Bodies Under the Action of Surface Tension.” Journal of Physics, 9(5), 385–390.

Goblirsch, G.M., Benson, S.A., Hajicek, D.R., and Cooper, J.L. (1982). “Sulfur Control and Bed Material Agglomeration Experience in Low-Rank Coal AFBC Testing.” Presented at the Seventh International Conference on Fluidized-Bed Combustion, Philadelphia, PA, October 1982. DOE/FC-1005, p. 15. Goblirsch, G.M., Benson, S.A., Karner, F.R., Rindt, D.K., and Hajicek, D.R. (1983). “AFBC Bed Material Performance with Low-Rank Coals.” Presented at the Twelfth Biennial Lignite Symposium, Grand Forks, ND, May 18–19, 1983, pp. 1–25. Gunderson, J.R., Anderson, C.M., Moe, T.A., Bolin, K., and Klosky, M. (1995). “Combustion Characteristics of RDF.” EERC 95-04-03, April. Hajicek, D.R., Zobeck, B.J., Mann, M.D., Miller, B.C., Ellman, R.C., Benson, S.A., Goblirsch, G.M., Cooper,

J.L., Guillory, J.L., and Eklund, A.G. (1985). “Performance of Low-Rank Coal in Atmospheric Fluidized Bed Combustion.” EERC Topical Report for the U.S.

Department of Energy.

DOE/FE/60181/1869, October. Henderson, A.K., Mann, M.D., Swanson, M.L., and Erickson, T.A. (1995). “Development of Methods to Predict Agglomeration and Deposition in FBCs.” Proceedings of the Advanced Power Systems 95 Review

20

S. A. Benson and E. A. Sondreal Meeting. DOE/METC-95/1018, Vol. 1, June, pp. 440–447.

Hurley, J.P., Benson, S.A., Erickson, T.A., Allan, S.E., and Bieber, J.A. (1992). “Project Calcium Final Report,” KERC Publication, Grand Forks, ND: University of North Dakota. Hurley, J.P., Watne, T.M., O’Keefe, C.A., Katrinak, K.A., Nowok, J.W., Roling, T.A., and Dockter, B.A. (1996). “Chemical and Physical Analyses of Tidd Hot-Gas Filter Ash,” Proceedings of the Pittsburgh Coal Conference, September, 129–134.

Jones, M.L., Kaimanovitch, D.P., Steadman, E.N., Zygarlicke, C.J., and Benson, S.A. (1992). “Application of SEM Techniques to the Characterization of Coal and Coal Ash Products.” H.L.C. Meuzelaar (Ed.), Advances in Coal Spectroscopy. New York: Plenum Press. Kalmanovitch, D.P., and Frank, M. (1990). “An Effective Model of Viscosity for Ash Deposition Phenomena.” R.W. Bryers and K.S. Vorres (Eds.), Mineral Matter and Ash Deposition from Coal, United Engineering Trustees, Inc. Karner, F.R., Schobert, H.H., Falcone, S.K., Benson, S.A. (1986). “Elemental Distribution and Association with Inorganic and Organic Components in North Dakota Lignites,“ Mineral Matter and Ash in Coal. K.S. Vorres (Ed.), ACS Symposium Series 30: Washington, DC.

Kong, L., Zygarlicke, C.J., Benson, S.A. (1996). “Computer-Controlled Scanning Electron Microscopy Analysis of Minerals in Coal,” Proceeding of the Thirteenth Annual Pittsburgh Coal Conference, September 3–7, 1996. Vol. 1, p. 241.

Mann, M.D. (1997). “Capture of Alkali During Pressurized Fluidized-Bed Combustion Using In-Bed Sorbents.” Ph.D. Thesis, University of North Dakota, May. Mann, M.D., Swanson, M.L., Ness, R.O., and Haley, J.S. (1995). “Hot-Gas Filter Testing with the Transport Reactor Demonstration Unit.” Proceedings of the Advanced Power Systems 95 Review Meeting. U.S. Department of Energy, DOF/METC-95/1O18, Vol. 1, June, pp. 89–97.

Manzoori, A.R. (1990). “Role of the Inorganic Matter in Agglomeration and Defluidization During the Circulating Fluid Bed Combustion of Low Rank Coals.” Ph.D. Thesis, The University of Adelaide. Manzoori, A.R., Lindner, E.R., and Agarwal, P.K. (1992). “Inorganic Transformation During the Circulating Fluid Bed Combustion of Low-Rank Coals With High Content of Sodium and Sulphur.” Inorganic Transformations and Ash Deposition During Combustion. S.A. Benson (Ed.), New York: American

Society of Mechanical Engineers, pp. 735–762. Manzoori, A.R. (1992). “The Fate of Organically Bound Inorganic Elements and Sodium Chloride During Fluidized Bed Combustion of High Sodium, High Sulphur Low Rank Coals.” Fuel, 71,513–522. Manzoori, A.R., and Agarwal, P.K. (1993). “The Role of Inorganic Matter in Coal in the Formation of Agglomerates in Circulating Fluid Bed Combustors.” Fuel, 72(7), 1069–1075. Manzoori, A.R., and Agarwal, P.K. (1994). “Agglomeration and Defluidization Under Simulated Circulating Fluidized-Bed Combustion Conditions.” Fuel, 73(4), 563–568. McCollor, D.P., Zygarlicke, C.J., and Benson, S.A. (1996). “Mechanisms of Ash Fouling During Low-Rank Coal Combustion,” E. Baxter and R. DeSollar (Eds.), Applications of Advanced Technology to AshRelated Problems in Boilers, New York: Plenum Press, pp. 223–235. Raask, E. (1985). Mineral Impurities in Coal Combustion. Washington: Hemisphere Publication Corporation. Raask, E. (1988). Erosion Wear in Coal Utility Boilers. Washington: Hemisphere Publication Corporation. Rosner, D.E. (1986). Transport Processes in Chemically Reacting Flow Systems. Stoneham, MA: Butterworth, 539 p. Schobert, H.H. (1995). “Lignites of North America.” Coal Science and Technology, 23rd ed. New York: Elsevier.

Schobert, H.H., Streeter, R.C., and Diehl, E.L. (1985). “Flow Properties of Low-Rank Coal Ash Slags,” Fuel, 64,1611. Skrivfars, B.J., Hupa, M., Moilanin, A., and Lundqvist, R. (1996). “Characterization of Biomass,” E. Baxter and R. DeSollar (Eds.), Application of Advanced Technology to Ash-Related Problems in Boilers. New York: Plenum Press. Swaine, D.J. (1990). Trace Elements in Coal. Boston: Butterworths.

Swanson, M.L., Ness, R.O., Jr. (1997). “Hot-Gas Filter Testing with a Transport Reactor Development Unit.” Presented at the Advanced Coal-Based and Environmental Systems ’97 Conference, Pittsburgh, PA, July 22–24. U.S. Environmental Protection Agency. (1996). Review of National Ambient Air Quality Standards for

Particulate Matter: Policy Assessment of Scientific and Technical Information. EPA-452\R-96-013, July. Watt, J.D. (1969). “The Physical and Chemical Behavior of the Mineral Matter in Coal Under the Conditions Met in a Combustion Plant, Part I I . ” BCURA Industrial, Leather Head Laboratories. Wilemski, G., Srinivasachar, S., and Sarofim, A.F. (1992). “Modeling of Mineral Matter Redistribution and Ash Formation in Pulverized Coal Combustion.” S.A. Benson (Ed.), Inorganic Transformations and Ash Deposition During Combustion. Engineering Foundation Press.

Ash-Related Issues During Combustion and Gasification

21

Williamson, J., and Wigley, F. (Eds.) (1994). The Impact of Ash Deposition on Coal Fired Plants: Proceedings

of the Engineering Foundation Conference. London: Taylor & Francis. Zygarlicke, C.J., Galbreath, K.C., McCollor, D.P., and Toman, D.L. (1996). “Development of Fireside Performance Indices tor Coal-Fired Utility Boilers,” L. Baxter and R. DeSollar (Eds.), Applications of Advanced Technology to Ash-Related Problems in Boilers. New York: Plenum Press, pp. 617–635. Zygarlicke, C.J., and McCollor, D.P. (1997). “Validation of Fireside Performance Indices.” EERC Report for the U.S. Department of Energy, July. Zygarlicke, C.J., Ramanathan, M., and Erickson, T.A. (1992). “Fly Ash Distribution and Composition: Experimental and Phenomenological Approach.” S.A. Benson (Ed.), Inorganic Transformations and Ash Deposition During Combustion. Engineering Foundation Press.

This page intentionally left blank.

MINERAL CHARACTERIZATION FOR COMBUSTION The Contribution from the Geological Sciences

Colin R. Ward School of Geology University of New South Wales Sydney 2052 Australia

1. INTRODUCTION The material referred to as “mineral matter” in coal has been to the combustion engineer a close to random association of chemical elements. These “random” elements are transformed in the combustion process to ash, react to form boiler deposits (slagging); abrade the internal parts of the boiler (erosion), and produce vapor phases that react with the metals in the boiler to cause metal loss (corrosion). Mineral matter in coal, however, is a far from random occurrence. The elements present in coal occur in predictable assemblages related to the processes that formed the coal. Knowledge of the way in which a coal formed, and of its post-depositional history, provides a guide to the forms and types of “mineral matter” that can be expected. This paper aims to develop such a theme, and to introduce the mineralogical tools that will allow the operator of combustion systems to characterize the non-organic component of coal, for it is this component that is the source of many problems in the combustion process. The knowledge base and techniques of the geological sciences have much to offer the combustion engineer in this particular field. Mineral characterization, often aided by the use of thermodynamic data to draw inferences as to the physical and chemical environment of formation, is the standard approach in the geological sciences. This applies not only to the relatively low-temperature processes that form the minerals actually found in coal, but also to the high-temperature processes associated with the formation and crystallization of magmas, which are similar in many ways to the behaviour of minerals in the combustion environment. When added to the generation of engineering indices, Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

23

24

C. R. Ward

there is the possibility to have predictive tools that are better based on the reality of the processes involved.

2. GEOLOGICAL ORIGIN OF MINERALS IN COAL Coal is made up of an organic and an inorganic component. The organic component is represented by the carbon-oxygen-hydrogen-nitrogen assemblages that are the coal macerals, the carbon analogues of the mineral material. The macerals in a coal reflect the parent plant material and the depositional and post-depositional histories of the coal deposit. There are many variations, which are reflected in the wide range of ranks and types that constitute the world’s coal resources. Mineral matter in coal encompasses three classes of material, namely:

• dissolved inorganic ions or compounds in the coal’s pore or surface water; • inorganic elements or compounds, with the exception of nitrogen or sulphur, incorporated in some way with the coal’s organic molecules; • crystalline mineral particles. Components of the first two types make up a large proportion of the mineral matter of lower-rank coals (brown coals and lignites). Changes in organic matter structure, however, reduce these to relative insignificance in higher-rank coals; crystalline mineral particles make up most if not all of the mineral matter in bituminous coals and anthracites. The minerals most commonly found in coals, including lower-rank coals, are listed

in Table 1. The relative abundance of these may vary considerably, depending on factors such as the depositional and hydrochemical environment in which the peat was formed, and the history of fluid migration, heating and alteration which the coal has suffered during and after burial. The minerals in coal (as opposed to the non-mineral inorganics) may represent particles of quartz, clay and other sediment washed or blown into the peat swamp. These are the detrital minerals. They may, however, also represent inorganic material produced by organisms, the biogenic minerals. Examples of the latter are shells of animals or diatoms that lived in the swamp and inorganic secretions (phytoliths) developed within the coal-forming plant tissues. A third possibility is that some of the minerals may represent compounds precipitated from solution. These are the authigenic minerals. They form in two ways, either precipitated in the pores associated with the plants that make up the peat in the early stages of coal formation, or deposited in cleats and other fractures during later stages of the coal’s geological history. The first, those associated with the early history, are the primary precipitates; the cleat fillers formed at later stages are called secondary precipitates. The processes of formation are important in combustion, because primary precipitates, particularly silica phytoliths, are normally fine grained and distributed throughout the coal in such a form that they are not unlocked when the coal is ground, even to the PF specification of 70% passing 75 microns. They remain in the PF particle and

subsequently reach temperatures beyond 2,000 °C, greater than the fuming temperature of silica.

Mineral Characterization for Combustion

25

In mined coals the minerals may also include components derived from intraseam bands or layers of non-coal material within the seam, or from dilution of the coal by associated roof and floor rocks. Igneous intrusions may also contribute to the mineral matter in mined coals, either directly if the intrusions themselves are extracted or indirectly through formation of mineral-impregnated heat-altered coal or natural coke phases. Much of the mineral matter introduced to mined products from intra-seam or inter-seam materials may be removed by washing. The minerals that can be washed from the coal are sometimes referred to as “extraneous” mineral matter. Such minerals are only significant if the coal is used as a run-of-mine (ROM) or un-washed product. The

mineral matter intimately associated with the coal itself, the “inherent” mineral matter, can possibly be removed only by chemical demineralization techniques. There is, however, a set of elements associated with the organic component of the coal that cannot be removed by washing or chemical demineralization. These elements, mainly locked in the organic compounds, are an unavoidable part of the coal when it is used in combustion processes.

3. METHODS OF MINERAL MATTER ANALYSIS Effective use of coal by modern technology requires definitive information on the nature and distribution of the minerals present, as opposed simply to the chemical composition of the ash. It also increasingly requires mineralogical information to be

26

C. R. Ward

supplied in quantitative form. This produces a need to assess the relative proportions of the different minerals or mineral groups present, together with the size of the mineral grains, the nature of any mineral associations, and possibly a range of mineral shape parameters. Quantitative mineralogical analysis, however, is an inherently difficult process, due to factors such as variability of composition and variations in crystallographic characteristics. The indices that have been used to assess coals for specific purposes have relied to date on elemental analysis, or ratios of elemental analyses. However, it is the minerals that react in the furnace, not the elements. Many of the minerals and other inorganics in coal undergo changes or interactions at high temperature, and as a result some of the mineral matter components may be lost from the solid phase during carbonization or combustion processes. The ash of the coal therefore usually represents only the nonvolatile or non-combustible residue of the mineral matter. By implication the elemental ash analysis is not representative of even the original inorganic elements in the coal, much less the mineralogy. The relation between mineral matter and ash is complex, particularly in low-rank coals. Kiss (1982), for example, indicates that up to one-third of the ash produced by combustion of Australian brown coals may be derived from fixation of organic sulphur by calcium, sodium or magnesium originally present in non-crystalline form, rather than as direct residues of the coal’s crystalline mineral particles.

A number of different techniques have been used for determination of the mineral matter (as opposed to ash) percentage in a coal sample. These include acid digestion of the mineral components (Radmacher and Mohrhauer, 1955), heating for a protracted period at around 370 °C, and calculations based on ash yield and other chemical data (King et al., 1936). Phong-anant et al. (1991), in their study of mineral matter in coal, recognized that no single analytical method was adequate for characterization of coal minerals and their transformation in boiler deposits, but noted that a combination of techniques was effective and desirable.

3.1. The Mineralogical Tools Mineralogical tools are diverse. They range from the human eye to sophisticated computer-controlled instruments. Each has a specific application, and the data delivered are limited or enhanced by the instrument used. In a typical investigation the tools can be thought of as qualitative or quantitative. Qualitative instruments can deliver descriptive data, and in some instances semiquantitative information. Such instruments are used to gain an overview; they are for reconnaissance. The optical microscope and the scanning electron microscope (SEM) are

two examples. Quantitative instruments deliver numbers. Quantitative systems are of two varieties, those that deliver precise determinations from single points and those that provide determinations, in some cases less precise, from many points on a specimen. 3.1.1. Optical Microscopy. Minerals in coal can be seen under the microscope in coal petrology studies (Stach et al., 1982), and volume percentages of the main mineral groups are typically included along with maceral percentages in the output from petrographic analyses. Many of the minerals are difficult to identify, however, partly because of a lack of distinguishing characteristics in reflected light and partly because they are fine-grained and often intimately associated with the organic components. Although useful in identifying the mode of mineral occurrence, petrographic analysis

Mineral Characterization for Combustion

27

based on point counting under the optical microscope gives only an approximate guide to the total mineral matter content. 3.1.2. Scanning Electron Microscopy. Many of the limitations on mineral matter identification using optical microscopy can be overcome if polished sections or other samples of the coal are examined using a scanning electron microscope (Stanton and Finkelman, 1979; Corcoran, 1989). SEM technology enables higher magnification to be used, and also allows the chemical composition of individual particles to be investigated by an accessory energy-dispersive X-ray spectrometer (EDS) unit. The EDS capability of the SEM gives the instrument a semi-quantitative capacity that can be a useful adjunct to its imaging and characterization role.

3.1.3. X-ray Diffraction Analysis. X-ray diffraction (XRD) analysis is a technique that has traditionally been considered to be qualitative, but with recent developments has become more quantitative. Identification of minerals by XRD is based on the crystallographic characteristics of the different components. Variations in crystallographic characteristics, however, coupled in some cases with preferred orientation or differential X-ray absorption effects, traditionally limit the technique to a qualitative or at best semi-quantitative role (Renton, 1986). Several techniques have been developed to improve the quantitative determination of mineral percentages by X-ray diffraction. One of these, based on principles detailed by Rietveld (1969), is the computer-based S IROQUANT system developed by Taylor (1991). Use of this technique on minerals isolated from coals of the Callide Basin, Australia, in comparison to estimates of mineral percentages based on normative evaluation of chemical analysis data from the high-temperature (815°C) ash (see below), has been investigated by Ward and Taylor (1996). Direct application of SIROQUANT to coal and similar organic-rich materials, without a preliminary low-temperature ashing step, has also been used by Mandile and Hutton (1995) to indicate both the overall mineral matter content and the relative proportions of the different minerals present. XRD techniques are particularly suited to identification of the clay minerals, which in many cases make up the bulk of the total mineral matter (Gluskoter, 1967; Rao and Gluskoter, 1973; Ward, 1977, 1978) and are difficult to identify positively by other methods.

3.1.4. Electron Microprobe Analyser. The electron microprobe analyser (EPMA) is a prime example of an instrument able to give precise measurements from a single data point. EPMA technology is well advanced (Reed, 1996), and is a mainstay for materials science in general. The data from the EM PA for minerals can be plotted on phase diagrams, and inferences drawn from these plots. It is a specialist tool that in certain cases can yield unique results, and has the advantage of being direct and non-destructive. Ramsden and Shibaoka (1982) have demonstrated well the application of the technique in a benchmark study where ash types were characterized in an exhaustive appraisal that related composition to morphology and erected a general classification system for fly ash particles. Electron microprobe analysis is most widely used for study of relatively heavy (high atomic number) elements, including components such as sulfur and chlorine in the organic matter of coal samples. Development of special measuring techniques in recent years, however, has also allowed it to be applied also to determination of light elements, including the carbon and oxygen in the organic matter of individual coal macerals (Bustin

28

C. R. Ward

et al., 1993; Ward and Gurba, 1997). Such a capacity can usefully complement investigations of mineral particles. 3.1.5. Computer-controlled SEM and Image Analysis Systems. Computer-controlled scanning electron microscopy (CCSEM) has been under development for the coal industry over some years, and incorporates many techniques common to other instruments that use electron beams. QEM*SEM (quantitative evaluation of materials by scanning electron microscopy) is perhaps the best example. The application of this

technology is the subject of several other papers in this volume, and so details are not listed here. QEM*SEM has been under development for the minerals industry and has become a prime tool in mineral processing. The object of QEM*SEM is to obtain statistically reliable information on the materials that are passing through the process. It also has image analysis capabilities, and can generate statistical information on textural features such as associations, rimming and simple shape functions, although the latter are still under development. The instrument’s application to coal and the products of coal combustion (Creelman et al., 1993) has been mainly in the field of erosion. It has been used in this capacity for erosion studies in fluidized-bed combustors. Creelman and Ward (1996) have recently applied (QEM*SEM) to the quantitative

evaluation of mineral matter in bituminous coal samples. The technique determines the association of chemical elements at individual points on a coal polished section from the output of several X-ray analysers directed at each point under the SEM. The element association at each point is then processed to identify the mineral species or group involved. Data from numerous such points in a scan of the sample are integrated using the image analysis unit, enabling volumetric assessment of the relative proportions of the different minerals or element-associations present. The system also has the capacity to map the individual mineral particles, and to provide a statistical evaluation of the size and shape of the different mineral occurrences in the coal sample.

3.2. Integrated Techniques for Quantitative Evaluation of Coal Mineral Matter 3.2.1. Low-temperature Oxygen-plasma Ashing. Oxidizing the organic matter and isolating the minerals without major alteration can be achieved by ashing coal at low temperature (120–150°C) in an electronically-excited oxygen plasma (Gluskoter, 1965). This probably represents the most reliable method for determining the percentage of total mineral matter (Standards Australia, 1990). The use of hot, concentrated hydrogen

peroxide to remove the organic matter and isolate an unaltered mineral fraction (Ward, 1974) may represent a useful substitute in some circumstances, but has more limited overall application. The nature of the minerals isolated by either technique can then be investigated by X-ray diffraction and similar methods. Precise determination of the mineral matter percentage in coal by low-temperature ashing also involves correction of the oxygen-plasma ash yield for un-oxidised organic carbon residues and for sulfur fixed in the LTA from the organic sulfur component (Standards Australia, 1990). These are often relatively small corrections, however, and for many purposes the proportion of low-temperature ash (LTA) gives an adequate indication of the percentage of mineral matter present.

Mineral Characterization for Combustion

29

3.2.2. Normative Interpretation of Ash Analysis Data. Estimates of the relative proportions of the different minerals in coal can also be made by interpretation of the chemical composition of the coal’s conventional (high-temperature) ash, as determined in routine analytical programs. Such techniques are based on the assumption that the various elements in the ash were originally partitioned between particular minerals, and hence represent an attempt to derive a “theoretical” mineralogy from the chemical analysis data. They are most effective if the actual minerals present are known from independent evidence, such as X-ray diffraction, and if the element proportions in minerals that may be of variable composition, such as interstratified clay minerals, can also be determined for that particular sample. An interactive computer-based procedure for normative analysis of sedimentary materials (SEDNORM), developed by Cohen and Ward (1991), has been applied to ash analysis data for a range of Australian and other coals, in the light of XRD data from the same coals’ LTA residues. The results were compared against quantitative mineralogical evaluations obtained from other techniques, such as XRD of the LTA using SIROQUANT (Ward and Taylor, 1996) and analysis of mineral matter in coal polished sections using the QEM*SEM technique (Creelman and Ward, 1996). Each method has its advantages and limitations, but relatively good agreement was achieved between SEDNORM and the other techniques for the main mineral groups involved. 3.2.3. Application to Low-rank Coals. Plasma ashing of lower-rank coals typically results in the formation of abundant calcium, ammonium and other sulfates, due to interaction of the principal non-mineral inorganics with the coal’s organic sulfur during the low-temperature oxidation process. The non-mineral inorganics that typically dominate

the mineral matter of such coals may therefore need to be removed and investigated separately by a process of selective leaching (Miller and Given, 1979; Benson and Holm, 1985; Ward, 1991; 1992) before low-temperature ashing can take place. One such procedure (Ward, 1992) involves: • Soaking the coal in distilled water, with analysis of the extract to identify the material dissolved, or at least potentially soluble, in the coal’s pore water; • Treatment of the water-washed coal with a concentrated ammonium acetate solution, with analysis of the extract to identify water-insoluble but exchangeable elements, such as those associated with carboxylate groups;

• Leaching of the -exchanged coal with hydrochloric acid, followed by analysis of the extract to identify the inorganics occurring in any acid-soluble organometallic complexes and in any otherwise-insoluble carbonate minerals. The acid-treated coal can then be subjected to oxygen-plasma ashing, to remove the organic matter (which has in the exchangeable-ion positions to prevent formation of mineral artifacts by organic sulfur fixation) and isolate the acid-insoluble mineral components.

3.3. Ancillary Techniques to Evaluate Coal Mineral Matter Ancillary techniques provide information that although strongly linked to mineralogical data does not necessarily in itself represent mineralogical data. Oxide determinations, the old standby, can be produced not only by conventional chemistry but also by newer and more rapid techniques such as inductively-coupled plasma (ICP) and X-

30

C. R. Ward

ray fluorescence (XRF) spectrometry. Chemical analyses are useful but they are in themselves imperfect as predictors. Elemental oxides are combined in minerals, singularly or with other oxides. It is the mineral that has the specific physical and chemical properties, not the component chemical part. Infra-red (IR) and Fourier Transform Infra-red (FTIR) spectrometry are related techniques that yield specific information on crystal structure and hence on mineral matter transformations that occur as minerals are heated. Application to mineralogy has been successfully attempted for a wide range of processes (Fredericks et al., 1985; Nyquist et al., 1990). FTIR has the potential to be an on-line method to monitor the development of certain glass/crystalline phases as they form in the furnace. Nuclear magnetic resonance (NMR) spectroscopy is also a specific technique for investigating mineral structure, able to quantify minerals with specific structures involving Al and Si (Phong-anant et al., 1991). This method is similar in concept and application to FTIR.

4. CONCLUSIONS The nature and distribution of the mineral matter have a fundamental effect on the behaviour of coals when used for various purposes. Quartz particles have long been recognised as a cause of erosion in grinding mills and on exposed furnace surfaces, and sulphur in various forms can give rise to furnace corrosion and stack-gas pollution. Mineral groups such as the iron sulfides and siderite or calcite, as well as inorganically-associated calcium, also appear to be major contributors to slagging in furnace operation. Interaction of sulphur trioxide, water vapor and coal ash can produce fouling deposits on the lower-temperature convective parts of a furnace system. Phosphate and sulphur-bearing minerals can also transfer P and S to coke when coal is heated in coke ovens. These elements are then transferred to the iron when the coke or when S and P-bearing pulverized coal is injected with iron ore into a blast furnace, producing a need for costly removal at a later stage in the metallurgical process. Evaluation of minerals and other inorganics in coal, based on techniques such as oxygen-plasma ashing and X-ray diffraction, provides a range of fundamental data essential to modern coal characterization, extending the information from traditional sources such as chemical ash analysis. Given the growing sophistication of mineral analysis technology, the relevant data can also increasingly be provided in a quantitative form. The behaviour of minerals and other inorganics during combustion depends to a significant extent on how the various components occur within the coal, the form in which they exist at combustion temperatures, and the opportunities that they have to interact with each other in the course of the combustion operation. Calcium occurring in pore water or in an organic association, for example, is typically released as independent Ca ions or atoms on combustion. These are relatively free to react with other ash-forming components. Calcium occurring as the carbonate mineral calcite, on the other hand, forms discrete CaO particles that are different aerodynamically, less reactive and therefore less likely to be associated with slagging or similar problems. Use of mineralogical analysis opens up a number of paths to engineers and others concerned with coal combustion, providing a better basis for assessment of potential problems with materials handling (through the presence of smectite and similar clay minerals), abrasion (due to liberated quartz particles), slagging, fouling, sulphur capture and ash handling characteristics. It is also relevant to the applicability of on-line quality

Mineral Characterization for Combustion

31

assessment systems. Understanding of the distribution of mineral matter in coal at the source (i.e. within the deposit being mined) can sometimes be used for better control of coal product quality through mining and preparation, and hence provide a basis for optimising the total resource utilisation process.

5. ACKNOWLEDGEMENTS The author would like to thank the Engineering Foundation, and Bob Creelman

as Session Convenor, for assistance in the presentation of this paper.

6. REFERENCES Benson, S.A. and Holm, P.L. (1985). ”Composition of Inorganic Constituents in Three Low-rank Coals”. Industrial and Engineering Chemistry, Product Research and Development, 24, 145–149. Bustin, R.M., Mastalerz, M. and Wilks, K.R. (1993). “Direct Determination of Carbon, Oxygen and Nitrogen Content in Coal using the Electron Microprobe”. Fuel, 72, 181–185. Cohen, D.R. and Ward, C.R. (1991). “SEDNORM—a Program to Calculate a Normative Mineralogy for

Sedimentary Rocks based on Chemical Analyses”. Computers and Geosciences, 17(9), 1235–1253. Corcoran, J.F. (1989). “The Use of SEM/EDS Techniques in Studies of Coal and Associated Mineral Matter”. Proceedings of Mineralogy-Petrology Symposium, MINPET 89, Australasian Institute of Mining and Metallurgy, Sydney, 85–87. Creelman, R.A., Agron-Olshina, N. and Gottlieb, P. (1993). “The Characterization of Coal and the Products of Coal Combustion using QEM*SEM”. Final Report, Project 1467, National Energy Research, Development and Demonstration Program, Australian Department of Primary Industries and Energy, Canberra. Creelman, R.A. and Ward, C.R. (1996). “A Scanning Electron Microscope Method for Automated, Quantitative Analysis of Mineral Matter in Coal”. International Journal of Coal Geology, 30, 249–269. Fredericks, P.M., Osborn, P.R. and Swinkels, D.A.J. (1985). “Rapid Characterization of Iron Ore by Fourier Transform Infra-red Spectrometry”. Analytical Chemistry, 57, 1947–1950. Gluskoter, H.J. (1965). “Electronic Low Temperature Ashing of Bituminous Coal”. Fuel, 44, 285–291. Gluskoter, H.J. (1967). “Clay Minerals in Illinois Coal”. Journal of Sedimentary Petrology, 37(1),

205–214. King, J.G., Maries, M.B. and Crossley, H.E. (1936). “Formulas for the Calculation of Coal Analyses to a Basis of Coal Substance free from Mineral Matter”, Journal of the Society of Chemical Industry, 55, 277–281. Kiss, L.T. (1982). “Chemistry of Victorian Brown Coals”. Australian Coal Geology, 4(1) , 153–168. Mandile, A.J. and Hutton, A.C. (1995). “Quantitative X-ray Diffraction Analysis of Mineral and Organic Phases in Organic-rich Rocks”. International Journal of Coal Geology, 28, 51–69. Miller, P.N. and Given, PH. (1978). “A Geochemical Study of the Inorganic Constituents of some Low-rank Coals”. Report, Contract EX-76-C-01-2494, U.S. Department of Energy, Coal Research Section, Pennsylvania State University (unpublished): 314 pp. Nyquist, R.A., Leugers, M.A., McKelvy, M.L., P-yen-uss, R.R., Putzig, C.L. and Yurga, L. (1990). “Infra-red Spectrometry”, Analytical Chemistry, 62, 223R–255R. Phong-anant, D., Pang, L.S.K., Vassalo, A.M. and Wilson, M.A. (1991). “Mineral Matter in Coal—the Characterization, Transformation and Effects on Boiler Deposit Formation and Boiler Erosion”. Final Report, Project 1227, National Energy Research, Development and Demonstration Program, Australian Department of Primary Industries and Energy, Canberra. Radmacher, W. and Mohrhauer, P. (1955). “The Direct Determination of the Mineral Matter Content of Coal”. Brennstoff-Chemie, 36, 236. Rao, C.P. and Gluskoter, H.J. (1973). “Occurrence and Distribution of Minerals in Illinois Coals”. Illinois State Geological Survey, Circular 476, 56 pp. Reed, S.J.B. (1996). “Electron Microprobe Analysis and Scanning Electron Microscopy in Geology”. Cambridge University Press, Cambridge, 201 pp. Renton, J.J. (1986). “Semiquantitative Determination of Coal Minerals by X-ray Diffractometry”. In: Vorres, K.S. (ed), “Mineral Matter and Ash in Coal”. American Chemical Society Symposium Series, 301.

32

C. R. Ward

Ramsden, A.R. and Shibaoka, M. (1982). “Characterization and Analysis of Individual Fly-ash Particles from Coal-fired Power Stations by a Combination of Optical Microscopy, Electron Microscopy and Quantitative Electron Analysis”. Atmospheric Environment, 16(9), 2191–2206.

Rietveld, H.M. (1969). “A Profile Refinement Method for Nuclear and Magnetic Structures”. Journal of Applied Crystallography, 2, 65–71. Stach, E., Mackowsky, M.Th., Teichmuller, M., Taylor, G.H., Chandra, D. and Teichmuller, R. (1982). “Stach’s Textbook of Coal Petrology”. Gebruder Borntrager, Berlin, 535 pp. Standards Australia (1992). “Higher Rank Coal—Mineral Matter and Water of Constitution”. Australian Standard 1038.22, 16 pp.

Stanton, R.W. and Finkelman, R.B. (1979). “Petrographic Analysis of Bituminous Coal: Optical and SEM Identification of Constituents”. Scanning Electron Microscopy, 1, 465–471.

Taylor, J.C. (1991). “Computer Programs for Standardless Quantitative Analysis of Minerals using the Full Powder Diffraction Profile”. Powder Diffraction, 6, 2–9. Ward, C.R. (1974). “Isolation of Mineral Matter from Australian Bituminous Coal using Hydrogen Peroxide”,

Fuel, 53, 220–221. Ward, C.R. (1977). “Mineral Matter in the Harrisburg-Springfield (No. 5) Coal Member of the Carbondale Formation, Illinois Basin”. Illinois State Geological Survey, Circular 498, 35 pp. Ward, C.R. (1978). “Mineral Matter in Australian Bituminous Coals”. Australasian Institute of Mining and Metallurgy Proceedings, 267, 7–25.

Ward, C.R. (1991). “Mineral Matter in Low-rank Coals and Associated Strata of the Mae Moh Basin, Northern Thailand”. International Journal of Coal Geology, 17, 69–93. Ward, C.R. (1992). “Mineral Matter in Triassic and Tertiary Low-rank Coals from South Australia”. International Journal of Coal Geology, 20, 185–208.

Ward, C.R. and Gurba, L.W. (1997). “Use of the Electron Microprobe in Chemical Analysis of Coal Macerals, with Special Reference to the Direct Determination of Organic Sulphur”. In: Boyd, R.L. and Allen, K. (eds), Proceedings of the 31st Symposium, Advances in the Study of the Sydney Basin, Department of

Geology, University of Newcastle, New South Wales, 115–122. Ward, C.R. and Taylor, J.C. (1996). “Quantitative Mineralogical Analysis of Coals from the Callide Basin, Queensland, Australia using X-ray Diffractometry and Normative Interpretation”. International Journal of Coal Geology, 30, 211–229.

THE DEVELOPMENT OF POWER TECHNOLOGIES FOR LOW-GRADE COAL

K. Basu Bharat Heavy Electricals Limited Corporate Research & Development Division

Vikas Nagar, Hyderabad-500093. INDIA

1. INTRODUCTION The need for Clean Coal Technology, CCT, is well understood for the sustainable development and growth of a country. It can be defined as “technology designed to enhance both the efficiency and environmental acceptability of coal extraction, preparation and use”. Various clean coal technologies are available and under development to enhance the acceptability of coal as a clean fuel. For conventional technology, pre and post combustion cleaning are generally retrofit systems to meet the pollution regulation. The new technologies promise to increase efficiency and to reduce pollutant generation, beyond the process capability of conventional power plants. In spite of Human Development Index, HDI, proposed by UNDP to measure the wellbeing of a society, supply of clean commercial electricity to the consumer at an affordable price still remains the vital link to socio-economic development of a country. It is particularly relevant for India, since its per capita electricity consumption is still about l/30th fraction of developed countries. The country’s development is further hampered due to chronic shortage of electricity supply. A GDP loss of Rs. 180 billion was attributed to shortage of electricity during 1996–97. Coal has been India’s mainstay fuel for power generation. At present, 65 per cent of total electricity of the country is generated by using non-coking coal in sub- critical

pulverized fuel (PF) fired boilers. The coal supplied to power stations contains about 40 per cent ash. Highly abrasive nature of ash and non-standard quality of coal supplied to power plants reduces thermal efficiency and increases forced outage. Average plant load factor (PLF) and operating availability of Indian plants are 64 per cent and 80.7 per cent respectively. To achieve projected power growth rate, improvement in performance of existing power plants and installation of 56,700 MW new capacity is proposed. Repowering and Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

33

34

K. Basu

refurbishing of old power plants and supply of beneficiated coal to distant power stations, are expected to improve their performance and PLF. Improvement of one per cent PLF today is equivalent to addition of 750 MW new capacity. Relaxation for import of coal also allows performance improvement by switching to or blending with low ash and high heat value imported coal. However, domestic fund constraint and limitation in infrastructural facilities for port handling and inland transportation can restrict bulk import of coal. India’s ambitious plan for addition of new capacity offers a large potential market for both conventional and emerging clean coal technologies. Other than domestic fund constraint, technical reliability and economic viability of emerging technologies are presently the major concerns for commercial investment. Consequently, proven sub critical PF boilers will continue to have the major share of electricity generation in the near future. India has already imposed a strict emission regulation for power plants. However, less efficient sub-critical technology cannot meet the expected emission lelvel of and other pollutants in the distant future. Also, at present, Indian power plants are not equipped with FGD and threfore, cannot use high sulphur Indian or imported coal. To address the future market need, BHARAT HEAVY ELECTRICALS LIMITED, BHEL, the major engineering and supplier of power plant equipment in India, has been

assessing and developing clean coal technologies for Indian coal.

2. PRESENT OPTION As cited earlier, majority of India’s new power plants will be based on conventional technologies and will use Indian coal. Due to presence of about 85 per cent silica and alumina in ash and medium volatile content in coal, the major challenges for this technology are erosion of, coal and ash paths and combustion stability of boiler. Oil support is required even at as high as 50 per cent of the design load. Beneficiation of Indian coal and use of low sulphur and high heat value coal offer best possibility for efficiency

improvement and emission reduction for the conventional technology.

2.1. Coal Other than poor characteristics of Indian coal, present mining practice further reduces the quality of coal. The total coal reserve of India is estimated to be 200 billion tonnes out of which 68.6 billion tonnes are proven reserves. The reserves are distributed in central and eastern sectors of the country. Some lignite deposits are also found in the southern sector. There are 10 coal bearing river basins with 58 coal and lignite fields consisting of about 450 coal seams and 60 per cent of coal deposits are within the depth of

300 meters. This shallow deposits encourage surface mining of coal. Indian coal is said to have 25 per cent inherent ash content similar to other Gondwanan coal. However, the quality is progressively deteriorating over the years due to emphasis on open cast mining. The share of coal produced from open cast mines

increased from 26 per cent in 1975 to 73 per cent in 1993. The non coking coal for power sector is graded based on their useful heating value, UHV between 26.0 MJ to 10.0 MJ per Kg. into seven categories, “A” to “G”. “E” “F” and “G” are graded as inferior quality which have UHV between 14.0 MJ to 10.0 MJ per Kg. While no inferior grade was produced in 1947, its present share is 70 per cent and

The Development of Power Technologies for Low-Grade Coal

35

it is expected to reach 78 per cent by the turn of the century under “Business as usual” scenario.

2.2. Beneficiation At present, about 33 MMT of coking and semi coking coal is beneficiated for industrial use. So far, beneficiation of non coking coal has not been taken up for regular use in power plants. High proportion of near gravity material in coal makes the separation difficult and expensive. Separation is further aggravated by the dispersion of minerals throughout the coal matrix. Hard and abrasive nature of alumina and silica which constitute more than 85 per cent of ash particles, increase the capital and running cost of the plant. The experts claim that to reduce ash from 45 per cent to 19.7 per cent, the generation of middlings and tailings will be around 85 per cent with high carbon content. Further use of these rejects in fluidized bed boiler will be necessary to fully utilise the fuel content of coal. To evaluate economic level of cleaning of mineral, number of technologies had been evaluated in India. Also, tests had been conducted with beneficiated coal to evaluate improvement in performance of a power plant. Contra Flow Separation technololgy could reduce 8 to 10 per cent ash from coal. The cost of a 100 TPD plant was estimated to be between Rs.5 million to Rs.55 million. The ROM Jig technology is expected to reduce ash from 45 to 29 per cent in deshaled portion and 31.6 per cent in total mix. The yield is expected to be 67 per cent. Tests at the Satpura power station recorded an improvement of plant utilisation factor from 73

per cent to 93 per cent and energy generation from 3.71 MU to 4.83 MU per day with coal, beneficiated from 44 per cent to 31 per cent ash content. A study indicates that the reduction of ash from 45 per cent to 34 per cent will be economically viable against transportation cost of coal to power plants situated beyond 1,000 Km. A cost benefit analysis at a macro level including all socio economic cost of each element can resolve the debate ensuing in India on coal beneficiation. However, the Ministry of Environment & Forests has given a directive to supply coal with less than 30 per cent ash to all power plants beyond 1,000 KM. from fuel source. At present, 13 per cent of total power coal is transported to power plants beyond 1,000 Km. and its share is expected to go up to 18.4 per cent by 2002.

2.3. Switching and Blending Since slagging and fouling is a combined effect of boiler design and coal property, some operators are concerned to use imported low ash, high heat value coal in existing boilers. High heat flux in burner area and slagging and fouling of boilers are their main

concerns even with characteristically low slagging coal.

3. EMISSION The major pollutants from Indian power plants are the generation of particulates and emission of Emission of is generically similar to other power plants. Low incidence of sulphur in coal reduces the burden of post combustion cleaning of gaseous

sulphur emission.

where required is cleaned using Ammonia.

36

K. Basu

Under “Business as usual scenario”, we will be generating 128 MMT of ash, 4.3 MMT of and 570MMT of to meet the energy demand of 475 billion Units in 2002 from coal based power plants. In 1996–97, the respective pollutant generations were 85 MMT of ash, 2.8 MMT of and 380 MMT of based on average specific coal consumption of 0.75 Kg/KWH. Hypothetically, if we use only imported coal with ash content of 15 per cent, sulphur content of 3 per cent and also improve specific coal consumption from 0.75 Kg/KWH to 0.35kg/KWH, then we will be generating 20MMT of ash and 10MMT of in 2002. This will require installation of FGD plants at an additional cost of about 25 per cent of generation cost. India’s total emission to global environment is about 6 per cent with over 16 per cent population. Consequently, our per capita emission is insignificant as compared to the developed countries. Also, the role of emission for global warming and its subsequent fallout as predicted by Inter Governmental Panel on Climate Change has been debated by World Energy Council and by other agencies. However, other than emission, the strongest compulsion for India to improve the efficiency of power plants is to conserve its precious fuel.

4. NEW TECHNOLOGIES BHEL had to modify design and constructional features of technology acquired for boiler to achieve desired performance with Indian coal. During 1970s, the major failures experienced were fuel pipes, burners and above all, boiler tubes. At one time, more than 56 per cent of plant forced outage was attributed to boiler tube failures. To address these problems, design changes were incorporated by introducing wear resistance material in critical paths, optimizing flue gas velocity and redesigning the surface to minimize erosion. Subsequently, the tube failure has been reduced to between 3–5% of total forced outage. Also, India’s experience with entrained bed gasifiers imported for fertilizer sector was below expectations. Thus, the assessment of new technologies for compatibility with Indian coal became a prerequisite for commercial use. In mid 1970s, BHEL had started evaluation of emerging and future clean coal technologies and planned a comprehensive development

programme. It was decided that near term and most promising emerging technologies applicable to Indian scenario would be given priority for initial development. Accordingly, Atmospheric Fluidized-bed Boiler (AFB) and Integrated Gsification Combined Cycle (IGCC) with relevant gasification processes were given priority. It was also decided that other upcoming technologies would be developed atleast upto the pilot plant level to generate experimental data with Indian coal. Under this category, Pressurized Fluidized Bed Combustion (PFBC) and Hot Gas Clean-up System (HGCS) were identified as next priority and Magneto Hydro-dynamic (MHD), Slagging Combustor and Fuel Cell for long term development.

4.1. AFB During development stage in late 1970s, three prototypes, two 2 MT/H, followed by one 10 MT/H, were constructed to evaluate wide range of Indian coals and to establish

The Development of Power Technologies for Low-Grade Coal

37

scale up methodology. The larger prototype had octagonal, water cooled and compartmentalized furnace connected to a square shaped freeboard section. It was a balance draft design, with finned tube economiser and spiral type bed coil for steam generation. Over the years, other than coal and lignite, coal water slurry, washery rejects and bio mass

fuels were tested in AFB. More than 30 boilers have been supplied for various ratings

ranging from 12 to 100 MT/H steam supply.

38

K. Basu

4.2. IGCC Priority was given to IGCC technology demonstraion for its potentiality to achieve process inherent cleanliness over other CCT, Fig. 1. and high thermal efficiency, Fig. 2. The selection and development of a gasification process in BHEL was an integral

and complementary part of the development of IGCC technology. It consisted of two phase gasifier development for integration to a 6.2 MW IGCC test plant. Techno-economic studies concluded that for Indian coal with 40 per cent ash content and for an ash melting temperrature of approx. 1,673K, use of oxygen-blown slagging or agglomerating type gasification process would not be economical. That left the choice between air blown moving bed or fludized bed gasifier with dry ash extraction system. The moving bed gasifier was the first choice in spite of its inherent process inadequacy and cost penalty for this application. The selection was primarily due to tech-

The Development of Power Technologies for Low-Grade Coal

39

nological maturity of this process and backup experience India had in operating and testing of Lurgi type gasifiers at Central Fuel Research Institute, CFRI, Dhanbad, and at Indian Institute of Chemical Technology, IICT, Hyderabad. Air-blown pressurised fluidized bed gasifier was identified as the most techno-economically attractive process for Indian coal and lignite. The first phase of IGCC programme consisted of concurrent development of a 150 TPD air blown moving bed gasifier integrated to a power block with 4 MW gas and 2.2 MW steam turbines, Fig. 3. Wet gas cleaning was adopted due to its technology readiness. The plant commissioned in the year 1987, was tested extensively for its operational capability and

40

K. Basu

performance with Indian coal. Under a joint programme sponsored by Department of Coal, Government of India, North Karanpura coal with 40 per cent of ash content was also successfully gasified in BHEL’s 150 MTPD and in IICT’s 24 MTPD gasifiers to generate data for scale-up. A 18 MT per day PFBG pilot plant was developed in 1986 and tested over 2,700 hours. It is a two diameter reactor with dry ash removal system. The process operates in “bubbling bed” region at a temperature of 1,230K to 1,313 K. The same configuration was scaled up to a 168 MTPD gasifier and retrofitted in the existing 6.2 MW IGCC plant in 1996, Fig. 4. At present, 168 MTPD PFBG is undergoing performance testing. BHEL’s achievement in gasification and IGCC programme is summarized in Table 1. To study the performance and carbon conversion capability of fluidized bed reactor

operating in a different fluidization region, a 100 mm. ash agglomerating gasifier, Fig. 5, and a 200mm gasifier coupled to a 100 mm. fluidized bed combustor, Fig. 6, were developed. The first phase of performance testing of both the gasifiers has been completed. Partial agglomeration of sub bituminous coal with enriched air was achieved with unburnt carbon between 0.5–1.5 per cent in discharged ash from the bottom. The coupled gasifier was tested for wide range of input parameters for optimization. Carbon in discharged ash from the combustor could be reduced to 0.5 per cent.

The Development of Power Technologies for Low-Grade Coal

41

4.3. PFBC & HGCS The other promising clean coal technology identified for near term application is combined cycle based on PFBC. The preliminary development was first conducted in

a 200 mm pseudo adiabatic reactor. A 440 mm. test facility was then established and integrated with a Circulating Bed Granular Filter (CBGF), Fig. 7. Over the years, various coals with ash content upto 52 per cent, lignite and agro based fuels have been tested to generate operational and performance parameters. The CBGF concept was selected for development since it offers the possibility of simultaneous removal of particulate and

4.4. MHD A 5 MW (thermal) MHD pilot plant was set up under Indo-USSR Science & Technology programme. The plant has a MHD channel with 2 Tesla iron core magnet

42

K. Basu

and seed recovery system. Tests were carried out with preheated Blue Water Gas and LPG. Out of 15 major trials conducted, 10 were for power generation.

4.5. Slagging Combustor A 3 MW (thermal) Slagging Combustor was set up in BHEL in order to evaluate this technology for Indian coal. Possible application perceived was MHD and direct firing

of coal in gas turbine. However, the flow of slag and heat reecovery from it were not satisfactory mainly due to high melting temperature of ash.

4.6. Fuel Cell Subsequent to development of 1 KW Phosphoric Acid Fuel Cell module and

scaling it up to 5 KW stack, BHEL is presently developing a 50 KW fuel cell. Under

The Development of Power Technologies for Low-Grade Coal

43

India’s Science & Technology Programme, demonstration of a 200 KW PAFC has also been taken up. Recently development of solid oxide fuel cell has been started for its potential use with coal gasifier.

5. FUTURE PROGRAMME Over the years, BHEL has established test facilities to evaluate wide ranging technologies for efficient utilization of Indian coal. These facilities can also be used for testing other types of coal.

44

K. Basu

With the experience gained from IGCC development, BHEL has proposed to set up an IGCC demonstration plant either for a capacity of 60 MW or 120 MW in association with CSIR under the Clean Coal Technology Mission. Site specific feasibil-

ity report has been prepared and submitted to Government of India . Another proposal to retrofit an existing steam plant with PFBC topping cycle has also been submitted to Government for its consideration.

6. CONCLUSION Beneficiation of Indian coal and operation of power plants with imported coal will improve the efficiency of power generation to some extent but they will not satisfy overall future requirements of pollution control and conservation of energy. Therefore, there is a need to adopt new clean coal technologies. BHEL has commercialized “bubbling” AFB based on in-house development and Circulating Fluidized-bed Boiler through technology acquisition.

Extensive theoretical and experimental data on compatibility of various clean coal technologies with Indian coal have been generated from the pilot scale development of PFBC, CBGF, MHD and Slagging Combustor. These data can be used for technology assessment in future. BHEL’s experience in designing and operating IGCC plant with both moving-bed and fluidized bed gasifiers spans over a decade. IGCC appears to be the most suitable emerging technology for Indian high ash low grade coal.

LOW-RANK COAL AND ADVANCED TECHNOLOGIES FOR POWER GENERATION Dong-ke Zhang’, Peter J. Jackson2, and Hari B. Vuthaluru 1,2 1

Department of Chemical Engineering The University of Adelaide Adelaide, SA 5005, Australia 2 CRC for New Technologies for Power Generation from Low-rank Coal Unit 8, 677 Springvale Road Mulgrave, Victoria 3170, Australia

1. INTRODUCTION

Substantial deposits of low-rank coals exist in southern Australia and a number of other countries including Indonesia, Thailand, Turkey, China, USA and Germany. Properties that mitigate against their widespread use for power generation are high moisture content, low ash-melting temperature and often highly fouling nature of the ash. South Australia and Victoria have large reserves of readily accessible low-rank coals. Victorian deposits have been used for many years for power generation and generally have high water contents (60–70%), but are low in ash (l–3%db), and low in total sulphur (<0.5% db) [Durie, 1991]. Sodium in ash is usually low in the coals currently utilised in Victoria (generally <5%). While the South Australian coal (Leigh Creek) which is presently utilised for power generation has a moisture content of <38%, other deposits have higher moisture (50–61%), higher ash contents (generally <20%db), and higher total sulphur (gen-

erally <5%db). Sodium in ash in the S.A. coals in particular is often relatively high (3 to 16%). The relatively high sodium and sulphur contents have mitigated against their use in conventional pulverised fuel combustion systems because of the fouling and hightemperature corrosion problems associated with the low-melting-point ash. At the present time, one-third of the electricity generated from coal in Australia uses low-rank coal as the primary energy source [Brockway and Zhang, 1997]. The Cooperative Research Centre for New Technologies for Power Generation from Low-rank Coal (CRC–Power Generation) was formally established in June 1993 to conduct research and technology development into those new generation technologies and processes that have the best prospects of overcoming the principal challenges (cost and environmental impact) facing the use of low-rank coal as a competitive energy source for electricity Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

45

46

D.-k. Zhang et al.

generation. New generation technologies are the key to the objectives of reducing both costs and emissions to the environment. Clearly, development of any new technologies (or new technology components) for efficient and effective utilisation of low-rank coals for power generation must address two major problems inherently associated with low-rank coal, viz, drying of the large amount of moisture and the solving of the ash-related problems. The present contribution focuses on the role of mineral matter in low-rank coals in the selection or development of new technologies for power generation from these coals. Recent developments in research on ash behaviour of low-rank coals in various reaction environments are reviewed. Several

technologies and processes (particularly those based on fluidised-bed technologies) that could potentially alleviate the ash-related problems associated with the use of low-rank coals are proposed and areas where further research and development are necessary are identified.

2. COAL MINERAL MATTER AND ADVANCED POWER TECHNOLOGIES

2.1. Background The role of mineral matter in coal during high temperature processing is governed by its nature and occurrence in the coal matrix, reaction temperature and reacting environment (oxidising or reducing). In conventional p.f. fired power plants, where peak flame temperatures in the neighbourhood of 1500°C are not uncommon, the mineral matter in the coal is directly associated with problems of slagging, fouling and ash deposition

as well as high temperature corrosion [Raask, 1985]. At these high temperatures, alkali and alkaline-earth metal constituents (eg. Na, Ca, etc.) vaporise into the gas [Raask, 1985, Wall, 1992]. When the gas containing inorganic vapour (referred to as fume) contacts relatively cool heat transfer surfaces (about 400 °C), a series of physical and chemical processes occur leading to the formation of an initial ash layer composed chiefly of alkali sulphates. The condensed sulphates form a “glue” to capture ash particles that are conveyed to the vicinity of the heat transfer tubes by the aerodynamics of the gas in the furnace and around the tubes. The particles will adhere to the tube or the deposit surface if either the particles or the surface are “sticky” enough to overcome the particle kinetic energy. As the ash deposits grow, a molten layer will eventually form as a result of reduced heat transfer flux and thus increased deposit surface temperature. When the proportion of liquid in the solid/liquid matrix is high enough to cause the matrix to collapse and the deposit surface becomes fully molten, and if the viscosity of the melt is low enough, the deposit will stop growing any further due to the liquid running off the surface and a dynamic steady-state of the deposit layer establishes. Sandwiched between the initial and

molten layers is a so-called sintered layer which is held together by partial melting of the arriving particles during the build-up of the deposit. This layer becomes more sintered as the deposit grows and its temperature increases. The sintering of the deposit depends

on the state of melting of the incoming ash particles and the proportion of molten phase in the sinter. The formation of the adhesive initial layer and subsequent growth of ash deposit can also be influenced to a certain extent by the local gas environment which determines the thermodynamic equilibrium of the alkali vapour and the chemical processes during the fume formation. Figure 1 illustrates the processes involved in formation of ash deposits on heat transfer tubes in conventional p.f. fired furnaces. These deposits reduce the heat transfer flux through the tubes and thus limit the steam tem-

Low-Rank Coal and Advanced Technologies for Power Generation

47

perature. Periodic shutdowns are therefore required for tube cleaning to restore boiler performance. High temperature corrosion of the steam super-heater tubes in particular, leading to tube failure, due to attack by the deposits is also a major problem with high alkali content low-rank coals. Based on the above brief overview, it is evident that three principal factors, viz, coal ash composition, process temperature and processing environment, govern the formation of ash deposits on the heat transfer surfaces and cause the operational problems in the p.f. fired power plants. For low-rank coals, the usually high levels of inorganic constituents are inherent problems in high temperature p.f. firing systems. Fluidised-beds provide an isothermal operating environment, are conducive to vigorous heat and mass transfer and thus enhanced chemical reaction rates, and are easy to control. Also,

fluidised-beds operating at much lower temperatures (about. 700–900 °C) offer better prospects in reducing ash-related problems associated with the utilisation of low-rank coals. Thus fluidised-bed combustion and/or gasification can be the bases of new technologies for power generation from low-rank coal to achieve a high overall process efficiency.

2.2. Fluidised-Bed Based Advanced Technologies Several emerging advanced power generation technologies that have better prospects for low-rank coals are at various stages of development. These include Integrated Gasification Combined Cycle (IGCC) incorporating a fluidised-bed gasifier and steam fluidised-bed drying (SFBD), pressurised fluidised-bed combustion (PFBC) combined cycle, pressurised fluidised-bed combustion combined with partial gasification (Advanced PFBC), and circulating fluidised-bed combustion (CFBC). McIntosh [1997] recently reported a large number of case studies based on these four principal advanced technologies using a thermodynamic analysis to evaluate various advanced cycles for power generation using low-rank coals, with an attention to drying of the coals. Figure 2 compares the efficiencies (power sent out based on HHV) for a typical low-rank coal and a typical bituminous coal when different technologies, including the conventional p.f.

48

D.-k. Zhang et al.

fired cycle, are used. It can be seen that significant improvements in the power generation efficiency for both low-rank and high-rank coals can be achieved by using different advanced power cycle technologies, in ascending order as follows conventional p.f. firing systems < CFBC < PFBC < IGCC < A-PFBC

It is worth noting that as a more advanced technology is applied, the difference in efficiencies between the low-rank and high-rank coals decreases. For example, for the conventional p.f. firing system, the efficiency is about 29% for the low-rank coal while 37% for the high-rank coal. For the IGCC process, the efficiencies are about 40 and 42%, respectively. However, with the A-PFBC technology, the low-rank coal efficiency can be

as high as 44%. This is very significant, because fuel costs for low-rank coals are generally substantially lower than higher rank coals. The advanced technologies can therefore place low-rank coals in a much more competitive position in the power generation market. Fluidised-beds are a key to the above advanced technologies. However, low-rank coals do not present trouble-free operation of fluidised-beds. Growth in the size of bed material (silica sand particles) due to ash deposition during fluidised-bed combustion has been reported [Williams, 1984; Manzoori, 1990] while burning Morwell (Victoria) and Lochiel (S.A.) coals, respectively. Bed material agglomeration and bed defluidisation leading to unexpected plant shut-down may also occur due to undesired ash behaviour in fluidised-bed combustion or gasification. Corrosion and erosion of heat transfer tubes immersed in a fluidised-bed burning Victorian low-rank coals have also been reported in the abovementioned literature. These problems are again inherently related to behaviour and transformation of the inorganic matter present in the low-rank coals. Although the behaviour of the inorganic matter in the p.f. firing systems has received considerable research attention, the knowledge is not directly transferable to fluidised-bed systems due to significant differences in operating conditions between the two systems. In order to realise the full potential of fluidised-bed based advanced tech-

Low-Rank Coal and Advanced Technologies for Power Generation

49

nologies for power generation from low-rank coal, there is a clear need for research leading to the understanding of inorganic matter transformations in fluidised-bed combustion and gasification environments.

3. FLUIDISED-BED AGGLOMERATION AND DEFLUIDISATION 3.1. General Behaviour of Ash in Fluidised-Bed Combustion and Gasification One of the key parameters with which fluidised-bed operation is generally characterised is the minimum fluidisation gas velocity below which the bed is not fluidised

and appears as a fixed-bed.

depends on a number of parameters including bed

particle size and density, fluidising gas viscosity and therefore the bed (operating) temperature and configuration of the bed. For a given fluidised-bed, is only slightly dependent on and is expected to decrease with increasing due to increased gas viscosity at higher temperatures. The superficial gas velocity defined as the ratio of the volumetric flow rate of fluidising gas to the cross-sectional area of the bed, is usually about 3 to 5 times of for a bubbling fluidised-bed and is up to 20 times for a circulating fluidised-bed. However, during fluidised-bed combustion and gasification of coal, the bed particles can grow in size due to deposition of coal ash on the particle

surface. With a low-rank coal with a low melting-point ash, the bed particles can become sticky so that particles may adhere together to form agglomerates, the latter phenome-

non is referred to as agglomeration. The increase in particle sizes due to ash deposition and agglomeration is then translated into an increase in the minimum fluidisation gas

velocity. As a result, the superficial gas velocity has to be increased in order to maintain the fluidisation state in the bed. If is not increased accordingly, or in the extreme

case of agglomeration where the whole bed material agglomerates, fluidisation can no longer be maintained and the bed is said to be defluidised. The above processes are apparently temperature-dependent and may be illustrated by a schematic diagram as shown in Fig. 3. The temperature corresponding to the turning point on this diagram is termed as the initial sintering temperature of the bed material [Manzoori, 1990], above which the fluidised-bed has to operate with a much higher fluidising gas velocity, or if this demand can not be met, defluidisation will occur. The on-set of initial sintering of the bed material can occur at much lower temperatures when firing low-rank coals due to the deposition of low melting-point ash on the particle surface. Defluidisation is one major challenge faced by fluidised-bed combustion and gasification based technologies for power generation. Understanding of physical and chemical mechanisms of the

50

D.-k. Zhang et al.

particle agglomeration and bed defluidisation processes is a key to any solutions of these problems.

3.2. Mechanisms of Agglomeration In either fluidised-bed combustion or gasification of coal, ash constituents evolve as coal organic matter is consumed. If the temperature of the coal particles is high enough, a molten phase of ash can form on the coal particle surface. The molten ash is then transferred onto the inert bed medium (usually sand) particle surface by collision between burning char particles and bed particles. It is highly probable that the ash layer coating on the bed particle surface will remain molten at typical fluidised-bed operating temperatures, particularly for low-rank coals with a low melting-point ash. Bed particles thus become sticky over prolonged periods of operation and often, in order to achieve

desired carbon conversion levels, fluidised-beds are operated at temperatures above the melting point of ash. As a consequence, agglomerates will form if the kinetic energy of particles cannot overcome the so-called sticky force between particles [Manzoori, 1990]. Figure 4 schematically illustrates this process. 3.2.1. Fluidised-bed Combustion.

In fluidised-bed combustion, the coal inventory in the bed is usually less than 5%, with the bulk of the bed being an inert material, typically

sand. It is this bed material which can agglomerate under certain conditions. Studies by Goldberger [1967] at bench and pilot scales indicated that ash could be retained in bed agglomerates, thus eliminating fly ash carryover into the combustion flue gases. Their tests showed that the sand particles were coated with molten ash and formed agglomerates. In another study, Goblirsch et al. [1983] performed tests on high-sodium coal and suggested that sulphated ash was responsible for the formation of agglomerates at lower bed temperatures. More recently, Manzoori [1990], Vuthaluru and Zhang [1997a] and Vuthaluru et al. [1997] studied the combustion of high-sodium, high-sulphur South Australian low-rank coals in a spouted bed using silica sand bed material at temperatures ranging from 700 to 850 °C. The bed material resulting from these tests consisted of coated sand particles as well as agglomerates of two or more bed particles. Examination of the agglomerates by SEM indicated that the characteristics of the ash coating and the material at the interface of agglomerated bed particles were identical. This sug-

Low-Rank Coal and Advanced Technologies for Power Generation

51

gests that agglomeration occurs by random collision between the coated bed particles. Spot analyses and X-ray maps of polished cross-sections of ash coated sand grains showed that the ash coating is rich in sulphated compounds of Na and Ca which act as a binder. Figure 5 shows typical SEM images of ash-coated sand particles, agglomerated sand particles, and cross sections of ash-coated sand particles, resulting from combustion of Bowmans coal in a spouted fluidised-bed. Based on extensive investigations, Manzoori [1990] proposed the following criteria for bed material agglomeration: 1. operating temperatures must be above the initial sintering temperature of the coating; and 2. minimum “critical thickness” of the coating must be deposited on the bed particles. Studies conducted so far suggest that agglomeration of bed materials in fluidisedbed combustion is a consequence of stickiness of particles resulting from the presence of a molten phase which acts as a binding agent. The molten phase is usually rich in sulphates of Na and Ca. This is particularly true when South Australian low-rank coals are used. In fluidised-bed combustion, the burning char particle temperature

can be 100–300 °C higher than the bed temperature, Although may be less than the ASTM ash melting-point of the coal, the formation of a binding matrix of lowmelting eutectics and subsequent transfer of this molten phase onto bed particles are inevitable. Thus, agglomeration of bed particles is an inherent operational problem for fluidised-bed combustion of low-rank coals, particularly with coals with high alkali and sulphur contents. 3.2.2. Fluidised-bed Gasification. In contrast to fluidised-bed combustion, fluidised-bed gasifiers using steam and oxygen or air as the gasifying agents usually operate with a bed of coal char, without any bed material. Thus agglomeration of char and ash particles, rather than bed materials is of concern. Owing to the relatively low reaction rates in gasification environments, fluidised-bed gasifiers require relatively high bed temperatures. In addition, the high coal/char inventory in the fluidised-bed coal gasifier means a relatively high ash content in the bed as well. Thus the formation of a molten ash phase and consequent agglomeration of char and ash particles are potential problems. Studies by Yerushalmi [1976] suggested that ash agglomeration in fluid bed gasifiers consists of several steps which include:

52

D.-k. Zhang et al.

1. Formation of smaller molten beads on the surface of char particles. 2. Growth of the molten beads (both in size and number) due to generation of additional molten ash and possibly due to collision with other char particles. 3. Separation of larger beads from the char surface due to high interfacial contacts between the molten ash and the char, thus forming sticky agglomerates. 4. Coalescence of ash agglomerates with small beads or other agglomerates. The ash formation chemistry in fluidised-bed gasification is presently less understood. Formation of sulphides and chlorides with low-melting points is most likely in the

gasification environment. Godel [1966] suggested that char agglomeration may be avoided in gasification systems by operating the gasifier at high fluidising velocities. However, the high flowrates of fluidising gases presents an economic penalty because the gasification agents (steam and oxygen) would be less efficiently utilised.

3.3. Mechanisms of Defluidisation In fluidised-bed combustion and gasification, defluidisation is believed to result from particle stickiness. Some researchers [Langston and Stephens, 1960] postulate that defluidisation is the extreme case of agglomeration and is governed by the same princi-

ples as agglomeration. However, studies by Basu [1982] indicate defluidisation may not necessarily have the same mechanisms as agglomeration. Gransden et al. [1970] proposed

that defluidisation can take place either by the formation of agglomerates or without the formation of agglomerates. The latter case is thought to be caused by interaction of sticky bed particles without the formation of agglomerates which results in the increase in the minimum fluidisation velocity [Manzoori, 1990]. The mechanisms of defluidisation are not yet fully understood. In a more recent study of fluidised-bed combustion of three different low-rank coals in a spouted bed combustor, He and Zhang [1997] confirmed that two different types of defluidisation can occur. These are so-called partial and com-

plete defluidisation. As shown in Fig. 6, at 700°C, the pressure drop across the bed decreased gradually with time over a period of 8 hours without defluidisation. As pres-

sure drop is a measure of the mass of particles being fluidised, Figure 6 suggests that the fraction of particles fluidised decreased with time as some particles became defluidised. This is termed partial defluidisation. Comparison of the ash coating thickness between the fluidised and stationary particles revealed that the defluidised particles had less coating than those which remained fluidised. This was expected since the fluidised particles were continuously coated during combustion, whereas the contact of the defluidised particles with the burning coal particles was significantly reduced. Consequently, ash

transfer from the burning coal particles to these bed particles became minimal. However, this observation contradicts the trend reported in the literature that particles with a

thicker coating had a higher tendency to agglomerate and defluidise [Manzoori, 1990; Manzoori and Agarwal, 1994; and Arastoopour, et al., 1988]. However, at 850°C, the whole fluidised-bed collapsed (defluidised) within 45 minutes (Fig. 6). Visual examination showed that there was no apparent particle agglomeration. SEM analyses also did not identify any significant differences in the composition of the ash coatings on particle samples collected from the experiments at 700 and 850°C. However, bed temperature seemed to play a major role, with the time for total defluidisation rapidly decreasing with increasing bed operating temperature, as shown in Fig. 7. When the surface of ash-coated particles become sticky, it is likely to form agglom-

Low-Rank Coal and Advanced Technologies for Power Generation

53

erates, resulting in defluidisation. In the case of coal combustion with sand particles as the bed media, the thickness of the ash coating on the sand particles increases with time. When the coating thickness reaches a certain value, defined as the critical coating thickness, the bed will defluidise [Manzoori and Agarwal, 1994]. It has been observed that the critical coating thickness decreases with increasing bed temperature [Vuthaluru and Zhang, 1997a; Vuthaluru et al., 1997; and Manzoori, 1990].

3.4. Mineral Matter Transformations Agglomeration, sintering, deposition and defluidisation have all been observed in a number of fluidised-bed combustion and gasification units ranging from laboratory scale to full industrial scale. Analyses of various samples from different studies indicate that the initial layers of ash coating on the bed materials, in the case of agglomeration and defluidisation, and on the heat transfer tubes, in the case of ash deposition, consist of sodiumand potassium-rich calcium sulphates. As the temperature of burning char particles in a fluidised-bed can be several hundred degrees higher than the temperature of the bed material particles, sodium, potassium, and calcium are readily volatilised and will condense on the cooler surfaces. These condensed inorganic species can also react with sulphur compounds evolved from the coal to form substances that act as a glue in binding bed

54

D.-k. Zhang et al.

particles, leading to agglomeration and defluidisation. In addition, sodium can also react with silica sand to form sodium silicate which sinters at relatively low temperatures. The formation of low melting point eutectics between sodium- potassium-, calcium-, and sulphate-rich components within coal readily occur during fluidised-bed combustion processes [Benson et al., 1995]. These eutectics have significantly lower

melting-points than their respective pure compounds. The melting points of typical inorganic compounds and eutectics found in combustion and gasification of low-rank coals are given in Table 1. In a study of fluidised-bed pyrolysis of Bowmans coal, Telfer and Zhang (1997) observed that transformation of organically bound Na and Ca into their respective sulphides occurred within the coal particles. It is possible that the sulphides are further converted into sulphates and the sulphate matrix may be molten at typical operating temperatures for fluidised-bed combustion. The molten phase can be subsequently transferred onto bed material surfaces and serve to glue particles together in the form of

agglomerates or to form thicker deposits. In the extreme case, defluidisation occurs. In fluidised-bed gasification, the behaviour of the inorganic matter in low-rank coal

is substantially different to that in combustion systems. It is reported from laboratory scale studies that sodium compounds are likely to vaporise under gasification conditions, with the extent of vaporisation increasing with increasing gasification temperatures [Mojtahedi and Backman, 1989; Bellin et al., 1987; and Kosminski and Manzoori, 1990]. Results of thermodynamic calculations also suggest that vaporisation of sodium chloride is more intense in gasification environments than in combustion conditions

Low-Rank Coal and Advanced Technologies for Power Generation

55

[Kosminski and Manzoori, 1990]. It is also reported that sodium in low-rank coals forms silicates, aluminosilicates, sulphates, and chlorides during fluidised-bed gasification [Kosminski and Manzoori, 1990; and Bellin et al., 1987]. Alkali and alkaline-earth chlorides can melt at fluidised-bed gasification temperatures. Table 1 also includes some of the pure chemical compounds formed during gasification. Calcium sulphide and other sulphides are the main sulphur containing compounds in ash from a fluidised-bed gasification process. In addition to forming sulphides, calcium also leads to the formation of oxides, silicates and aluminosilicates [Laughlin and Reed, 1991]. However, magnesium mainly forms magnesium oxide under gasification conditions. Major problems arise from the iron compounds during gasification as these are known to initiate the formation of agglomerates in fluidised-bed gasification systems [Marinov et al., 1992]. Pyrite in coal decomposes to iron sulphide and a part of the sulphur is liberated. The iron sulphide undergoes further decomposition to form pyrrhotite, which oxidises to FeO. Iron oxide combines with silica to form silicates, which are precursors to agglomeration in fluidised-beds. FeS-FeO eutectics and metallic iron also initiate agglomeration, as these eutectics melt below 900 °C. Their formation is accelerated by the presence of alkali. During gasification processes the reactions involving inorganic matter occur not only on the char surface, but also extensively inside the char. This may lead to coalescence or sintering of inorganic matter into larger and consolidated ash particles [Kosminski and Manzoori, 1990]. Low-rank coals are generally rich in alkali and alkaline-earth elements that can be librated during combustion and gasification. Organically bound elements have been shown to be the major precursors to bed particle agglomeration and in-bed tube ash deposition. However, competing reactions with other coal inorganic matter may reduce the alkali availability. The fate of these potential deposit- and agglomerate-forming minerals will ultimately influence the extent of deposition and agglomeration. Therefore, it is important to understand the nature of inorganic components in the original coal and their chemical and physical transformation during various fluidised-bed combustion and gasification processes so that improvements can be made in the prediction of reactor performance.

4. TECHNIQUES FOR AGGLOMERATION AND DEFLUIDISATION CONTROL As discussed above, fluidised-bed combustion and gasification of low-rank coals suffer from agglomeration and defluidisation problems, in which sodium, potassium, calcium, and sulphur compounds, and their reactions, as well as the sodium—silica reaction, play a major role. Consequently, in order to control agglomeration and defluidisation, the formation of the low melting-point eutectics of these constituents and their reactions have to be limited. The following sections discuss several techniques that are currently under investigation.

4.1. Additives Mineral additives may be used to control ash related problems in fluidised-beds by limiting the formation of low melting point eutectics. An additive with specific characteristics can be employed in fluid bed combustors to suit a particular problem. The testing of four additives in a laboratory-scale spouted bed reactor to determine

56

D.-k. Zhang et al.

their effectiveness in controlling ash deposition under CFBC conditions has been reported [Linjewile et al., 1995]. The additives investigated were dolomite, CW clay (a kaolinite and sillimanite-rich clay), DV clay (a kaolinite and quartz-rich clay) and gibbsite (a hydrated alumina referred to as AH). In this study gibbsite was found to be most effective. Evidence obtained from analyses of the bed materials from the runs with the use of these additives suggests that depending on the type of additive used, interference in the process of ash deposition may be due to either a chemical reaction or physical dilution. These interactions apparently occur at the surface of burning char particles. Although previous investigations have identified clays to be effective in sodium capture, care must be taken in the selection of clay additives since the mineralogical composition of the clays appears to play a major role in the extent of deposit formation control. Gibbsite and the DV clay have demonstrated potential for in-situ sodium capture which may facilitate use of low-rank coals with high sodium contents in advanced FBC based power generation plants.

4.2. Alternative Bed Materials In recognition of the chemical reactions between silica and alkali metals (eg. sodium) that contribute to the formation of low-melting point eutectics, it has been proposed that aluminium-rich minerals may be used as alternative bed materials for control

of agglomeration and defluidisation. The technique works on similar principles to the use of additives. Bauxite and calcined sillimanite have been trialed in a laboratory spouted bed combustor burning Bowmans coal. The bauxite is rich in (60.7%) and (26.4%), while the sillimanite contains 53.3% and 43.7% SiO2. Combustion runs were carried out at bed temperatures of 800 °C and 850°C with the two alternative bed materials as well as silica sand until defluidisation occurred. Table 2 shows the observed times to defluidisation for each bed material. The results indicate that the operation can last for at least 7–10 times longer with calcined sillimanite and bauxite than with sand as the bed material. This is a significant improvement in terms of

combustion operation and could lead to longer run lengths and fewer outages when these alternative bed materials are employed in fluid bed combustors. The total bed mass recovered at the end of each combustion run (after defluidisation) was weighed. Irrespective of the bed material used, all the samples showed an ash coating build-up on the surface of the bed particles. The ash coating thickness on sillimanite and bauxite bed material particles was greater than for the sand. This is mainly due to the extended operation with these two bed materials. The composition of the inorganic constituents in the bed material coating was determined by chemical analysis. The analyses for the runs using sand as the bed material are reported on a silica-free basis. Figure 8 shows the average compositions of the

Low-Rank Coal and Advanced Technologies for Power Generation

57

58

D.-k. Zhang et al.

bed material coating (bed ash), cyclone ash (fly ash) and the feed coal ash for the different different bed materials. The results show that the coating on particles of all the bed materials is enriched in Na and S compounds. It may be noted that in comparison to the feed coal ash, a significant proportion of the Ca and Mg compounds was not retained by the bed particles and escaped from the bed into the cyclone. It is also evident that the ash coatings are enriched in alumina for the sillimanite and bauxite runs. Electron microprobe spot analyses and X-ray maps of polished cross sections of coated bed particles showed the presence of a highly sulphated ash dissolved in an aluminium-rich ash. XRD analyses indicated that low melting point eutectics were not formed, at least not in any significant quantity. It is postulated that, in the presence of aluminium-rich phases, the eutectics are less sticky at typical fluidised-bed combustion temperatures. Hence the use of bauxite and calcined sillimanite as the bed materials may result in longer operating periods (Table 2). XRD analyses of the ash-coated sand particles indicated the prevalence of thenardite, Na-Ca-sulphate and anhydrite. Thenardite and Na-Ca-sulphate are lowmelting point compounds capable of forming low-temperature eutectics with other ash compounds and hence are considered to be the basis of ash binder in the coatings. However, the bed material coatings from the combustion runs with calcined sillimanite and bauxite showed the presence of aluminium rich phases (such as nepheline, mayen-

ite, nosean) and complex silicates with Fe, Mg and Ca (Augite). It appears that the increased proportion of aluminium rich phases with calcined sillimanite and bauxite have prolonged the combustion operation. XRD analyses of the fly ash samples showed several mineral phases, such as quartz, maghemite, hematite, thenardite, in different proportions. The presence of aluminiumrich phases in the fly ash can lead to less sticky ash in downstream heat transfer sections. Hence the ash characteristics of the deposits on heat transfer surfaces may also be modified using these alternative bed materials.

4.3. Pretreatment of Coal Pretreatment of coal is one of the other options for controlling ash related problems in fluid bed combustors. This approach involves either addition or removal of specified elements in the coal to alter the coal ash properties before firing into the com-

bustor. The pretreatment is aimed at producing coal ash rich in high melting-point ash and lean in low melting-point ash. Several pretreatments have been conducted on Bowmans coal, including: water-washing (WW) (with demineralised water until the washing shows no traces of chloride), acid-washing (AW) (with 3M HC1 solution, followed by water-washing until the washing shows no traces of chloride), and Al enrichment (Al treated) (soaking of acid-washed coal in 1 M solution, followed by water-washing until the washing shows no traces of chloride). Combustion experiments using the pretreated coal samples were carried out on a spouted bed combustor at temperatures ranging from 700 to 850 °C. Silica sand was used as the bed material [Vuthaluru and Zhang, 1997b,c]. The bed material recovered at the end of a combustion run was weighed and the rate of ash deposition on the bed particles was estimated from the increase in the mass of bed inventory. Figure 9 shows the ash build-up with operating time for the samples

collected from the raw and the pretreated Bowmans coals. For the raw and acid-washed Bowmans coal runs, the bed defluidised after 80 and 15 minutes, respectively. However, for the water-washed and Al treated coals, no defluidisation occurred during the operat-

Low-Rank Coal and Advanced Technologies for Power Generation

59

ing periods shown in Fig. 9. The level of ash build-up for the raw coal was greater than for the pretreated coals. This is attributed to the small amounts of low melting point species in the pretreated coals. Despite the lower proportions of low melting species in the acid-washed coal, defluidisation did occur after a few minutes of operation at 800°C. This is in contrast to earlier observations by Manzoori and Agarwal [1994] with Lochiel

coal (SA) in the same fluidised bed; the raw Lochiel coal caused defluidisation after 2hr combustion at 800 °C, while the acid-washed Lochiel coal did not result in defluidisation for much longer operating periods (>10hrs). SEM analyses of bed particles of acidwashed Bowmans coal runs showed significant proportions of sulphur in the ash coating layer (in combination with Na). It appears that the difference in the proportions of sulphur in AW Bowmans coal and the AW Lochiel coal used by Manzoori and Agarwal [1994] may be responsible for the different behaviour of the two coals. This requires further study. Figure 10 shows the rate of ash deposition versus sodium content in different tests.

Comparison of results for different coals indicate that the ash deposition increases with the sodium content of the coal. The results suggest that sodium is transformed into low melting-point compounds in the ash. These compounds appeared to have enhanced the

rate of deposition on bed particles. Figure 11 shows the distribution of the inorganic constituents in the bed particle coating, cyclone ash and coal ash for the tests with the raw coal and the pretreated coals. The results show that the coating on the bed particles is enriched in Na and S compounds. A large proportion of Ca is retained in the ash coatings in runs with WW and AW coal samples. Al retention in the ash coating and the cyclone ash for the Al treated coal

runs is quite substantial compared to other coal runs. SEM analyses of ash-coated bed particles and cyclone ash samples from the Al treated coal runs indicated an Al-rich fluffy ash, similar to that reported by Vuthaluru et al. [1996]. This suggests that the increased levels of Al in the coal would lead to the formation of Al-rich compounds in the ash, rendering them less sticky when deposited on bed material particles. In addition, the formation of Al-rich fluffy ash in the cyclone ash in the Al treated coal runs would lead to fewer ash related problems in the downstream heat transfer sections of fluid bed combustors.

60

D.-k. Zhang et al.

X R D analyses [Vuthaluru and Zhang, 1997b] of ash-coated bed particles obtained from the raw and pretreated coal runs indicated that several minerals (maghemite, thenardite and hematite) were present in various proportions in the bed particles for raw and acid-washed Bowmans coal runs. However, for the water-washed and Al treated coal runs, those mineral phases were only present in trace quantities. In the latter cases, it appears that the absence of thenardite which acts as a binder may have prolonged the operation without defluidisation. In addition, periclase, maghemite, halite, magnetite, forsterite, anhydrite, and thenardite were also present in ash coatings with all the raw, WW and AW coal runs. For the Al treated coal runs, beta-aluminium oxide was detected, as a result of the increased levels of Al in the coal. The three techniques for agglomeration and defluidisation control discussed in this paper are effective to different extents, however continuing research and development are still required to determine the most effective and cost-efficient technique. Of the three techniques described, the use of sillimanite as an alternative bed material appears to be the most promising. Coal pretreatment is unlikely to be economic.

5. ALTERNATIVE ADVANCED TECHNOLOGIES Three principal factors, viz, coal ash composition, process temperature and processing environment, govern the behaviour of coal ash in fluidised-bed combustion and gasification processes and consequently determine the limits on operation due to agglomeration and defluidisation. The process temperature and coal residence time in fluidisedbed gasification are generally selected to achieve the desired level of coal conversion. Low temperature operation, while offering the potential for alleviation of ash related problems, will lower the extent of coal (and char) conversion, and possibly result in the formation of tars. Hence processes based on low-temperature operation, such as low temperature pyrolysis, will require different process configurations to conventional gasification processes, and may include co-production. In a low-temperature pyrolysis process, low-rank coal undergoes thermal

Low-Rank Coal and Advanced Technologies for Power Generation

61

62

D.-k. Zhang et al.

decomposition (pyrolysis) in an inert environment to produce volatile matter and char at a moderate temperature, say, at about 600 °C. The char can then be burnt in a combustor with sub-stoichiometric air, again at relatively low temperatures (say, 650 °C), to generate an inert gas (flue gas) and heat for the endothermic pyrolysis reactions. The excess heat from the combustor can be readily used for power generation. The combustible volatile matter may be utilised in two ways, viz, direct combustion for power generation, or separated into condensables (fuel oil) and non-condensables (a low-specific-energy gas mixture), the latter can be utilised for power generation. Utilisation of volatile matter in low-rank coals is feasible because of their high volatile matter contents (35–50%db.). A number of similar processes for low-rank coal utilisation can also be proposed. Partial gasification of the pyrolysis char with steam and air (or oxygen) may be used to replace the char combustion. Activated carbon production from the pyrolysis char is also an option, given the highly porous and reactive nature of the char. Tars from pyrolysis of low-rank coal may also be used to produce high-value organic chemicals. Processes for co-production of power, methanol, and ammonia and/or fertiliser from low-rank coal have also been proposed. Laboratory-scale experimental research and process simulation studies are being conducted at the University of Adelaide and the CRC—Power Generation to evaluate the full potential of these processes.

6. CONCLUSIONS Fluidised-bed based advanced power generation technologies offer higher efficiencies than conventional pulverised fuel fired power plants and better prospects in reducing ash-related problems associated with low-rank coal in such plants. However, bed

material agglomeration and bed defluidisation present significant operational difficulties for the utilisation of the low-rank coal in fluidised-bed processes. Alkali and alkalineearth elements and sulphur compounds, often found in low-rank coals, form low melting point eutectics at typical fluidised-bed combustion and gasification operating temperatures. These low melting-point materials are subsequently transferred onto the bed material particle surfaces, and the ash-coated particles then become adhesive and agglomerate. Defluidisation can occur either as an extension of agglomeration as a rate process gradually leading to defluidisation or as an instantaneous event without agglomeration. A critical thickness of the ash coating layer on the particle surface exists, above which defluidisation occurs. This critical thickness decreases with an increase in bed temperature. Several mineral additives, alternative bed materials and pretreatment of coal have been shown to suppress, to different extents, particle agglomeration and bed defluidisation when burning a high sodium, high sulphur low-rank coal in a spouted fluidised-bed combustor. Sillimanite as an alternative bed material is found to be most effective for defluidisation control. Alternative advanced technologies such as low-temperature pyrolysis and co-production are proposed for future investigation.

ACKNOWLEDGEMENTS The authors gratefully acknowledge the financial and other support received for this research from the Cooperative Research Centre for New Technologies for Power Generation from Low-Rank Coal, which is established and supported under the Australian Government’s Cooperative Research Centres program. Thanks also go to Dr. Yinghe He,

Low-Rank Coal and Advanced Technologies for Power Generation

63

Mr. Adam Kosminski, and Ms Marnie Telfer for their contributions to the work reported in this keynote paper.

REFERENCES Arastoopour, H., Huang, C.S. and Weil, S.A. (1988). “Fluidisation behaviour of particles under agglomerating conditions.” Chemical Engineering Science, 43, 3063–3075. Basu, P. (1982). “A study agglomeration of coal-ash in fluidised-beds.” The Canadian Journal of Chemical Engineering, 60, 791–795. Bellin, A., Bruck, U. and Schrader, L. (1992). “Report on the results of the HTW-pilot plant gasification tests with Bowmans coal.” Journal of the Institute of Energy, 65, 55062. Benson, S.A., Sondreal, E.A., and Hurley, J.P. (1995). “Status of coal ash behaviour research.” Fuel Processing Technology , 44, 1–12. Brockway, D.J. and Zhang, O.K. (1997). “Cooperative Research Centre for New Technologies for Power Generation from Low-rank Coal.” in World Coal Combustion Research Centres (Ed. L.D. Smoot), Progress in Energy and Combustion Science, in press . Dean, J.A. (1985). Lange’s Handbook of Chemistry, New York: McGraw Hill Book Co. Durie, R.A. (1991). The Science of Victorian Brown Coals: Structure, Properties and Consequences for Utilisation. London: Butterworths, Heinemann. Gluckman, M.J., Yerushalmi, J. and Squires, A.M. (1976). Fluidisation Technology, (Ed. D.L. Keairns), Washington: Hemishpere Publishing Corporation, Goblirsch, G.M., Benson, S.A., Karner, F.R., Rindt, D.K.. and Hajicek, D.R. (1983). “AFBC bed material performance with low-rank coals.” a paper presented at the 12th Biennial Lignite Symposium, N.D. USA. Goblirsch, G.M., Vander Molen, K.H., Wilson, K. and Hajicek, D.R. (1980). “Atmospheric fluidised-bed testing of North Dakota lignite.”, 6th I n t . Conf. of Fluidised-bed Combustion, 850. Godel, A.A. (1966). (in French) Rev Gen Therm , 5, 349. Goldberger, W.M. (1967). “Collection of flyash in a self-agglomerating fluidized-bed coal burner.” a paper presented at ASME Winter Annual Meeting and Energy systems Exposition. Gransden, J.F., Sheasby, J.S. and Bergougnou, M.A. (1970). “An investigation of detluidisation of iron or during reduction by hydrogen in a fluidised-bed.” Chemical Engineering Progress Symposium Series, 66 (105), 208–214. He, Y. and Zhang, D.K. (1997). “Types of detluidisation in a spouted bed” CRC Power Generation Internal Report. Kosminski, A. and Manzoori, A.R. (1990). “Inorganic matter behaviour in the gasification of South Australian coals for combined cycle power generation”, Technical Services Department, The Electricity Trust of

South Australia, Report submitted to SENRAC. Langston, B.G. and Stephens, F.M. (1960). “Self-agglomerating fluidised-bed reduction.”, J Metals, 312. Laughlin, K. and Reed, G. (1991). “Behaviour of coal ash in a pressurised fluidised-bed gasifiers”, 6th International Conference on Coal Science, Newcastle upon Tyne, 396–399. Levin, E.M., Robbins, C.R. and McMurdie, M.F. (1979). Phase diagrams for Ceramists. The American Chemical Society, Ohio, USA. Linjewile, T.M., Linard, D. and Manzoori, A.R. (1995). “Mechanistic evaluation of the role of additives in controlling bed material ash deposition and agglomeration during fluidised-bed combustion of low-rank coal.”, Proc. 5 th Japan-Australia Joint Technical Meeting on Coal, Adelaide. Manzoori, A.R. and Agarwal, P.K. (1994). “Agglomeration and defluidisation under simulated circulating fluidized-bed combustion conditions.” Fuel, 73, 563–568. Manzoori, A.R. (1990). Role of the inorganic matter in agglomeration and defluidisation during the circulating fluid bed combustion of low-rank coals. PhD. Thesis, The University of Adelaide. Marinov, V., Marinov, S.P., Lazarov, L. and Stefanova, M. (1992).” Ash agglomeration during fluidised-bed gasification of high sulphur content lignites.“ Fuel Processing Technology, 31, 181–191. McIntosh, M. (1997). ”Assessment of advanced power cycles with high moisture content low-rank coals.“ 5th Annual Technical Seminar of APEC Experts’ Group on Clean Fossil Energy, Nevada, October 27–30, 1997. Mojtahedi, W. and Backman, R. (1989). “Release of alkali metals in pressurised fluidised-bed combustion and gasification of peat.” Technical Research Centre of Finland (VTT) Publication No.53, 48.

64

D.-k. Zhang et al.

Raask, E. 1985. Mineral Impurities in Coal Combustion. New York: Hemisphere Publishing Corporation.

Souto, M.A., Rodriguez, J.C., Conde-Pumpido, R., Guitian, Gonzalez, J.F. and Perez, J. (1996). “Formation

of solid deposits in the gas circuit of a pressurised fluidised-bed combustion plant”, Fuel, 75, 675–680. Telfer, M. and Zhang D.K. (1997). “X-ray Mapping of Sulphur Constituents in Low-Rank Coal Char: The Influence of Pyrolysis Conditions.”, To be presented at Australian Symposium on Combustion and 5 th Australian Flame Days, Sydney, November, 1997. Vuthaluru, H.B. and Zhang, D.K. (1997a). “Ash characteristics of low-rank coals during fluidised-bed combustion.” CRC—Power Generation Internal Report. Vuthaluru, H.B. and Zhang, D.K. (I997b). “Role of inorganics during fluidised-bed combustion of low-rank coals.” A paper to be presented at the Engineering Foundation Conference, November 2–7, Hawaii.

Vuthaluru, H.B. and Zhang, D.K. (1997c). “Effect of pretreatments on ash characteristics of low-rank coals during fluidised-bed combustion.” CRC—Power Generation Internal Report.

Vuthaluru, H.B., Domazetis, G., Wall, T.F. and Vleeskens, J.M. (1996). “Reducing fly ash deposition by pretreatment of brown coal: effect of aluminium on ash character.'Fuel Processing Technology, 46,117–132. Vuthaluru, H.B., He, Y, Zhu, J.N., Zhang, D.K., Yurismo, H. and Sastrawinata, T. (1997). “Fate of inorganic

constituents during fluidised-bed combustion of low-rank coals: effect of temperature on ash characteristics”, Proceedings, 9th International Conference on Coal Science, 1171–1174. Wall, T.F. (1992). “Mineral matter transformations and ash deposition in pulverised coal combustion.” Proceedings, the 24th Symposium (International) on Combustion, The Combustion Institute, Pittsburgh, 1119–1126. Williams, J.E.. (1984). “Fluidised combustion of brown coal.” State Electricity Commission of Victoria,

Research and Development Dept., Report No. ND/84/006. Yerushalmi, J., Gluckman, M.J. and Graff, R.A. (1976). Fluidisation technology. (Ed D.L. Keairns),

Washington: Hemishpere Publishing Corporation, 437.

THE THERMAL CONDUCTIVITY OF ASH DEPOSITS Particulate and Slag Structures

R. P. Gupta, T. F. Wall1, and L. Baxter 2 1

The Department of Chemical Engineering The University of Newcastle, Australia 2 Sandia National Laboratories Livermore, CA

1. INTRODUCTION The thermal properties of the ash deposits have a significant influence on the heat absorption rate in the pulverised fuel fired furnaces. Heat transfer rates depend primarily on surface absorptivity, the effective thermal conductivity, and thickness of the deposit. The radiative properties and thermal conductivity of deposits have significant influence on the boiler performance deposit (Mulcahy et al., 1966). The radiative properties of ash deposits and similar materials were reviewed earlier (Wall et al., 1994). This review is limited to the thermal conductivity of ash deposits and materials having similar chemical and physical character. Thermal conductivity of glassy ash like materials is a function of it constituents, however, it is in the range of for most solid glasses. The reported low thermal conductivities of some particulate ash deposits (typically less than for porous materials) can produce temperature gradients of or larger in the deposit resulting in accelerated ash deposition. The deposit thickness, however, depends indirectly on the temperature distribution in the deposit, which in turn is governed by the effective thermal conductivity. The thickness of deposits can vary from zero for clean tube surfaces to a maximum value, where the slag starts flowing, determined by viscosity and thermal properties of the deposit (Mulcahy et al., 1966). In that investigation the authors noted that a deposit thickness of about 1 mm can be expected to decrease the heat transfer significantly. The thermal properties are strongly influenced by the physical structures of the deposits: particle size, porosity and sintered condition. Sintering is a function of Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

65

66

R. P. Gupta et al.

chemical composition and temperature, which has already been discussed and causes

irreversible changes in the structure. As sintering increases, contact increase and the structure of ash deposits vary from granular packed bed for loose deposits to porous slags with solid phase continuous as shown in Fig. 1. The particles in particulate phase are also connected at a very minimum level not shown in Fig. 1. The ash deposits can be characterised broadly into three types of structures and all the three types can exist in a given deposit as seen in Fig. 2. The proportion of different types of structure however depends on the mineral matter in coal and the environment surrounding the deposit. The initial layer adjacent to the wall or a heat transfer surface is usually particulate in nature with continuous gas phase and discrete particles embedded in the continuous phase. This is characterised by low thermal conductivity resulting in increased deposit temperature. The outermost layer, at very high temperatures can be completely fused slag and has very high thermal conductivity depending on its porosity. In slags, the continuous phase is the solid ash with voids resulting in a porous structure. In porous ash deposits, apart from conduction, heat transfer also takes place via radiation mechanism as well, whose contribution increases as temperature increases. At high temperatures encountered in pulverized fuel fired furnaces, radiation plays an important role in heat transfer through the deposit. Since conductive and radiative heat transfer in the ash deposit occur both in parallel and in series, the concept of an “effective” thermal conductivity is commonly used. This quantity is defined such that the following form of the heat transfer equation for one-dimensional, steady-state conduction is satisfied.

Since the heat flux, q, is constant, the local effective thermal conductivity can be defined as;

The Thermal Conductivity of Ash Deposits

67

Depending on the temperature, physical state, and surface condition at the wall and deposit surface, the ratio of conductive to radiative heat fluxes will change with position in the deposit. An overall effective thermal conductivity calculated for the entire ash

deposit thickness, can be defined by integrating the local value with respect to location (y) and dividing by the deposit thickness. This effective value will increase rapidly at the temperatures when radiative transfer through the deposit becomes significant.

2. MEASUREMENT OF THERMAL CONDUCTIVITY This section reviews the thermal properties of ashy materials, their measurement techniques. Tye (1969) has compiled several reviews on methods for determining thermal conductivity of solids. Pratt (1969) has discussed specifically the experimental techniques for low conductivity materials ie. glasses, slags and powdered materials. Most of the techniques have been used under steady state, however, some of the methods utilise transient analysis of unsteady state measurements. The unsteady state methods were designed to

save time in experiments, however, the analysis is slightly more complex. Laubitz (1969) has discussed the effect of design parameters for these techniques on the accuracy of measurements. He has also discussed the possible sources of error involved in terms of: • measuring heat flux from power of heater,

• temperature measurement, • contact resistance and • departure from steady state. The thermal conductivity measurement techniques are of two types: radial heat flow technique mostly for powders and axial heat flow technique for discs

2.1. Radial Heat Flow Technique for Powders Most of the radial flow techniques consist of two concentric ceramic cylinders. The test material, which is in the form of a powder is filled in the annular space between these cylinders. The inner ceramic tube is heated electrically by a carbon rod. Temperatures are measured at several radial locations through the powder. The thermal conductivity of the particle bed is determined from temperatures at two radial locations by the following equation:

The typical experimental apparatus is described in Fig. 3. A number of variations are discussed below: Diessler and Boegli (1958) determined thermal conductivity of magnesia powder in different gases. The effective thermal conductivity for the powder in helium was found to be 4–5 times that in the air. A reasonable good correlation was obtained for the ratio of with the ratio The thermal conductivity of gas, rather than that of solid, was found to have more influence over the effective thermal conductivity of powders.

68

R. P. Gupta et al.

Laubitz (1959) measured thermal conductivity of powders. The experimental apparatus essentially consisted of a long cylindrical furnace of about 25mm diameter inside which was a centrally located heater of 4mm diameter and 30cm long. The furnace is filled with powder. Two thermocouples located within the powder at radial distances of 5.6mm and 10.1 mm give the temperatures at two radial positions. The thermal conductivity is then determined from equation (3). Thermal conductivity of several powders of uniform particle size such as alumina, magnesia and zirconia were determined at temperatures up to 1,200°C. Flynn (1963) also used an apparatus based on radial flow method for measurements on powders and granular materials over a wide range of temperatures. The inner ceramic tube contained a central heater (approximately 41 cm long). There were heaters at both ends of the main heater, whose temperature is adjusted to that of the central heater. The outer ceramic tube has 3.2cm diameter and about 60cm long. One of the thermocouples is inserted into the central ceramic tube and the other is inserted into a thermocouple well in the outer ceramic tube. The accuracy of these methods is influenced by two factors: thermal resistance at the interface and temperature measurement or thermocouple placement. The possible sources of error are thermal resistance between sample and thermocouple, uncertainties in the exact location of the thermocouple. The location of the thermocouple may alter if the samples shrink. If the thermocouples are bare, they may be contaminated by reaction with test material. A combination of all the uncertainties showed that a maximum error was no greater than 3% at 1,000°C in Flynn’s experiments. Godbee and Ziegler (1966) used radial heat flow in a hollow cylinder to measure thermal conductivity under steady and unsteady conditions. The powdered sample is held between two concentric cylinders. Heat produced by an axially located electrical resistance is conducted radially outward through the sample. Power to the central heater was determined from the measured voltage and current. The transient method was mostly used at temperatures close to the ambient. The steady state method was used up to temperatures of 1,000°C. The thermal conductivity was determined for well-characterised powders. The accuracy of k e by both the steady state and unsteady state techniques was estimated to be within 10%, where as the reproducibility was within 3%.

2.2. Axial Heat Transfer Technique for Slabs/Discs A number of methods for the measurement of thermal conductivity of slags or glasses utilise axial heat transfer technique. The test material is in the form of two identical circular or square slabs, which are placed one on each side of a flat heater. A steady state is established by adjusting the power to the heater and flow rate of the cooling fluids

The Thermal Conductivity of Ash Deposits

69

of the sink. The temperature of the hot guards is maintained the same as the central metering section to eliminate the radial heat losses. The thermal conductivity can then be determined from temperature difference, across the thickness of the sample, by the following equation:

Ratcliffe (1963) determined thermal conductivity of 22 glasses using an electrically heated disc method of determining thermal conductivity of solids. Each glass sample was in the form of a pair of identical three inch diameter discs (10mm thick) with lapped faces. He derived some additive correlations for thermal conductivity as a function of chemical composition of glasses by weight. Boow and Goard (1969) used the axial heat transfer technique to measure the thermal conductivity and emissivity of an ash layer. The ash layer of 3–5 mm thickness is kept on a flat-bottomed stainless steel crucible sitting on top of a furnace. The surface

temperature of the layer and its emissivity are measured by a surface pyrometer. The bottom temperature is determined by means of a thermocouple. The heat flux used in equation (4) can be determined from top surface temperature and emissivity. The measurements proved to be remarkably reproducible. The measurements demonstrated the effect of particle size, chemical composition. They also demonstrated the irreversible rapid increase in the thermal conductivity at higher temperatures due to sintering. For very fine ashes, the thermal conductivity was observed to be less than that of air at that temperature. Anderson et al. (1987) developed an experimental facility to study steady state heat

transfer through coal ash deposits with an objective to determine local thermal conductivity of fly ash, slagging deposits and fouling deposits. The facility consisted of a large cavity kiln capable of continuous operation at 1,500 °C. Six silicon carbide electrodes irradiate uniformly a thin ash deposit sample. The irradiation level was similar to those in coal fired furnaces. A calorimeter was located directly under the center of the sample. The temperature profile across the calorimeter surface was uniform. Botterill et al. (1989) used the unsteady state two-dimensional heat transfer technique to measure thermal conductivity of alumina powder. A large amount of ash sample was uniformly heated by fluidising the powder in a cylinder of 70cm diameter by 70cm length. After attaining the required uniform temperature, a carefully set up array of thermocouples was located in the bed and the bed defluidized. As it cooled over a period of days, the radial temperature profiles were logged. The data was analysed to determine the thermal conductivity of the powder at different temperatures. Nowok and Steadman (1992) measured thermal conductivity of coal ashes and slags using the axial one-dimensional heat flow method. The method uses two reference material of known thermal conductivities as a function of temperature. Similar technique has been used at the University of Newcastle (Gupta et al., 1997) to investigate the effect of porosity and sintering on the effective thermal conductivity of several coal ashes and synthetic slags. The axial heat flow is assumed constant across the three materials. The schematic diagram is shown in Fig. 4. The temperature difference across one reference material is used to determine the heat flux, whereas, the temperature difference across the other reference material is used to estimate the errors in the measurements. It must be noted that mostof the techniques described above may not be suitable for highly porous deposits. One of the in-situ technique is described by Robinson and Baxter (1997).

70

R. P. Gupta et al.

3. THERMAL CONDUCTIVITY OF PURE MATERIALS (ZERO POROSITY) It is important to consider the thermal conductivity of each individual pure phase before one delves into modeling the effective thermal conductivity of complex structures. The fundamental mechanism of heat transfer in material like dense and pure refractory oxides, glasses and ashy material can either be understood as propagation of elastic waves due to lattice vibration or as the interaction between quanta of thermal energy known as phonons. The thermal conductivity resulting from phonon transfer is represented by a function of mean free path of phonon, which decreases with decreasing order in molecular structure. For this reason, the thermal conductivity of crystalline material (alumina and magnesia) is much higher than the amorphous material (silica) and glasses by one to three orders of magnitude (Kingery, 1976) as seen in Fig. 5. This was best illustrated by Berman (Kingery, 1976) where the micro-structure order at molecular level in quartz crystal is randomized by neutron irradiation. Increased irradiation resulted in a

decrease of the thermal conductivity by three orders of magnitude.

The Thermal Conductivity of Ash Deposits

71

The mean free path of phonon in glasses and dielectric is limited to the order of inter-atomic distances by random structure resulting in lower thermal conductivity. The effect of composition of glasses or ashy material, has very small effect on their thermal conductivity resulting from phonon transfer. The temperature dependence of this process is well understood. The phonon mean free path in crystalline materials and subsequently, the thermal conductivity decreases rapidly with increase in temperature, whereas the phonon conduction does not vary significantly with temperature in amorphous materials like ash and glasses. In addition to heat conduction by lattice vibrations in such solids, a much smaller energy transfer results from higher-frequency electromagnetic radiation. This component is termed as photon conductivity and becomes important at high temperatures as it is proportional to the fourth power of temperature. The photon conductivity becomes important in Alumina above 1,600 K, whereas it is important in glasses even above 800 K. The photon conductivity is dependent on the optical properties of the material and is shown to be inversely proportional to the absorption coefficient of the material. The photon conduction becomes important at temperature of 800–900 K, whereas for materials with high absorption coefficient, the photon conductivity becomes important only at very high temperatures. There are two ways, in which temperature increases the photon conductivity. Firstly, more radiant energy being concentrated in shorter wavelengths and secondly, most of the materials have lower absorption index at smaller wavelength resulting in rapid increase in photon conductivity at higher temperatures (Kingery, 1976). It must be noted that the effect of macro-structure (porosity and pore/particle size) can overshadow the effect of the microstructure discussed in this section. Even half a

percent of porosity can bring down the effective thermal conductivity of pure materials at high temperatures by an order of magnitude.

4. THERMAL CONDUCTIVITY OF ASH AND ALIKE MATERIALS This section discusses the experimental values for thermal conductivities of coal ashes and ash like materials ie oxides and glasses. These results are discussed in order to understand the effect of main parameters (such as temperature, structure, porosity, chemical composition and particle or pore size) influencing the effective thermal conductivity of ash deposits.

4.1. Effect of Temperature In an earlier review (Wall et al., 1979), the thermal conductivities of pure oxides

and mixtures were compared with those of coal ashes. The thermal conductivity of dense and MgO decreases with increasing temperature, however, for silica (SiO2), the thermal conductivity increases continuously with temperature (Fig. 5). The rate of increase is rapid at high temperatures due to the radiation contribution. The thermal conductivity of shows a minimum at about 1,600K. At higher temperatures the increase is due to the radiative contribution. The effect of radiation becomes appreciable above 500 °C for glasses. It is seen that the variation with temperature for ashes is less than that for pure oxides and pure phases. The thermal conductivity measurements are seen to be nearly independent of temperature usually below a temperature of 900–1,200 °C, at which sintering occurs. A slow

72

R. P. Gupta et al.

and steady increase in thermal conductivity up to this temperature is attributed primarily to changes in the thermal conductivity of gas phase. A steep rise in thermal conductivity is either due to appreciable increase in the radiative component as mentioned earlier or due to sintering that softens the sample and improves the contact between particles. The steep rise due to radiation is reversible, whereas the sintering causes irreversible changes in the structure responsible for higher thermal conductivity as observed by Anderson et al. (1987) in Fig. 6. After fusion, the sample has a higher thermal conductivity than that of the original particulate sample as it was heated. The chemical and physical changes result in the thermal conductivity measurements being irreversible. The higher the temperature at which the cooling cycle starts, the higher the thermal conductivity on cooling. However, on cooling of particulate (unsintered), the thermal conductivity is reversible.

4.2. Effect of Chemical Composition It is of interest to note that for mixtures of oxides, the thermal conductivity is approximately weighted to the mass proportions of the component oxides. The refractories having less than 50% silica have a decreasing thermal conductivity with temperatures. The dense mixtures with more than 50% silica show a continuous increase in thermal conductivity with temperature (Wall et al., 1979). According to Anderson et al. (1987) silica ratio is important at high temperatures, where strong sintering and fusion can occur. Ratcliffe (1963) investigated the thermal conductivity of glasses with varying composition at temperatures between –150 °C and 100°C. Additive formulae were prescribed, which are claimed to estimate the thermal conductivity of most glasses within 5% accuracy in the given range. The effect of chemical composition in this manner does not influence the thermal conductivity of the material significantly. However, glasses of varying chemical composition are expected to have different optical constants resulting in significant differences in the radiative component of the effective thermal conductiv-

ity at high temperatures. The glasses containing high iron oxides are expected to have lower thermal conductivity. However, Boow and Goard (1966) also observed higher thermal conductivity for iron containing ash deposits, which is attributed to improved sintering of the deposits rich in iron. Further, the chemical composition determines the structure of the deposit at a given temperature. There is a vast difference in the thermal conductivity of particulate ash and a sintered deposit.

The Thermal Conductivity of Ash Deposits

73

4.3. Effect of Particle Size The experimental data from Laubitz (1959) shows an increase in thermal conductivity with increase in particle size. Nasr et al. (1994) also measured higher effective

thermal conductivity for larger particle sizes in the size range of 3mm to 12mm alumina, glass and silica particles in a packed bed. In both the cases the effect of particle size was observed to be more evident at higher temperatures suggesting that the particle size plays an important role in radiative component of the effective thermal conductivity. Godbee and Ziegler (1966), somehow failed to observe a significant effect of particle size on the effective particle size of magnesia. Boow and Goard (1969) clearly show an increase of thermal conductivity with an increase in mean particle size. The published data for a number of ashes of varying mean particle size at 700 °C is presented in Fig. 7. Least square regression of their experimental data yielded a logarithmic equation:

The ash samples with low median diameters such as present in primary layers of boiler deposits or the laboratory ash, have very low thermal conductivity, even less than that of air (Boow and Goard, 1969). The thermal conductivity of a

gas is due to molecular motion, which decreases significantly when the inter-particle space is comparable to the mean free path of the gases. The same can happen at low gas pressures, where the mean free path of the gas increases and becomes comparable to interparticle space resulting in very low thermal conductivities. This behaviour was also observed by Diessler and Boegli (1958). The thermal conductivity of such layers has been observed to be less than that of the air.

4.4. Effect of Porosity The thermal conductivity of particulate and slag structures increases with a decrease in void fraction or porosity. This effect has been observed by several researchers. Krupiczka (1967) analysed the effective thermal conductivity of granular material for cylindrical and spherical grains embedded in a continuous gas medium available in literature. The resulting correlations predicted the effective thermal conductivity of particulate beds within 30% of the experimental data from several researchers for a number of powders in different gases. The correlation is given by the following equation:

74

R. P. Gupta et al.

In more recent measurements, Nowok and Steadman (1992) measured thermal conductivity of coal ashes and slags of different porosity, sintered at different tempera-

tures. The thermal conductivity of gehlenite sintered at 1,300°C increased from about at room temperature to

at 450°C.

4.5. Effect of Deposit Structure The deposit structure is characterised by porosity, pore/particle size distribution, and state of sintering. Sintering is a function of chemical composition and temperature, which has already been discussed and causes irreversible changes in the structure. As sintering increases, contact increase and the structure changes from particulate nature to partially sintered to that of completely fused slag or foam type structure with solid phase continuous as shown in Fig. 1. The data in Fig. 6 shows that the thermal conductivity of powdery ash is quite low but will increase by up to a factor of ten for sintered ash. This agrees with data for pure oxide powders, bricks and cast refractory materials reported by Wall et al. (1979). This emphasises that the physical structure of the ash deposit has a strong influence on thermal conductivity.

5. MODELS OF THERMAL CONDUCTIVITY (MACRO-STRUCTURE) There have been a number of reviews on thermal conductivity of granular materials (Krupiczka, 1967, Hadley, 1986, Botterelli, 1989 and Bauer, 1993). This section discusses several approaches towards estimation of thermal conductivity of porous or particulate materials as shown in Fig. 1 at low temperatures where radiative heat transfer can be neglected. Treatment of radiation becomes essential in estimating the effective

thermal conductivity of the deposits at high temperatures and is discussed in the following section. The slagged deposits correspond to the porous material similar to the structure of foamed material with discrete bubbles or pores of entrapped air. The loose deposits formed from high melting point ash are similar to particulate structure with discrete solid particles embedded in continuous gas phase. Thermal conductivity of particulate or slag type structures depends on the following: • thermal conductivity of solid phase ks • thermal conductivity of gas phase kg • porosity and • the size distribution of pores or of the particles • connectedness of particles The thermal conductivity of solid and gas phase is influenced by its chemical composition. There does not appear to be any model that considers particle size in the estimation of the effective thermal conductivity. The effect of particle size is significant at high temperatures and is considered in the next section. Hence this section essentially considers the effective thermal conductivity models as a function of porosity. The porosity of ash deposits can vary from 0.1 for molten slags to 0.95 for particulate deposits

The Thermal Conductivity of Ash Deposits

75

depending on conditions of sintering and time. There has been an attempt to obtain the initial estimate of deposit by using a stochastic model. The initial estimate can be modified by using some sintering model (Rezaei, 1997). The simplest approach is to approximate the porous deposit structure consisting of parallel layers of solid and gas. The effective thermal conductivity is given by Equation (7) if the direction of heat transfer is parallel to these layers as shown Fig. 8a. The correlation is given by Equation (8), when the heat transfer is in a direction normal to these layers as in Fig. 8b.

where e is the porosity of the material. The heat transfer is higher when the heat trans-

fer is in the direction parallel to the layers. The equations (7) and (8) correspond to upper and lower limits on thermal conductivity for a given porosity in practice. Rayleigh (1892) derived an exact theoretical equation for the effective thermal con-

ductivity for a cubic array of uniform spheres (limits porosity greater than crete spheres are embedded in a continuous phase as in Fig. 9a.

The dis-

where

= thermal conductivity of continuous phase, = thermal conductivity of discrete phase, r = p = is the volume fraction of discrete phase. The simplified model assumes the temperature isotherms normal to the direction of heat transfer. The equation can be used for slags by considering air bubbles as discrete phase and for particlulate by considering air as continuous phase. Russell (1935) calculated the thermal conductivity for an array of cubes in a cubical array (Fig. 9b), assuming again the isotherms being planes normal to the direction of flow. This model has a definite advantage over Rayleigh’s model, as it can accommodate

the complete range of porosity.

76

R. P. Gupta et al.

Diessler and Boegli (1958) solved the Laplace equation for a cubic lattice of spheres and presented results in the form of thermal conductivities normalised to the thermal

conductivity of gas. They approximated the heat transfer in the direction normal to heat flow direction is extremely small. Woodside (1958) also considered a cubic lattice of spheres and obtained an expression for the case, where isotherms are planes normal to the direction of flow. Leach (1993) considered cubic cells stacked together in an attempt to determine the thermal conductivity of foam like materials. He considered two ways of stacking the cubic cells of air. Air is considered as discrete phase. The cubic parallel series (CPS) model considers the layers gas and of the composite material stacked normal the direction of heat transfer (Fig. 10a). This model is the same as that of Russell’s model for a cubicle array of uniform cubes, except in this case the discrete phase consists of air bubbles. The cubic series parallel (CSP) model (equation 11) first considers solid and gas in series across which the heat transfer takes place, and then these series are considered as a composite material stacked in parallel with gas layers as shown in Fig. lOb.

Rayleigh’s model for an array of spheres was also used to determine the thermal conductivity of foams. It can be shown that equation for low density foams or slags all these three relations can be reduced to one single correlation given by Equation (12).

N being 1 , 2 , 3 for CSP, CPS and spherical models. For all the models the first term is sufficient to fit the experimental data on foams. These models are valid for slags, however, these can be used for particulate deposits by interchanging the discrete and continuous phases. Figure 11 compares these three models CPS (also similar to Russell’s model), CSP based on array of cubical cells, and array of spheres (Rayleigh’s Model) for particulate and slag type deposits. It should be noted that Rayleigh model can predict the thermal conductivity over a limited range of porosity as discussed earlier. The suffixes -S and -P in Fig. 11 relate to slags and particulate type structures, respectively. The thermal conductivity for slag structures is considerably higher than that for the particulate structures,

especially in the porosity range of 0.2–0.8. All these models lie within the two extreme limits given by Equations (7 and 8).

The Thermal Conductivity of Ash Deposits

77

The CPS model, considering layers of composite materials stacked in series with layers of fluid (isotherms normal the heat flow, or infinite conductivity in direction normal to the heat flow), predicts higher thermal conductivity compared to the other models. The figure also suggests that the predictions from Rayleigh’s model based on the array of spherical particles lie between the two models: CSP and CPS based on arrays of cubes. Sintering produces a solid continuous phase resulting in higher thermal conductivity. The reduction in density during sintering also results in increased thermal conductivity. The solid thick line in Fig. 11 shows how the porosity and effective thermal conductivity is expected to increase for a maturing ash deposit. Confirmation of the trends awaits future experiments. This increased thermal conductivity of the deposit due to sintering is very well observed in the experimental data of Mulcahy et al. (1966), as shown in Fig. 6.

78

R. P. Gupta et al.

5.1. Thermal Conductivity of Complex Structures These models can not be used to predict the thermal conductivity for complex materials, where both the fluid continuous (Fig. 1a) and solid continuous (Figure 1b) regions co-exist as shown in Fig. 12. The next three models are related to the thermal conductivities of such materials, where both phases are continuous in some parts of the total

structure. The structure is therefore defined by some structural parameteres in all these models; the parameters were determined by best fit to their experimental data. Brailsford and Major (1964) considered sandstone as a random mixture of solidcontinuous-phase and fluid-continuous-phase for determining its thermal conductivity. The two-phase assembly can be regarded as having regions of the above mentioned phases embedded in a random mixture of the same two phases. The concept does not favour the spatial continuity of one single phase. The resulting relation is given by the following equation:

where ;

and and being the volume fraction of the two phases. The thermal conductivities and in equation correspond to the thermal conductivities of fluid-continuous phase and solid-continuous phase, respectively and can be determined using Rayleigh’s correlation for an array of spheres. It was assumed that the porosity in the two phases is similar. The equation (13) needs the volume fractions for the two phases for determining and for the solid and fluid continuous phases, respectively using equation (8). The assumption that the fraction of fluid-continuous phase is same as that of the overall volume fraction of the fluid phase resulted in an overestimate of of sandstone. Figure 13 compares the random distribution of the two phases with two continuous phases. The random model gives results closer to the solid phase continuous for low porosity and results closer to the fluid phase continuous for high porosity. They also suggested an empirical correlation to estimate the volume fraction of fluid continuous phase from fraction of fluid phase ‘p’ given by the following equation:

where factor is related to the structure of the deposit. The results for sandstone agreed with predictions for The proportion of solid phase continuous, a

The Thermal Conductivity of Ash Deposits

79

measure of sintering, can be an input parameter for the model. An increase in sintering results in a higher proportion of solid continuous phase and the from random phase distribution model would move closer to the slag model as shown by the arrow in Fig. 13. Hadley (1986) introduced an extra parameter and modified Rayleigh’s correlation for fluid phase being continuous, as the previous correlation was with the assumption of particles being too far apart. He accounted for increase in conductivity due to establishments of contacts between the particles by assuming a small fraction a (a measure of consolidation) of the matrix with solid phase continuous and by modifying the porosity

of fluid phase. The two phases exist in parallel to the flow of heat.

Nimick and Leith (1992) also used a mixture of fluid and solid continuous phases to determine of different plagioclase (fused silica). They considered discrete spheres

of fluid-continuous-phase embedded in a medium characterised as solid-continuous phase. They determined two empirical correlations for the porosity in the two phases and the fraction of fluid-continuous-phase. Another model is suggested by Nozad et al. (1985), where both the phases are continuous to account for contacts. Such structure can result, once the particulate phase starts sintering, establishing contacts between the particles resulting in solid phase

continuous. At the same time, the continuity of the fluid phase is not destroyed. This model can be used to determine the effective thermal or electrical conductivity, as the contact radius (neck growth) grows. The porosity of such a structure is given by equation (16).

where, a and c are ratios of particle size and neck size (contact) to the unit cell size. A simple equation for derived from electrical resistance analog for cubicle particles and square contacts (Fig. 14), is given below:

80

R. P. Gupta et al.

All these correlations (equations 7–17) do not consider the particle size distribution of the grains that influences the thermal conductivity of granular beds at high temperatures due to radiation. It was assumed that the thermal conductivity component is not influenced by the particle size. The effect of sintering can be accommodated only by the last three models. Research is needed to evaluate such structural parameters as function of sintering.

5.2. Contribution of Radiation at high Temperature The radiation contrbutes to the effective thermal conductivity in two ways: • radiation across the solid transparent medium • radiation across voids

The contribution of first type known as photon conductivity is usually lumped into the thermal conductivity of solid medium with zero porosity and can be determined by solving radiation and conduction simultaneously (Viskanta, 1965). This contribution is always included while determining the same experimentally for glasses and pure oxides with zero density. The contribution of second type is dependent on porosity and the pore size in case of slags and of particle size in particulate materials. The following review corresponds to the contribution of the second type. The effect of radiation is additive to the effective thermal conductivity. McAdam (1954) suggested such term as given by the following correlations:

This term does not account for the emissivity of the particles. Yagi and Kunii (1957) obtained theoretical formula for estimating the effective thermal conductivity of packed beds with different size packing. Their model considered the heat transfer by conduction across solids, across gases in the voids, conduction across contacts, by convection, and by radiation. Their correlations expressed heat transfer in packed beds with motionless gases, especially at high temperatures of upto 800 °C.

The Thermal Conductivity of Ash Deposits

81

Some models (Godbee and Ziegler, 1958, Laubitz, 1959) calculate the overall effective thermal conductivity by adding the radiative conductivity to the conduction component calculated from the unit cell. Other models (Zehner, 1970; Kunii and Smith, 1960) calculate radiative conductivity as a resistance in the network of resistances representing the unit cell. Diessler and Boegli (1958) investigated effective thermal conductivity of powders in various gases at room temperatures. Accordingly, the free convection does not influence the effective thermal conductivity of powders of The effective thermal conductivity is strongly dependent of Larkin and Churchill (1959) determined the radiative heat transfer through porous insulations. The equations account for the porosity and pore size and temperature on the radiative heat transfer. Experimental results were successfully interpreted by a simple theoretical model. Laubitz (1959) measured for several powders from 100°C to 1,000°C and was able to estimate it from the correlations He modifies (doubles) the effective thermal conductivity determined by Russell’s Equation (9) by the radiative conductivity which is a function of particle size (d), its emissivity and the porosity (p) (Equation 19).

Kunii and Smith (1960) derived equations for predicting the effective thermal conductivity of beds of unconsolidated particles containing stagnant fluid. This is a function of porosity. At high temperatures, it is also a function of emissivity, mean temperature and the particle size. The equations correctly predict the for sandstone and sintered metal systems. A new relationship for radiation contribution to the thermal conductivity was modified by Schotte (1958) as given below:

Predicted thermal conductivities resulted in good comparisons with experimental

values. The above equation assumes that the composite material of solid and void is stacked in parallel to layers of voids. Botterelli et al. (1989) tested a number of published models and compared the predictions. They reviewed the radiative correction for thermal conductivity from several researchers and presented the correction in the following form:

and the parameter c varied from one model to another model. This parameter, also known as exchange factor, includes the effect of particle size and emissivity of the pores or particles. The average pore size or particle size is very important parameter, as the radiative component is directly proportional to it. According to Botterill the models of Godbee and Ziegler (1960) and Kunii and Smith (1960) gave reasonable predictions. The

radiative models of Zehner (1970) and of Kunii and Smith gave good estimates for the thermal conductivity for packed bed of particles at high temperatures with some modifications (Nasr and Viskanta, 1994). The steep increase in the thermal conductivity in Fig. 6 in temperature range of

82

R. P. Gupta et al.

900–1,250°C is expected due to contribution from radiation. Figure 15 compares the thermal conductivity of ash deposits at 100°C and 450 °C using slag model for structure and radiative heat transfer model with exchange factor, The comparison is quite encouraging.

6. ESTIMATION OF THERMAL CONDUCTIVITY OF ASH DEPOSITS The predicted trends of thermal conductivity with structure and temperature in Fig. 11 and with porosity in Figure 15 can be interpreted on the basis of fundamental components ie conduction across solid and gas phase radiation across solid phase (if it is transparent like glass) and across voids. The estimation of thermal conductivity of ash deposits in practical situations would then involve following steps: Stepl: Estimate the thermal conductivity of zero density material at room temperature. Ratcliffe (1963) has discussed several additive formulae for estimating the thermal conductivity of glasses. As most of the deposits contain silica, alumina, calcium and iron oxide, determine the thermal conductivity by using the additive formula on the basis of volume or weight fraction. Step 2: Estimate the conduction part of thermal conductivity of zero density material at the given temperature. This information may not be available easily. However, as a first estimate, one can use values for alumino-silicate from Kingery (1976) for zero density as a function of temperature. The experimental determination of such thermal conductivity is expected to be more reproducible. This includes the radiation across solid phase in case of slag type structure. Step 3: Estimate the effective thermal conductivity of the slag or particulate material using a model for accommodating the effect of porosity (Russell’s model, equation 10) for particulate or slag structure. The effective thermal conductivity of partially sintered layer is more complex as it requires more structural parameters for the deposit. In absence of these parameters, the Brailford and Major’s model (equation 13) of random dispersion of the solid and gas continuous phases can be used. It is important to develop structural parameters such as the proportion of solid phase continuous in the partially sintered deposit as a function of sintering.

The Thermal Conductivity of Ash Deposits

83

Step 4: Take into account the effect of radiation through voids using equation (21). This effect takes into account the particle or pore size distribution.

7. CONCLUSIONS The thermal conductivity of granular or slag type deposits can be determined by correlations using two phase medium equations. At room temperature, the effective thermal conductivity appears to be independent of pore size or particle size. The influence of particle size or pore size is significant at higher temperatures when radiative transfer becomes important. Most of the equations for determining the radiative contribution indicate this transfer to be directly proportional to particle size and to the cube of the temperature of the deposit. The estimates of the thermal conductivity for a partially sintered deposit require

the knowledge of extra structural information describing pore size distribution and the connectedness of the particles. This information can be reduced to structural parameters, such as proportions of purely particulate and slag phases and their respective porosity. These parameters can then be used in the hybrid models described for complex structures. The structural parameters, would also help in providing the anisotropic nature of thermal conductivity of deposits (Ramer, 1996). There is little information available in literature on structural parameters of coal ash deposits and particularly on anisotropic nature of ash deposits. Hence it is proposed to develop the physical structural parameters as function of chemical character of ash and sintering history.

8. BIBLIOGRAPHY Anderson, D. W. et al., J. of Engg. for Gas Turbines, 109, 215, (1987). Bauer, R., Int. J. Heat Mass Transfer, 17, p4181, (1993). Boow, J. and Goard, P. D., J. Inst. of Fuel, 42, pp412, (1969). Botterill, J. S. M. et al.., Int. J. Heat Mass Transfer, 32, pp585, (1989).

Brailsford, A. D. and Major, K. G., Brit. J. Appl. Phys., 15, 313, (1964). Diessler, R. G. and Boegli, J. S., Trans. ASME, 1417, (1958).

Flynn, D. R., J. of Research of NBS, 67C(2), pp129–137 (1963). Godbee, H. W. and Ziegler, W. T., J. Appl. Phys., 37, pp40, (1966). Gupta, R. P., et al., The Heat Transfer Properties of Ash Deposits in PF Fired Furnaces, Project Report, Inst.

of Coal Res., Uni. Of Newcastle, (1997). Hadley, G. R., Int. J. Heat Mass Transfer, 29, pp909–920, (1986).

Kingery, W. D., Introduction to Ceramics (1968). Krupiczka, R., Int. Chem. Engg., 7, 122, (1967). K u n i i , D. and Smith, I. M., AICH.E.J., 6, pp71, (1960). Larkin, B. K. and Churchill, S. W., AICH.E.J., 5, pp467, (1959). Laubitz, M. J., Cah. J. Phys., 37, pp798–808, (1959). Leach, A. G., J. Phys. D: Appl. Phys., 26, 733, (1993). McAdams, W. H., Heat Transmission, 3rd Ed. McGraw Hill, New York, (1954). Mulcahy, M. F. R., J. Inst. Fuel, 39, pp385,394, (1966). Nasr, K., Viskanta, F. and Ramdhyani, S., J Heat Transfer. 116, pp829, (1994). Nowok, J. W. and Steadman, E. N., Proceedings Of the 12th Annual Gasification And Gas Stream Cleanup Systems Contractors Reveiw Meeting, R. A. Johnson and S. G. Jain Ed, Morgantown, West Virginia,

pp154–164, Sept (1992). Nozad, I., et al., Chem. Engg. Sci., 40(5), pp843–855, (1985). Ramer, F. R. and Martello, D. V., Applications of Advanced Technology to Ash Related Problems in Boilers, Ed.

L. Baxter and R. DeSollar, Plenum Press, (1996). Ratcliffe, E. H., Glass Tech., 4, 116, (1963).

84

R. P. Gupta et al.

Rezaei, H. R., et al., Proceedings of EF Conf., Impact of Mineral Impurities in Solid Fuel Combustion, Kona, (1997). Russell, J. Am. Ceramic Soc., 18(1), (1935). Schotte, W., AICH.EJ., 6, pp63, (1958). Tye, R. P., Thermal Conductivity, Vol 1, Academic Press (1959). Viskanta, R., J. Heat Transfer, pp143–150, (1965). Wain, S. E. el al., Inorganic Transformations and Ash Deposition During Combustion, Ed. S. A. Benson, ASME, (1992). Wall, T. F., Mai-Viet T., Becker H. B. and Gupta R. P., Fireside deposits and their effect on heat transfer in p. f. boilers: The emmisivity and thermal conductivity of deposits and their components., Proceedings Pulverised Coal firing— The Effects Of Mineral Matter, University Of Newcastle, L8.1–L8.16, August (1979) Wall T. F., Bhattacharya S. P., Zhang D. K., Gupta R. P. and He X., The properties and thermal effects of ash deposits in coal-fired furnaces., Prog. Energy Combust. Sci., Vol. 19, pp487–504, (1993). Woodside, W., Can. J. Phys., 36, pp815, (1958) Yagi, S. and Kunii, D., AICH.EJ., 3, pp377, (1957). Zehner, P. and Schunder, E. U., Chemie. Ingr. Tech., 42, pp933, (1970).

THE DEVELOPMENT OF ADVANCED CLEAN COAL TECHNOLOGY IN JAPAN Mineral Matter Issues

Sadayuki Shinozaki Center for Coal Utilization, Japan

Tokyo, Japan

1. INTRODUCTION As shown in Table 1, Japan’s power facility plan in 1995 was a total of 20GW, and 42 GW of coal-fired thermal power is planned by 2010, an increase of about double the

1995 level. In 1995, Japan’s coal consumption was about 130 million tons, of which 45 million tons were used for power generation, 25 million tons for industry and 60 million tons for steel-making. The area whose consumption is expected to increase in the future is steaming coal for power generation.

On the other hand, however, because of the increased concern about global environmental issues, it is pointed out that coal generates higher than other fossil fuels. Therefore, in coal utilization, it is important to develop high efficient power generation technology and utilization technology and apply them as control measures for generation. Japan is a coal importing country and its import coal accounts for about 95% of the total consumption. Japan imports coal from a wide range of countries, including Australia, the United States, South Africa, Indonesia, China and Russia, and the quality of coal (steaming coal) that Japan uses ranges from sub-bituminous coal (low melting point and rather higher Ca content) to bituminous coal whose fuel ratio is around 2.5. For example, any power plant uses 20 to 30 types of coal. Therefore, the development of coal utilization technology faces difficulties because coal-fired boilers and gasified furnaces in Japan must be designed to fit coals having different quality. The thermal efficiency of coal-fired thermal power plants in Japan is high with an average of 38%, and the thermal efficiency of new coal-fired thermal power plants of 1,000 MW class use USC technology reaches 41%. Regarding environmental characteristics, on the other hand, less than 50ppm of SOx and NOx and less than 10mg/Nm3 of dust have been achieved by the introduction of the high performance flue gas desulfurization system Impact of Mineral Impurities in Solid fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

85

86

S. Shinozaki

(wet method), low NOx combustion technology, and denitrification system (SCR). Therefore, coal-fired thermal power plants can be said to be high efficient and excellent in environmental characteristics even today, but the development of further clean and high efficient coal utilization technology is in progress to create better global environment.

2. DEVELOPMENT OF CLEAN COAL TECHNOLOGIES Because Japan promised to stabilize its generation at the 1990 level in “Frame Work Convention on Climate Change” it carries out “Higher Efficiency Power Generation Technology Development” under the CCT development target as shown in Table 2. The outline of the CCT development is introduced in this report.

3. OUTLINE OF CLEAN COAL TECHNOLOGIES DEVELOPMENT THEMES Table 3 shows the development system of advanced clean coal technology in Japan. Of the development themes of advanced clean coal technology, themes that are

The Development of Advanced Clean Coal Technology in Japan

87

unique and typically affected by the behavior of mineral matters are outlined and introduced. As the concept of development themes indicate, in most themes, clarifying behavior of and measures for mineral matters are important factors for the continuous and stable operation as well as scale up of plants.

3.1. Pressurized Fluidized Bed Combustion Technology3 Establish a technology to commercialize a pressurized fluidized bed boiler with economic superiority by carrying out a series of demonstration test. As compared to conventional pulverized coal fired boilers, the boiler will have distinctive features as given below.

• Scope of plant 71 MW demonstration plant (1) Development issues 1) Demonstration assessment test • Grasping of environmental and combustion characteristics • Demonstration of high efficiency and economic advantages • Verification reliability • Demonstration of optimum operation • Solutions to problems of the behavior of mineral matters (2) Development targets 1) Improvement in thermal efficiency by combined cycle power generation Net thermal efficiency: 42 · 43% (Pcf: 38·40%) Combustion efficiency: 99% over

88

S. Shinozaki

2) Environmental characteristics SOx: 30ppm Max. NOx: 150ppm Max. (30ppm Max. at outlet of stack) Dust: Max. (3) Subjects related to Mineral behaviors • Development period: April 1992 to March 1998

3.2. Pressurized Internally Circulating Fluidized Bed Combustion Technology In order to cope with global environment, research will be carried out on pressurized fluidized bed boiler with high reliability and excelling in load responsibility by using the characteristics—corresponding to advanced low SOx and low NOx environ-

The Development of Advanced Clean Coal Technology in Japan

89

ment type—of revolving floe type internally circulating fluidized bed that can burn many kind of coal grades at high thermal efficiency. • Scope of 4 MW Pilot plant (1) Development issues 1) Verification of higher efficiency by 4MW pilot plant The study of reliability and operatability of plant with high environmental characteristics 2) Technology coping with variety coals 3) Verification of load capability 4) Verification of environmental characteristics 5) Solutions to problems of the behavior of mineral matters 6) Scale up technology • Development targets 1) Thermal efficiency target at 150 MW Trial design Gross thermal eff.: 42% Net thermal eff: 40.74% 2) Environmental characteristics and system design NOx: 30 ppm Max. CO: 50 ppm Max. Desulfurizing rate: 95% (1% S coal: about 40 ppm) 3) Load capability 2.5·3.0%/min at MCR 100% • Subject related to Mineral behaviors • Development period: April 1992 to March 1998

S. Shinozaki

90

3.3. Advanced Pressurized Fluidized Bed Combustion Technology In order to reduce a global environmental problem, the development of highly efficient power generation technologies is a matter of utmost urgency. To that end, the fluidized bed gasification technology that were researched in the past in combination with the pressurized fluidized bed technology which is under research at present. The research is aimed at the development of highly efficient power generation technology by next generation technology.

• Scope of plant: 5 MW, PDU (1) Development issues 1) Establishment of hot desulfurization system 2) Turning CaS into harmless substance 3) Checking the controllability of system and stabilized operability 4) Multi-coat utilization technology (Possible to apply to coal grades with high melting point) 5) Solutions to problems of the behavior of mineral matters 6) Process control technology 7) Scale up technology (2) Power generation efficiency: (Trial calculation values by 409 MW trial design)

Gross thermal: 49.7% Net thermal: 46.2%

The Development of Advanced Clean Coal Technology in Japan

91

(3) Environmental characteristics at desulfurizing furnace outlet) dust: (4) Carbon conversion Calorific value of gas Unburnt • Subject related Mineral behaviors. • Development period: April, 1996 through March, 2001

3.4. Pressurized CPC Combustor Technology Development of a coal partial combustion furnace to convert boilers for general industrial use to coal, while establishing an application system technology for boilers of coal partial combustion furnace, and also develop pressurized CPC combustion technology that can be applied to highly efficient conbined cycle power generation.

• Scope of plant: 25 t/d pilot plant ( 1 ) Development issues 1) Development of pressurized CPC furnace 2) Establishment of high temperature dedusting technology 3) Solutions to problems of the behavior of mineral matters 4) Demonstration of continuos operability—One month continuos and stable operation

S. Shinozaki

92

5) Expansion of applicable coal grades To make it possible to utilize a wide range of coal types from low to high fuel ratios. 6) Collection of scale up data: 25t/d . 200t/d 7) Scale up technology (2) Development Targets 1) Gasification performance under pressure (air blow) Generated gas Cold gas efficiency: 69% (pilot), 72% (actual plant system) 2) Collection of oxygen rich combustion data Generated Cold gas efficiency: 73% (pilot), 80% (actual plant system) 3) High temperature dedusting technology Outlet

4) Power generation efficiency: 45% (3) Subject related to Mineral behaviors • Development period: April, 1990 to March, 1999

3.5. Coal Gas Production Technology for Fuel Cell (IGFC) Aimed at the establishment of a technology to manufacture gas for supply to the production of fuel cells, conduct tests on a total system of oxygen blown coal gasification to develop an optimum coal gasification furnace for fuel cells together with a purifying system capable of generating gas of composition best suited for fuel cells (MCFC and SOFC) • Scope of plant: 150t/h pilot plant ( 1 ) Development issues 1) Development of an optimum gasification furnace

The Development of Advanced Clean Coal Technology in Japan

93

2) Establishment of technology for purifying product gas of composition best suited for fuel cells Highly efficient dedusting technology High performance gas purifying technology 3) Demonstration of high efficiency 4) Solutions to problems of the behavior of mineral matters

(2) Development Targets Reflect at-site investigation and research results on the fundamentals and detailed design of pilot testing facilities (150 t/d ) 1) Determine the type of coal gasification furnace 2) Carbon conversion rate: 98%min Low gas efficiency: 78% Gas (3) Predicted technological targets of a commercial plant 1) Power generating efficiency (Gross thermal: 50%min) 2) Gas purity Sulfur compounds: 1 ppm max Halogen compounds: N • D (0.1 ppm max) Trace element: N • D (0.1 ppm max) • Development period: March, 1995 to March, 2005

3.6. Fluidized Bed Cement Production Technology A cement calcination technology by the fluidized bed system which is completely different from the present kiln system. The characteristics of fluidized bed technology

94

S. Shinozaki

is utilized to aim at the development of a highly efficient environmentally-friendly manufacturing technology, making it possible to lower energy consumption, reduction in the consumption of finishing mill pulverizing power, lower NOx, improvement of combustion efficiency, expanding the usage of coal grades, and making it possible to manufacture a greater variety of cement products. • Scope of plant: 200 t/d scale-up plant (1) Development issues 1) Establishment of stable operation and control technology 2) Cement calcination technology that can also address cement-related

products and diversify cement production 3) Development of coal combustion technology that can handle a large

variety of coal grades Calorie: 5,000kcal/kg, Fuel ratio: 8–10 4) Collection of scale-up data 5) Scale up technology (2) Development Targets 1) In order to address the global environment, the kiln type was compared to develop a cement calcination technology based on the following.

—Energy saving type: Heat consumption, 750–760 kcal/kg clinker (Reduction of approximately 10–12% as compared to the kiln type) • Heat efficiency: 54–55% • Low type: reduction of approximately 10% • Low NOx type: 200 ppm max

• Pulverizing power consumption: Improvement by 6–7% • Production cost: Development of technology being capable of reducing production cost by about 15–18% • Development period: From 1987 through March 1998

The Development of Advanced Clean Coal Technology in Japan

95

3.7. Metal Melting System by Oxy-Coal Combustion The metal melting technology is different from the present electric furnaces. Energy obtained at high temperatures by burning pulverized coal directly with oxygen is used for melting scrap aluminum, copper, iron, and other metals. The development aims to enhance recovery by improving overall energy efficiency and raising concentration. The technology is useful in lowering the melting cost. Therefore, it will be completed as a system technology that is capable of protecting the global environment. • Scope of plant: 5t/charge pilot plant (1) Major development issues 1) Continuous manufacturing technology 2) Optimization and automation of overall process 3) Extended life of refractories 4) Collection of scale-up data (2) Performance targets — Heat efficiency: Final target (Commercial plant: 60%) — Yield: Fe: > 94% Cu, Al > 96% — NOx: < 180ppm —SOx: Less than emission standard of K value 3 region —Dust: < 0.2g/Nm3 (Approximately 1/10 of electric furnace —CO2 concentration of emission gas: 90% min, easy to recover • Development period: April, 1992 to March, 1998

3.8. Multipurpose coal conversion utilization technology— Coal Flash Pyrolysis With a view to assure the availability of energies in Japan in the area of expected petroleum shortage in the 21st century and beyond, and to address the global environmental problems, develop a technology to manufacture at high efficiency gas of intermediate calorie (approximately by flash pyrolysis of carbon at comparatively low temperature, and contribute to the construction of a foundation for coal chemical in the future. • Scope of plant: 100t/d pilot plant (1) Development issues 1) Development of optimum char gasification and pyrolysis furnaces 2) The grasping of flow characteristics of pulverulent body and therma equipment that are involved with the scale up of pyrolysis and char gasification furnaces. 3) Performance check of optimum gas production process 4) Establishment of stable operation technology (2) Development targets 1) Total heat efficiency: 85%min 2) Co-gas and liquid production yield: 70%min 3) Production gas calorie: 4) Product tar: Raise the recovery rate of acid tar. • Development period: March, 1996 to March, 2001

96

S. Shinozaki

The Development of Advanced Clean Coal Technology in Japan

97

3.9. Production Technology of Coal Advanced Conversion Cokes Due to obsolescence of existing coke ovens, development of a new coke manufacturing technology which excels in environmental characteristics such as smokeless, technology odorless, low NOx emission, and low dust production, high productivity, and allows the use of general coal. • Scope of pilot plant: 12t/d (1) Development issues

1) Coal pretreatment technology 2) Plug transportation technology of coal particles under high temperatures 3) Coke improving technology at low to medium temperatures—optimum design of coking chamber 4) Establishment of new coke manufacturing technology 5) Summary of effects on heat efficiency, economic advantages, and working coal grades (2) Development targets 1) General alternating rate: 50% 2) NOx emission: 150ppm max 3) SO2 emission: 10% reduction 4) 5) Energy saving: 20% reduction 6) Productivity: 3 times that of conventional coke ovens • Development period: April, 1994 to March, 2002

4. EXAMPLE OF HINDERED STABLE PLANT OPERATION AND DEVELOPMENT DUE TO MINERAL MATTERS IN COAL The major CCT development condition in Japan is described above, and the reaction behavior of combustible organic compounds has been clarified to a certain extent in the development and plant operation stages and has been coped with through optimum design of plants.

However, as shown in Table 4, a large portion of the reaction behavior of mineral matters in coal under high temperatures and high pressure is still unknown, causing difficulties with stable plant operation and scaling up in advanced CCT development. On the other hand, plants in Japan must effectively utilize the discharged coal ash, and because the chemical and physical properties of coal ash differ for each coal lot and for each boiler design and operating condition, the performance of coal ash varies, therefore forming one of the factors that significantly hinders effective utilization of coal ash when it is used as civil engineering material. Table 4 gives examples of difficulties for stable operation of coal utilizing power plants due to mineral matter in coal. Particularly, because Japan is a coal importing country and each plant must use many types of imported coal, unlike on-site plants near origins of production, trouble due to mineral matters seems to occur more frequently and clarifying the behavior of mineral matters tends to be more difficult. Figure 9 shows problems for studies on reaction behavior of coal ashes under high temperatures that are assumed from the examples of hindered stable operation due to ashes described above.

98

S. Shinozaki

The Development of Advanced Clean Coal Technology in Japan

99

S. Shinozaki

100

5. IMPORTANCE OF STUDIES ON CLARIFYING BEHAVIOR OF MINERAL MATTERS IN COAL As explained, mineral matters show many patterns of unknown reaction behavior under high temperatures, pressurized and reduced conditions. As reaction behavior of mineral matters in coal is presently being clarified by the

Brain-C Program of Japan, conclusions have not been reached. But types of main mineral matters in coal (refer to Table 5). and reaction behavior under high temperatures are roughly expected as shown in Fig. 10. On the other hand, the difference in chemical and physical properties of fly ashes generated from coals containing ashes with high melting point (herein after referred to as coal with high melting point ash) and coal containing ashes with low melting point (herein after referred to as coal with low melting point ash) was surveyed in order to assume the reaction behavior of mineral matters in pulverized coal-fired boilers, and the results are shown in Table 6, and Fig. 11 shows examples of SEM photos. Of the SEM photos, J-05 of Table 5 presents domestic coals (low melting point), and it is assumed that the shapes and physical property of fly ashes generated from different types of mineral matters vary significantly. Table 6 shows that fly ashes discharged from coals with low melting point have

small average radius and higher content of glassy matters (Ca-Al-silicate) than that of coals with high melting point. Applying micro analysis to these ashes will useful for the clarification of reaction behavior. Fly ash from PCF boiler (Domestic Coal)

Fly ash from PCF boiler (Imported Coal)

Fly ash from FBC (Imported Coal)

In the aspect of chemical properties, fly ashes from coals with low melting point contain larger amount of CaO compared with that in fly ashes from other coals. This may be due to the difference in the content of Ca compounds that is originally contained in coal, and Ca compounds are responsible for the generation of glassy matters. As a result, the strength of mortar when adding ash with low melting point to cement is greater. However, since slagging may occur depending on the design conditions (heat load, fluidizing condition of particles, etc.) of a boiler when a melting point is low, melting point design for coals fitting for the design conditions of each boiler will be necessary in order to obtain fly ashes having high utilization value for the purpose of preventing slugging.

However, it is quite difficult to use only the optimum coal in the current condition for Japan which must use many types of imported coal. The reaction behavior of mineral matters will be clarified accurately by correlating the data on mineral matters in coal and the micro analysis data on fly ash, eventually contributing to the establishment of boiler design and optimum operating conditions for slagging prevention and obtaining fly ashes with high utilization values. It will also be useful to the design of boilers and gasifiers and scaling up technology for advanced power generation if research progresses.

6. CONCLUSION Although high efficient power generation technology and utilization technology of coal are important to control generation, design of boilers and furnaces will take high temperature into consideration for economy and high efficiency.

The Development of Advanced Clean Coal Technology in Japan

101

102

S. Shinozaki

The Development of Advanced Clean Coal Technology in Japan

103

Mineral matters contained in coal differ depending on their types, namely composition and origin of production. Therefore, they may cause ash troubles unless they fit the design and operation conditions of boilers and furnaces. This represents obstacles to the development of advanced coal utilization technology. In the case of countries like Japan where diversed types of coal must be used, the

design of boilers and furnaces that are capable of handling many coal types becomes important for the stable operation of plants. Since there are still many unknown factors in the reaction behavior of mineral matters in coal, clarifying reaction behavior of mineral matters under high temperatures

and pressure is quite important for the stable operation of plant as well as for the smooth scale up of technologies under development, and in Japan, both public and private sectors

have jointly worked to clarify the behavior and to create simulation models. However, we are still in the research stage and unable to report the conclusion in this report. As the clarification of reaction behavior of mineral matters is considered as a common subject to many countries, it seems effective to clarify the reaction behavior of mineral matters with the cooperation of participating countries. Therefore, participating countries have planned a research scheme and proposed this to the technical meeting of IEA-CCS.

We hope to clarify this matter soon as the clarification of reaction behavior of mineral matters will be useful in many areas and will help promote global environmental measures that are common interest of all people.

7. REFERENCES 1) “Outline of Power Development,” Public Project Department of the Ministry of International Trade and

Industries, Oct. 1996. 2) Unpublished material, EPDC technical material. 3) “Proceeding of the Rheology Symposium,” CCUJ, Feb. 1997. 4) “Coal Fundamental Property Section of 1996 Report on Coal Utilization Fundamental Technology Development and Research,” CCUJ/NEDO, unpublished data, Mar. 1997.

This page intentionally left blank.

FACTORS CRITICALLY AFFECTING FIRESIDE DEPOSITS IN STEAM GENERATORS Richard W. Bryers Retired Flemington, New Jersey 08822

INTRODUCTION Fireside slagging, fouling, and corrosion due to impurities in steam-raising fuels has been under investigation for over a hundred years. From the very beginning impurities in steam-raising fuels has been responsible for costly maintenance problems, a reduction in heat transfer, derating, costly alterations or the use of more expensive fuels. For

the first eighty years scientists and engineers have taken an empirical approach to understanding and dealing with these problems. Although complex models have been proposed for various phases or steps in the deposition process, it was not until the development of sophisticated analytical techniques for characterizing fuels and the expanded use of the computer in the late seventies and early eighties that comprehensive modeling was attempted. Consequently a large data base critically outlining the various steps in the complex process of deposition already existed awaiting development of more comprehensive economic models. Fouling and slagging are very complex phenomena which depend upon the release of inorganic constituents in fuels during combustion and their transformation into troublesome compounds during the post-combustion heating and cooling process. Fouling, slagging, and corrosion also depends upon the juxtaposition of individual inorganic ash-forming species during comminution, combustion, and quenching while recovering heat; the chemical reactions between gas, liquid, and solid phases in motion and at rest; the kinetics of transformation of minerals and fly ash; the existence of nonequilibrium conditions frequently associated with super-cooling; and the attachment of impurities to surfaces and the detachment or re-entrainment of deposited solids (1). Figure 1 is a flow diagram illustrating these changes prior to deposition or emission from the steam generator. The complexity of fouling and slagging is compounded by the fact that fireside problems cannot be simply represented by a single rate of deposition on a target by ash characterized by a single elemental analysis. The inorganic components—minerals and organically associated cations—in coal and the fly ash it generates are heterogeneous with Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

105

106

R. W. Bryers

regard to size and composition. Therefore, individual species behave differently during combustion and their subsequent flight through the steam generator. The degrees of fouling and slagging vary throughout the steam generator depending on local gas temperatures, tube temperatures, temperature differentials, gas velocities, tube orientation, and local heat flux. Furthermore, fireside problems manifest themselves in a variety of ways, some of which include:

Factors Critically Affecting Fireside Deposits in Steam Generators

• • • •

plugging of slag taps formation of agglomerates and clinkers burner eyebrows loss in furnace absorption due to excessive accumulation—be it molten or sintered deposit or due to a change in thermal properties of the slag • damage due to excessive slag fall from a specific location or type of heat exchange surface and it’s orientation • undersize furnace hopper due to a weak slag-to-tube bond, causing excessive sloughing of sintered ash • freezing of slag on hopper slopes • slagging of the heat recovery zone due to unpredictable low melting phases in the ash • fouling of convective heat recovery surfaces by condensation of volatile fumes or the sulfation of vulnerable submicron species at the tube surface • tube wastage due to flame impingement, the presence of CO due to striation of flue gases as a result of poor mixing, or excessive concentration of corroding elements in the fuel • surface oxidation and exfoliation of inside tube surfaces due to high gas temperatures created by excessive fouling • slagging and/or fouling due to the size distribution of the fuel • slagging and/or corrosion due to imbalanced air distribution • pluggage of gas passes due to excessive accumulation of fly ash in zones of low gas velocity • formation of agglomerates in fluidized beds and clinkers on grated (1,2). Some of these problems may be resolved by altering operating conditions, modifying surface or improving fuel handling techniques while others require a more thorough understanding of the deposition process. The wide variation in type, form, and location in which a deposit may form indicates slagging and fouling are very much dependent upon design and operating parameters. Boilers of identical design firing the same fuel have been reported to encounter very different slagging and/or fouling conditions, suggesting the deposition process is sensitive to operating procedures. Consequently slagging, fouling, and corrosion are also dependent upon design and operating parameters. These include:

1. Design a) furnace exit temperature b) furnace absorption c) furnace configuration d) burner arrangement e) burner size f) tube size, spacing, orientation, and temperature g) air distribution—steam conditions 2. Operation a) coal size b) air distribution between burners c) burner operation d) excess air, level e) flame impingement

107

108

R. W. Bryers

f) soot blower operation g) boiler load Since economics dictate that steam generators must be designed to minimize capital cost, operating cost and maintenance costs, with maximum efficiency and fuel flexibility, the philosophy used to characterize fuels must change from one of predicting the slagging or corrosion potential of fuels based on elemental composition and fluid temperature of the ash to one permitting the selection of a matrix of engineering parameters that will allow proper location and orientation of heat transfer surface in the combustion and heat recovery process and the cost of doing so. The option of selecting an alternative fuel as a means of solving a fireside problem is rapidly disappearing. Increased trading activ-

ity with fuels will require increased flexibility and a need for rapid assessment of the cost of changes. Empirical indices already in place are limited to the conditions from which they were devised. Mechanistic models must be developed to expand or eliminate these restrictions of applicability to accommodate changes in loads and blending of fuels. The introduction of very sophisticated microanalytical techniques such as the scanning electron microscope complete with microanalytical capabilities and the computer, has made it possible to characterize fuels and their impurities in greater depth. Models may now be developed employing the analytical techniques along with engineering design parameters and configuration to reflect all aspects of ash deposition.

FUEL CHARACTERIZATION Although the impurities in steam-raising fuel include most of the known elements only 15 occur in sufficient quantities to contribute to slagging, fouling, and/or corrosion. Fortunately, these elements may be categorized by the role they play in fireside problems,

by the fuel type and the rank as shown in Table 1. This significantly reduces the number of elements one must deal with for any given problem. In addition to those listed in Table 1, barium and arsenic have been known to contribute to fouling problems on rare occasions. All but aluminum contribute directly to fireside slagging, fouling, and/or corrosion. Aluminum tends to inhibit the deposition process by raising the melting temperature of the ash. Aluminum may also simply be entrapped in deposits as innocent aggregate by virtue of its high concentration in the coal. Historically fuel impurities have been characterized by the elemental composition of a sample of coal completely ashed at 1,500°F (815°C). Traditionally the elemental composition is reported as the elements oxidized to their highest oxide state and in some cases, normalized to 100%. 100% closure may not always be achieved due to unaccounted for elements, excess residual carbon, elements in the ashed coal not completely oxidized to the highest state and loss of some elements during the ashing process. Based on a large data base of prior fireside problems, elemental analysis alone may provide a good starting point for predicting the potential for slagging, fouling or corrosion. In fact, fuels are purchased and used, and steam generators are bid, designed and constructed today on the basis of the elemental composition of high temperature ash fuel analysis and the estimated deviation or limits imposed on the individual elements alone. This simple evaluation, however, leaves room for substantial error due to unaccounted for behavior of the compounds these elements may form. Fireside problems with few exceptions are associated with a molten phase formed as the result of a low melting compound or condensation of a vapor phase. Consequently, the melting temperature of the high temperature ash from a fuel sample has been used

Factors Critically Affecting Fireside Deposits in Steam Generators

109

as a second criteria for the slagging or fouling potential of the impurities in fuels. The ASTM ash fusion temperature procedure was accepted in 1924. Since then seven other procedures have been accepted worldwide. Details of the terms and procedures have been described in an ASME publication prepared by Hough, Sanyal and Davis (3). It would seem reasonable that a correlation exists between the melting temperature and the ash chemistry. Consequently, numerous investigators have tried to interpret the ash composition and its melting temperature by plotting the data on phase diagrams, or developing fusibility diagrams. Such correlations provide a good test for the validity of the individual analysis and provide a basis for what must be done to alter the melting temperature of the ash and thereby reduce its potential for fireside problems. Others have developed empirical indices using elemental composition or ash fusion temperatures as independent variables for establishing the limit to trouble free operation. Winegartner assembled a classic lexicon summarizing these indices which are still used commercially today (4). This empirical approach, unfortunately, assumes that coal ash is homogenous and remains intack during combustion. As we shall see, this is not an accurate assumption in many cases. For specific designs such as a wet bottom furnace, or cyclone where it is essential that molten conditions be maintained, viscosity of the ash was selected as the most appropriate property for characterizing the impurities in the fuel. Since viscosity measurements are tedious and expensive, numerous correlations were developed on the basis of temperature and elemental composition. Viscosity as a property of molten ash is not only important to fuel selection of specific steam generator types where molten ash is essential, but also plays a significant role in the growth and size of deposits that may be initiated by the sintering process. Viscosity is also being used as a property for predicting the

sticking potential of fly ash which is either molten or encapsulated by a wet surface. Consequently, it is essential to understand the viscous behavior of fuel ash.

Corey reported that completely liquid slag behaves as a Newtonian fluid and that the rate of change in viscosity with a change in temperature at a given viscosity was essentially constant (5). Consequently, the functional relation of viscosity vs. temperature in

110

R. W. Bryers

the Newtonian range is the same for all coal slags. By making an appropriate parametric correction for composition, the viscosity of all slags can be represented by a single curve. A typical curve for velocity vs. temperature appears in Fig. 2. Reid indicated slags may be categorized into three different types according to their behavior as they are cooled below the crystallization point, also illustrated in Fig. 2 (6). Reid shows that slags rich in silica appearing as quartz remain glassy over the entire viscous temperature range. No divitrification occurs and viscosity depends only on temperature, not the previous thermal history. In a second group of slags, a solid phase enriched with alumina calcium or iron precipitates out abruptly at some given temperature in the cooling process forming a pseudoplastic fluid. Upon reheating, the solid phase is gradually redissolved, but at a much higher temperature than that at which the crystal was formed. The temperature difference between crystal formation and crystal melt depends on cooling rate. At high rates of cooling the freezing temperature is considerably below the critical melting temperature. Corey indicates the transition from Newtonian to plastic flow is due to the formation of crystals in the slag and corresponds to the softening temperature of the ash. A third group includes those slags in which the crystal precipitation, upon cooling, occurs at almost the same as the temperature as the crystals dissolve upon heating. Reid’s observation of a potentially large difference between melting temperatures of ash upon heating and solidification temperatures upon cooling is very important as ash is characterized for

its slagging potential while being heated while fly ash is deposited under a rapid quenching process. Consequently ash may deposit at much lower temperatures than predicted from fuel ash analysis.

Factors Critically Affecting Fireside Deposits in Steam Generators

111

Kalmanovitch and Williamson examined the divitrification of eastern and western North American coal-type ashes quenched over a period of several hours. Their results show that the crystallization is well represented by the quarternary. (7) Huffman and Moza, however, both using somewhat different techniques but much higher quenching rates, found that freezing or solidification temperatures were 200– 400°C below the initial deformation temperature occurring during the heating cycle (8,9). Huffman's tests were performed in a muffle furnace whereas Moza’s experiments were performed in a drop tube. The solidification temperature rather than melting temperature of ash is a critical factor influencing fireside problems. Presently there are at least five expressions for viscosity in terms of temperature and the five major constituents of coal ash. All are claimed to be accurate over a given

range of slag composition. None of these formula, however, include modest amounts of the alkalis believed to be influencing the surface fusibility of fly ash and the fusibility of furnace wall slag formed by high sodium-bearing lignites. A plot of viscosity using the

Urbain and Watt-Fereday equations applied to various gravity fractions of a bituminous coal in Fig. 3 indicates substantial disagreement between correlations as iron concentration increases. This suggests that correlations developed to reduce time and cost of determined viscosity should be verified for slag whose composition fall out of the range used to derive the correlations. The discussion on viscosity also suggests that the liquid phases may be present at temperatures much below those predicted by melting temperatures for fuels containing significant quantities of free quartz. Consequently, biomass fuels should be much more vulnerable to slagging and conventional techniques used to assess slagging potential of coal may not apply for biomass.

112

R. W. Bryers

MINERAL COMPOSITION The three fundamental properties used to characterize fuel ash are all based on a composite sample assuming homogenous distribution of impurities. In reality, however, impurities in steam-raising fuels exist as minerals and mineral matter each of which is composed of several elements which have been oxidized, chlorinated, sulfated, or organically bound. Minerals refer to the completely inorganic constituents whereas mineral matter refers to the organically bound inorganic elements. Anyone familiar with coal cleaning operations, for example, will clearly recognize that minerals are heterogeneously dispersed throughout coal by size and juxtaposition with regard to the coal and other mineral species. Each mineral or group of minerals contributing to a single ash particle has its own physicochemical properties and goes through its own peculiar transformation, thereby subjecting the steam generator to a wide variety of ash particles with

Factors Critically Affecting Fireside Deposits in Steam Generators

113

different physicochemical properties. Further partitioning of the mineral species may occur during the communition process depending upon the fuel, its hardness, moisture, and particle size distribution. The final degree of variability in fly ash properties will depend on the initial size distribution of the minerals in the fuel and the size reduction operation. There are about 125 minerals found in coal. Twenty-five are reported in Table 2 (10). Nine have been recognized as occurring in significant amounts. The remaining 14 are listed as accessory minerals. During the process of coalification, the minerals have been altered by time, heat, moisture, and pressure producing a variation in mineral composition by coal rank already implied by the variation in elemental behavior in Table 1. An examination of the genesis of coal in Fig. 4 indicates the elemental composition of coal and the minerals within which they are contained is dependent on the original peat swamp and the age of the coal. As the coal ages pressure and temperature convert the

114

R. W. Bryers

organometalic compounds, sulfides, sulfates, and carbonates into oxides associated with the silicates. Consequently, each coal bed has its own characteristic mineral distribution by size and composition which is influenced by its age and the surrounding geology. Coal beds are formed over long periods of time representing various periods in the history of the peat swamp. Therefore one should expect to encounter a variation in mineral composition by layer or lateral location in a fuel bed. In extreme cases one might encounter a variation in coal rank. The largest difference in mineral composition occurs by rank. Coals of similar rank generally contain minerals of similar composition and fireside behavior. The combustion characteristics of the fuel are also altered by age. With increased coalification all macerals become enriched with carbon, eventually forming anthracite. As coals age they become harder, contain less volatile, contain less moisture, and are more difficult to ignite. These factors all influence the steam generator size and configuration and have a direct impact on the final mineral composition of the fuel with regard to quantity and composition. Variability in the mineral composition of coal due to variability in mineral composition at the mine, unexpected inclusions and mining procedures are a major source of slagging and fouling problems. Wood and biomass, which are precursors of the peat swamp, are the primary source of the organometallic and water soluble salts in coal. A typical analysis for wood and biomass appear in Table 3. Wood is low in ash which is enriched with calcium and potassium. Biomass, which includes grasses, shells, pits, and manure, etc., is also low in ash

but rich in silica, calcium, potassium, and occasionally, phosphorous. All but the silica

are ionically bound to the plant life. Moisture level is high in both fuels resulting in a low flame temperature. Both fuels are most frequently fired on spreader stokers and more recently in fluidized beds. Petroleum and petroleum coke derivatives originate from decayed animal life and

Factors Critically Affecting Fireside Deposits in Steam Generators

115

are frequently associated with sea beds. Thomas identified as many as 25 elements intrinsic to the crude (11). The principal ash forming constituents were reduced to 9 and tabulated by Bowden in Table 4 (12). Both inorganic and organically-bound oil soluble forms have been observed for each element. The mineral matter is not destroyed during processing and thus appears in heavy oil in the same form. Sodium is not intrinsically found in crudes. Consequently occurrence in oil must be due to intrusion during shipment.

However, it should not exceed 60 PPM if properly desalted at the refinery. Asphaltic base crudes are generally rich in vanadium, particularly Venezuelan oil, which often contains 900 PPM reported as Paraffinic base crudes are usually free of vanadium. Petroleum derivative fuels such as fluid coke, delayed coke, etc. contain the same mineral forms as the crude but in much higher concentrations. Huffman, using X-ray absorption nearedge structure analysis, has confirmed that the vanadium porphyrin are not destroyed in the coking process (2). Nickel concentrations may be somewhat higher in cokes than crude oils due to contamination by cataylst. Petroleum products are moisture free and burn with very high flame temperatures. The rates of combustion of the liquid and solid fuels are decidedly different as is their ignitability. Consequently, they are fired in very different furnace configurations. Oil is fired in a conventional front wall furnace with a very high heat release rate where as petroleum coke is fired in an arch-fired anthracite type furnace. Municipal solid waste (MSW) is a man made fuel composed of all types of waste products from human consumption including grasses, leaves, and food materials. The impurities in MSW fuels include all the elements found in coal plus modest amounts of zinc and lead. Vanadium fortunately is not found in significant amounts. The ash is highly basic and hence resembles the ash from lignite enriched with sulfur. Moisture levels are high and the fuel is most frequently fired on a grate in a furnace lined with refractory. Numerous techniques are now available for characterizing minerals in samples. Xray diffraction performed on low-temperature ash is a powerful tool and the most widely used technique for qualitatively identifying the presence of minerals in their crystalline form in concentrations of a few weight percent or greater. Thermal analytical techniques such as differential thermal analysis (DTA) and thermogravimetric analysis (TGA)

116

R. W. Bryers

have been used as a signature analysis based on changes in physical properties with temperature. Microanalytical techniques including scanning electron microscopy (SEM) equipped with energy dispersive x-ray (EDX), electron probe microanalysis equipped with EDX, and transmission electron microscopy (TEM) and scanning transmission microscopy have very good resolution to very small particle sizes. The techniques may be applied to low temperature ash as well as raw coal and provide visual images of the

morphology of the mineral structure. Extended x-ray analyses (XANE) is another signature analytical technique recently applied to coal and petroleum coke. Mossbauer spectroscopy has been widely used for characterizing the mineral forms of iron in coal as well as in slags and deposits. Computer automation of the scanning electron microscope (SEM) and energy dispersive x-ray (EDAX)–commonly referred to as a computer controlled scanning electron microscope (CCSEM)—has been developed to locate dispersed inorganic particles in coal and quantify their size, shape, and composition. When coupled with automated digital image analysis, morphological data can be stored, examined, and used to determine differences between the included and excluded mineral matter as well as species sharing partially common boundaries i.e. juxtaposition of minerals with coal. The two systems may be applied to raw coal, partially spent char and fly ash or deposited ash, in effect revealing changes in mineral chemistry at various stages of combustion. Figure 5 illustrates the distribution of minerals in a coal sample as determined by CCSEM (13). Presently these techniques, although commercially available, require the skills of a

highly trained scientist and are expensive. The services for determining analysis via these techniques are provided by only a few laboratories, limiting their availability. Widespread practical application requires more extensive demonstration of their utility in predicting and solving problems. Their availability, however, is a critical factor in the transformation of slagging fouling and corrosion from an empirical art to an applied science. Chemical fractionation, a technique developed by Miller and Given and put to practice by Benson and Holm, uses selective extraction of elements based on solubility which reflects their association in the coal or biomass (14,15). The process consists of three successive extractions, first by water to remove water soluble elements such as

Factors Critically Affecting Fireside Deposits in Steam Generators

117

sodium in sodium sulfate which was most likely associated with ground water, second by ammonium acetate to remove elements such as sodium, calcium, and magnesium that

may be associated as salts of organic acids and finally with 1M HCl to remove acid-soluble species such as iron, magnesium, aluminum, and calcium in the from, of hydroxides, oxides, carbonates and organically coordinated species. Table 5 illustrates the application of the widely accepted technique to coal and biomass. The elemental concentration of the fuel ash is insufficient to predict fireside problems. The mineral form is critical in determining the type and extent of fireside ash deposition.

TRANSFORMATION OF MINERAL MATTER DURING COMBUSTION During combustion, the inorganic elements in coal are transformed into inorganic gases, vapors, molten particles and solids depending upon their association in the fuel as included or excluded minerals, their juxtaposition with other mineral grains or their occurrence as organically bound species such as sulfur, chlorine, and the alkalis, or salts such as the alkalis, heavy metals, and chlorine. The excluded minerals include mineral grains released during size reduction (including minerals found in small partings, fractures, and fissures) and dirt picked up during mining, storage, and fuel handling. The excluded minerals generally constitute 90% of the ash in high rank coals and contribute primarily to the coarse size fractions of the fly ash. In low rank coals the mineral grains

118

R. W. Bryers

may constitute 60% or less of the total inorganic content of the coal. These mineral grains will under go further size reduction and partitioning during combustion attributed to ash shedding, char fragmentation, mineral fragmentation, decomposition and volatilization. The sulfides and carbonates decompose while free quartz and calcite will partially vaporize under reducing conditions. The organically bound cations decompose and form gases or a vapor which subsequently are oxidized, sulfated, or chlorinated, depending upon the levels of concentration of sulfur, chlorine, and oxygen. These elements form noncondensable gases, condensable vapors, or sublimated super-fine particles which interact with entrained fly ash, the tube surface, deposit surface or are discharged from the steam generator. Each mineral specie undergoes a unique thermal transformation as illustrated in Figure. Melting temperatures for all but illite biotite and pyrite generally exceed 2,552°F. (1,400°C.) Ulrich et al. examined the transformation of volatile species generated from minerals during combustion under reducing conditions using JANAF tables to predict vapor equilibrium compositions at high temperatures under varying degrees of oxygen (16). They found that substantial amounts of silicon, aluminum, iron, calcium, magnesium, sodium, and potassium appear as pure metal cyanides depending upon oxygen levels. The volatile species are formed at the carbon flue gas interface. The degree to which they are released is diffusion controlled. Vapors are oxidized and precipitated during the

migration process forming submicron particles. Nucleation generally occurs rapidly

forming a cloud of active particles that coagulate at a decreasing rate. Brownian collision and coalescence appear to be a major growth phenomena. Growth ceases in the case of oxides and metals when temperature drops below the solidification temperature. Via juxtaposition, groups of minerals will form eutectics with melting temperatures below

that of the pure compound. Phase diagrams indicate that below 1,900 to 2,000°F (1.038 to 1,093°C) the oxidized mineral species are generally dry and not vulnerable to unassisted capture. These are equilibrium diagrams and do not take into account supercooling due to a rapid quench. Illite and biotite contain small concentrations of iron and/or potassium and form a glassy phase at 1,742°F (950°C) and 2,012°F (1,100°C) respectively (2). Depending upon its fluidity, this glassy phase could be responsible for surface deformation at a relatively low temperature and provide the necessary sticking potential to permit retention on hot surfaces. These minerals are stable to high temperatures; however Cl in the environment will release potassium as a volatile species. Pyrite is a combustible component of the fuel with a specific gravity of about 5, approximately twice that of other minerals, a heating value of 3,000 BTU/lb, and a combustion rate comparable to that of anthracite coal of similar size. The rate of decomposition is governed by reaction rates, oxygen level, pore diffusion, and bulk or stream diffusion as well as the adventitious impurities (2). Although the bulk of the particle does not shrink during combustion, some fragmentation occurs altering the mineral size distribution. The melting temperature of the skeleton ash composed of or is 1,540°C and is of little consequence in ash deposition except as or a fluxing agent to other mineral species. An intermediate phase comprised of FeO and FeS may form a melt as low as 1,724°F (950°C). under reducing conditions pyrrhotite is formed generating a melting phase with temperatures reported as low as 1,300°F (2), significantly reducing the rate of combustion. At least a portion of the extraneous particles of pyrites fracture and disintegrate when thermally shocked, yielding a fume of 0.1–1 mm diameter particles (17). The inherent pyrites were found to generate an iron fume. The extent to which the iron fume originates as inherent or extraneous ash is not entirely resolved. However,

Factors Critically Affecting Fireside Deposits in Steam Generators

119

the fume is estimated to be 4–7% of the iron in the coal. Approximately 90% of the iron in North American coals occur as pyrites. The remaining iron usually occurs as the siderite, a carbonate which decomposes at 1,085°F (585°C) with the release of Mineral matter includes the organically bound sodium, calcium, and magnesium found in coal, the vanadium found in oil, and the potassium, calcium, and phosphorous found in biomass. One might also want to include the organically bound sulfur and chlorine found in coal in this group. Although strictly speaking they are minerals the salts of sodium, potassium, zinc, and lead found in the low rank fuels and MSW may also be included in this group on the basis of their thermal behavior. Linder investigated the release of organically bound sodium and included sodium chloride from coal which are the major contributors of fouling for most coals (18). He found the sodium is released under reducing conditions as metallic sodium. The rate of release was dependent on pore size, pore diffusivity, and bulk diffusivity. The organically bound sodium is evaporated as the volatiles are released, intimating that the char is depleted of sodium once the volatiles have escaped. Formation of Na(OH) is completed in the gas stream. Sodium chloride melts at 1,473°F (801°C) and begins to volatilize at about 1,382°F (750°C). Above 1,382°F (750°C) the rate increases significantly. Its rate of escape, as in the case of organically bound sodium, is dependent upon pore and bulk diff-usivity and instant properties of the char affecting either. Evaporation of crystals 0.1

120

R. W. Bryers

in diameter was bulk diffusion controlled whereas particles 0.3 were pore diffusion controlled. Evaporization rates were comparable to devolatilization. Once released the dominant sodium bearing gas species may be either NaCl, Na, or Na(OH) depending upon the oxygen level and the presence of chlorine. As the temperature levels drop below 802°F (1,477 °C), equilibrium favors the formation of silicates. Below 525°F (977°C) sul-

fates begin to form. Precise temperature levels depend on and levels. The reaction rates of the sodium with silica to form silicates changes as the process proceeds in five stages: • • • • •

Transportation of gaseous reactants and products to the silicate particle Diffusion of gaseous reactants and products through the boundary layer Gas Phase Reactions at the molten surface Sodium silicate forming reactions at the sodium/silica interface Diffusion through the silica melt (18)

Organically bound calcium occurs substantially in low rank coals. It is known to

form a fume of reactive submicron CaO. Its release from coal has not been examined, however, like sodium has. The fume initiates furnace wall deposits as well as low temperature convection bank deposits by forming calcium sulfate bonded deposits. Calcium silicate may perpetuate the growth of partially developed slag as semi-molten amorthite or gehlinite.

Potassium is most frequently found as a mineral rather than mineral matter in coal. In wood and biomass the concentration of ionically bound potassium is high. Potassium is found in regions where plant growth is most vigorous. Consequently, tree branches and forest wood slash, as well as straw, grasses, etc. are generally rich in potassium compared to core wood. During combustion the potassium is released as a vapor either as a hydroxide or chloride. Like calcium, the release mechanism has not been explored as thoroughly as sodium. Vanadium in petroleum is rapidly and completely oxidized to vanadium pentoxide at very high temperatures, probably as it is released at the droplet surface. Organically

bound vanadium in solid coke derivatives behaves quite differently as complete oxidation of the vanadium is suppressed to very low levels despite virtually complete burnout of the char. Solid fuels burn slower at lower temperatures. In addition, the vanadium must migrate through an atmosphere enriched with CO and counter current to the flow of oxygen. Exposure to the maximum level of oxygen occurs at a lower temperature in the flame front removed from the surface of the particle. The rate of cooling is faster than the rate of oxidation of vanadium to Like calcium and potassium, details of

the release of the vanadium have not been investigated. Unlike the other elements, the practical impact of the release of this element has the greatest significance on slagging, fouling, and corrosion as vanadium in solid fuels forms innocuous or rather than the highly corrosive low melting

formed from vanadium in oil.

Nickel, lead, and zinc all are found in the combustibles of manmade products in a variety of compounds. During combustion, they volatilize forming vapors which react with the HCl or present to form low melting chlorides, sulfides, or sulfates with very high vapor pressures.

Approximately 60–90% of the sulfur in coals occur as pyrite. The remaining portion is organically bound. The release mechanism and rate depends on the source of the sulfur, i.e. pyrite or coal. In the case of coal, the mechanism is influenced by a shrinking core

Factors Critically Affecting Fireside Deposits in Steam Generators

121

whereas with pyrites the particle core size remains constant. Initially, it was felt was formed in the flame. Hedly showed that cooling rates were too fast for equilibrium to be achieved. Consequently, only a small portion of the was converted to (20) Levy and Merryman showed the level in the boundary layer could be many times higher than in the gas stream due to longer residence times and catalytic conversion by Recent aqueous leaching experiments by Gluskoter et al. and Chen et al. along with XANE analysis by Huggins and Huffman have led these investigators to change their original thinking about chlorine in North American coals from chlorine occurring predominantly as inorganic alkali chlorides to a mixture of soluble chlorides and unspecified organic chlorides (22,23,24). With the chlorine occurring as anions of moisture in fine

cracks and pores, it would appear that a high level of removal could be achieved by fine grinding and aqueous removal. NaCl will precipitate from the chlorine rich solutions upon drying (25). Gibbs reports rapid dechlorination of coal during the early stages of devolatilization releasing and HCl (26). Equilibrium favors formation of HCl at the low temperature oxidizing conditions to which heat transfer surface is exposed. Engdahl indicates Cl has never been detected in incinerating steam generators (27) Once in the gas stream, particle fragmentation and coalescence of particles may occur altering the composition and solidification temperatures of some of the particle groups. During the quenching operation in the temperature range of 1,900°F and 1,400°F (1,038°C and 760°C) absorption of alkalis by silicates can produce a low melting eutectic making the particle sticky and vulnerable to unassisted capture or adherence to a contacting surface. The precise temperature at which the surface becomes sticky is dependent on cooling rates, rates of diffusion of alkalis to the surface, reaction rates, and diffusion rates through a viscous substrate. The absorption is also dependent on the HCl and Na(OH) levels of the gas stream. The mineral occurrence of the elements and their juxtaposition with regard to carbon and other mineral species are critical factors in predicting their contribution to fireside problems.

TRANSPORT MECHANISMS The mode of transport of fly ash to the heat-transfer surface is primarily inertial impaction for particles over 10 microns and thermophoresis and diffusion for particles 10 microns and smaller. The mechanism of transport by which the impurities reach the tube surface is strongly influenced by tube orientation in the gas stream, point of contact on the tube, wall effects, and location in the bundle. The bulk stream profile has gross effects on local turbulence also influencing rates and degrees of accumulation. The process of transporting particles of various sized to the surface via inertia impaction, thermophoresis or diffusion under laminar or turbulent conditions is reasonably well understood. Adherence or re-entrainment of impacting particles, and all the factors influencing particle retention are not well understood and certainly not quantified in meaningful parameters, i.e. tube temperature, surface roughness, impact angle, impact velocity, particle temperature, physiochemical state, size or body force (mechanical, Van der Waals, electrostatic). These forces may change with time as slags of many tons may be supported by a very small surface bond. Knowledge of sticking and subsequent bonding or adherence has a strong influence on ash removal and subsequent tube cleanliness.

122

R. W. Bryers

SLAGGING Slagging is defined as the formation of fused or sintered deposits on heat transfer surface or refractory in the furnace cavity subject to radiant heat exchange. Agglomeration in fluid beds or clinkering on stokers are special cases of slagging in the absence of heat exchanger surface. Designers try to avoid slagging by selecting a furnace exit

temperature below the initial deformation temperature of the ash. The furnace is sized to accommodate the moisture level of the fuel and local absorption rates are empirically adjusted to account for expected fouling factors based on the elemental fuel composition. The nature, degree and composition of slags may vary significantly throughout the furnace depending upon tube surface temperature, flame temperature, instant particle temperature, local gas temperature, absorption rate, direction of gas flow, mineral composition (i.e. coal rank and fuel type), mineral and coal size and distribution, concentration of mineral matter, and oxygen level at the flue gas boundary. Therefore an understanding of slagging requires a complete profile of the furnace gas temperature and composition with particular emphasis on the boundary layer. One must also be

able to predict changes in this profile as the deposit develops and changes from a dry sinter state through the plastic range and to a running fluid. Knowledge of the change in conductance and emittance of slag as a function of its composition and physical state is necessary. A generic conceptual model for slag formation on refractory surface evaporating surface or superheater surface is proposed in Figure 7 based on the data for emittance and conductance by Mulcahy et al. (28). Mulcahy’s data indicates the greatest resistance to heat flow occurs at the inner layer and tube slag interface where the layers of ash may be dry or sintered and quite thin. He indicates the inner layers may have a thermal conductance one quarter that of the fused or molten deposit. Consequently, once the initial layer has been developed and the heat absorption rate significantly reduced, the deposit can grow to substantial thickness with little further loss in absorption.

Evaporating and superheating surfaces run between 750 and 1,100°F (400–594°C) while refractory surfaces run 1,700°F (927°C) or higher depending upon refractory thickness. Consequently, slagging would be expected to be most severe on refractory surface exposed to low melting fly ash. Slagging is initiated with the attachment of particles in

a dry, semi-molten or molten state depending upon the composition of the fly ash particle, its particle size, the local heat flux, local gas composition, degree of super cooling, gas phase reactions with the tube surface or fly ash particle, or condensation of volatile constituents. Small dry particles may adhere to the surface when the tube-particle body forces exceed gravitational and gas dynamic shear forces. Semi-molten and molten par-

ticles form as the result of fluxing of quartz by alkalis, partial decomposition of pyrites, the chlorination or sulfation of heavy metals and alkali earth, metals, or supercooling due to a fast quench below the melting temperature of the ash particle. These semi-molten particles may bond to the tube surface by surface tension prior to freezing. Submicron particles and products of condensation generally deposit uniformly over the furnace wall, whereas larger semi-molten or molten deposits develop in irregular patterns depending upon the fluid dynamics of the gas stream and variations in distribution of fly ash in the

gas stream. Partitioning and segregation of minerals does occur within the gas stream due to differences in size and gravity of individual mineral species. The segregation may

begin as far back as the milling operation. Table 6 illustrates the different types of initial slag layers which may form by fuel

Factors Critically Affecting Fireside Deposits in Steam Generators

123

type and ash composition. A sharp increase in surface temperature of the slag is experienced with the formation of a minimal layer due to the thermal resistance between the ash and the tube and the low thermal conductivity of sintered ash. The surface becomes more vulnerable to capture of low melting mineral species in the fly ash, the fluxing of dissimilar mineral species resulting in the formation of a low melting eutectic, or mechanical entrapment of otherwise innocuous fly ash.

The rate of growth depends upon the rate of impact of troublesome constituents, their physical size and state and the change in surface temperature of the deposited slag.

As the ash deposit develops, the absorption rate decreases, substrates cool below the solidification temperature of individual species in the deposited ash and the surface temperature increases at a reduced rate. The surface temperature of the slag may eventually reach a level at which melting of the composite sample begins. A plastic state is achieved. The rate of entrapment increases. Depending on the composition of the ash, the semimolten phase may remain plastic over a large temperature range as crystals dissolve and the higher melting species go into solution. The slag surface will remain plastic until all the deposited ash goes into solution and the surface exceeds the softening or liquidous temperature of the deposited ash. At slightly higher temperatures, the viscosity of the ash decreases significantly and the slag begins to flow. Equilibrium is achieved and the slag is fully developed when the rate of deposition is equal to the rate of flow of the fireside layer. Rates of deposit growth change as the percent re-entrainment changes with surface conditions. Flame temperature, surface temperature, local absorption rates, and heat transfer coefficients of deposited ash are critical factors influencing slagging for a fuel with a given ash characterization. Diagnosing and predicting slagging problems requires knowledge of the furnace temperature profile and gas composition as well as the composition and temperature profile of the slag. The composition of the deposited ash may be quite different from that of the bulk fuel ash. The composition of the slag may also vary throughout the accumulated deposit as well as from location to location in the furnace or about the tube surface. The variation in composition depends on partitioning of mineral matter

124

R. W. Bryers

in the gas stream followed by further partitioning in the deposit process. An understanding of how and why the initial layers of deposit are formed is critical to the understanding of the slagging process. Knowledge of the thermal characteristics of the deposit is critical to predicting the impact of deposition on the steam generator performance.

FOULING Fouling is generally used as a term to define ash deposition in convective heat transfer zones downstream of the furnace cavity where radiant heat exchange is minimal. Selecting a furnace exit temperature below the melting temperature of the lowest melting

Factors Critically Affecting Fireside Deposits in Steam Generators

125

fly ash specie is critical to avoid slagging or catastrophic fouling of convective heat transfer surface. This means that furnace exit temperatures for oil may be in-excess of

2,300°F (1,260°C), for coal 1,800°F (982°C)–2,300°F (1,260°C) and as low as 1,250°F (676°C) for low rank fuels. Fouling may be due to condensation of volatile species on fly ash particles resulting in a low melting phase on the particle surface thereby increasing its sticking potential, condensation of low melting species directly on the tube surface, or the reaction of gases such as CL, HO, or CO with fly ash particles at the tube surface to form low melting species such as chlorides, sulfates, and sulfides. At the high velocities and low gas temperatures encountered in convection banks, the re-entrainment rate is very high for the bulk of the fly ash composed of coarse refractory oxides. Superfine particles with a

high surface to volume ratio and vapors which migrate to the tube surface by means of diffusion, thermophoresis or Brownian motion are considered the primary source of fouling. The vapors are formed by species which have high vapor pressure and low melting

temperatures in the range of 910–1,600°F (510–871°C). These are low compared to the oxides of Si, A1, Fe, Ca or Mg, which melt at temperatures around 1,750°F (954°C) or higher. Deposition may begin at the point of separation of flow from the tube surface, as in the case of superfine particles released from oil or the stagnation point for larger

particles released by ash from solid fuels. The submicron particles are held in contact with the surface by mechanical, electrostatic or Van der Waal forces as shown in Fig. 8 (29). Vapors such as HCl, etc. diffuse to the—surface and react with the particulate to form a low melting phase. As demonstrated by Levy, the concentration of gases such as at the tube surface maybe many times that in the bulk gas stream due to the catalytic conversion of to which proceeds at the tube surface where residence times are high and gas temperatures low. Within the inner layers conversions from silicates or chlorides to sulfates may also occur. The precise compounds that form will depend upon the local gas composition and the preferred equilibrium. In any event, a sticky surface is created reducing the rate of re-entrainment of impacting solid particles and increasing the rate of fouling. As in the case of slag, the substrate are subcooled as the deposit develops. Sintering occurs as the gases within the deposit react with the deposite surface

forming surface melts which subsequently freeze hard. Depending on bulk gas stream temperature, composition and residence time, some condensation or gas-solid reactions may occur prior to contact with the tube surface making the fly ash more vulnerable to capture once in contact with a stationary surface. Deposition does occur between tube rows and in dead gas space, causing unique fouling problems. With some of the low grade fuels rich in heavy metals, sulfur and chlorine it has been necessary to go to parallel flow convective heat transfer surface at very low gas temperatures to avoid severe fouling. Fouling may be broken down into various types depending upon fuel type and the mineral matter inclusions. They generally include alkali-bonded deposits, (calcuim sulfate-bonded) deposits, phosphorus-bonded deposits, vanadium, pentoxide, and heavy metals, as described in Table 7. The critical factors affecting fouling include the levels of concentration of the troublesome mineral matter in the fuel, the tube metal temperature, gas temperature, oxygen level, tube orientation and spacing. As in the case of slagging a model defining deposit thermal properties, coefficients of heat exchange, and strength are essential for and engineering solution to the fouling problem rather than an empirical assessment of the fouling potential of the fuel. Unfortunately, not all the mechanisms of fouling are clearly understood.

126

R. W. Bryers

CORROSION Fireside corrosion in steam generators has been attributed to sodium, sulfur, chlorine, CO, and salts of heavy metals. Although material selection has helped reduce

Factors Critically Affecting Fireside Deposits in Steam Generators

127

high oxidation rates at high steam temperatures, the solution to most of the severe corrosion rates involves limiting metal temperature, limiting concentration of the trouble-

some impurity in the fuel and control of the CO level at the tube surface. Alkali trisulfates and pyrosulfates formed on superheater surfaces and furnace walls respectively are corrosive when in contact with tube surface in a molten state. The pyrosulfates melt at 732°F (398°C) and require 150 PPM

to 2,000 PPM

for

the formation of potassium and sodium pyrosulfate respectively under reducing conditions (30). For pyrosulfates to exist at tube metal temperature of 899°F (432°C), there must be at least 1,000 PPM present for to form and 2% for to form. Consequently, pyrosulfates are confined to evaporating surfaces. The alkali trisulfates melt at 1,000°F (537°C) and decompose above 1,295°F (701°C) (31). The molten ash acts as a fluxing agent for the iron in the tube and provides the media for counter current flow of and corrosion product sustaining the process on a continuous basis. Although chlorine does not enter into the process, it may promote the release of potassium from otherwise stable illite, permitting the

R. W. Bryers

128

formation of the potassium trisulfate. The alkali trisulfates and pyrosulfates are controlled by limiting superheater temperature and avoiding flame impingement on furnace walls. For years it was believed corrosion due to chlorine could be expected to occur once the level of chlorine in the coal exceeded 0.3%. Recent experience with coal and municipal solid waste have shown that corrosion associated with chlorine will only occur in the presence of reducing or cycling reducing-oxidizing conditions. Under oxidizing conditions, chlorine occurs as HCl in the gas stream, which does not become seriously corrosive until temperatures exceeding the evaporation temperature corresponding to the steam generator pressure are exceeded. Corrosion of Cl becomes excessive at about 400°F (204°C) (32). Under reducing conditions, HCl decomposes releasing which can react with the iron in the tube to form which has a low melting temperature and high vapor pressure volatilizes and passes through the protective layer of iron oxide counter current with In the presence of it decomposes releasing fresh and forming or Since has a specific volume many times that of the reacted metal, its formation and subsequent decomposition will induce a nonprotective porous scale allowing the process to proceed linearly rather than parabolically. Brooks and Meadocraft have shown in the laboratory that the corrosion most frequently associated with Cl is indeed CO attack enhanced by the presence of Cl by virtue of the

formation of a porous scale (33). Lee clearly showed that “Cl” attack only occurred in the presence of CO at the boiler tube surface (34). Attack in the presence of Cl in the fuel appears to be an operating or design problem which permits exposure of portions of furnace wall to reducing conditions. It is quite possible that some of the corrosion problems associated with alkali pyrosulfates were indeed Cl enhanced CO attack as the tube surface was inadequately analyzed to detect the presence of Cl. Chlorides of heavy metals such as zinc and lead may also cause corrosion by forming low melting eutectics with alkali and iron chlorides acting as a fluxing agent of the protective surface of the tube. These problems are generally associated with low sulfur low grade fuels such as MSW. Blending of MSW with a high sulfur fuel has reduced the corrosion potential. Vanadium in oil, when completely oxidized at very high temperature, forms vanadium pentoxide which is catastrophically corrosive in the molten state. melts at 1,245°F (675°C). In the presence of alkalis such as sodium occasionally introduced

as NaCl in the oil during shipment, the melting temperature drops to 950° (51O°C). The problem begins to occur in steam generators once the vanadium in the oil exceeds 80 PPM and the sodium levels exceed 40 PPM. In the presence of less than 3% excess air, the lower oxide states of vanadium are formed, i.e. or which have melting temperatures in excess of 3,500°F (1,926°C) and hence are innocuous. Vanadium attack is controlled by controlling metal temperatures. Where that is not acceptable, additives such as magnesium are used to form intermediate compounds which raise melting temperatures of the ash. Tube supports are frequently fabricated from 60/40 chrome nickel alloy. At low temperature steam conditions, oils with vanadium concentration as high as 950 PPM can be fired providing there is a provision for removal of furnace slag.

Oil derived fuels such as fluid coke contain very high levels of vanadium, i.e. as high as 140,000 PPM V. They can be fired corrosion free as the lower flame temperature and slow rate of burning from a solid fuel inhibit the formation of

Factors Critically Affecting Fireside Deposits in Steam Generators

129

SUMMARY AND CONCLUSIONS The overview of the critical factors influencing high temperature fireside problems while firing various grades of steam-raising fuels and the various mechanisms that have been proposed suggest that a good many of the problems do not require detailed modeling to arrive at a solution. Variability in coal impurities in the mine, inclusions picked up during mining, transportation or handling, purchasing coal on the spot market can all be the source of unexpected fireside problems. Selection of fuels on the basis of elemental composition of the coal ash without due regard to the mineral form, size, or juxtaposition with regard to other minerals present can be the source of a drastic change in fireside slagging or fouling and unexpected furnace deposition. Improper milling of coal and burner adjustment or selection of burner size can be the source of fireside deposits. Flame impingement and low levels can cause unnecessary corrosion, slagging, and fouling problems. All these problems can be resolved by an adjustment in procedure or operation. Most corrosion problems can be handled by restricting metal and or gas temperatures and controlling levels at the tube surface. MSW steam generators, despite high excess air, require a good deal of furnace turbulence and flue gas mixing as well as imposed temperature limits to avoid corrosive situations. Modeling of air distribution in plenums through grates and into the furnace may be helpful. By itself wood does not often present a problem. Mixing with other fuels containing sulfur will introduce severe slagging and fouling problems. Biomass, including a large group of fuel types, must be treated on a case by case basis due to a large variation in fuel ash chemistry. Because of the large concentration of volatile mineral matter, characterization of the fuel ash is difficult. Users of this fuel type may have to develop new procedures peculiar to the fuel. Furnace temperatures must be kept low and designs requiring an intermediate convective furnace void of suspended surface between the radiant furnace and convective heat recovery may have to be used. On the other hand when switching, blending, penetrating new coal reserves, or dealing with a fuel known to be difficult to fire, it would be highly desirable to have a model which could predict steam generator performance based on changes in fuel ash chemistry. This model would be particularly helpful in dealing with coals for which there are no empirical data. Such a model must be capable of predicting complete velocity, temperature, and composition of combustion gas profiles, deposition rates and the thermal and compositional profile of the deposits formed. With such a model, one can not only predict performance, but also make an economic assessment of the problem and examining alternative approaches. Evaluation of slagging and fouling requires an understanding of the economic tolerance for the degree of ash deposition from a given fuel. It is not simply a matter of the degree of ash deposited. Development of a successful model requires an understanding of each step in the process a mineral must go through from the storage pile to the tube surface. In addition each step must be formulated and quantified. The brief discussions on mechanisms suggest that furnace partitioning, tube bonding, re-entrainment, deposit hardness, and a comprehensive model for deposit development require some attention. It is imperative that the industry accept solidification temperature rather than melting temperature as a standard for evaluating fouling, slagging, and corrosion, as the process occurs under a quenching rather than a heating process. Correlations for such properties as viscosity should be reevaluated for ash whose composition is out of the

130

R. W. Bryers

range for which the correlation was developed. The role of Cl in high temperature corrosion must be resolved.

REFERENCES 1 ) Benson, S.A., Jones, M.L., and Bryers, R.W., (1993), “Practical Measures To Minimize Ash Deposition On Coal Fired Plants”, Proceedings of The Engineering Foundation Conference, England, 657–678. 2) Bryers, R.W., (1996), “Fireside Slagging, Fouling and High Temperature Corrosion of Heat Transfer Surface Due To Impurities in Steam Raising Fuels”, Prog. Energy Combustion Sci., 22, 29–120.

3) Hough, D.C., Sanyal, A. and Davis, D.G., (1990), ASME Ash Fusion Research Project, American Society of Mechanical Engineering Research Committee on Corrosion and Deposits, CRTD-18, New York.

4) Winegartner, E.C., (1974), “Coal Fouling and Slagging Parameter Book”, No. H86, ASME, New York. 5) Corey, R., (1964), “Measurement and Significance of the Flow Properties of Coal Ash Slag”, Bureau of Mines Bulletin, 618. 6) Reid, W.T., (1976), “Corrosion and Deposits in Combustion Systems”, Penn. State Fuels Seminar, College Park Pennsylvania. 7) Kalmanovitch, D.P., and Williamson, J., (1986), “Crystallization of Coal Melts”, ACS symposium Series

301, American Chemical Society, Washington, D.C., 234–255. 8) Huffman, G.P., Huggins, F.E., and Dunmyer, G.R., (1981), “Investigation of The High Temperature Behavior of Coal Ash In Reducing and Oxidizing Atmospheres”, Fuel, 60.

9) Moza, A.K., and Austin, L.G., (1981), “Studies On Slag Deposit Formation In Pulverized Coal

Combustion Part 1 - Results Of The Wetting and Adherence Of Synthetic Coal Ash Drops On Steel”, Fuel, 60. 10) Jackson, P., (1979), “From Mineral Matter to Deposits In P-F Fired Boilers; Part 1 The Behavior Of Minerals in The Flame; Part 11. The Basic Physics And Chemistry of Deposit Formation and High

Temperature Corrosion”, Pulverized Coal Firing—The Effect of Mineral Matter, T. Wall, (ed.), University of Newcastle, Australia, 136–139. 1 1 ) Thomas, W.H., (1938), “Inorganic Constituents Of Petroleum Part 11”, Science of Petroleum, Oxford University Press, London. 1 2 ) Bowden, A.T., Draper, P., and Rowlins, H., (1953), The Problem of Fuel Oil Ash Deposition In OpenCycle Gas Turbines”, Proc. Inst. Mech. Eng. (A) 167. 13) Shah, H., Huffinan, G.P., Higgins, E.E. and Shah, A., (1991), “Graphical Representation of CCSEM Data for Coal Minerals and Ash Particles”, Seminar on Fireside Fouling Problems, Brigham Young University Center for Advanced Combustion Engineering Research, Provo, Utah. 14) Miller, RN.. and, Given, PH., (1978), “A Geochemical Study of The Inorganic Constituents in Some Low Rank Coals”, U. S. Department Of Energy Report FE-2494 TRI. 15) Benson, S.A. and Holm, PL., (1985), “Comparison of Inorganic Constituents in Three Low Rank Coals”, Ind. Eng. Chem. Prod. Res. Dev., 24, 145. 16) Ulrich, G.D., Riehl, J.W., French, B. Rand Desrusiers, R, (1977), “The Mechanism of Submicron Fly Ash Formation in a Cyclone Coal Fired Boiler”, Engineering Foundation Conference On “Ash Deposition and Corrosion from Impurities in Combustion Gases”, R.W. Bryers (ed.). New England College, New Hamphire, 253–268. 17) Srinivasachar, S., Helble, J., Katz, C.B., Worench, J.R. and Boni, A. (1988), “A Transformation of Inorganic Coal Constituents in Combustion Systems”, PSI Technology Quarterly Report No.7, PSI1024/SR-386.

18) Lindner, E.R, Manzoori, A.R., and Wall, T.F., (1991), “A Theoretical Analysis of Sodium Silica Reaction During Pulverized Coal Combustion”, Proceedings of The Engineering Foundation Conference in Inorganic Transformations and Ash Composition During Combustion, S.A.Benson, (ed.), Shereton Palm Coast, Florida, 565–581.

19) Jackson, P., (1963), “The Physicochemical Behavior of Alkali Metal Compounds In Fireside Boiler Deposits”, “Proceedings Of The International Conference on The Mechanism of Corrosion by Fuel Impurities”, H . R . Johnson, and D.J. Littler (ed.), Marchwood England, 484–495. 20) Hedly, A.B., (1967), “Factors Affecting The Formation of Sulfur Trioxide in Flue Gases”, J. Inst. Fuel, 40, 142. 2 1 ) Levy, A. and Merriman, E.L., (1967), “Mechanisms of Formation of Sulfur Oxides In Combustion”, Trans. ASME J. Eng. Power, 89 Series A, 297.

Factors Critically Affecting Fireside Deposits in Steam Generators

131

22) Gluskoter, H.J., and Rees, O.M., (1964), “Chlorine In Illinois Coal”, Geological Survey Circular 372, Urbana Illinois. 23) Chen, H.L., and Pagano, M., (1986), “The Removal of Chlorine From Illinois Coal by High Temperature Leaching”, Fuel Proc. Tech., 13, 261–269. 24) Muggins, F.E., and Huffman, G.P., (1991), “An XAFS Investigation of The Form of Occurrence of Chlorine in U. S. Coals", “Proceedings of CRSC-EPRI First International Conference on Chlorine in Coal”, J. Stringe (ed.), 11–26. 25) Chou, C.L., (1991), “Distribution and Forms of Chlorine in Illinois Basin Coal”, Proceedings of The CRSC-EPRI First Intentional Conference On “Chlorine In Coal", J. Stringer (ed.), 11–26. 26) Gibb, W.H. and Angus, J.G., (1983), “The Release of Potassium From Coal During Bomb Combustion”, Journal of The Institute of Energy, 149–157. 27) Engdahl, D.B., Dartoy, J. and Bebte, P., (1979), “European Refuse Fired Energy Systems; Evaluation of Design Practices”, U.S. Department of Commerce, PB 80-115322. 28) Mulcahy, M.F.R., Boow, J. and Goard, P.RC. (1966), “Fireside Deposits and their Effect on Heat Transfer

In a Pulverized Fuel Fired Boiler, Part 1; The Radiant Emittance and Effective Thermal Conductance of the Deposits, Part 11; The Effects of the Deposit On Heat Transfer from The Combustion Chamber Considered as a Well Stirred Reactor”, Journal of The Institute of Fuel, 39, 308, 385–394.

29) Benson, S.A., Steadman, F.H., and Kalmanovitch, D.P., “Studies of the Formation of Alkali and Alkaline Earth Aluminosilicates During Coal Combustion using a Laboratory Scale Furnace, Conference on “The Effects of Coal Slagging on Power Plants”, EPRI, Atlanta, Georgia. 30) Coats, A.W., Dear, .J.A., and Pendfold, D., (1968), “Phase Studies on the System and J. Inst. Fuel, 41, 129–132. 31) Nelson, W. and Cain, C, (1960), “Corrosion of Superheaters and Reheaters of Pulverized Coal Fired Boilers”, Trans. ASME, 82, 194–204. 32) Brown, M.H., Delong, W.B. and Auld, J.R, (1947), “Corrosion by Chlorine, Hydrogen, Chloride at High Temperature”, Ind. Eng. Chem., 39, 7, 834–844. 33) Brooks, S. and Meadowcroft, D.B. (19??), “Corrosion Resistant Materials For Coal Conversion Systems”, Applied Science Publishers, 105.

34) Lee, D.J. and Whitehead, M.E., (1983), “The Influence of Gas and Deposit Chemistry of The Fireside Corrosion of Furnace Wall Tubes In Coal Fired Boilers”, Proceedings of The Engineering Foundation

Conference on “Fouling Of Heat Exchangers”, R.W. Bryers (ed.), White Haven Pennsylvania, 69–104.

This page intentionally left blank.

ADVANCED ANALYTICAL CHARACTERIZATION OF COAL ASHES—AN IDEMITSU KOSAN— ELSAM COOPERATION PROJECT Ole Hede Larsen 1 , Flemming J. Frandsen2, Lone A. Hansen 2 , Signe Vargas2, Kim Dam-Johansen2, Karin Laursen 3, Takeo Yamada4, and Tsuyoshi Teramae5 1

Fælleskemikerne ELSAM, I/S Fynsværket DK-5000 Odense C Denmark 2 Department of Chemical Engineering, Technical University of Denmark DK-2800 Lyngby, Denmark 3 Geological Survey of Denmark and Greenland Thoravej 8, DK-2400 København NV Denmark 4 Energy Development Department Idemitsu Kosan Co., Ltd. No. 3-6, Kita-Aoyama 1-Chome Minato-ku, Tokyo 107, Japan 5 Coal Research Laboratories Idemitsu Kosan Co. Ltd. 3-1 Nakasode, Sodeguara Chiba, Japan

1. INTRODUCTION The fly ash from coal combustion can lead to operational disturbances and damage of boiler components due to deposition on the boiler heat transfer surfaces—slagging in the furnace and fouling on superheaters. Evaluations of the slagging and fouling propensity of a given coal or coal blend are based on different empirically indices calculated

from traditional coal analyses such as ash composition and ash fusion temperatures. However, these indices have a limited applicability (Couch, 1994). To improve the predictive capability, Elsam is running a R&D-project on slagging and fouling, targetting towards gaining a better understanding of the processes through Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

133

134

O. H. Larsen etl al.

improved methods of characterization and modelling combined with full scale tests on boilers (Laursen 1997a; Laursen et al., 1997). At the same time, Idemitsu Kosan is working on improving methods for characterization of ash and ash melting behavoir, allthough following a different approach than Elsam. Idemitsu Kosan has a very well-equipped laboratory for coal analyses including comprehensive facilities for coal research (Teramae et al., 1997).

Due to these circumstances, and due to the fact that both in Japan and Denmark a large part of the electricity production is based on pf-boilers using coals of worldwide origins there was mutual interest to set up a collaborative project on advanced ash characterization. A parallel investigation scheme including Japanese and Danish analyses of 9 coals, 3 coal blends and 3 fly ash samples from actual boilers was carried out. The Japanese analyses were carried out by the Coal Research Laboratory at Idemitsu Kosan. The Danish analyses were carried out by the Geological Survey of Denmark and Greenland (GEUS), the Combustion and Harmfull Emission Control (CHEC) group at the Department of Chemical Engineering at the Technical University of Denmark and by the coal laboratories of the power plants within Elsam.

2. PRETREATMENT AND ANALYTICAL METHODS The investigation scheme is shown in Fig. 1. There are two types of samples, 12 pulverized coal (9 single coals and 3 blends) and 3 fly ash samples. The mineral contents of the coal samples are analyzed after three different pretreatment methods: no ashing, low temperature ashing and standard ashing.

2.1. Pretreatment The not-ashed pulverized coal samples were mounted in carnuaba wax and polished sections were prepared for Computer Controlled Scanning Electron Microscopy (CCSEM) analysis of the not-transformed mineral content at GEUS. Low temperature ashing (LTA-ash) of the coal samples were performed by Idemitsu Kosan in order to make an ashing with minimum transformation of the minerals. For

low temperature ashing, the pulverized coal was thinly placed on a vessel and put into a plasma reactor. The coal was ashed in plasma at a temperature of 150–200°C until the weight of the sample was constant. It took about 2–3 days for ashing the 2–3 g of coal. Standard laboratory ashing at 815°C (HTA-ash) to constant weight was performed by Idemitsu Kosan according to the Japanese Industrial Standard (JIS), which is very close to the ISO standard used by Elsam.

2.2. Analytical Methods The not-ashed pulverized coal samples were analyzed by CCSEM at GEUS (Laursen 1997b). The low temperature ashes (LTA) are analyzed by bulk ash composition (X-Ray Flourescence) by Idemitsu, X-Ray Diffraction (XRD) by Idemitsu and GEUS, High Temperature Light Microscopy (HTLM) by Idemitsu and Simultaneous Thermal

Advanced Analytical Characterization of Coal Ashes—An Idemitsu Kosan—Elsam Cooperation Project

Analysis (STA) by CHEC. The LTA-ashes and the corresponding ashes after STAanalyses were analyzed by Scanning Electron Microscopy (SEM) at GEUS. In HTLM a heating stage microscopy was used for measuring the linear shrinkage rate of a cylindrical ash pellet (5 mm in diameter, 1 mm in heigth) formed by hand press. During heating in atmosphere by 25°C/min, the pellet shrunk concentrically and the linear shrinkage rate was measured as the decrease in diameter in percentage of the original diameter. In the STA technique a sample and an inert reference are heated in atmosphere at a heating rate of 10°C/min. During heat up the sample weight and sample temperature are continously measured. The changes in weight in percentage give the Thermo Gravimetric Analyses (TGA) and differentiation of this curve by time provides the Dif-

ferential Thermo Gravimetric analyses (DTG) in %/min. The measurements of the sample temperature compared with the temperature of the inert reference sample give the Differential Thermal Analysis (DTA) or Differential Scanning Calorimetry (DSC).

In this case the DSC analysis is used and this shows the heat in W/g produced or consumed by chemical or physical processes during heat-up. A procedure for quantification of melt as a function of temperature based on the DSC-analysis was developed by Hansen (1997). From the onset of melting untill completion of melting, the melt fraction is quantified from integration of the DSC deviation from the base line. Further details are given in Hansen (1997), Hansen et al. (1997) and Frandsen et al. (1997). The high temperature ashes (HTA) were analyzed by bulk ash composition by Idemitsu (XRF) and by Elsam (Inductive Coupled Plasma-Atomic Emission Spectroscopy—ICP-AES), Ash Fusion Temperatures (AFT) by Idemitsu and Elsam and viscosity by Idemitsu. The ash fusion temperatures by Idemitsu were made according to JIS M8801 in oxidizing (air) and reducing atmospheres with a heat rate of 5°C/min.

135

136

O. H. Larsen et al.

The AFT analyses by Elsam were made according to ASTM D1857 in reducing atmosphere with a heat rate of 8°C/min. Sample geometries are similar but sizes different for JIS and ASTM. Critical temperature points (Initial deformation Tempetature—IT, Hemispherical temperature—HT and Fluid Temperature—FT) are based on almost similar criteria, but ASTM includes an additional Softening Temperature— ST. The available maximum temperatures are 1,500°C and 1,650°C for the

Idemitsu and Elsam analyses respectively. Viscosity analyses were carried out in a Haake RV2 Rotovisco unit in a atmosphere, and decreasing temperature at 10–50°C intervals after the ash sample had been slowly heated to the desired temperature, typically about 1,500–1,700°C. The fly ash samples from power plants were analyzed by CCSEM at GEUS, ash composition by Idemitsu and Elsam and particle size distribution (PSD) by Idemitsu. The PSD was analyzed with a laser diffraction type CILAS Granulometer 715 after dispersion in water by use of ultrasonic waves.

3. RESULTS The results of all the analyses carried out on the 12 coal and 3 fly ash samples are

given in internal reports. The ranges of composition, mineralogy and ash fusion temperatures are given in Table 1 and some interesting results and findings are presented in more detail in the following section.

Advanced Analytical Characterization of Coal Ashes—An Idemitsu kosan—Elsam Cooperation Project

137

4. DISCUSSION The evaluation of results will be discussed from the following characteristica: • Chemical and Mineralogical Composition of Coal Minerals and Fly Ash • Particle Size Distribution of Coal Minerals and Fly Ash • Sintering, Melting and Viscosity Characteristics of Coal Ash

4.1. Chemical and Mineralogical Composition of Coal Minerals and Fly Ash Analyses of bulk ash composition of HTA from coals and on fly ash were carried out by Idemitsu and by Elsam. A comparison of these two sets of data showed rather similar results on most oxides. The largest discrepancy on absolute basis is the content, where Elsam underestimates with appr. 4% (appr. 10% relative). This only lead to a slight overestimation of some of the traditionally used empirical slagging and fouling indices, such as those based on base to acid ratio. It has more effect on some types of modelling, primarily modelling of ash viscosities, where the content is an important figure. The largest discrepancy on relative basis was seen for with appr. 25%. This

has a larger effect on fouling indices. Idemitsu also determined the bulk ash composition of the LTA ashes and found almost identical composition of LTA and HTA ashes, with the exception of S. The sulfur content is higher in the LTA than in the HTA ash, probably due to that less

has evaporated. Based on the CCSEM analyses, bulk ash compositions were also calculated. All values are calculated on free basis. Comparison with standard ash analyses shows that the CCSEM analyses can reproduce the variations of chemical composition but with a general overestimation of and and underestimation of and

This is probably related to sectioning of samples, omission of ZAF-correction in the CCSEM analyses, small particle sizes not analyzed by CCSEM, organically associated elements and other factors. For many applications of CCSEM (i.e. comparing chemical/mineralogical composition of individual particles in coal, fly ash and deposits),

these deviations are of moderate importance, as they in general are systematic. In other applications, such as use for modelling of viscosity, these deviations can lead to erroneous interpretations and for such cases it should be considered to upgrade the CCSEM analyses, e.g. to include ZAF correction. The mineralogy of coal inorganics was determined in two ways: XRD analyses of LTA ashes and CCSEM analyses of the pulverized coal. In Table 2, the results for two coals and a blend of these are shown. XRD is semiquantitative and gives a ranging of minerals according to their dominant peak height. The CCSEM results will in this case be given the same way—in order of decreasing abundance. From the results, it is seen that the two laboratories range the minerals almost similarily based on XRD, but one discrepancy is, that GEUS never detects illite, whereas this clay mineral is generally found by Idemitsu. Comparison of XRD of LTA with CCSEM of pulverized coal for mineralogical analyses shows that the minerals detected by XRD, generally are also found as major phases with CCSEM, allthough the ranging is different. CCSEM generally ranges the clay minerals (kaolinite, illite, montmorillonite) higher and quartz lower than the XRD analyses. In addition to the minerals found by XRD, CCSEM find some minor phases

138

O. H. Larsen et al.

of silicates and oxides. However, some of these might be two minerals placed very close to each other so that they are analyzed as one mineral. Thus, it seems that there are possibilities to be able to get quantitative mineralogical composition data from CCSEM analyses, but it probably requires, that the CCSEM is improved on element quantification (e.g. ZAF correction), data reduction and on differentiating individual particles.

4.2. Particle Size Distribution of Coal Minerals and Fly Ash The particle size distribution of the coal minerals was only analyzed by CCSEM, whereas the fly ash samples were analyzed with CCSEM and laser diffraction. In Fig. 2, a typical plot of the particle size distributions as analyzed by CCSEM and by laser diffraction is shown. The CCSEM can not analyze reliable in the few-micron size and below, which is mainly due to the spot size in the SEM. At higher particle sizes the curves cross

over, and the CCSEM allways measures a smaller particle mean size, than the laser diffraction. The most important reason for this is probably the cross-section effect for the CCSEM technique, where the minerals are generally not cut at their largest diameter. The difference found in particle mean size from the two techniques seems very systematic for the samples analyzed. This could indicate, that it might be possible to correct the CCSEM

results in order to give more reliable particle size distributions above appr. but the analyses are too few to verify this. These comparisons indicate, that it might also be possible to measure the coal mineral particle size distribution with the CCSEM analyses after proper correction of the values allthough it needs further verification. This is of great importance for mechanistic modelling of fly ash formation in the boiler, as the particle size distribution of coal minerals can not be analyzed by other methods.

4.3. Sintering, Melting and Viscosity Characteristics of Coal Ash Several analyses provide information on the sintering, melting and viscosity changes at heat-up or cooling of coal ashes. Mentioned in order of increasing temperature ranges,

Advanced Analytical Characterization of Coal Ashes—An Idemitsu Kosan—Elsam Cooperation Project

139

these analyses include: 1) linear shrinkage rate (LSR) by HTLM as a measure of sintering characteristics, 2) STA analyses as measure of weight and temperature changes due to reactions, evaporation and melting, 3) ash fusion temperatures as measure of fusion and melting and 4) viscosity measurements of the melted ash.

In Fig. 3, all these characteristics for one coal are presented on the same figure in order to get a total picture of these characteristics. Sintering characteristics of LTA ash were measured as linear shrinkage rate (LSR) by use of High Temperature Light Microscopy (HTLM). In figure 3 the LSR is shown

for coal A in the low temperature range. It was found for all ashes, that sintering started in the temperature range of 600–750°C. This onset of sintering is much lower than the ash fusion temperatures. The sintering characteristics differed much between the coal

140

O. H. Larsen etl al.

types. Sintering phenomena occur in ash deposits on boiler walls and tubes and affect build-up and removal of ash layers. Ash fusion temperatures of HTA ash for coal A are shown as AFT in Fig. 3. The 3 temperatures shown are Initial deformation Temperature (IT), Hemisphere Temperature (HT) and Fluid Temperature (FT). The ash fusion temperatures of all 12 HTA ashes measured by Idemitsu and by Elsam are shown in Fig. 4. The dotted lines are best fit through the data points and the solid lines represents same values from both laboratories. It shows, that there is some scatter of values between the laboratories as has also been shown by other workers (Coin et al., 1996). It is also seen, that allthough the correlation coefficient is very high for the IT and decreases a bit to HT and to FT, the latter two scatters around the solid line, whereas the initial deformation temperature (IT) shows a significant deviation in the way, that Elsam measures systematically higher IT-values than Idemitsu, and the deviation increases at lower temperatures. The reason for this is not clear, but it is important, as the IT-temperature is considered the most important of

Advanced Analytical Characterization of Coal Ashes—An Idemitsu Kosan—Elsam Cooperation Project

141

the ash fusion temperatures in relation to slagging and this temperature also is included in many boiler manufacturers design criteria. In Fig. 5, the result of STA analysis of LTA ash from coal A is shown. Supported by the XRD analyses, this STA analysis has been interpretated as follows: clay minerals dehydroxylate below 150°C. Kaolinite, pyrite and bassanite are transformed to metakaolinite, hematite and anhydrite, respectively, between 400°C and 650°C. Carbonate (calcite, dolomite and siderite) decomposition occurs around 400–700°C. Melting starts just below 1,000°C and is completed at 1,390°C. SEM analyses of the STA ash confirms complete melting at 1,390°C. SEM/EDX analyses of all ashes after STA analyses show, that most ashes are totally fused glass showing complete melting. Some ashes were fused glass with iron oxide or iron sulfate particles embedded indicating remnants of partially fused pyrite. One ash showed partly fused glass showing only partial melting. These features are important in

the interpretation of melting. From the DSC curve in the STA analysis, a melting curve has been calculated as described by Hansen (1997) and Hansen et al. (1997). The result is shown as Melt % for coal A in Fig. 3. For this coal, the IT temperature corresponds to 33% melt, HT corresponds to 74% melt and FT corresponds to 98% melt. Figure 6 show graphs of the melt % versus IT-temperature and HT-temperature for all 12 samples. The data seem to divide into two groups. One group with the major part of the samples shows low melt % at the IT-temperature (0–15%) and moderate melt % at the HT-temperature (10–40%) as would also be expected. The other group has high melt % at the IT-temperature (50–80%) and at the HT-temperature (70–100%). It is also clear, that the latter group has higher ITand HT-temperatures than the group with low melt %. The reasons for this picture is not clear but part of the explanation might be, that ash with high fusion temperatures devel-

142

O. H. Larsen et al.

ops a more rigid skeleton by sintering before the melting occurs and hereby can contain more melt phase before deforming. Viscosity of the HTA ashes was measured by Idemitsu and the results for coal A are shown in Fig. 3. Six models designed to estimate viscosities of completely molten silica mixtures have been tested against the measurements by Idemitsu (Vargas et al. 1997). These models are categorized in 3 groups according to their temperature

relation: Tabulated values Weymann form: Arrhenius form:

Bottinga-Weill (1972) Urbain (1981), Kalmanovitch & Frank (1988), Streeter(1984) Watt (1968), Greenberg (1984)

The Bottinga-Weill model will not be treated herein, as surplus of alumina can not be accounted for in this model. The models are only evaluated for temperatures, where

the ash is completely molten. In Fig. 7 the results of the Weymann and Arrhenius form models are shown together with the measurements of Idemitsu for one coal. In this case, the Greenberg and the Urbain models give the best results, the Streeter model gives reasonable results, whereas the Kalmanovitch & Frank and the Watt models give poor results. When comparing results of all 12 coals, the Urbain model generally gives satisfactory estimates, the Greenberg model gives good estimates for ashes with more than 25 wt% the Kalmanovitch & Frank model gives good estimates for ashes with more than 50 wt% and the Streeter and the Watt models generally give poor estimates.

Advanced Analytical Characterization of Coal Ashes—An Idemitsu Kosan—Elsam Cooperation Project

143

4.4. Application of the Results Figures such as Fig. 3, where different analyses on ash sintering, melting and viscosity characteristica are collected might turn out to be useful in getting a more com-

plete picture of how different coal types behave with respect to slagging and fouling. However it still requires, that operational experiences with these coals in different boilers

are collected systematically and compared with the analyses. Also, more information might be gained upon a more detailled comparison of the individual analyses, e.g. rela-

tions between sintering and melting characteristica and the detailled chemical/miner-

alogical analysis by CCSEM. This kind of presentation also might enable better and faster evaluations of effects of coal blending. An example is shown in Fig. 8, where the sintering, melting and viscosity analyses are shown for two coal types and a blend of these. In this case the ash fusion temperatures of the blend is between those for the individual coal types. The same accounts for the viscosity, whereas the sintering and melt % analyses deviate from this

tendency. Other coal blends behave very differently from this picture.

144

O. H. Larsen et al.

Some of the abovementioned analyses also can be of great value with respect to mechanistic modelling. CCSEM analyses of chemical and mineralogical composition and particle sizes of individual coal minerals are neccessary for modelling of fly ash formation. Measurements of sintering, melting and viscosity of coal ash are applied for the modelling of fly ash and deposit stickiness (fly ash capture), slag flow and strength build-up in deposits. In order to test the empirical predictions and mechanistic models, full scale tests and experiences are neccessary and should be included in further investigations along this line.

5. CONCLUSION A collaborative project between Idemitsu Kosan, Japan and Elsam, Denmark on advanced characterization of coal ash has been carried out. A worldwide suite of 9 coals, 3 coal blends and 3 fly ash samples were analyzed by traditional and advanced analyses. CCSEM is valuable as it can enable insight in chemical and mineralogical composition and grain size of individual coal minerals. These informations are important for gaining a better understanding of the processes and also in mechanistic modelling of fly

ash formation. Comparison with traditional bulk chemical analyses and with XRD analyses on low temperature ashed coal samples shows, that CCSEM is reasonably good for chemical and mineralogical analyses, but could probably be improved significantly by

upgrading to include ZAF correction. Comparison of fly ash particle size distribution measurements by CCSEM and by laser diffraction shows, that CCSEM can not analyze particle size distribution reliably below appr. 5 µm, but it is possible, that with proper correction enabled with laser diffraction analysis, the CCSEM can measure realistically above 5 µm, but this needs further verification. In case of succes it might also be possible to analyze particle size distribution of coal minerals, which is of great importance for mechanistic modelling of fly ash formation. Measurements of coal ash characteristics related to sintering, melting and viscosities were compared and further studies of these, together with comparison with the CCSEM analyses might enable a better understanding of processes in heating up and cooling of the coal ash in the boilers. Sintering characteristics were analyzed as linear shrinkage rate of pressed pellets by High Temperature Light Microscopy. The different coal types showed very different sintering characteristics, which is of importance for strength build-up in ash deposits and removal of deposits. STA analyses were carried out on all coal samples and a melt fraction as function of temperature was calculated. Melt quantification at the final high temperature of the analyses has been verified by SEM analyses of the STA ashes. Comparison with traditional ash fusion temperatures shows different relations between ash fusion temperature and melt % for low melting and high melting ashes respectively. The reasons for this are not yet clarified. Melt quantification is important in evaluations of fly ash deposition on heat transfer surfaces. Viscosities of the coal ashes has been analysed and compared with different viscosity models. The Urbain model generally gave satisfactory estimates, whereas the other

Advanced Analytical Characterization of Coal Ashes—An Idemitsu Kosan—Elsam Cooperation Project

145

models showed poorer performances. This has importance for modelling of fly ash deposition on heat transfer surfaces and slag flow characteristics. Coal blends respond very different to the applied analyses and it is possible, that

the combination of the traditional and advanced analyses can enable a better understanding and prediction of coal blend performance. In order to test the empirically predictions and mechanistic models, full scale tests and experiences are neccessary and should be included in further investigations along this line.

REFERENCES Bottinga, Y. & Weill, D.F. (1972). “The Viscosity of Magmatic Silicate Liquids: a Model for Calculation.” American Journal of Science, 1972, 272, 438–475. Coin, C. D.A., Kahraman, H. & Peifenstein, A.P., (1996). “An Improved Ash Fusion Test”. In: Baxter & deSollar (eds.) “Applications of Advanced Technology to ash-Related Problems in Boilers”, Proceed-

ings Engineering Foundation Conference, July 16–21, 1995, Waterville Valley New Hampshire, 187–200. Couch, G. (1994). “Understanding slagging and fouling in pf combustion”. IEA Coal Research/72, 118pp.

Frandsen, F.J., Hansen, L.A., Dam-Johansen, K. & Laursen, K. (1997). “Elsam-Idemitsu Kosan Cooperative Research project: STA- and SEM/EDX-Analyses of Low temperature Coal Ashes.” Department of Chemical Engineering, Technical University of Denmark, CHEC Report, August 1997.

Greenberg, S. (1984). “Viscosity of Synthetic and Natural Coal Slags.” Project Review METC 1984, Viscosity. Hansen, L.A. (1997). “Sintering of Ash Deposits”, Ph.D.-Thesis, Department of Chemical Engineering, Technical University of Denmark, To be finished November 1997. Hansen, L.A., Frandsen, F.J., & Dam-Johansen, K. (1997). Ash Fusion Quantification by Means of Thermal Analysis. Paper presented at the Engineering Foundation Conference “Impact of Mineral Impurities in

Solid Fuel Combustion”, Kona, Hawaii, November 2–7, 1997.

Kalmanovitch, D.P. & Frank, M. (1988) “An Effective Model of Viscosity for Ash Deposition Phenomena.”

in Bryers, R.W. & Vorres, K.S.: Mineral Matter and Ash Deposition from Coal, Engineering Foundation Conference, Santa Barbera, California, February 22–26, 1988, p.89. Laursen, K. (1997a). “Characterization of Minerals in Coal and Interpretations of Ash Formation and Deposition in Pulverized Coal Fired Boilers.” Ph.D.-Thesis, Geological Survey of Denmark and Greenland,

1997. ISBN-87-7871-022-7. Laursen, K. (1997b). “Advanced Scanning Electron Microscope Analyses at GEUS.” Geological Survey of

Denmark and Greenland Report 1997/1, 24pp. Laursen, K., Frandsen, F.J., & Larsen, O.H. (1997). “Ash Deposition Trials at Three Power Stations in Denmark.” Paper presented at the Engineering Foundation Conference “Impact of Mineral Impurities in Solid Fuel Combustion”, Kona, Hawaii, November 2–7, 1997. Streeter, R.C., Diehl, E.K. & Schobert, H.H. (1984) “Measurement and Prediction of Low-Rank Coal Slag

Viscosity.” in: Schobert, H.H.: The Chemistry of Low-Rank Coals. ACS Symposium series 264, p. 195. Teramae, T. (1997). “ELSAM/ldemitsu Coal Research Cooperation—Results and Interpretation of Analyses.”

Idemitsu Internal Report, April 1997. Teramae, T. & Yamashita, T. (1997). “Behavior of Inorganic Materials during Pulverized Coal Combustion.”

Paper presented at the Engineering Foundation Conference “Impact of Mineral Impurities in Solid Fuel Combustion”, Kona, Hawaii, November 2–7, 1997.

Urbain, G., Cambier, F, Deletter, M. & Anseau, M.R. (1981). “Viscosity of Silicate Melts.” Trans. J. Br. Ceram. Soc. 1981, 80, 139 141. Vargas, S., Frandsen, F. & Dam-Johansen, K. (1997). “Elsam-Idemitsu Kosan Cooperative Research project:

Performance of viscosity models for high-temperature coal ashes.” Department of Chemical Engineering, Technical University of Denmark, CHEC Report 9719, August 1997. Watt, J.D. & Fereday, F. (1968). “The flow properties of slags formed from the ashes of British coals: Part I . Viscosity of homogeneous liquid slags in relation to slag composition.” Journal of the Institute of Fuel

1968, 41, 99–103.

This page intentionally left blank.

A NOVEL APPLICATION OF CCSEM FOR STUDYING AGGLOMERATION IN FLUIDISED BED COMBUSTION

Mika E. Virtanen, Ritva E. A. Heikkinen, H. Tapio Patrikainen, and Risto S. Laitinen Department of Chemistry, University of Oulu Linnanmaa FIN-90570 Oulu, Finland Bengt-Johan Skrifvars and Mikko Hupa Åbo Akademi University, Combustion Chemistry Research Group Lemminkäisenkatu 14-18 B, FIN-20520 Abo, Finland

1. INTRODUCTION Computer controlled scanning electron microscopy (CCSEM) combined with an energy dispersive X-ray spectrometer (EDS) has widely been used in examining coal and its mineral content [Huggins et al., 1982, Straszheim et al., 1986, 1988, Galbreath et al., 1995]. Automated image analysis (AIA) has many advantages over the manual point analysis or X-ray maps of the sample. This method of making a large number of individual point analyses which produces a vast tabular file demands a sophisticated way of representing the results. One way of doing this is to categorize the point analyses into certain classes in an analogous way to the classification of mineral matter in coal. When inspecting ashes derived from the original fuel, the mineral classification may be questioned, since the

classification may produce information related to the sampling artifacts rather than to a realistic problem. The problems involved in the classification have been circumvented by representing the data graphically, using binary and ternary diagrams [Shah et al., 1988]. During the recent years, fluidised bed combustion has become very popular. At the same time the environmental regulations have become more stringent. Although the combustion of lower rank coals and other low rank fuel like peat and biomass favors the use of fluidised-bed, the technique has its disadvantages. The coating and agglomeration of the bed material may lead to total defluidisation of the bed causing an unscheduled shut down of the boiler. Bed agglomeration and defluidisation have attracted substantial interest during recent years, since defluidisation seems to occur relatively often [Anthony et al., 1995, Nordin et al., 1995, Skrifvars, et al., 1992, Skrifvars, et al., 1997, Kauppinen et al., 1997]. Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

147

148

M. E. Virtanen et al.

A key process is the formation of coatings on the bed particles. If the properties of the coatings can be identified, it is possible to obtain first hand indications about the bed agglomeration process itself. CCSEM has advantages over conventional analytical methods when investigating the nature of the coating and agglomerates in the bed. The use of image processing and extended ternary diagrams (so called quasi-ternary diagrams) [Heikkinen et al., 1997] provides the possibility of identifying different phases present in the coatings and the adhesive material between the agglomerated bed particles.

2. EXPERIMENTAL 2.1. Sampling Numerous samples were collected from a small pilot-scale (5kW) fluidised bed reactor at ETC Piteå Sweden during an extensive series of controlled agglomeration tests [Nordin et al., 1995]. To eliminate the influence of the particle temperature variation in the bed, every test was started by burning the fuel under examination in the reactor at 600°C u n t i l the bed/ash ratio was 6 wt-%. Subsequent to this “ashing procedure” the fuel

feeding was stopped and the bed temperature was externally raised by 3°C per minute. The onset of bed agglomeration was indicated by measuring differential pressures and temperatures of the bed as well as by video filming the process. The detection of initial bed particle agglomeration by considering all bed related variables simultaneously was performed by an online multivariant statistical process control (MSPC). In all the tests, pure quartz was used as the bed material and the fuel was varied.

During the tests, several different fuels were burned in a bed consisting of pure quartz sand (99.9%). From each test, one sample was taken at the beginning of the external heating, one or two samples during the external heating, and the final sample after

agglomeration occurred.

2.2. CCSEM All bed and agglomerate samples were mounted in epoxy resin, Struers Epofix ( 1 5 : 2 ) . Mounts were cross-sectioned by grinding with 200, 800 and 1,000 mesh SiC

powder in glycerol, and polished with 3 and

diamond pastes. Finally, the samples

were coated with a thin carbon layer to eliminate the electrostatic effects.

The measurements were carried out with a Jeol JSM-6400 scanning electron microscope combined with a Link ISIS energy dispersive x-ray analyzer. The analyzer included the Link ISIS software package. 1 5 k V acceleration voltage and a beam current of 120* A were used. The sample distance was 15mm. The SEM-analysis was started by searching for a suitable field from the sample and collecting a BE- (backscattered electron) image from it at 256*208 resolution. In the BEimage the epoxy is seen a as black background, the bed material as the darkest of the grey areas, and the coatings and adhering material are seen as brighter areas (see Figs. 1 and 2). The original image was smoothed by strengthening the contrast and by removing roughness in order to facilitate further treatment. One or two binary images were extracted from the smoothed BE-image, depending on the nature of the sample. In the case of bed material coatings, only one binary image was extracted, which included the bright areas on the surface of the bed particles. When examining bed

A Novel Application of CCSEM for Studying Agglomeration in Fluidised Bed Combustion

agglomerates with a possibility of obtaining different grey levels on the adhesive material, two binary images were extracted. Depending on the nature of BE-image, there is a possibility of variation in the elemental composition over large areas of a seemingly uniform grey level. Therefore, after producing the binary images, the large areas were divided into smaller domains; the

resulting images are shown in Figs. 1 and 2. This procedure provides information from over two hundred discrete domains over a large area of a coating or adhesive material in

149

M. E. Virtanen et al.

150

the agglomerate. For convenience, each of these domains could be considered as different particles, and their elemental composition could be determined separately. The actual determination was performed as a point analysis from the center of each domain. It is clearly seen from the Figs. 1 and 2, that the bed particles are easily excluded in all agglomerated samples. The ZAF-corrected relative amounts of sodium, magnesium, aluminum, silicon, phosphorus, sulfur, potassium, calcium, titanium, and iron were determined at each point. At least one thousand points from about ten fields were analyzed for each sample in order to gain statistical validity.

3. VISUALIZATION OF THE ANALYTICAL RESULTS Ternary diagrams have increasingly been used in the visualization of large number of point analyses [Laursen et al., 1995, Wigley et al., 1995]. The traditional ternary diagrams, however, are limited in their capability to visualize results, when ten elements are

present. To overcome this limitation, so called quasi-ternary diagrams have been used. They are a logical extension to the conventional ternary diagrams. In a corner of a conventional ternary diagram, the amount of an element is 100%,

and in the opposite side it is 0%. In a quasi-ternary diagram two or more elements can

be placed on one corner, and therefore, the sum of the contents of these elements in that corner is 100%. For a point analysis to appear in the diagram, the sum of the contents

of the chosen elements have to exceed a minimum limit, usually 80%. This means, that if the point analysis appears in a corner marked it contains at least 80% Si and Al. The relative amounts of silicon and aluminum, however, is not fixed as long as the

sum exceeds 80% and there can be up to 20% of other elements. If the point is exactly halfway towards the opposite side, the combined content of Si and Al is within a range of 46%–50%. Let us next consider a side of a quasi-ternary diagram, where

is marked on

the other end and a combination of is on the other (see diagram (b) in Fig. 3). Along the side, the sum of the contents of these six elements is 80-100%, and the contents of the elements on the third corner ( in diagram (b)) is 0%. The rest (up to 20%) consists of the analyzed elements that are not shown, e.g. P and S in the case of the diagram (b) of Fig. 3. A series of these diagrams is shown in Fig. 3. They show a distribution of 1,037 point analyses of a real sample. In the first diagram [Fig. 3(a)] all the ten analyzed elements are included in the corner definition and all point analyses are shown in the diagram. The corners are defined in order to deduce the presence of silicates or aluminosilicates, phosphates and sulfates. Because the distribution of point analyses does not spread towards the phosphorus and sulfur corner, it can be inferred, that the analyzed matter in the sample does not contain significant amounts of phosphates or sulfates. All the maxima in the diagram are along the side between the and corners, this implies the presence of silicates or aluminosilicates,

where there is none or a very small amount of phosphorus or sulfur present. There is also a maximum with a composition of 50–60% of and a smaller maximum with 75–90% of these will be discussed below. Since there are no phosphates or sulfates in the sample, phosphorus and sulfur are omitted in the next diagram. The corners in the diagram (b) of Fig. 3 are now defined in such a way as to reveal the presence of iron or titanium in the silicates and alumi-

A Novel Application of CCSEM for Studying Agglomeration in Fluidised Bed Combustion

151

nosilicates. The majority of the particles still appear along the side between and corners, but in few cases these silicates or aluminosilicates contain small amounts of iron or titanium. Also there are a few point analyses exhibiting mainly iron or titanium. It should be noted that results from three of the analyzed points are not present in this diagram. These three points obviously contain at least 20%

In the next diagram [Fig. 3(c)] the corners are redefined again. The arrangement is now chosen to separate silicates and aluminosilicates. Referring to the previous diagram [Fig. 3(b)], the amount of iron and titanium was found to be negligible, and only six elements (Si, Al, Na, Mg, K, and Ca) are chosen for the diagram (c). The diagram still includes the majority of the analyzed points (1,015 of 1,037) according to the 80% rule. Now the diagram reveals the presence of a minor aluminosilicate phase, but the majority of points do not contain aluminum.

152

M. E. Virtanen et al.

The elements in the corners are rearranged to examine the relation between alkali and alkaline-earth metals. In the diagram (d) silicon and aluminum are combined on one corner again, alkali metals are on the second corner, and alkaline-earth metals on the third. In this diagram three major phases are easily distinguished. In order to identify maxima 1 and 3, one more diagram [Fig.3(e)] is drawn, where the only difference to diagram (d) is the absence of aluminum. By comparing the diagrams (d) and (e), the maximum 1 can be identified as silicate and maximum 3 as aluminosilicate. In fact, the silicate phase can be presented with only three elements: silicon, sodium, and calcium. This is illustrated in the diagram (f) in Fig. 3, where 76.6% of analyzed particles are included (i.e. in 76.6% of analyzed particles The maximum 2 appears to be pure silicon derived from the bed material, as discussed below. It is also possible to estimate the amounts of different elements in a given phase. The diagram (f) in Fig. 3, shows the Na, Ca- silicate phase. The maximum in the diagram can be defined within the limits of and and when the possibility of having 20% of other elements (80% rule) is considered, the real limits for this phase can be expressed as: and

After a graphical inspection it is possible to go back to the original tabular file and extract the rows (point analyses) that belong to the major phases. By estimating the range

of different elements, criteria for extraction of information can be produced. Criteria would be defined by the ranges of different elements determined above and is analogous to the mineral classification. Closer examination of the extracts can reveal the role of minor elements like phosphorus or sulfur in the sample, which is not negligible in the formation of coatings and agglomerates.

4. DISCUSSION The CCSEM-technique is well suited for the characterization of bed material coatings and agglomerates. Using this method, more than a thousand points can be analyzed from the adhesive material of the agglomerates, as well as from coated bed particles. The quasi-ternary diagrams always show a peak of pure silicon, which implies quartz in the sample. This is a consequence of the use of the automated image processing. In some cases, the smoothing of the grey scale image can result in the inclusion of the smallest quartz particles in the adhesive material. Another reason is the relative roughness of the binary images, when the domain boundaries are not accurate enough. These problems can be solved by scanning the BE-image at a greater resolution (512*416 or 1024*832), but then the time required for image scanning and processing is radically extended. In all samples tested in this work, the bed material was easily excluded, and the division of the large areas of adhesive material into smaller domains was always successful. In the case of bed material coatings, however, the following considerations should be taken into account. The bed material coatings could only be analyzed, if they were thick enough. Then the bed particles could easily be separated and the determinations then focused on the coatings. A very thin layer of coating on the bed particles was hard to distinguish from its grey level in the BE image, resulting in difficulties in the image processing. Further,

the determinations were not focused solely on the coatings, but also on the bed particle itself, giving rise to the large amounts of silicon.

A Novel Application of CCSEM for Studying Agglomeration in Fluidised Bed Combustion

153

The extension of traditional ternary diagrams to quasi-ternary diagrams is very informative. By defining the corners in terms of all elements, these diagrams can include all analyzed points. Furthermore, by systematically excluding elements from the corners, it may be possible to deduce the phases present in the sample. Some elements like iron and phosphorus can be important in the agglomeration process, even in small amounts. Thus, they have to be observed even in small quantities and this may require a more rigorous inspection of the original data files.

5. CONCLUSIONS CCSEM-technique is a valuable method when characterizing the bed material coatings and agglomerates. It is comparatively easy to carry out more than a thousand point

analyses and obtain valid statistical information by using automated image processing. The use of quasi-ternary diagrams provides a good insight into the nature of the sample and can even reveal the phases present in a coating or an agglomerate.

6. ACKNOWLEDGEMENTS This work was performed under LIEKKI 2-program, which was financed by the Ministry of Trade and Industry. We would also like to thank Dr Anders Nordin and Mr

Marcus Öhman ETC/Piteå for bed samples.

REFERENCES Anthony, E. J., Iribarne, A. P. and Iribarne, J. V. (1995). “A new mechanism for FBC agglomeration and fouling when firing 100% petroleum coke.” Proc. of the 13th ASME FBC Conference, Vol. 1, Orlando, Florida. Galbreath, K., Zygarlicke, C., Casuccio, G., Moore, T., Gottlieb, P., Agron-Olshina, N., Huffman, G., Shah, A., Yang, N., Vleeskens, J. and Hamburg, G. (1995). “Collaborative Study of Quantitative Coal Mineral Analysis using Computer Controlled Scanning Electron Microscopy.” In L. L. Baxter and R. DeSollar (Eds.), Application of Advanced Technology to Ash-Related Problems in Boilers, New York, Engineering

Foundation. Heikkinen, R., Laitinen, R. S., Patrikainen, T., Tiainen, M. and Virtanen, M. (in press). “Slagging Tendency of Peat Ash.” Fuel Processing Technology.

Huggins, F. E. Huffman, G. P. and Lee, R. J. (1982). “Scanning Electron Microscope-Based Analysis (SEMAIA) and Mössbauer Spectroscopy, Quantitative Characterization of Coal Minerals.” In E. L. Fuller, Jr (Eds.), Coal and Coal Products: Analytical Characterization Techniques, Washington DC, American

Chemical Society. Kauppinen, E. (1997). “Tuhkan muuntuminen Leijukerroskaasutuksessa ja -poltossa: Haitallisten hivenmetallien vapautuminen ja alkalien käyttäytyminen.” In M. Hupa, J. Matinlinna and M. Ljung (Eds.), LIEKKI 2 Year book 1997, Turku, Åbo Akademi Förbränningskemiska forskargruppen. Laursen, K., Frandsen, F. and Larsen, O.H. (1995). “Slogging and Fouling Propensity: Full-Scale Tests at Two Power Stations in Western Denmark.” In L. L. Baxter and R. DeSollar (Eds.), Application of Advanced Technology to Ash-Related Problems in Boilers, New York, Engineering Foundation. Nordin, A., Öhman, M., Skrifvars, B.-J. and Hupa, M. (1996). “Agglomeration and Defluidization in FBC of Biomass Fuels—Mechanisms and Measures for Prevention.” In L. Baxter and R. DeSollar (Eds.) Application of Advanced Technology to Ash-Related Problems in Boilers, New York, Engineering Foundation. Skrifvars, B-J., Backman, R. and Hupa, M., (in press). “Characterisation of the sintering tendency of ten biomass ashes in FBC conditions by a laboratory test and by phase equilibrium calculations.” Fuel Processing Technology.

154

M. E. Virtanen et al.

Skrifvars, B-J., Hupa, M. Anthony, E. J. (1997). “Mechanisms of bed agglomeration in the cyclone and the return leg of a petroleum coke fired circulating fluidised bed boiler.” Proc. of the 14th ASME FBC Conference, Vol. 2, Vancouver, Canada, May 1997. Skrifvars, B-J., Hupa, M. Hiltunen, M. (1992). “Sintering of Ash During Fluidized Bed Combustion.” Ind. Eng. Chem. Res 31(4), 1026–30. Shah, N., Huffman, G. P., Huggins, F. E. and Shah, A. (1988). “Graphical Representation of CCSEM Data for Coal Minerals and Ash Particles.” In S. A. Benson (Eds.), Inorganic Transformation and Ash Deposition During Combustion, New York, The American Society of Mechanical Engineers. Straszheim, W. E., Yousling, J. G. and Markuszewski, R. (1986). “Analysis of Ash-Forming Mineral Matter in Raw and Supercleaned Coals by Automated Image Analysis-Scanning Electron Microscopy.” In K. S.

Vorres (Eds.), Mineral Matter and Ash in Coal, Washington DC, American Chemical Society. Straszheim, W. E. and Markuszewski, R. (1988). “Characterization of Mineral Matter in Coal for Prediction of Ash Composition and Particle Size.” In S. A. Benson (Eds.), Inorganic Transformation and Ash Deposition During Combustion, New York, The American Society of Mechanical Engineers. Wigley, F. and Williamson, J. (1995). “Modeling Fly Ash Generation for UK Power Station Coals.” In L. L. Baxter and R. DeSollar (Eds.), Application of Advanced Technology to Ash-Related Problems in Boilers, New York, Engineering Foundation.

THERMOMECHANICAL ANALYSIS AND ALTERNATIVE ASH FUSIBILITY TEMPERATURES S. K. Gupta, R. P. Gupta, G. W. Bryant, L. Juniper 1 and T. F. Wall CRC for Black Coal Utilisation Department of Chemical Engineering University of Newcastle Callaghan, NSW, 2308 Australia 1 Ultra System Technology Indooroopilly QLD, 4068, Australia

1. INTRODUCTION Ash deposition on furnace walls in pf (pulverised fuel) fired boilers is termed slagging when it occurs in the high temperature areas of furnaces directly exposed to flame radiation, and fouling in other regions such as tubes in the convection section of the boiler. These deposits not only reduce the heat transfer efficiency, but they can lead to corrosion and significant damage to the boiler tubes. Therefore, the nature of slagging properties of ash derived from coal is very crucial for the boiler designer and consequently for the coal exporter. The ash fusibility test is the most widely used method of assessing whether a coal ash will deposit on the heat transfer surfaces of boilers. The ash fusibility test is based on observing the temperatures at which successive characteristic stages of fusion occur in a moulded ash specimen when heated in a laboratory furnace under specified conditions. For instance, the Australian Standard involves heating of a cone of laboratory ash (prepared at 815°C in a muffle furnace) at approximately 5°C to 10°C/minute from 1,000°C to 1,600 °C. Four temperatures that are recorded in the conventional ash fusibility test which are believed to indicate different stages of melting are given below and shown in Fig. 1. • The deformation temperature (DT) is noted when ash just begins to fuse or shows the first sign of deformation or rounding of the apex of the cone. • The sphere temperature (ST) is recorded when the height of the ash becomes equal to the width of sample. • The hemispherical temperature (HT) is noted when the height of fused ash becomes equal to half of its width. Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

155

156

S. K. Gupta et al.

• The final temperature, known as flow temperature (FT) is noted when the height

becomes a sixteenth of the width. I t may be noted that terminology, sample preparation and configuration for this test

differ from country to country. The deformation and sphere temperature are normally used in coal specifications in the belief that these temperatures may indicate ash behaviour in the furnace. For instance, deformation temperature has been accepted as the temperature where the ash first softens and therefore becomes sticky. The hemispherical and flow temperatures are considered invariably by most of the design engineers to avoid slagging,

whereas deformation, hemispherical and flow temperatures are believed to be more important to control fouling [Barret, 1987]. There are well-documented shortcomings of poor precision of ash fusibility temperatures, which is mainly related to the subjective nature of the test [Juniper, 1995]. Therefore, there is a strong need for an alternative method to char-

acterise the ash fusibility of coal, which would have greater reliability of the measurements. To achieve this objective, a new technique-thermo-mechanical analysis (TMA) has been

developed. Thermo-mechanical analysis is frequently used in the metallurgical and ceramic industries to investigate thermal behaviour of materials at high temperatures. Many studies have been reported in past, which used changes in electrical resistance and shrinkage properties to characterise the melting characteristics of coal ashes [Raask, 1979, Ellis. 1979, Doolan, 1984 and Sanyal, 1993]. However, none of these studies could provide a satisfactory alternative. The TMA technique utilised a similar approach but in a different manner. The technique measures the percentage shrinkage and first derivative of the shrinkage as

a function of temperature. The main aim of this paper is to propose alternative ash fusibility temperatures from the TMA shrinkage measurements so that various stages of ash melting can be reliably characterised.

Thermomechanical Analysis and Alternative Ash Fusibility Temperatures

157

2. EXPERIMENTAL The ash fusibility was characterised using thermomechanical analysis of laboratory ashes of a large number of thermal coals from Australia, USA, UK and other countries. The sample selection was primarily based on ash chemistry. The new technique involved the measurement of the penetration of graphite ram into ash (~50 mg) placed in a cylindrical graphite crucible with a flat bottom, and referred as TMA shrinkage. As the assembly containing ash is heated under Argon from room temperature to 1,600°C at a rate of 5°C/minute, the penetrating ram sank into ash and eventually the slag flowed into the gap between the penetrating ram and the base of crucible as shown in Fig. 1 [Saxby et a/., 1996]. The rapid shrinkage events were obtained from the derivative of TMA shrinkage. Various “peaks” (representing the maximum rate of shrinkage with temperature), of various intensity and at progressively higher temperatures, were related to sintering and melting events in the ash sample. No significant shrinkage was observed in most of ashes up to 900°C. The temperatures corresponding to particular shrinkage levels were noted as TS%. Multiple measurements on the same ash indicated that major fusion events (rapid shrinkage) are reproducible to within [Saxby et al., 1996], The ash fusibility temperatures from the conventional test were also obtained under reducing conditions as per Australian Standards. The ash pellets were heated under simulated ash fusibility

test conditions and then quenched rapidly in air from various temperatures of interest. The extent of melting at various temperatures was quantified by visual examination of residual solid particles in the scanning electron microscope (SEM) images of quenched ash pellets.

3. SAMPLE GROUPING It is well known that ratio and different fluxing contents of ash are mainly responsible for differing melting characteristics [Huggins et al., 1981, Bryers, 1996], The extent and ratio of various oxides in ash may approximately relate to the proportion of the original minerals present in coal [Raask, 1986]. For instance, high ratio may indicate the presence of large amounts of as free quartz in ash, which may decrease the temperature range for the completion of melting. The from 1 to 3 wt% in ash may be related to 10 to 30% illite content of the total mineral matter in coal which is quite common in bituminous coals [Raask, 1986]. Therefore, a sample grouping scheme, based on silica/alumina ratio and various fluxing oxides, is designed to understand the TMA shrinkage measurements (Table 1). The melting and crystallisation behaviour of ashes is commonly explained by simplifying the ash compositions to three or four oxides where or [Huggins et al., 1981 and Kalmanovitch, 1986], Therefore, normalised ash composition of some of the ashes is also shown in Fig. 2 based on a ternary diagram (Levin et al., 1964).

3.1. Group A: Ashes with High Silica and Alumina Contents Group A ashes contain very high levels of refractory components and a limited amount of fluxing components. Quartz and clay are the major minerals expected to be present in the parent coals of Group A ashes. The ashes are further divided in to three series according to the difference in extent and nature of fluxing

158

S. K. Gupta et al.

159

Thermomechanical Analysis and Alternative Ash Fusibility Temperatures

components. Series A1 and A2 contain low and high levels as the major fluxing oxide respectively, whereas Series A3 contain ~5 % iron oxide as an additional fluxing component. The TMA shrinkage and melting behaviour of these ashes will be primarily governed by the quartz and clay interactions in ash. Figure 2 shows that composition of these ashes lie in mullite primary phase field of the system and their liquidus temperatures are expected to be high. However, for a given series, liquidus temperature would fall as the ratio of ash increases. These ashes are likely to have large error in observing the deformation temperatures in the existing ash fusibility test [Wall, 1995].

3.2. Group B: Ashes with Moderate Iron and Calcium Oxide Contents Group B ashes are characterised by the presence of reasonable levels of fluxing components so that content is less than 90%. The ashes are further divided

in three series based on different

ratio and various fluxes. Series B1 and B2

contain iron oxide as the major fluxing component, while ratios are greater and less than 2.2 respectively. Series B3 contain both iron oxide and calcium oxide as the fluxing components. In addition to quartz and clay, the pyrite and calcite contents in the parent coal would also affect the melting behaviour of these ashes significantly. The ash composition of most of these samples lie in the mullite primary phase field in phase diagram except for AEC18 (anorthite), OPC03 (hercynite)

and OPC04 (tridymite). Consequently, the melting and hence the TMA shrinkage behaviour of ashes AEC18, OPC03 and OPC04 may be slightly different than the other ashes. The liquidus temperatures of Group B ashes are usually low, however these may be higher if ratios is lower (Fig. 2). The detailed ash composition and fusibility temperatures for all ashes are given in Table 2.

4. EXPERIMENTAL RESULTS AND DISCUSSION

4.1. Typical TMA Results of Ash Groups Figure 3 shows typical TMA shrinkage curves of selected ashes, which contain various levels of fluxing components and silica/alumina ratios. Ash AEC06 (Series A1: high ratio) did not indicate any significant shrinkage up to 1,500°C due to very low levels of fluxing oxides that is whereas, PN6 (Series A2) showed substantial shrinkage at a temperature less than 1,400°C due to high Thick curves show the TMA shrinkage of ashes which contain of iron oxide as the variable fluxing oxide for example AEC12, OPC01 and OPC03 ash. Figure 3 shows that the increasing fluxing levels moves the TMA shrinkage curves to a low temperature region. Figure 4 shows that low shrinkage occurs if melting is low (AEC06: 5% shrinkage and melting) and large shrinkage occurs if extent of melting is high (OPC01: 85% shrinkage and melting). Therefore, it is clear that the TMA technique appears to distinguish the effect of ash chemistry on various amounts of ash melting in the form of shrinkage curves.

4.2. Alternative Ash Fusibility Temperatures Ideally, the ash melting may begin at as low as 950°C due to the presence

which may form low temperature eutectics with

and

and continues up to

160

S. K. Gupta et al.

Thermomechanical Analysis and Alternative Ash Fusibility Temperatures

161

162

S. K. Gupta et al.

greater than 1,600°C. Practically, the amount of liquid at low temperatures may be very low. SEM analysis of ash pellets indicated that greater TMA shrinkage levels were associated with greater levels of melting (Fig. 4). The melt phase composition of various ashes is quite different for various ashes, and also changes as the melting progresses. Therefore, it was difficult to specify a unique shrinkage level to a unique level of melting. Therefore,

four temperatures from the TMA technique were selected corresponding to shrinkage levels of 25%, 50%, 75% and 90%. These temperatures are proposed as alternative ash fusibility temperatures, which may be considered to represent initial, intermediate, completion of melting and slag flow characteristics for practical purposes. 4.2.1. Initial Melting Temperature: T25%. Figure 5 shows the relationship between extent of melting and various shrinkage levels. SEM examination of ash pellets quenched

Thermomechanical Analysis and Alternative Ash Fusibility Temperatures

163

at the T25% temperature indicated the presence of 25% liquid at this stage of melting. In general, the amount of liquid required to cause 25% shrinkage is low for

refractory ashes (Group A) when compared to that required for samples containing reasonable fluxing components (Group B). It was observed that, T25% for Group A samples range from 1,100°C to 1,300°C and for Group B range from 1,050°C 1,150°C. It is

believed that this difference of temperature range, which changes the melt phase viscosity significantly, is mainly responsible for this deviation in predicting the extent of melting. In few ashes, which contain very high refractory oxides (series A1), indicated exceptionally low extent of melting which occurred by softening or sintering of particles as shown in Fig. 4B, and consistent with past observations [Vassilev et al., 1995]. Irrespective of mechanism that is either by sintering or reaction, substantial particle

agglomeration and melting was observed at T25%. The viscosity at the T25% would be sufficient to relate this stage of melting to sticky nature of particles which caused this shrinkage. Therefore, it is reasonable to propose T25% as the initial melting temperature of ash. No specific correlation between T25% and ratio is observed. However, a slight presence of fluxing component, or and CaO, was found to decrease the T25% values significantly. It appeared that at T25%, a small amount of liquid is formed by the low melting or decomposition point minerals (e.g. illite and anhydrite), and is detected in the TMA test due to the high sensitivity of measurement. The liquid formed at this stage is not sufficient to cause a visible physical change, and therefore can not be observed in a conventional ash fusibility test.

4.2.2. Intermediate Melting Temperature: T50%. T50%, defined as the temperature related to 50% shrinkage, corresponds to approximately 60% melting as shown in Fig. 5. Group A ashes indicated a high temperature range for T50%, being greatest for series A ashes that contain least amounts of fluxing oxides. For the Group B ashes, T50% occurred at low temperature and was associated with a higher extent of melting in most of the ashes. In general, ashes with higher ratio and higher fluxing contents would give lower values of intermediate melting temperature.

164

S. K. Gupta et al.

4.2.3. Melting Temperature: T75%. For all practical purpose, 80% (mass fraction) or greater extent of melting may be considered as an approximation for the completion of melting. Therefore, T75% is proposed as the melting temperature of ash for practical purpose. The SEM examination of ash pellets quenched at T75% indicated the presence of 80% melt phase i.e. 80% mass fraction of the particles were melted (Fig. 5). In subsequent discussion, the melting temperature will refer to this state of melting. For Group A ashes (refractory ashes) T75% did not occur up to 1,600°C except for the ashes with high ratios (>3.5). Generally, the T75% for Group B ashes was low, ranging from 1,100°C to 1,600°C. Similarly to T50%, the T75% is also influenced by the ratios and amounts of fluxing components. Ashes with high ratios will result in low values of T75% for Group A ashes. For iron rich ashes from Group B, the significant fluxing effect of iron oxide appears to be responsible for very low values of T75%. The results will be further discussed in section 4.3. 4.2.4. Flow Temperature: T90%. T90% is proposed to represent the final stage of melting (melt phase >80%) and to indicate the slag flow characteristic of ash. Flow temperatures of Group A ashes were found to be usually greater than 1,600°C except for AEC10 and EN3 ashes, which contain significant amount of In general, the T90% of Group B ashes were less than 1,600°C. T90% decrease consistently with an increase in ratio and fluxing levels except for some ashes that contain very high concentration of fluxing components (especially CaO).

4.3. Comparison of Alternative TMA and Existing AFT Temperatures T25% and DT represent different events in the process of ash sintering and melting. T25% indicates the presence of small amounts of liquid phase, which can be detected by the sensitive penetrating ram. At the deformation temperature, an extensive melt phase is expected to cause a visible change in the shape of ash specimen. Therefore, the initial melting temperature from the TMA test did not find any equivalent temperature in the existing ash fusibility test. T25% was invariably less than the deformation temperature.

4.3.1. Intermediate Melting Temperature (T50%) and Deformation Temperature (DT). The Australian Standard [1995] allows a maximum acceptable difference of up to for repeatability and up to for reproducibility of deformation temperature. In practice, this difference was found to vary up to 300°C for some ashes [Wall, et al., 1995]. The deformation temperature for all the ashes vary from 1,100°C to 1,600°C, and is higher for group A ashes. In general, more than 60% melting was observed at the deformation temperature, however in exceptional cases, it might vary from 20% to 75% as shown in Fig. 4. Similar melting range has also been related to deformation temperature in past. [Huffman, 1981; Vassilev et al., 1995]. The T50% and deformation temperature were observed to represent a similar range of melting, which is approximately 60% as shown in Fig. 5. In addition to precision problem associated with the observation of deformation temperature, it does not represent the beginning of ash melting. Therefore, the common belief that deformation temperature is associated with surface stickiness may not be totally justified [Ely and Barnhart, 1963]. Figure 6 shows that apparently there is no direct correlation between deformation temperature and intermediate temperature from the TMA technique. However, in more than 60% of ashes, T50% was observed

Thermomechanical Analysis and Alternative Ash Fusibility Temperatures

165

within

of the deformation temperature. T50% of some refractory ashes (CaO + < 10%) were very high compared to their respective deformation temperature and very low for a few Group B ashes which contain combined calcium and iron oxides greater

than 10%. The exact reason of this deviation is not clear at this moment. It appears that this deviation is not related to the normal inaccuracies associated with measurements of deformation temperature. However, the T50% from the TMA technique would still be a

better choice for measuring the intermediate melting due to a greater precision of repeatability 4.3.2. T75% and Hemispherical Temperature. The physical changes in ash specimen that occur at hemispherical are clearly visible, and are often used to indicate the completion of melting for practical purposes. SEM analysis of ash pellets in present study also indicated the presence of approximately 80% liquid at hemispherical temperature and T75%. Figure 7 compares T75% and hemispherical temperature. Ashes, OPC01 and AEC09, indicate very low values of T75%, which may be due to the high levels of fluxing components. Figure 8a and 8b show the effect of iron oxide on T75% and hemispherical temperatures for series Bl and B2 ashes, which have their ratios greater and less than 2.2 respectively. Both temperatures are found to decrease with an increase in iron oxide content provided iron oxide is less than 35%. The ash composition of samples, which contain iron oxide greater than >35%, is shifted to hercynite primary field that results in higher liquidus temperature. T75% of series Bl ashes is usually higher than series B2 ashes. In general, T75% and hemispherical temperatures followed the liquidus temperatures from phase diagrams (Fig. 8). Therefore, it is reasonable to define the T75% as the melting temperature of ash for a practical purpose.

4.3.3. T90% and Flow Temperature. Figure 9 shows that the correlation between T90% and the flow temperature is least satisfactory. For instance, T90% is very high for ashes

166

S. K. Gupta et al.

that contain very low levels of fluxing oxides from Group B (e.g. AEC11 and AEC18) and very low for those containing high levels of fluxing contents when compared with

their respective flow temperatures. T90% is very low especially if CaO content in ash high (series B3). This large deviation may be related to the large error associated with measurements of flow temperature or due the amphoteric nature of iron oxide that may not affect the fluidity of the melt consistently, and hence the T90% values. The T90% may be used to measure the slag flow characteristics of the ash.

Thermomechanical Analysis and Alternative Ash Fusibility Temperatures

167

5. CORRELATION OF MELTING WITH TMA AND AFT TEMPERATURES In the preceding section, it was shown that substantial melting occurred at the deformation temperature as the sample matrix did not deform due to the high viscosity of silicate slag. Therefore, any observer has to wait to record the first sign of deformation till an extensive melting is developed in the sample. Sphere, hemispherical and flow temperatures from conventional test are further believed to denote a similar degree of

melting, and differing only in the viscosity of the melt phase [Huffman, 1981 and Bryers, 1996]. The subjective nature of conventional test makes it further difficult to assign a specific level of melting with various ash fusibility temperatures.

The progressive shrinkage in the TMA technique basically measures the ease of relative movement of ash particles and the penetrating ram. At low temperatures, the ram movement into ash is primarily affected by various stages of ash melting i.e. sintering,

decomposition of clays and chemical reactions among various fluxing oxides. At higher temperatures, the TMA shrinkage is more dominated by the melt composition that changes by the continues dissolution of the refractory components. Therefore, TMA shrinkage depends on the proportion of melt, melt composition and temperature. The relationship between the TMA shrinkage and extent of melting can be expressed as

where

K is defined as the constant of proportionality, and is a function of ash chemistry and temperature. Evaluation of the extent of melting indicates that K is greater for refractory ashes compared to other ashes those contain reasonable amounts of fluxing oxides. Thus, we can see that the TMA technique has a potential for quantifying the various levels of melting with an improved accuracy, which is a crucial requirement for a reliable prediction of ash deposition tendencies of coal. For instance, in a pf combustion system, the proportion of melt phase from any ash fusion test would indicate the proportion of molten particles in the furnace at a particular temperature. Therefore, the T25%, which is significantly less than the existing deformation temperature, would be the temperature

168

S. K. Gupta et al.

at which sufficient proportion of particles might be sticky and cause ash deposition

problem. It is worthwhile mentioning that slagging performance of some pf combustion systems indicated a good correlation with this level of shrinkage [Wall et al., 1995). T90%

is expected to indicate the slag flow characteristics of ash and hence provide a basis for another temperature of interest often required in other applications such as entrained flow slagging gasifier. For instance, T90% may be related to the temperature at which

ash will transform to a completely molten layer in a gasifier. However, the exact viscosity at which this molten layer may flow is not clear. The correlation of the extent of melting with the extent of sticky particles and slag viscosity in various combustion systems will enable the new temperatures to characterise slagging and fouling with improved reliability.

6. CONCLUSIONS The traditional measurement of the temperatures of fusibility of coal ash, which are called ash fusibility temperatures (AFT), are shown in correspond to the existence

of more than 60% of melt phase in the samples. These temperatures do not appear to correspond to the ash melting characteristics often associated with them. It was found that the deformation temperature is not the temperature at which initial melting begins as normally perceived and the hemisphere temperature is below the liquidus temperature. A new technique—Thermomechanical Analysis (TMA)—measures the progressive shrinkage of ash and is shown to be capable of characterising the sintering and melting behaviour at temperatures lower than the traditional technique. Shrinkage temperatures may be defined which corresponding to particular shrinkage levels are denoted as TS% (for S of 25%, 50%, 75% and 90%), and these are proposed as alternatives to the existing ash fusibility temperatures. The new temperatures are measured with much better accuracy than the traditional temperatures and are suggested as indicators of initial melting (25%), intermediate melting (60%), completion of melting (80%) and slag flow characteristics. Correlation between the alternative shrinkage temperatures, the traditional ash fusibility temperatures and the measured extents of melting indicated that the shrinkage temperatures provide an improved and objective method of quantifying the various stages of melting. The correlation of the extent of melting with the extent of sticky particles and slag viscosity will allow the application of the new shrinkage temperatures to characterise ash effects in combustion systems.

7. ACKNOWLEDGMENTS The authors wish to acknowledge the support of the CRC for Black Coal Utilisation which is funded in part by the CRC Program of the Commonwealth of Australia. We are most grateful to Mr. Dick Sanders of Quality Coal Consulting Pty Ltd for managing the project, selecting samples and providing insights in to the AFT procedures. The study is also supported by the Australian Coal Association Research Program with Mr Grant Quinn of BHP Coal Pty Ltd as industry monitor. The assistance of Dr John Saxby at CSIRO and Dr R. A. Creelman is greatly appreciated.

Thermomechanical Analysis and Alternative Ash Fusibility Temperatures

169

8. REFERENCES Australian-Standard AS1038-15-1995. Coal and Coke Analysis and Testing, Part 15: Higher Rank Coal Ash and Coke Ash Fusibility. Barret, R. E. (1987). “Slagging and fouling in pulverised-coal-fired utility boilers—Volume 2: A survey of boiler design practices for avoiding slagging and fouling”, EPRI Report CS-5523. Bryers, R. W. (1996). “Fireside slagging, fouling and high temperature corrosion of heat transfer surface due to impurities in steam raising fuels”, Progress in Energy Combust. Science, Vol. 22, pp. 29-120. Ellis, G. C. and Ledger, R. C. (1989). “The Thermomechanical, electrical conductance and chemical characteristics of coal ash deposits”, In “SECV (R&D) Report SO/89/164, NERDDP Project No. 1181, SECV Project No. 2411, Australia. Ely, F. G. and Barnhart, D. H. (1963). In: H. H. Lowry (Ed.), Chemistry of coal utilisation, Supplementary volume, Wiley, New York, p. 820. Huggins, F. E., Kosmack, D. A. and Huffman, G. P. (1981). “Correlation between ash-fusion temperatures and ternary equilibrium phase diagrams”, Fuel, 60, 577–584. Huffman, G. P., Huggins, F. E. and Dunmyre, G. R. (1981). “Investigation of the high-temperature behaviour of coal ash in reducing and oxidizing atmospheres”, Fuel, 60, 585–597. Juniper, L. (1995). “Applicability of ash slagging “Indices” Revisited”, Australian Combustion Technology Centre, February, Combustion News p. 1–4. Kalmanovitch, D. P. and Williamson, J. (1986). “Crystallization of coal ash melts”, In Karl Vorres (Ed.), Mineral Matter and Ash in Coal, ACS Symposium series 301, August, 1984, pp. 234-255. Levin, E. M., McMurdie, H. F. and Hall, H. P. (1964), “Phase Diagrams for Ceramists” Am. Ceramic. Soc., Inc., Columbus, Ohio and references therein. Raask, E. (1979). “Sintering characteristics of coal ashes by simultaneous dilatometry-electrical conductance measurements”, Thermal Analysis, 16, p. 91. Raask, E. (1986). “Flame vitrification and sintering characteristics of silicate ash”, In “Mineral Matter and Ash in Coal”, ACS Symposium series 301, in (Ed. Karl Vorres), p. 139. Sanyal, A. and Mehta, A. K. (1994). “Development of an electrical resistance based ash fusion test”, In Williamson, J. and Wigley, F. (Ed.), The Impact of Ash Deposition in Coal Fired Furnaces, Taylor and

Francis, Washington, p. 445. Saxby, J. D. and Chatfield, S. P. (1996). “Fusion of coal ash by thermo-mechanical analysis”, Proceedings of the 7th Australian Coal Science Conference, Australian Institute of Energy, p. 391, Australia. Vassilev, S. V., Kitano, K., Takeda, S. and Tsure, Tl. (1995). “Influence of mineral and chemical composition of coal ashes on their fusibility”, Fuel Processing Technology, Vol. 45 pp. 27–51. Wall, T. F, Creelman R. A., Gupta, R. P., Gupta, S. K, Sanders, R. H. and Lowe, A. (1995), “Demonstration of the true ash fusion characteristics of Australian thermal coals”, ACARP Project Final Report C3093, Australia.

This page intentionally left blank

ASH FUSIBILITY DETECTION USING IMAGE ANALYSIS Klaus Hjuler dk-TEKNIK Energy & Environment Gladsaxe Moellevej 15 DK-2860 Soeborg

Denmark

INTRODUCTION In the recent years the use of biofuels for power production has gained increasing

importance as a substitute for coal or by being co-fired with coal. This development is particularly forced by the international concern about antropogeneous carbon dioxide emissions. However, the use of biofuels for steam raising or in gas turbine cycles is so far restricted by the fact that biofuels generally have a higher content of potentially deposit

forming and corrosive elements than coal, especially on a heat value basis. The international standard method of estimating the deposit propensity of solid fuels, of which a number of variants exist (e.g. ISO, ASTM, AS, DIN), was originally proposed to estimate the suitability of a particular coal for grate firing. The result from the test is valuable for comparing new coal qualities with known coals that behaves satisfactorily in a specific plant. However, the standard fusion test has shown to be unsuitable for ashes from biomass fuels as the ash test body may “blow up” like a baloon or melt may flow out from it without overall changes in shape, leaving a “skeleton” composed of e.g. silicon and calcium. This complicates the interpretation and reporting of the test, especially for automatic equipments where the characteristic temperatures are determined only from the height and width of the test specimen. Moreover, the standard fusion test is more or less based on a subjective evaluation of the change in shape—i.e. the silhouette—of a test body (cone or cube) while this is being heated. Consequently the reproducibility as well as the repeatability is poor; even a skilled operator cannot obtain a repeatability better than about (ISO/TC, 1991). Another serious problem is that the appearance of the first melt is not detected because it takes place “inside” the test body and does not necessarily affect the shape of the body. The appearance of the first melt is important because the presence of a molten phase increases the probability of ash sticking significantly. Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al.

Kluwer Academic / Plenum Publishers, New York, 1999.

171

172

K. Hjuler

The literature has been reviewed elsewhere [Coin et al., 1996; Hansen et al., 1997; Wall et al., 1996] and is therefore not discussed in this paper. However, it should be mentioned that it is well known that visual observation of an ash sample using light microscopy and a heating stage may reveal details in the process of ash shrinkage and melting during heating (e.g. [Vassilev et al. (1995]). Similar equipment is used in metallurgy for studying recrystallization and sintering processes, where the sample are observed using incident or transmitted light or a combination of both. The main task of the present work has been to develop a method to quantify the information.

EXPERIMENTAL

Apparatus The apparatus is commercially available: the microscope consists of an Olympus SZ 1,145 TR stereo microscope body with eyepieces for magnification, a base illuminator, and a fibre optic illuminator. The camera is a Sony RGB XC-711 CCD and the heating stage is a Leitz 1350, which is suitable for examinations in transmitted and incident light (Fig. 1). The heating stage is powered by a Heinzinger LNG 16–30

power unit. The image grabber is a Neotech colour video digitizer for personal computers. The software was programmed specifically for the image analysis and data acquisition, and for the operation of the heating stage.

Samples More than 80 samples have been subjected to the HTLM fusion test. As the main focus in the present work was posed on developing a method which is well suited for

ashes from biofuels, mainly fly ash, bottom ash, and deposits from straw firings have been investigated as well as laboratory prepared biofuel ash. Repetitions of runs with the above mentioned samples and tests with an inert (below 1,250°C) quartz fiber sample, a geological standard, and analysis quality salt mixtures have also been performed. All runs were performed at a rate of 10°C/min in a nitrogen atmosphere.

Analysis Procedure Before ashing, coal samples are milled to a maximum particle size of whereas biofuels are milled to a particle size of less than The milled fuel sample is ashed in a Pt-boat in air at 815°C for coal and at 550°C for biofuels. The residual

Ash Fusibility Detection Using Image Analysis

173

carbon content of laboratory prepared ash is lower than 5%(w/w). Ash samples originating from a combustor are normally glown for determination of the residual carbon content, that is, the glown ash is used for fusion testing. Fly ash samples do normally not require milling, while most other ash types do. The maximum particle size should be less than about in order to observe a representative number of particles/agglomerates. Coals are milled to a maximum particle size of in an agate mortar as described in the standard ISO 540 method. Ashes from biofuels are milled in the same way as coal ashes. The quantity of sample needed is small, typically about (or about 1– ). In special cases information about the local melting behavior of eg. a deposit is needed. In this case a small piece, say may be broken off and analysed directly. The sample is placed randomly on a 7mm diameter sapphire specimen disc by means of e.g. a small spatula. The sapphire disc with sample is placed in the sample holder of the heating stage. When a glass cover is mounted the heating stage is gas tight and can be operated in a controlled atmosphere (normally nitrogen). The microscope is focused on a random part of the ash, the magnification is set at typically and the light conditions are optimized. The digital camera and the heating stage is operated from a personal computer. The temperature of the heating stage is ramped at typically 10°C/min from an initial temperature of 550°C for ashes from biofuels and 800°C for coal ashes. The ash sample is photographed initially and then every fifth second (0.2Hz). The image acquisition rate

can be adjusted according to the rate of heating/melting, but 0.2Hz has been found appropriate for 10°C/min. Each image is processed before the next image is acquired. The results are written to a text file. The power for the heating stage is automatically interrupted when the sample is completely melted or when the temperature exceeds 1,250°C (present maximum operation temperature).

Image analysis The image analysis is based on grey level images and binary images. In the binary images, the ash sample appears black and the background white. The resolution is 288 pixels, which gives a manageable file size of 1 1 1 kB. At a maximum magnification

of the corresponding specimen area is and the resolution is At the typical magnification of l00×, the corresponding figures are and respectively. Each grey level image that is acquired during heating of the specimen is converted to a binary image using either a constant threshold value or a subroutine, which calculates a proper threshold value. By using the last method one accounts for an eventual varying light intensity. In the binary image the solid part of the ash appears black and the melt and background appears white. This allows the total area covered by the solid part to be calculated. In addition a difference image is created, which contains information about areas in the actual image that have changed from black (B) to white (W). The

analysis result comes as the following two fractions: = Sample area in the actual image/Initial sample area, and = Total area of B-to-W changes/Initial sample area. Areas are measured in pixels and t is the actual temperature in °C. Ideally, the area fractions

are complementary, i.e.

174

K. Hjuler

Examples of a grey scale image of a laboratory prepared straw ash and the corresponding binary images acquired during melting are shown in Figs. 2 and 3, respectively.

Calibration The heating stage is calibrated with materials, which do not form a protective oxide layer. Suitable materials are (394°C), (593°C), LiF (870°C), Ag (961°C) og Au (1,063°C), where the numbers indicate melting points. The heating stage

applied has a maximum operation temperature of 1,350°C measured at the specimen glass holder. Due to heat transfer primarily through the top glass seal it has been found that the maximum operation temperature corresponds to a specimen temperature of about 1,250°C. 5-point calibrations performed using the above mentioned materials show that the temperature measured at the specimen holder is comparable to the specimen temperature within up to about 830°C. Above 830°C the specimen temperature is corrected using a third order polynomial approximation.

RESULTS As mentioned previously tests were conducted using inert quartz glass fibers at a heating rate of 10°C in nitrogen. Figure 4 indicates that the parameter increases

significantly in the range 500–600°C and is not further increased above about 800°C. The cause of this error is movements of the specimen holder relative to the microscope

Ash Fusibility Detection Using Image Analysis

175

objective caused by thermal stresses during heating. The area fraction is not as sensitive as in this respect due to the fact that the area of fibers moving into the image field

is approximately equal to the area moving out of the field as the fibers are placed randomly. In principle, does not detect sample movements, whereas detects any movements. Therefore, the melting curves discussed in the following are based on the area fraction by plotting versus the temperature. Also seen in Fig. 4 are the melting curves for 100% analysis quality KCl and (melting points 774°C and 1,069°C, respectively). Initially the melt fractions increase slowly to about 8–10%, which is believed

as being neither due to melting and particle shrinkage, nor particle movements. It is not yet clear why area reductions are detected initially with pure crystalline materials. Figure 5 shows the measured melting curves of a mixture (approx. 25%w/w KCl) and a mixture (approx. 79%w/w KCl). The mixtures were prepared from analysis quality reagents, which were milled in a mortar. Also shown are predictions obtained from the relevant phase diagrams. It is seen that the temperatures of significant melt formation are correctly identifyed. Initially the melt fractions increase slowly as discussed above. When more than about 60% melt has formed the measured results deviate from the predicted. These deviations can be reduced by optimizing the method of threshold value calculation. The threshold value has particularly importance when the melt is not totally transparent. Also shown in Fig. 5 for comparison is a melting curve for fly ash sampled from a straw fired combined heat and power plant. The ash contains about 58%(w/w) KCl, 18%(w/w) 9%(w/w) Si as and 5%(w/w) Ca as CaO. The initial part of the curve is quite similar to that of the salt mixture, but in this case visual inspection shows clearly that the ash shrinks before melting. The “tail” of the melting curve is caused by the presence of Si and probably also Ca. The sample was completely melted at 1,015°C (in nitrogen). This particular sample was also subjected to a Round Robin standard

176

K. Hjuler

method test (ISO 540). As expected, the results (Table 1) confirmed some of the problems about using the standard test for ashes from biofuels. An important point is the repeatability of the method that may be influenced by the quantity and distribution of sample on the specimen disc, the magnification, and lighting conditions. Figure 6 shows results obtained when melting a laboratory straw ash,

a superheater deposit from straw firing, and a geological standard (BCR-l).Each run was repeated two, two, and three times, respectively. BCR-1 is considered as being very homogeneous, and with this material the magnification was varied from 50 to to enhance any magnification effects on the results. The resulting melting curves of BCR-1 are nearly identical. Generally, the deviations are greater, but still acceptable, with ashes as exemplified by the straw ash and the deposit. The decrease in the melt fractions in some

parts of the curves are caused by rearrangement of solid particles due to flow of melt.

Ash Fusibility Detection Using Image Analysis

177

This is primarily observed when the quantity of melt produced initially is large as

observed with the deposit. Finally, an interesting feature of the method is that material evaporating from the sample may condense on the colder glass cover of the heating stage. This phenoma has frequently been observed when analyzing samples with relatively high salt contents. The resulting reduction in light intensity is tolerated by the method, except when the quantity of condensated material is very large. In this case the top glass seal can be rotated to a clear part. The condensing particles can be analysed using SEM-EDX directly on the top glass seal. Figure 7 shows an example of KC1 particles condensed in the temperature range 700–800°C during analysis of a fly ash from straw combustion.

DISCUSSION Two fundamental features of the ash fusion determination method presented in this paper is that the sample is placed randomly, and that the fusion characteristic is detected by changes in the area covered by the sample. It is implicitly assumed that changes in the area fraction are related to changes in the solid volume fraction or mass fraction. In the following this is shortly discussed. If A is the measured area, h m the mean height and e the porosity, then the solid volume of ash observed in the image field at a given time may be given by:

Solid volume at time t: Initially during heating a reduction in to about 0.95–0.90 due to ash shrinking is typically observed. However, the solid volume is unchanged, and that means that

178

K. Hjuler

the porosity must decrease (the height does not increase). It is also observed that melting mainly takes place from below due to heat transfer from the specimen disc. In this case A is constant but h decreases, approaching zero as the material is melted. In reality shrinking and melting from below take place simultaneously, and consequently the effect

of height reduction may be cancelled by a reduction in the porosity. Based on these qualitative considerations it can be argued that changes in the volume fraction may be approx-

imated by changes in the area fraction. The volume fraction is again related to the mass fraction. These findings are supported by comparison of results obtained with the HTLM metod presented in this paper with results obtained using STA (combined TGA and DSC) [Hansen, 1997].

CONCLUSION The ash fusibility detection method using image analysis which has been presented in this work is very simple and is suitable for routine analysis. The equipment is robust and stable. The number of samples which can be run per day is high due to the uncomplicated sample preparation and a high cooling rate of the heating stage, which has a low heat capacity. Moreover, the resulting fusion characteristics are produced automatically directly from the measurements without using any expert knowledge or corrections. The method of image analysis can be refined depending on the application. If the sample is highly agglomerated as biomass ashes typically are, an analysis based on detection of changes in the area covered by the sample is suitable. If agglomeration is not predominant, single particles can be counted, measured, and observed when melting. In this way it is possible to evaluate the melting behavior as function of particle size. Normally details of ash morphology are not reported, but it is also possible to quantify such information. This work was primarily initiated due to the “problematic” melting behavior of biomass ashes with respect to the standard method. Consequently a maximum heating stage operation temperature of 1,250°C was not totally unacceptable. But for coal ashes it is, and work is in progress to extend the operation temperature to about 1,500°C. In addition, more samples will be analyzed in order to establish a data base of fusion characteristics and in order to obtain more knowledge about the relationship between mineral composition and melting behavior.

Ash Fusibility Detection Using Image Analysis

179

ACKNOWLEDGMENTS This work was made possible by financial support from ELKRAFT, ELSAM, and the Danish Energy Research Programme. This support is gratefully acknowledged.

REFERENCES Coin C. D. A., Kahraman H., and Peifenstein A. P. (1996). “An Improved Ash Fusion Test”. Proceedings of the Enginering Foundation Conference on Applications of Advanced Technology to Ash-Related Problems in Boilers, July 16–21, Waterville Valley, New Hampshire. Eds. L. Baxter and R. DeSollar (1996). Hansen L., Frandsen F., and Dam-Johansen K. (1997).

This conference proceedings. Vassilev S. V., Kitano K., Takeda S., and Tsurue T. (1995). “Influence of Mineral and Chemical Composition of Coal Ashes on their Fusibility”. Fuel Processing Technology 45, 27 51. 1SO/TC 27/SC 5/WG 6 (1991). “International Round Robin for Determination of the Fusibility of Coal Ash”. ISO 540 (1981). “Coal and Coke—Determination of Fusibility of Ash”. International Standard Organisation. Wall T. F., Creelman R. A., Gupta R. P., Gupta S., Coin C., and Lowe A. (1996). “Coal Ash Fusion Temper-

atures—New Characterization Techniques, and Associations With Phase Equilibra”. Proceedings of the Enginering Foundation Conference on Applications of Advanced Technology to Ash-Related Problems in Boilers, July 16–21, Waterville Valley, New Hampshire. Eds. L. Baxter and R. DeSollar (1996).

This page intentionally left blank

ASH FUSION QUANTIFICATION BY MEANS OF THERMAL ANALYSIS Lone A. Hansen, Flemming J. Frandsen, and Kim Dam-Johansen Department of Chemical Engineering Technical University of Denmark 2800 Lyngby, Denmark

1. INTRODUCTION The amount of melt present in an ash as a function of its temperature greatly influences the ash deposition propensity in thermal fuel conversion systems. The appearance of melt is believed to increase both the tendency for ash particles to stick to heat transfer surfaces [Srinivasachar et al., 1990; Walsh et al., 1990; Benson et al., 1993;

Richards et al., 1993] and the rate of strength build up in ash deposits [Skrifvars et al., 1996; Benson et al., 1993]. For years, laboratory tests have been carried out on fuel ashes to estimate their melting behaviour, and results have been used to estimate the slagging and fouling propensity of the ashes in full scale combustion systems. Laboratory tests used to estimate the melting behaviour of ashes include a variety of methods. Commonly used are the conventional ash fusion tests of which many variants appear [ISO 540, 1981; DIN 51730, 1984; ASTM D1857, 1987; AS1,03815, 1987]. These methods all imply the controlled heat up of an ash specimen of well defined shape,

and the simultaneous determination of temperatures corresponding to specified geometrical shapes. The main criticisms of these tests have been their low reproducibility and unreliability in the subsequent prediction of the ash behaviour in real boilers [Coin et al., 1996]. It has been emphasized that the initial deformation temperature is not the temperature at which the ash melting begins, and many coal ashes have been found to start melting at temperatures far below the initial deformation temperatures [Huffman et al., 1981; Huggins et al., 1981; Coin et al., 1996; Wall et al., 1996]. Alternatively, the ash melting behaviour has been estimated based on electrical resistivity [Raask, 1979; Sanyal and Cumming, 1981; Gibson and Livingston, 1992; Sanyal and Mehta, 1993] or conductance measurements of the ash [Cumming and Sanyal, 1981; Conn and Austin, 1984; Cumming et al., 1985] during heat up. These electrical quantities reflect the conduction path through the ash sample, and thereby the particle-particle contact and fusion. Both methods detect the onset of fusion in the ash as the temperature at which the electrical properties of the ash is drastically changed. The electrical conductance methods have higher repeatabilities than the standard ash fusion tests Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

181

182

L. A. Hansen et al.

and give better predictions of field slagging performance [Sanyal and Mehta, 1993]. However, these methods include some practical difficulties since satisfactory contact between ash and electrodes is hard to achieve and maintain [Wall et al., 1996]. Furthermore, the results contribute primarily with information on the onset of fusion and sintering, whereas the further melt quantity increase in the ash is harder to evaluate based on these methods. Recently, an improved ash fusion characterisation method based on dimensional changes of ash pellets during heating has been reported [Coin et al., 1996]. In this test four ash cylinders are used as pillars to separate two alumina disks. As the assembly is heated, the ash pellets shrink and the distance between the tiles is measured. Significant tile movement over a narrow temperature range is interpreted to correspond to melting of distinct chemical species, and the repeatability and reproducibility of the method is reported to be high, with reproducibilities below for significant tile movement Finally, ash melting behaviour can be estimated based on calculations. The fusion temperatures may be estimated by combining and weighting the effects of several compositional variables [Winegartner and Rhodes, 1975; Vorres, 1979; Gray, 1987; Lloyd et

al., 1989; Vassilev et al., 1995], or by use of chemical equilibrium calculations [Backman, 1989]. As indicated above, the estimation of melting behaviour of coal ashes, (and the subsequent prediction of ash behaviour in real boilers) is not a simple job. Still more problems arise, when trying to do the same job for biomass ashes. The chemical composition of biomass (i.e. in Denmark mainly straw) is very different from that of coal and thus the same kind of analyses that are useful for characterising coal ashes do not necessarily apply for biomass ashes. The standard AFT has shown to be unsuitable for ashes from biomass combustion and biomass laboratory ashes [Westborg, 1995]. Thus, the present work was initiated to generate a new method to quantify the melting behaviour of biomass ashes in order to improve the understanding and prediction of ash deposition propensities during firing of biomass. In a longer term, the aim was to apply the method also to coal ashes.

2. EXPERIMENTAL

2.1. Apparatus The new method for estimation of ash melting behaviour is based on Simultaneous Thermal Analysis, STA, and the results presented in this paper were obtained using a NETZSCH STA409. STA implies continuous measurement of sample weight (Thermogravimetric Analysis, TGA) and temperature (Differential Scanning Calorimetry, DSC) during heat treatment. The weight measurement reveals any mass changes taking place in the sample and by comparing the sample temperature to the temperature of an inert reference material, any heat producing or heat consuming (chemical or physical) processes occurring in the sample is detected, and the involved energy subsequently quantified.

2.2. Test Method STA was carried out on the ash samples, while heating them from 20 to typically atmosphere. On the resultant STA curves, melting is

1,390°C at 10°C/min in a

Ash Fusion Quantification by Means of Thermal Analysis

183

detected as an endothermic process involving no change in mass. Melting of a pure substance would be seen as a single endothermic peak in the DSC signal, while for “real” ashes, the melting results in several endothermic peaks overlapping each other, corresponding to melting of the different chemical species in the ash, which melt at different temperatures. Conversion of the STA curves into a melting curve is done based on a DSC signal reflecting only melting energies, which implies that energies related to other processes than melting are first subtracted from the (raw) DSC signal. Evaporation is typically occurring simultaneous to part of the melting, and evaporation enthalpies thus typically need to be quantified and subtracted. Evaporation energies are quantified as the product of a reasonably estimated evaporation enthalpy and the derivative of the TG-curve. After subtraction of evaporation enthalpies, the melting curve calculation can be carried out in one of the following two ways.

The total area below the melting curve, i.e. the area below the DSC curve from the first point where melting is detected, [Backman, 1987], and to the temperature where the melting is completed, reflects the total energy consumed for melting of that ash (Fig. 1). Calculating the area below the DSC curve from any temperature to and dividing this area, AA–B by the total area below the melting curve, the fraction of “total energy used for ash melting” which has been used in the specific temperature interval, is obtained. This energy fraction is a simple quantitative estimate of the mass fraction of ash melted in the specified temperature interval. This estimate is only correct, if the melting enthalpy of all species in the ash are alike, which is not necessarily true. The presented method thus is a simple way of determining the melting behaviour of an ash and the result expresses the melting behaviour in what could be termed an “energy-percentage” of melt as a function of temperature. Alternatively, a method based on quantitative determinations of the peak areas can be used. Any given peak below the melting curve corresponds to an absolute quantity of energy used for melting. The position of the peak (onset and peak temperature) gives an indication of the identity of the melting substance(-s), i.e. a reasonable estimation of the relevant melting enthalpy can be made. Based on these two figures, the mass of material melted in the given temperature range can be calculated, and by relating this mass to the

L. A. Hansen et al.

184

total mass of ash analysed, the mass fraction of ash melted in the given temperature range is obtained. The latter method gives the most correct estimates of “ mass fraction melted”, but this method implies that the substances melting at the different temperature intervals can be identified, so that a reasonable estimation of the involved melting enthalpies can be made, unless the species present in the ash have got equal melting enthalpies. The latter method thus typically implies an identification of the chemical species present in the ash (as provided by e.g. CCSEM) and detailed knowledge on the chemistry between the ash species.

3. RESULTS

3.1. Simple Systems First, the melting behaviour of two simple mixtures consisting each of only two chemical species will be presented. 3.1.1.

A sample of approximately 85 mole-% KC1 and 15 mole-% (

was prepared and analysed in the STA409. The resultant STA curves are shown in Fig. 2. The STA curves show evaporation (of crystal water from ) at 20–175°C. This is detected by 1) a decrease in mass (TG) and 2) a consumption of energy (the upward DSC peak). At 594°C, a new endothermic peak starts, peaking at 600°C, but also having a long “tail” behind it, so that the peak is not quite ended until the temperature has reached 696°C. The peak corresponds to the formation of a considerable

quantity of melt at the eutectic temperature, and the “tail” corresponds to the continuous increase in melt quantity as the temperature is raised from the eutectic temperature and to the liquidus temperature. Above 700°C, the DSC-signal is greatly increased due to the evaporation of KCl (seen as the decrease in the TG curve). This experiment was repeated three times. In Fig. 3, a comparison is made between the melting behaviour obtained by using the lever rule in the phase diagram [Levin et al., 1964] and the ones obtained when quantifying and comparing the areas below the DSC curve for each of the three experiments. It is seen that the “theoretical” curve predicts the melting to take place instantaneously, which is not happening in reality, but except for this, a very good correlation between

Ash Fusion Quantification by Means of Thermal Analysis

185

“theoretically” and experimentally obtained melting curves is found. The deviation between the two types of curves is 10% at maximum (neglecting the temperature range, 594–608°C, just above the eutectic temperature). Melting onset and completion deviate less than respectively 3 and 15°C from the phase diagram values. The method repeatability data is given in Table 1. Repeatability is seen to be very good, deviances are 4°C for melting onset, 4% for the melt fraction obtained at the eutectic temperature (calculated at 608°C, which is the end of the large peak), and 11°C for melting completion. Concerning the completion temperatures, the phase diagram shows that increasing KC1 fractions increases the liquidus temperature, and this tendency is also found in the experimental results. Concerning the measured melting enthalpies, a very good correlation to theoretical calculations is seen: the measured energies correspond to between 99.0 and

99.7% of the theoretical value. The theoretical and experimental description of the melting behaviour are thus judged to correlate well in this case. 3.1.2. _ The same analysis and comparison was made for a mixture of KCl and the result of which is shown in Fig. 4. As can be seen, the experimental curve overestimates the quantity of melt formed at the eutectic temperature with approximately 4% compared to the “theoretical” prediction. At temperatures above 700°C, the deviation between experimental and the “theoretical” prediction varies between 4 and 7%, until the curves meet at 860°C. The temperature differences between the transition temperatures given in the phase diagram and the experimentally determined ones are judged to be acceptable. This example therefore confirms the above and supports the assumption that STA measurements are able to describe melting behaviour of simple systems.

L. A. Hansen et al.

186

3.2. Ash Samples Several fly, bottom and deposit ash samples from 1) grate fired units firing pure straw and 2) PF-fired boilers co-fired with straw and coal have been investigated. In this paper, examples of the results will be given, representing melting behaviour results for a

fly ash and a bottom ash collected during a test run of different straws at a grate fired

boiler, and a fly ash collected during (pure) coal firing at a PF-fired boiler. The chemical composition of the ashes is given in Table 2, and a reduced data set of CCSEM results showing the dominant species in the ashes in Table 3. The STA curves for the fly ash collected at the straw fired boiler is shown in Fig. 5, where the increase in curve complexibility (compared to Fig. 2) is obvious. Referring to Fig. 5, the DSC curve is seen to show a distinct endothermic peak from 641°C to 712°C corresponding to an energy consumption of 176.1 J / g . As there is no simultaneous decrease in mass (TG curve), this peak corresponds to the onset of the ash melting. For increasing temperatures, a general increase in the DSC signal is seen. On top of this general increase, two distinct peaks are seen, one ranging from app. 920°C to 1,050°C and one starting at app. 1,150°C, which is not completely finished at 1,250°C. The first of these peaks is seen to

Ash Fusion Quantification by Means of Thermal Analysis

187

occur simultaneously to a large decrease in mass, and since the shape of the DSC peak

and the d(TG) peak are quite—but not totally—alike, a large fraction—but not all—of the energy corresponding to this peak is used for evaporation of material rather than melting. As described earlier, the evaporation energies are estimated as the product of an estimated evaporation enthalpy and the d(TG) curve. In this case the evaporation enthalpy of KCl has been used, since KCl constitutes a large part of this ash (app. 40% (w/w)) and is assumed to evaporate at these temperatures. For the last DSC peak, the simultaneous mass decrease is very low, and thus the energy corresponding to the area

of this peak is predominantly used for melting of ash. The general increase of the DSC curve is caused by the fact that when great mass losses are occurring (as for this sample) the DSC baseline is shifted upwards [Netzsch, 1995]. This explanation is supported by the slope of the DSC curve, which is quite higher around the temperatures of rapid evaporation (800–1,050°C) compared to above 1,100°C, where the fast evaporation is ended, and the DSC baseline has found a new level at app. 1.5mW/mg. A typical set of STA curves for a silicate rich ash—e.g. a fly ash produced during coal combustion—is shown in Fig. 6. Nothing significant seems to occur until the DSC curve starts increasing at 1,180°C, which represents the melting onset. The DSC curve continuously increases until the termination of the experiment at 1,390°C, at which point

188

L. A. Hansen et al.

the DSC peak is not ended; i.e. the ash is not completely melted at 1,390°C. Identification of chemical species melting in different temperature ranges above 1,180°C is not possible, why the melting curve is calculated based on area comparison (method no. 1), as is typical for silicate dominated ashes. The fraction of melt formed at 1,390°C is determined by studying the sample structure (after cooling) in the SEM. The material which has not been molten has maintained its original structure, and the melt fraction determination is made based on an area evaluation of material of original structure to that of fused material.

Ash Fusion Quantification by Means of Thermal Analysis

189

In Fig. 7, the melting behaviour calculated on the basis of the STA curves for the three ashes is shown. For the straw derived fly ash, two curves are shown, representing respectively the comparison of areas under the DSC curve (method 1), and the calculation that includes the absolute energy represented by the first melting peak and the relevant melting enthalpy (method 2). The first melting peak for this fly ash is supposed to represent melting in a salt system containing large amounts of KCl and only minor quantities of “other” K- and Ca-salts, since the temperature agrees with the eutectic temperature for these systems and since these species have been found in the ash by means of CCSEM. As it is seen, the curve including the melting enthalpy corresponding to the first melting peak shows a larger fraction to melt at 653 °C than the curve based on area comparisons This reflects that the melting enthalpies for the potassium and calcium silicates melting at the higher temperatures are higher than the melting enthalpy for KCl. Comparing the three melting curves, it is seen that generally the fly ash from straw combustion is “lower melting” than the bottom ash from straw combustion, which is again “lower melting” than the fly ash from coal combustion. Comparing the melting curve for the straw-derived fly ash with that for the straw-derived bottom ash, it is seen that at temperatures between 600 and 1,100°C, the melt fraction is considerably lower for the bottom ash than for the fly ash. This is due to the high content of simple salts in the fly ash. For temperatures between 1,100 and 1,250°C, the bottom ash shows higher melt fractions than for the fly ash. This is probably due to the fact that the bottom ash is highly dominated by K- and Ca-silicates (i.e. K-, Ca-, and Si-rich compounds) whereas the silicate part of the fly ash contains larger fractions of the more refractory quartz [Sørensen, 1997]. The coal derived fly ash starts melting at the highest temperatures, with initial melting at app. 1,180°C, and only partly melting at 1,390°C. This is due to the ash consisting almost entirely of various alumino silicates and quartz [Sørensen, 1997] which melt at relatively high temperatures. In conclusion, the melting curves are thus seen to reflect the different chemical composition of the ashes.

3.3. Repeatability of Melting Curves Figure 8 shows melting curves for the two fly ashes as obtained during repeating experiments. Starting with the straw-derived fly ash, it is seen on the curves, that the reproducibility of the melting onset is very good, within 5°C, as well as is the slope of the first part of the curve. The first curve part corresponds to the very distinct peak occurring at low temperatures (641-712°C in Fig. 5), and since this peak is dependent on the salt chemistry of the sample, and this chemistry is quite simple (i.e. includes only a few possible reactions between the present species), the melting peak occur at precisely the same temperature every time. The melt fraction obtained at these first peaks do deviate slightly, though; in this case the level obtained is 50, 52 and 47% melt respectively. This deviation is caused both by method uncertainty but may also be influenced by the inhomogeneity of the sample. For the rest of the melting curve, the uncertainty is somewhat larger, but still within 10% melt. The larger uncertainty for the last part of the curve is due to the less distinct peaks corresponding to the melting of the silicate part of the ash. Since the peaks corresponding to silicate melting are less distinct, a precise characterisation is dependent on a very well known baseline. For higher temperatures, drift in the DSC baseline cannot be avoided [Netzsch, 1995]. This leads to larger uncertainties for the melting curve. As stated above, the uncertainty is still within 10% melt, though. For the coal-derived fly ash, which consist mostly of silicates, the problem with not very dis-

190

L. A. Hansen et al.

tinct peaks may generally lead to a larger uncertainty for the melting onset, but as can be seen, the repeatability is still quite good: melting onset varies within 30°C, and melt fractions at given temperatures are within 10% melt. To reduce uncertainties from sample inhomogeneity, repeatability was also tested

by analysing a well-characterised and homogenized geological standard material (BCR1). Melting curves for these measurements are also shown in Fig. 8, and reveals that onset is determined with a deviation of 15°C, melting completion temperature with a deviation of 5°C, and the melt fractions at given temperatures deviate at maximum 14% (melt). Based on this, method repeatability is generally judged to be quite high.

4. DISCUSSION 4.1. Correlation to Standard AFT The results of the standard AFT (DS/ISO540) are marked on Fig. 7. For all ashes, melt is formed at temperatures considerably below the IDT, respectively 105, 42 and 65°C for each of the three ashes. This is consistent with previous criticism of the standard ash fusion tests [Huffman et al., 1981; Huggins et al., 1981; Coin et al., 1996, Wall et al., 1996]. For the straw fly ash, the STA predicts significant melt formation (51%) below the IDT. On the other hand, the melt fraction does not increase much for the next characteristic temperatures, since the hemispherical and the fluid temperature corresponds to a melt fractions of respectively 53 and 62% melt. For the bottom ash, the AFT temperatures seem to give a better description of the increasing melt fraction, with the three characteristic temperatures corresponding to respectively app. 3, 14 and 43 (energy) percent melt. For the coal-derived fly ash, the initial deformation temperature corresponds to 5% melt, whereas the hemispherical and the fluid temperature are both higher than the present maximum analysis temperature of 1,390°C and the only relation that can be given is that they correspond to more than 60% melt in the ash. Overall, the standard AFT and the STA melting behaviour curves thus seem to correlate qualitatively, as the three characteristic temperatures are all located in the temperature range corresponding to 3 to 65% melt in the ash. A further and more detailed comparison between the standard AFT and the STA melting behaviour results is presented else where [Hansen, 1997; Hansen et al., 1997].

Ash Fusion Quantification by Means of Thermal Analysis

191

4.2. Correlation of Results with Ash Chemistry (Mineralogical Changes) The DSC signal is closely related to the chemical composition of the ash, since the melting peaks indicate the melting point of either single or mixtures of ash species. Identification of the species melting in the different temperature intervals can thus be made based on CCSEM and/or XRD analysis of the ash and relevant phase diagrams, supplied with STA of simple synthetic ashes, if necessary. For the less distinct “peaks”, the interpretation is correspondingly less certain, but typically wide melting (temperature) ranges can be correlated to the species melting. A detailed comparison between STA melting curves and CCSEM compositional data for ashes collected at straw fired boilers is provided elsewhere [Hansen et al., 1997].

4.3. Method Limitations At present, the described method works as an expert tool. The melting behaviour measurements are easy and simple to perform, the repeatability of the results is good, and the measurement procedure can easily be standardized. The interpretation of the STA signal, that is, the conversion from the STA curves to the melting curve, is on the other hand not simple and can not at the moment be standardized. The “certain” interpretation requires detailed knowledge on the chemical species in each ash sample (as provided either by CCSEM or XRD analysis) and the chemistry between these. This is necessary for expressing the melt fraction as a mass percent (method no. 2), but is also important to avoid that energies related to solid phase transitions occurring simultaneously to the melting will be wrongly detected as so (melting).

4.4. Result Applicability The new method provides an improved and more detailed characterisation of the melting process occurring in ashes during heating. The results reveal/provide the temperature for which the first melt is formed in the ash, and gives therefore important information for boiler designers. Furthermore, the melting curves can be used as input to mechanistic modeling of ash deposit formation by inertial impaction, which will hope-

fully improve the understanding of ash deposit formation mechanisms during biomass and/or coal combustion.

5. CONCLUSIONS A new experimental method for quantification of ash melting has been developed. Using the new method, a conventional STA apparatus is employed, and the melting is detected as endothermic reactions involving no change in mass. The DSC signal is transferred into a melting curve (showing the melt fraction in the ash as a function of temperature) either by simple comparison of the areas below the melting curve or by

accounting for the relevant melting enthalpies. The execution of the measurement is simple and the repeatability of the results is very good. The subsequent conversion of the STA curves to a melting curve requires knowledge of the identity of chemical species in the ash and the involved chemistry. The method has so far been tested on a number of simple salt mixtures, for which the measured melting behaviour agrees with the predictions from phase diagrams, and a

192

L. A. Hansen et al.

number of ashes collected during combustion of pure straw, co-combustion of straw and

coal, and coal combustion, for which melt was detected between 40 °C and 110°C below the corresponding IDT. Characterising the fusion by STA provides a more detailed description of the ash fusion as compared to conventional methods, and the onset of ash fusion is precisely determined. Furthermore, the method typically enables identification of the chemical species melting in different temperature ranges. As ash melting has a major impact on the deposit formation tendency, the presented detailed ash fusion determination improves the prediction of ash deposition propensities.

ACKNOWLEDGEMENTS This work was carried out as part of the Combustion and Harmful Emission Control (CHEC) research program at the Department of Chemical Engineering, Technical University of Denmark. The CHEC research program is cofunded by ELSAM (The Jutland-Funen Electricity Consortium), ELKRAFT, the Danish Technical Research Council, the Danish and the Nordic Energy Research programs. Mrs. Gurli Mogensen (from Haldor Topsoe A/S) is acknowledged for her valuable contributions to the interpretation of the DSC curve behaviour during the long and troublesome running-in of the apparatus.

REFERENCES AS1038.15 (1987). “Methods for the analysis and testing of coal and coke—fusibility of higher rank coal ash and coke ash”. Australian Standards Association ASTM D1857-87 (1987). “Standard test method for fusibility of coal and coke ash”. American Society for

Testing and Materials Backman, R. (1989). Sodium and Sulfur Chemistry in Combustion Gases. Academic Dissertation, Åbo Academy University, Abo, Finland Backman, R., Hupa, M., and Uppstu, E. (1987). “Fouling and Corrosion Mechanisms in the Recovery Boiler Superheater Area”. Tappi Journal 70 (6) 123–127 Benson, S.A., Jones, M.L., and Harb, J.N. (1993). “Ash Formation and Deposition”. In L. D. Smoot (Eds.), Fundamentals of Coal Combustion for Clean and Efficient Use. New York: Elsevier Coin, C.D.A., Kahraman, H., and Peifenstein, A.P. (1996). “An Improved Ash Fusion Test”. In L.L. Baxter and R. DeSollar (Eds.), Proceedings of the Engineering Foundation Conference. New York and London:

Plenum Press Conn, R.E., and Austin, L.G. (1984). “Studies of Sintering of Coal Ash Relevant to Pulverised Coal Utility

Boilers. 1: Examination of the Raask Shrinkage-Electrical Resistance Method”. Fuel (63) 1664 Cumming, I.W., and Sanyal, A. (1981). “An Electrical Conductance Method for Predicting the Onset of Fusion in Coal Ash”. In R.W. Bryers (Eds.), US Engineering Foundation Conference on Slagging and Fouling from Combustion Gases. New York: Engineering Foundation Cumming, I.W., Joyce, W.I., and Kyle, J.H. (1985). “Advanced Techniques for the Assessment of Slagging and Fouling Propensity in Pulverised Coal Fired Boiler Plant”. In R.E. Barrett (Eds.), Proceedings of the Engineering Foundation Conference on Slagging and Fouling Due to Impurities in Combustion Gases. New York: Engineering Foundation DIN 51730, (1984) “Determination of Fusibility of Fuel Ash”. German Standard Gibson, J.R. and Livingston, W.R. (1992). “The Sintering and Fusion of Bituminous Coal Ashes”. In S.A. Benson (Eds.), Proceedings of the Engineering Foundation Conference on Inorganic Transformations and Ash Deposition During Combustion. New York: ASME Gray, V.R. (1987). “Prediction of Ash Fusion Temperatures from Ash Composition for Some New Zealand

Coals”. Fuel (66) 1230–1239 Hansen, L.A. (1997). Melting and Sintering of Ashes. Ph.D. Thesis; Department of Chemical Engineering, Technical University of Denmark

Ash Fusion Quantification by Means of Thermal Analysis

193

Hansen, L.A., Frandsen, F.J., Sørensen, H.S., Rosenberg, P., Hjuler, K., and Dam-Johansen, K. (1997). “Ash Fusion and Deposit Formation at Straw Fired Boilers”; these conference proceedings Huffman, P.H., Huggins, F.E., and Dunmyre, G.R. (1981). “Investigation of the High-Temperature Behaviour of Coal Ash in Reducing and Oxidizing Atmospheres”. Fuel (60) 585 Huggins, F.E., Kosmack, D.A., and Huffman, G.P (1981). “Correlation between Ash-Fusion Temperatures and Ternary Equilibrium Phase Diagrams”. Fuel (60) 577–584

ISO 540 (1981). “Determination of Fusibility of Ash”. International Standard Organisation Levin, E.M., Robbins, C.R., and McMurdie, H.F. (1964). Phase Diagrams for Ceramists, Vol. I . Columbus,

Ohio: The American Ceramic Society Lloyd, W.G., Riley, J.T., Risen, M.A., Gilleland, S.R., and Tibbits, R.L. (1989). “Estimation of Ash Softening Temperatures Using Cross Terms and Partial Factor Analysis”. Energy and Fuels (4) 325 Netzsch, (1995). Personal Communication

Raask, E.J. (1979). “Sintering Characteristics of Coal Ashes by Simultaneous Dilatometry-EIectrical Conductance Measurements”. Journal of Thermal Analysis (16) 91 Richards, G.H., Slater, P.N., and Harb, J.N. (1993). “Simulation of Ash Deposit Growth in a Pulverised CoalFired Pilot Scale Reactor”. Energy & Fuels (7) 774–781 Sanyal, A., and Cumming, I.W. (1981). “An Electrical Resistivity Method for Detecting the Onset of Fusion in Coal Ash”. In R.W. Bryers (Eds.), US Engineering Foundation Conference on Slagging and Fouling from Combustion Gases. New York: Engineering Foundation Sanyal, A. and Mehta, A.K. (1994). “Development of an Electrical Resistance Method based on Ash Fusion Test”. In J. Williamson and F. Wigley (Eds.), Engineering Foundation Conference on Impact of Ash

Deposition on Coal Fired Plants. Washington: Taylor & Francis Skrifvars, B.-J., Backman, R., and Hupa, M. (1996). “Ash Chemistry and Sintering”. Preprints of Papers Presented at the 211th ACS National Meeting, New Orleans, LA, March 24–28, 1996 Srinivasachar, S., Helble, J.J., Katz, C.B., and Boni, A.A. (1990). “Transformations and Stickiness of minerals during pulverised coal combustion”. In R.W. Bryers and K. Vorres (Eds.), Proc. of the Engineering Foundation Conference on mineral matter and ash deposition from coal. New York: Engineering Foundation Sørensen, H.S.(1997). “Computer Controlled Scanning Electron Microscopy of Straw Ash”. These conference

proceedings Vassilev, S.V., Kitano, K., Takeda, S., and Tsurue, T. (1995). “Influence of mineral and chemical composition of coal ashes on their fusibility” . Fuel Processing Technology (45) 27–51 Vorres, K.W. (1979). “Effect of Composition on Melting Behaviour of Coal Ash”. Journal Eng. Power (101) 497 Wall, T.F., Gupta, R.P., Polychroniadis, P., Ellis, G.C., Ledger, R.C., and Lindner, E.R. (1989). “The Strength, Sintering, Electrical Conductance and Chemical Character of Coal Ash Deposits”. NERDDC Project No. 1181— Final Report, Vol. 1, Summary Report Wall, T.F., Creelman, R.A., Gupta, R.P., Gupta, S., Coin, C., and Lowe, A. (1996). “Coal Ash Fusion Temperatures: New Characterisation Techniques and Associations with Phase Equilibria”. In L.L. Baxter and R. DeSollar, Proceedings of the Engineering Foundation Conference on Applications of Advanced Technology to Ash-Related Problems in Boilers. New York: Plenum Press

Walsh, P.M., Sayre, A.N., Loehden, D.O., Monroe, L.S., Beér, J.M., and Sarofim, A.F. (1990). “Deposition of Bituminous Coal Ash on an Isolated Heat Exchanger Tube: Effects of Coal Properties on Deposit Growth”. Prog. Energy Comb. Sci. (16) 327–346 Westborg, S. (1995). Round Robin Test—Analysis of Straw and Straw Ash. Internal report (in Danish) Biomass

Ash Characterisation Project, dk-TEKNIK, Soeborg, Denmark Winegartner, E.C., and Rhodes, B.T. (1975). “An Empirical Study of the Relation of Chemical Properties to Ash Fusion Temperatures” Journal Eng. Power (97) 395

This page intentionally left blank.

STICKING MECHANISMS IN HOT-GAS FILTER ASHES John P. Hurley, Bruce A. Dockter, Troy A. Roling, and Jan W. Nowok University of North Dakota Energy & Environmental Research Center PO Box 9018 Grand Forks, ND 58202-9018 John Hurley: Phone (701) 777-5159, e-mail [email protected]

1. INTRODUCTION Large-scale hot-gas filter testing over the past 10 years has revealed numerous cases of cake buildup on filter elements that has been difficult, if not impossible, to remove.

At times, the cake can bridge between candle filters, leading to filter failure. Physical factors, including particle-size distribution, particle shape, the aerodynamics of deposition, and system temperature, contribute to the difficulty in removing the cake, but

chemical factors such as surface composition and gas–solid reactions also play roles in helping to bond the ash to the filter and to itself. In order to develop methods to predict the formation of sticky ash in hot-gas filtration systems, the University of North Dakota Energy & Environmental Research Center (EERC) worked with EPRI and a consortium of companies in partnership with the U.S. Department of Energy (DOE) to determine the factors causing hot-gas cleanup filters to develop deposits that can bridge the filters and cause them to fail. The primary deliverable was the Filter Bridging Index Code, a graphics-driven computer code to tie all of the knowledge together and make possible the prediction of rates of filter bridging based on coal, sorbent, filter, and system parameters. The objectives of this project were threefold: • Determine the mechanisms by which a difficult-to-clean ash forms and how it bridges hot-gas filters

• Develop a method to predict the rate of bridging based on analyses of the feed coal and sorbent, filter properties, and system operating conditions • Suggest and test ways to prevent filter bridging The research took place over the years 1994–1997 and comprised five tasks. Task 1 involved detailed sampling at large-scale, operating hot-gas filter test units and gatherImpact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

195

196

J. P. Hurley et al.

ing representative archived samples from completed programs, then subjecting the

samples to intensive physical and chemical analyses. Task 2 concentrated on thermochemical equilibrium modeling to determine possible chemical contributions to ash stickiness, along with laboratory measurements of the rates and mechanisms of tensile strength development in ash cakes. In Task 3, bench-scale testing was employed to determine factors affecting the formation of ash in pressurized fluidized-bed combustors (PFBCs) and the factors affecting the rates of residual cake development under both fluidized-bed combustion (FBC) and gasification conditions. Under Task 4, a graphical user interface computer code was created to tie all of the knowledge together and make possible the prediction of rates of filter bridging based on coal, sorbent, filter, and system parameters. Task 5 involved reporting. EPRI is the prime contractor to the other sponsors. All research was carried out by the EERC. The EERC, through its Cooperative Agreement with the DOE Federal Energy Technology Center (FETC), received funds to approximately match those contributed by the sponsors. The sponsors were EPRI, Lurgi Lentjes Babcock (LLB), a British consortium led by PowerGen plc, Schumacher America, Westinghouse, Electricité de France, the Netherlands Energy Research Foundation (ECN), a Swedish group led by ABB Carbon, and the Electric Power Development Company of Japan. In this paper, we report measurements of several ash- and system-related factors affecting the stickiness of the ash in hot-gas particulate filters used in PFBC systems. Data are presented on the relative effects of temperature, cake porosity, and the presence of surface liquids on the bridging propensities of ash cakes.

2. BACKGROUND Many factors control the adhesiveness of ash particles that leads to the formation of ash deposits or bridges between filter elements in hot-gas particulate filtration systems. The following is a short evaluation of some of the most significant factors affecting adhesion within particle cakes.

2.1. Surface Area and Particle Attraction Predicting the ways in which a powder will agglomerate and form cakes is a matter of understanding not only all of the ways in which particles can attract one another, but also the way in which the particles fit together in the bulk. Any attempt to sum the microscopic contributions made by the particles themselves must take into account three factors: the interparticle forces, the three-dimensional shapes of the particles, and the way in which the particles interact geometrically to form a packed structure. Most powders are not easily characterized in terms of these properties. For example, most particles in PFBC ashes have complex shapes so that both the distributions of equivalent spherical diameters and the distributions of shape factors must be used in describing particle-size distributions. The situation is no less complicated with interparticle forces, which are

subject to change with environment and time. In hot-gas filter ashes, these forces can be of a variety of types, including the following from weakest to strongest:

• Mechanical forces caused by interlocking of irregular particles • Electrostatic forces, particularly for surfaces that easily become charged

Sticking Mechanisms in Hot-Gas Filter Ashes

197

• Molecular (or van der Waals) forces, particularly significant for particles of small diameters, especially those less than 10µm in diameter • Surface tension forces caused by bridging between surface liquids • Solid bridge forces, where solidification at contact points causes joining of the particles Mechanical forces are those created as the particles pack together to form the powder cake. Even if the particles were all of the same size and shape and interparticle attractions were absent, several three-dimensional packing structures would still be available to the particles. Particle size and shape and interparticle attractive forces both modify this situation, resulting in a structure of fine, irregular particles with an extremely complex geometric configuration. Electrostatic surface forces also cause particles to stick. They can develop if the particles have been in contact with an ionic solid or if electrons transfer between the surfaces of particles by contact charging. These forces differ from van der Waals forces, which occur from random motion of the electrons in the surface molecules. Even if a particle were perfectly dry, clean, and smooth and showed no tendency toward solid bridge formation, forces on the surfaces would still cause interparticle attraction, the van der Waals, or molecular, forces [Parfitt and Sing, 1976; Israelachvili, 1991; Funk and Dinger, 1995]. Equations for van der Waals forces predict that larger particles are more strongly attracted to each other than are small particles, although the ratio of van der Waals to gravitational forces is much greater for smaller particles. Additionally, these equations are valid only when the particles are close to each other. When any two particles are separated by more than a micron, van der Waals forces are negligible. There-

fore, these forces are reduced in comparison to the force due to gravity for larger, highly irregular particles since relatively small parts of their total surface area are actually touching. If the surfaces are deformable, the real area of contact and hence the adhesion between the particles will increase with the force pressing them together. In addition to van der Waals forces, the surface energy of the material must be considered. This becomes an especially great factor if a liquid layer coats the particles, because if the liquid wets the particles a strong bond forms. Solid bridge formation can occur by either direct or secondary interaction. Solid bridges can form through chemical or physical vapor deposition of gases, chemical reaction between liquids, crystallization of dissolved substances, freezing of liquids, or viscous flow sintering. For particles that are already touching or are in close proximity, mainly liquid and solid bridges and van der Waals forces are responsible for adhesion. At temperatures below 150°C, van der Waals forces and the presence of liquid (usually aqueous) bridges play the main role in particle adhesion. At longer contact times and higher temperatures (above 150°C), the influence of interface reaction increases greatly. If the temperature rises above 600°C and the appropriate atmosphere is present, it is possible for glassy ceramics to crystallize. This would result in very strong particle adhesion. Also, the most dominant interparticle forces may change with time as the particles interact. The magnitudes of the forces described above ultimately depend on the temperature and size and composition distributions of the particles that make up the cake. Measuring the relative effects of these factors on cake tensile strengths allows us to understand their influence on the formation of ash bridges. With this knowledge, existing computer codes that predict the size and composition distributions of ash particles on the basis of analysis of the inorganic matter in a coal can be modified, creating a tool to predict the

198

J. P. Hurley et al.

relative propensity for ash bridges to form for a given coal and system temperature. In

this way, fuels or conditions that will likely produce ash bridges can be avoided.

3. RESULTS AND DISCUSSION Tensile strengths of ash cakes at elevated temperatures in simulated combustion atmospheres were determined with the high-temperature tensile strength tester (HTTT).

A schematic of the HTTT is shown in Fig. 1. The tester consists of a split cylinder with a porous metal bottom. One side of the cylinder is fixed, and the other side is suspended like a pendulum. The force needed to swing the suspended side away from the fixed side divided by the cross-sectional area of the cake filling the cylinder gives the tensile strength of the cake. To determine the effects of gas composition and temperature on tensile strength, the tester is placed in a split oven and preheated gas is passed upward through the porous bottom and the ash cake. Because relative humidity may affect the cohesive properties of the filter ash, the samples are stored overnight at 200°C. In earlier tests, a weighed amount of ash was placed in the split cylinder and a weight was placed on top to compress the ash with a force similar to that experienced by a filter ash cake due to the pressure drop across the cake. However, the weight primarily compressed the top layers of the cake, leaving the bottom uncompressed so that the cake breaks were very uneven. In later tests, the ash was sifted onto the porous substrate while air was pulled through the cake at face velocities approximating those in a filter system. This preparation method resulted in much more realistic cake porosities (around 70%), although the porosity is still somewhat lower than that measured in actual cakes because the ash particles are less sticky at room temperature and so tend to settle into more compact cakes [Snyder and Pontius, 1996].

3.1. Critical Thickness Index In related EERC research on baghouse efficiency, it was determined that an ash cake tensile strength of between 0.5 and is recommended to ensure the cake

Sticking Mechanisms in Hot-Gas Filter Ashes

199

does not disintegrate during backpulsing, allowing the ash particles to reattach to the filters (Miller et al., 1993). This lower strength range is not expected to change greatly at

pressurized hot-gas temperatures because the effects of higher pressure on the density of the gas that would reentrain the ash are largely canceled by the reduction in density caused by the higher temperature. In testing of ash collected from the Westinghouse filter vessel used at the American Electric Power Company (AEP) Tidd plant [Newby et al., 1995], it was determined that at strength levels of 2.5 to 3.0g/cm2, the Tidd ash begins to sinter slightly and the cake begins to slide along the porous frit bottom of the splitcylinder chamber. This means that the capabilities of the HTTT are being exceeded at this level of strength for this particular source of filter ash. It also indicates the strength necessary to form a bridge for this ash. In later tests with other ashes, higher strengths could be measured with some cakes, different cakes had different plasticities, and tensile strength did not always correlate directly with bridging propensity. Figure 2 shows the tensile strength of two ashes designated C-l and hot-gas ash provided by ABB Carbon which were collected from a PFBC system. The tensile strength measurements were made at 700° and 750°C. The graph shows that although strength generally increases with temperature, the error bars are broad enough that there is not a large difference in tensile strengths of the two ashes. ABB Carbon experience is that the hot-gas ash is much more likely to bridge between filters than is the C-l ash. Since many bridges must support themselves against gravity, tensile strength alone does not indicate the likelihood for a cake to form a bridge. What is more important is the specific strength of the cake, or the strength in relationship to the density of the cake. The following equation gives the definition of specific strength, a parameter that we prefer to call the cake critical thickness index (CTI): CTI =

tensile breaking force/cake cross - sectional area cake weight/cake volume

The CTI has units of length and indicates the relative thickness of a cake or deposit that can form before it will shed under its own weight. A high CTI indicates that the cake

200

J. P. Hurley et al.

is relatively sticky for its weight and is more likely to form a bridge than a cake with a low CTI. Although by definition a CTI that is greater than half the distance between filters implies that an ash bridge can form, other forces caused by system vibration and friction may be involved in determining the likelihood of an ash bridge forming. Therefore, the CTI should not be viewed as an exact number, but as a relative number for comparing bridging propensities.

Measurements of the bulk densities of the cakes showed that the C-l ash is much more dense than the hot-gas ash. Figure 3 shows the measured CTIs for the ABB Carbon

samples. The data match the field observations of the relative bridging propensities of the ashes, showing that the hot-gas ash is more likely to bridge than the C-l ash and that between 700° and 750 °C, the relative bridging propensity of the hot-gas ash increases dramatically.

3.2. Effects of Cake Porosity and Temperature Figure 4 shows the CTIs measured at different temperatures and degrees of compaction for a filter hopper ash collected from the AEP Tidd plant during a period when filter bridging had been minimized [Newby et al., 1995]. The data show that as the temperature and level of compaction of the Tidd ash increase, the resulting tensile strength also increases. The effect of compaction is much more pronounced at higher temperatures, but generally causes a doubling in the CTI for every 6% decrease in void fraction (porosity) for temperatures above ambient. This increase in CTI indicates that strength increases relative to bulk density as void fraction decreases at higher temperatures, but much less at room temperature. However, for a given void fraction, the temperature has a much greater effect, increasing the CTI by an order of magnitude as temperature is increased from ambient to 700°C at a void fraction of 62%.

3.3. Effects of Surface Liquids To help determine the likelihood that liquids may form on the surfaces of ash par-

ticles, thereby increasing their stickiness, thermochemical equilibrium modeling of ash

Sticking Mechanisms in Hot-Gas Filter Ashes

201

and gas compositions was performed with an ideal solution model called the Facility for the Analysis of Chemical Thermodynamics (FACT) code. Using the FACT code, the quantities of alkalies in the vapor phase, ash slag, and liquid salt that may occur in ash and cause its sticking problems under combustion conditions have been estimated. As shown in Fig. 5, the code predicts the formation of liquid alkali metal salts in Tidd ash at temperatures as low as 560°C, with a rapid increase in concentration as temperature rises to as much as 6% of the total ash-forming material. The code predicts that potassium and sodium sulfates, carbonates, and hydroxides make up this liquid material. However, the flatness of the curve above 820°C indicates that the code is extrapolating at higher temperatures, so the data are suspect above that point. Experimental verification of the phases predicted also has shown that a sulfate–carbonate blend with a composition similar to the predicted liquid salt is stable for only a short time (about 10 minutes) in the presence of (5,000 ppm–20 vol%– balance) atmosphere at 727°C. After cooling to room temperature, the liquid crystallizes and forms

202

J. P. Hurley et al.

intermediate phases:

and

Over longer times at temperature

though, the material converts to complex sulfates, which may have somewhat higher, but

still troublesome, melting points. The transition to a higher melting point may also cause the material to solidify at a constant temperature, thereby making the deposit much harder. In order to determine the effects of surface liquid on the strength of an ash cake, tensile strength measurements were performed at room temperature on a bench-scale filter hopper ash by wetting it with ethylene glycol and glycerin. These liquids were chosen

because of their low vapor pressures and because they have viscosities approximating the lowest (ethylene glycol at 0.2 poise) and highest (glycerin at 10 poise) viscosities expected for the molten salts.

The filter hopper ash was sieved, and the fraction smaller than 270 mesh (53 microns) was used for the analysis. The liquid was then added to the ash and mixed thoroughly. Before the treated ash was placed in the split cell of the HTTT, it was passed through a sieve to break up any agglomerates. The ash was compacted in the split cell using the vacuum method. During the addition of the sample, very little change in pressure drop occurred. The sample was vacuumcompacted for a period of 10 minutes, and then the cake was broken. This was repeated four times to give an average and standard deviation. The CTIs for the different weight percentages of ethylene glycol are shown in Fig. 6. The graph shows that with the addition of 1% ethylene glycol to the filter ash, the CT1 increased by approximately 75% on average. This increase could turn a nonbridging ash into a bridging ash. The strength

variations in the l%-to-10% liquid range may be experimental error, although the literature suggests that some of these variations may be real, caused by changes in the mechanisms by which the particles are bonded [Kia, 1988]. The results obtained using the higher-viscosity glycerin were quite different, as shown in Fig. 7. The graph shows that with the addition of 2.8% glycerin to the filter ash, the cake strength increased by approximately 9% on average. This is a much smaller increase than with ethylene glycol, which has a much lower viscosity. However, the ethylene glycol appears to wet the ash much better than the glycerin, indicating that the ability of the liquid to wet the ash is much more important than the viscosity of the liquid

in increasing the strength of an ash cake.

Sticking Mechanisms in Hot-Gas Filter Ashes

203

The question then becomes, Does a liquid alkali salt wet coal ash that is primarily silicate-based? To answer that question, a pressed pellet of sodium sulfate powder was placed on the surface of a button of fused Illinois No. 6 coal ash slag. The materials were

heated in a furnace in air to the melting point of the sodium sulfate and photographed. Since sodium sulfate is a pure compound, it completely converts to liquid at its melting point. Also, the liquid is of a very low viscosity. As shown in Fig. 8, the molten sodium

sulfate does not immediately wet the silicate slag. However, over periods of tens of

minutes, it does react with the silicate to form an intermediate phase at the surface of the slag, which is wetted. This means that alkali salts condensing on silicate-based ash will

204

J. P. Hurley et al.

not immediately increase the strength of the ash cake, but that over tens of minutes, strength will continuously increase as the intermediate phase is formed.

4. SUMMARY AND CONCLUSIONS Tensile strength of an ash cake alone is not a good indicator of the propensity of an ash to form a bridge between filter elements in a hot-gas filter system. Since gravity is pulling down on the bridges, the tensile strength of the cake must be divided by the bulk density of the cake to give the critical thickness index, or CTI. The CTI has units of length and indicates the relative thickness of a cake or deposit that can form before

it will shed under its own weight. A high CTI indicates that the cake is relatively sticky for its weight and is more likely to form a bridge than a cake with a low CTI. CTI increases as cake porosity decreases, indicating that the strength of the cake increases relative to its weight as porosity decreases. The effect is much more prominent at higher temperatures. In addition, the strength of ash cakes is much higher at higher temperatures. Finally, the presence of wetting liquids on the surface of ash particles can dramatically increase the strength of a cake even at concentrations as low as 1%. However, the liquid must wet the ash in order to increase the strength and alkali salts do not wet

silicate ashes until they have had time to react with the silicate material, indicating that strength will increase significantly over periods of tens of minutes in situations where liquid alkali salts are depositing.

5. REFERENCES Funk, J. E., and Dinger D. R. (1995). “Flocculation by VDW Energy/Particle Size Distribution.” Journal American Ceramic Society, 74(1).

Israelachvili, J. N. (1991). Intermolecular and Surface Forces. London: Academic Press. Kia, S. F. (1988). “Modeling of the Retention of Organic Contaminants in Porous Media of Uniform Spherical Particles.” Water Research, 22(10), 1301–1309. Miller, S.J., Laudal, D.L., and Heidt, M.K. (1993). “Cohesive Properties of Fly Ash and How They Affect Particulate Control Optimization” presented at the 10th Particulate Control Symposium and 5th International Conference on Electrostatic Precipitation, April 5–8, 1993, Washington, DC, Volume 1: Session A l .

Newby, R. A., Lippert, T. E., and Mudd, M. J. (1995). “Tidd Experience Prepares Hot Gas Cleaning Technology for Commercialization.” Power Engineering, 99(09), 20–24. Parfitt, G. D., and Sing, K. S. W. (1976). Characterization of Powder Surfaces. London: Academic Press. Snyder, T. R., and Pontius, D. H. (1996). “Particle Characteristics and High Temperature Filtration.” Proceedings of the 13th Annual International Pittsburgh Coal Conference. University of Pittsburgh, Volume 1. Chiang, S.-H., Ed. pp. 116–121.

CLASSIFICATION SYSTEM FOR ASH DEPOSITS BASED ON SEM ANALYSES Karin Laursen 1 and Flemming J. Frandsen2 1

Geological Survey of Denmark and Greenland Thoravej 8, 2400 Copenhagen NW, Denmark Phone: +45 3814 2000 Fax: +45 3814 2050

2

Technical University of Denmark Building 229, 2700 Lyngby, Denmark

Phone: +45 4525 2883 Fax: +45 4588 2258

1. INTRODUCTION During the last 15 years scanning electron microscope (SEM) and energydispersive x-ray (EDX) has been used extensively for solving problems related to coal combustion, especially in relation to the ash forming components of the coal. SEM and EDX have been used for analyzing the inorganic components of the coal, the fly ash and ash deposits and various automatic techniques have been developed to analyze these materials (i.e. CCSEM and SEMPC) [Lee et al., 1978; Huggins et al., 1980; Straszheim et al., 1988; Zygarlicke and Steadman, 1990; Jones et al., 1992; Skorupska and Carpenter, 1993]. These automatic techniques are strong tools for analyzing ash deposits. However, a major limitation of these techniques is their lack of capability to provide information on the appearance (i.e., texture and morphology) of the deposits, which can have major influence on the physical properties of the material. Hatt (1990) suggested a classification system of slags based on their macroscopic appearance, but no such attempt has been made to classify deposits based on microscopic appearance and microanalyses. Thus, this paper includes a suggestion to a classification system for deposits based on their texture. Additionally, the appearance is combined with the information achieved from automatic SEM-EDX analyses (i.e., SEMPC) of the deposits.

2. EXPERIMENTAL As part of a Danish collaborative project on “Mineral transformation and ash deposition in pulverized coal fired boilers” six full-scale trials were conducted at three Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

205

206

K. Laursen and F. J. Frandsen

power stations in Denmark [Larsen et al., 1996; Laursen, 1997, Laursen et al., 1997].

Four of the coals burned during these trials were fuels with low ash deposition propensities (e.g., an Indonesian, a Colombian, a South African and a Polish coal). One coal had medium ash deposition propensities (e.g. a US high-S coal) and the last coal was a blend of the US and the Polish coal [Laursen, 1997]. During the full-scale trials deposits were collected on air and water-air cooled

probes with adjustable metal surface temperatures. The probes consist of a 1.5m long stainless steel pipe with a diameter of 3.8cm. The deposits are collected on a 10cm exchangeable piece of tube (test element) located approximately 10cm from the tip of the probe. In addition, some deposits from and an un-cooled ceramic protection cap of a suction pyrometer were analyzed. The test elements from the cooled probes were generally covered with a loose fly ash deposit on the downstream side (shelter side) and to a lesser extend on the sides. These loose deposits were often blown off the probes when the probes were retracted from the boiler due to air pressure from the probe and the pressure from the boiler. In addition, some of the fly ash was lost during dismantling of the test tube. Thus, no test elements contain a complete, intact deposit that indicates the magnitude of the deposition rate on the downstream side. Hardly any deposits were visible on the upstream (windward side) of the test elements from the combustion of the fuels with low ash deposition propensities,

except for a thin, black scale and some small irregularities (e.g. islands). The upstream side of the test elements exposed during combustion and the US coal and the blend were normally covered with a hard-bonded, rough deposit. Video recordings of one of the probes taken shortly before the probe was removed from the boiler showed that most of the deposits on the upstream side of the probe were lost during retraction. These deposits were significantly thicker (approximately 3–4cm) than the deposits found on the test elements (maximum 2mm after 8 hours) especially during combustion of the US coal. Thus, the deposits collected on the probes only represented the hard-bonded deposits.

3. CLASSIFICATION OF ASH DEPOSITS BASED ON TEXTURE Based on the texture, the deposits collected on the probes can be classified into five main types 1) porous deposit; 2) powder deposit; 3) iron-rich deposit; 4) semi-fused slag; and 5) fused slag (Fig. 1).

3.1. Porous Deposit The porous deposits are dominated by deformed, large iron-rich particles (10–250 surrounding mainly smaller and some larger Al-silicate fly ash particles (Fig. la). The deformed shape of the iron-rich particles is caused by collision with surrounding particles indicating that the iron-rich particles had low viscosity during impact. Locally, the iron-rich particles and large Al-silicate fly ash particles are arranged in “finger-structures” perpendicular to the tube surface. In some areas, several large particles create a very compact deposit. Holes or “pockets” between the large particles are normally loosely packed with smaller Al-silicate fly ash particles. These Al-silicate particles are only loosely connected to each other but locally they are totally incorporated into iron-rich particles. The porosity of the porous deposits varies from approximately 20 to 30% (based on image analysis). However, locally the porosity is close to zero in islands of iron-rich particles and up

Classification System for Ash Deposits Based on SEM Analyses

207

to 60% in “pockets” filled with small Al-silicate particles. Porous deposits were only collected during combustion of the US coal and the coal blend of US coal and Polish coal. Porous deposits were mainly seen on the upstream side of the probes and partly on the sides. The thickness of the collected porous deposits varies from to a maximum observed thickness of 2mm (after 8 h). Walsh and others (1990) reported deposits of a similar appearance as the porous deposits on a probe inserted at the exit of a pilot-scale furnace. Walsh and others (1990) termed the deposits “hard bonded deposits”, and related their morphology and occurrence to preferential deposition of iron-rich fly ash particles.

208

K. Laursen and F. J. Frandsen

SEMPC analyses of the porous deposits reveal that two main phases are present: iron-rich particles (i.e. iron oxide) and Al-silicates (i.e. illite, kaolinite and montmorillonite derived). The two chemical phases, iron-rich and clay-derived, are also clearly distinct on the ternary diagrams illustrated in Fig. 2, represented by the peak in apex and the peak on the respectively.

3.2. Powder Deposit Powder deposits consist of loosely bonded fly ash particles (Fig. 1b). The fly ash particles are mainly Al-silicates but some Fe-rich and Ca-rich particles are also present. The fly ash particle sizes varies from to The individual fly ash particles are very loosely bonded and only few neck formations between adjacent particles are seen. Sub-micron particles are often located between larger particles and these small particles are probably important for the strength of the deposit. Spot analyses and x-ray mappings of neck-formations revealed that Ca-sulfate often is the bonding material between two particles. The individual particles within the deposit are clearly only loosely connected and no neck formations between large particles are seen. Some of the larger

fly ash particles have a rough surface that appears like “scales” Spot analyses of these “scales” reveal that they often are rich in Ca-sulfate but Al-silicate “scales” are also seen. The “scales” are not spherical as the small fly ash particles but they posses a more irregular shape. The powder deposits were observed on probes exposed during combustion of the four coals with low ash deposition propensities (e.g. the Indonesian, the Colombian, the South African and the Polish coal). Powder deposits were mainly present on the downstream side of the probes but locally also on the wings. The total thickness of the deposits varies from to (after 6.5h) but the thicker deposits are only seen

Classification System for Ash Deposits Based on SEM Analyses

209

locally on the probes. As previously discussed most of these powder deposits were lost as the probes were retracted from the boilers. Deposits of similar appearance as the powder deposits were reported by Hurley and others (1994; 1995) as “downstream powder deposits”. These deposit were collected on the downstream side of probes and their deposition could be caused by eddy deposition of small fly ash particles [Hurley et al., 1994; 1995]. As indicated by the SEM images, SEMPC analyses reveal that the majority of the fly ash particles in these deposits are Al-silicates but some Fe-Al silicates are also present

210

K. Laursen and F. J. Frandsen

(Table 1). The Al-silicates (represented by the peak located on the and Fe-Al silicates (represented by the numerous particles located on the tie line between the Al-silicate peak and the apex) are also apparent on the ternary diagrams (Fig. 2). SEMPC analyses of powder deposits showed very low concentrations of calcium sulfate (Table 1). Manual SEM investigations reveal that Ca-sulfate is present mainly where individual particles in the powder deposits stick together and it is believed that Ca-sulfate is responsible for some of the strength of the deposits. The Ca-sulfate is not present as single particles but as a coating on larger particles. This phenomenon explains why few points in the SEMPC analyses are classified as Ca-sulfate. If a phase is present as a thin coating on another particles, a spot analysis of the coating will be influenced by the chemical composition of adjacent and underlying particles. The small amount of Ca-sulfate and the few observations of particles bonded by Ca-sulfate indicate that this bonding mechanism cannot solely be responsible for the strength of the deposits. Often very small fly ash particles (sub-micron) are seen between the larger fly ash particles. It is possible that the bonding of the powder deposits is not only a chemical phenomena but also a physical phenomena controlled by surface forces including van der Waals, electrostatic and surface-tension forces [Raask, 1985].

3.3. Iron-Rich Deposit The iron-rich deposits consist of Al-silicate fly ash particles embedded in a dense iron-rich matrix (Fig. 1c). Iron-rich deposits are always located adjacent to the tube surface or the oxide layer and it is difficult to identify the boundary between the oxide and the deposit. In general, the presence of fly ash particles indicates that the layer is deposited. The Al-silicate fly ash particles are characteristically spherical but irregular particles are also seen (mainly quartz). On the upstream side of the tubes, the contact between the fly ash particles and the matrix is less distinct compared to the downstream side. These diffuse contact zones could be reaction rims indicating dissolution of the fly ash particles on the hotter upstream side. The fly ash particles are considerable larger on the downstream side of the probes compared to on the upstream side whereas the deposits are thicker on the upstream side compared to on the downstream side (exposure time 6 to 36h.). The fly ash particles in these deposits are significantly smaller than the mean diameter of the fly ashes from the electrostatic precipitator, which was measured to be in the range of 11 to (based on CCSEM). A noticeable difference between the upstream and the downstream side deposits is the texture of the matrix. On the downstream side, the matrix appears homogeneous, whereas on the upstream side it appears more non-homogeneous. Locally, in the matrix on the upstream side, remnants of fly ash particles can be seen. Iron-rich

deposits were, in addition to the powder deposits, the only type of deposit observed on the probes during combustion if the coals with low ash deposition propensities. During combustion of the US coal and the coal blend this type of deposit was only observed locally on the probes. Lee and Whitehead (1982) have reported deposits of similar appearance as the ironrich deposits on corroded steel tubes. Lee and Whitehead (1982) termed these deposits “iron oxide/sulfide scale” and related the deposits to a sudden increase in the steel corrosion rates. Due to the very limited thickness of the iron-rich deposits only very few SEMPC

Classification System for Ash Deposits Based on SEM Analyses

211

analyses have been conducted on these deposits. As expected, these deposits are dominated by iron-rich phases mainly categorized as iron-oxide (Table 1). Due to the very small particle size of the Al-silicate fly ash particles in the deposits no “pure” Al-silicate particles were detected by the SEMPC technique. When analyzing such small particles, the matrix (iron) will always be detected, too. However, spot analyses in the center of large particles indicate that the fly ash particles mainly are iron-free. The ternary diagram also illustrates the mixing of the iron-rich matrix and the Al-silicate fly ash particles (Fig. 2).

3.4. Semi-Fused Slag Semi-fused slags consist of fly ash particles which are sintered together forming a deposit with high porosity (50–70%) (Fig. 1d). In some parts of the deposits, the fly ash particles are fused into a compact mass whereas in other parts only a neck formation between the individual particles is seen. The visual appearance of the slag is like loosely bonded sand grains. Semi-fused slags were only collected on the un-cooled ceramic probes during the combustion of the Colombian, the South African and the Polish coal. The thickness of these semi-fused deposits was approximately 1.5 to 2cm after an exposure time of 10 minutes. Hurley and others (1995) have reported deposits of a similar appearance on the upstream side of steam tubes located in the secondary superheater region. Hurley and others (1995) termed the deposits “conventional high temperature fouling” and related the bonding of these deposits to viscous sintering of silicate material. In the classification system suggested by Hatt and Rimmer this type of deposit would be classified as a “sintered deposit” which normally is found in the upper furnace and convective pass [Hatt, 1990]. The majority of the phases present in these deposits are classified as clay-derived (e.g. illite, kaolinite, and montmorillonite derived) (Table 1). The deposits collected during combustion of the South African coal differs significantly from deposits collected during combustion of the Colombian and Polish coal. The deposits from South African coal has a higher content of Ca-Al silicates and a low content of Fe-Al compared to the other two deposits. The high content of Ca-Al silicates and the characteristic white color of the deposits from South African coal compared to the other two deposits supports that the Ca-Al silicates are the main reason why South African coals produce a light grey fly ash and deposits [Laursen, 1997]. In general, the phases present in the semi-fused slags are the same phases as found in the fly ashes. However, there is a higher content of particles classified as unknown in the semi-fused slag compared to the fly ashes. These unknowns probably represent the fused phases in the deposit. The ternary diagrams reveal the high content of clay-derived phases in the deposits indicated by the dominant peak on the -binary in both the and the ternary diagram (Fig. 2).

3.5. Fused Slag Fused slags are highly sintered and only remnants of single quartz fly ash particles can are seen (Fig. le). Large pores are common. Crystallization of various minerals with crystal textures indicating growth during super-cooling (hopper- and dendrite crystals) is common. Fused slags were never sampled directly on any of the test elements and the few samples of this type of deposits were collected on the un-cooled tip of a

K. Laursen and F. J. Frandsen

212

probe exposed in the furnace during combustion of the high-S US coal. Fused slags are similar in appearance to the dense slags collected as bottom ash. In the classification system suggested by Hatt and Rimmer fused slags are classified as “visicular-glassy slag” which is related to the furnace and high temperature regions of the convective pass [Hatt, 1990]. The textural analyses of the few samples of fused slags indicated that glass constitutes the main phase in these deposits. In addition, a diversified fraction of newly formed crystals and few remnants of fly ash particles (mainly quartz) are present. Table 1 lists the result of a SEMPC analysis of a slag collected during combustion of the high-S US coal on the tip of a probe exposed in the furnace. A larger fraction of the phases in the fused slag from the probe are classified as unknown compared to the phases in the bottom ashes (not shown). This lower content of unknowns in the bottom ash is due to a higher amount of new-formed crystals in the bottom ash. The ternary diagrams for the bottom ash also indicate the crystallization of new minerals. These new crystals are indicated by the chemical compositions located on tie lines from the peak around the bulk chemical composition (i.e. the glass) towards the binaries of the diagrams (Fig. 2). The fraction of points located at a distance from the bulk chemical composition is indicators of crystalline phases present in the slag. The fraction of crystals is important for the physical state of a deposit as increasing fraction of crystals will increase the resistance to fracture formation and thus increase the strength of the deposit [Wain et a/.,

1992]. Thus, a deposit containing a high fraction of points located at a distance from the bulk chemical composition would indicate a stronger deposit compared to a glassy deposit with the chemical compositions clustered around the bulk chemical composition (i.e. the glass).

4. ASH DEPOSIT MECHANISMS The build-up of the porous deposits on the upstream side of the probes is most likely controlled by preferential inertial deposition of iron-rich particles (Fig. 3). The deformed shape of the large iron-rich particles clearly indicates that the majority of these

particles were molten when they impacted the tube surface or other ash particles in the

deposit. The viscosity of iron-rich particles is largely controlled by the oxidation state of the iron [Bool and Helble, 1996]. In case the fly ash particles are in a partly reduced state upon impact on a heat transfer surface they are likely to stick to the surface, however, if the particles are totally oxidized before impact they will have a high viscosity and likely to rebound [Moza and Austin, 1981; Abbott and Austin, 1982; Bool and Helble, 1996]. The build up of a “finger-like” structure in a porous deposit is probably initiated

by deposition of relatively large iron-rich particles. The outer part of such large particles may penetrate a thermal boundary layer around the tube to an area where the tempera-

ture is considerably higher. The temperature might even exceed the melting temperature of depositing particles and larger Al-silicate particles are likely to stick upon impact on the molten surface.

Local eddies between the “finger-like” structures are probably important for the deposition of the “pockets” of small Al-silicate fly ash particles. These small particles can be caught by eddies between the “fingers”. The velocity of these particles will be reduced and they will loose all their kinetic energy. As the porous deposits build-up, the temperature on the outside of the deposit gradually increases due to the low thermal conductivity of the deposit. Finally, the deposit sinters into a semi-fused slag and eventually

Classification System for Ash Deposits Based on SEM Analyses

213

into a fused slag. Fluxing occurs at the surface of a deposit between 815°C and 1,300°C (lowest eutectic in the system) [Bryers, 1996]. The deposition of the semi-fused slag and the fused slag is controlled by inertial impaction. However, it is no longer a preferential deposition of iron-rich particles, but a deposition of any particle reaching the hot surface of the deposit. This type of deposition leads to a chemical composition in the slag which is closer to that of the fly ash compared to the initial deposited layers. Crystallization of phases within the fused slags occurs when the melt reaches the liquid us for the crystals. Quenching of the melt can either occur when the slag is still bonded to a heat transfer surface due to insulation from radiation by newly deposited

material, or after detachment from the surface when the slag falls into the ash hopper. The crystallization may also take place as a re-crystallization from the glass. The fraction of crystals is important for the physical state of a deposit as increasing amount of crystals will increase the resistance to fracture formation and thus increase the strength of the deposit [Wain et al., 1992]. The deposition of the powder deposits is probably controlled by eddy deposition. Particles can be caught by eddies behind the tubes, loose their kinetic energy, and be more likely to stick upon impact. The powder deposits are significantly finer grained than the majority of the fly ash because the medium and large fly ash particles were not caught by the eddies but followed the main gas flow lines. The chemical analyses show that Casulfate may be a contributor to the bonding of the powder deposits. Sulfates are recognized as important bonding phases for low-temperature fouling deposits [Walsh et al., 1992; Osborn, 1992]. However, the mechanism of sulfate formation has not yet been clarified, especially regarding whether the sulfation takes place before or after deposition. Recent studies have revealed that for high-calcium coals the sulfate is fixed in-situ after the deposition [Hurley and Benson, 1995; Richards et al., 1996]. Richards et al., 1996, reported that the sulfation only occurred on Ca-rich fly ash particles and not

214

K. Laursen and F. J. Frandsen

on Al-silicates. These results do not agree with observations in this study where Casulfates also are seen on Al-silicates. The small amount of Ca-sulfate and the few observations of particles bonded by Ca-sulfate indicate that this bonding mechanism cannot

solely be responsible for the build-up of the strength in the deposits. Often, very small fly ash particles (sub-micron) are seen between the larger fly ash particles. It is possible that the bonding of the powder deposits is not only a chemical phenomena but also a physical phenomena controlled by surface forces including van der Waals, electrostatic and surface-tension [Raask, 1985]. The matrix of the iron-rich deposits may have two origins: 1) deposition of ironrich fly ash particles; and 2) diffusion of iron from the tube due to oxidation (corrosion) [Cutler et al., 1975]. Remnants of iron-rich particles can be identified locally in the deposits, mainly on the upstream side, indicating that impacting iron-rich particles contributes to the matrix. However, this does not rule out that iron diffusion was an active mechanism. The Al-silicate fly ash particles embedded in the iron matrix of the iron-rich deposits on the upstream side of the probes are numerous and relatively smaller than on the downstream side, where fewer but relatively larger fly ash particles are seen. This differential particle size distribution is probably controlled by various mechanisms. The deposition of the small particles on the upstream side is probably controlled mainly by thermophoresis, whereas the larger particles on the downstream side are controlled

by eddy deposition. On the upstream side, larger particles impact and bounce off due to their higher kinetic energy. On the downstream side, these particles are caught by eddies behind the tubes. The particles will have lower velocity and are more likely to stick on impact. Larger fly ash particles which comprise the majority of the fly ash, will not be caught by the eddies but they will follow the main gas flow.

5. SUMMARY AND CONCLUSIONS Based on SEM analyses of deposits collected during full-scale trials at three power stations in Denmark it has been possible to classify the deposits into five textural: 1) porous deposit; 2) powder deposit; 3) iron-rich deposit; 4) semi-fused slag; and 5) fused slag. These five textural types not only have district textural characteristics, but they also posses characteristic micro-chemical features, which especially are apparent by the differences in the distribution of point analyses (SEMPC analyses) of the deposits in the two ternary diagrams: and In the development of a deposit the porous deposit and the iron-rich deposit represent initiation stages for coals with medium and low ash deposition propensities, respectively. The semi-fused slags and especially the fused slags represent consolidation or maturation stages of the build-up of an ash deposit. The powder deposits will develop on downstream sides of superheater tubes, either directly on the oxide layer or on an iron-rich deposit.

ACKNOWLEDGMENTS This project was funded by ELSAM (The Jutland-Funen Power Consortium, Denmark) and the Danish Research Academy. The three power stations: Ensted, Funen and Nordjylland-Vendsyssel are acknowledged for allowing and for helpful support during the full-scale trials. The Energy and Environmental Research Center, University

Classification System for Ash Deposits Based on SEM Analyses

215

of North Dakota, is acknowledged for support with development of SEM-EDX techniques at GEUS.

REFERENCES Abbott, M.F. and Austin, L.G. (1982) Studies on slag deposits formation in pulverized-coal combustors. 4. Comparison of sticking behavior of minerals and low-temperatures and ASTM high-temperature coal ash on medium carbon steel substrates. FUEL, 61 (8), 765–770. Bool I I I , L.E. and Helble, J.J. (1996) Iron oxidation state and its effect on ash particle stickiness. In L.L. Baxter

and R. DeSollar (Eds.), Applications of advanced technology to ash-related problems in boilers. Plenum Press. Bryers, R.W. (1996) Fireside slagging, fouling, and high-temperature corrosion of heat-transfer surfaces due to impurities in steam-raising fuels. Prog. Energy. Combust. Sci., 22, 29–120. Cutler, A.J.B. and Grant, C.J. (1975) Corrosion of iron and nickel base alloys in alkali sulfate melts. Conf.

Metal-Slag-Gas Reactions. Electroch. Soc. Princeton. N J. Hatt, R.M. (1990) Fireside deposits in coal-fired utility boilers. Prog. Energy Combust. 235–241.

Huggins, F.E., Kosmack, D.A., Huffman, G.P. and Lee, R.J. (1980) Coal mineralogy by SEM image analysis. Scanning Electron Microscopy. 1, 531–540. Hurley, J.P., Benson, S.A. and Mehta, A.K. (1994) Ash deposition at low temperatures in boilers burning high calcium coals. In J. Williamson and F. Wigley (Eds.), The Impact of ash deposition on coal fired plants. Ed. Taylor and Francis. Hurley, J.P., Benson, S.A., Erickson, T.A., Allan, S.E. and Bieber, J. (1995) Project calcium. Final report.

DOE/MC/10637-3292. Hurley, J.P. and Benson, S.A. (1995) Ash deposition at low temperatures in boilers burning high-calcium coals. 1) Problem definition. Energy and Fuels. 9, 775–781. Jones, M.L., Kalmanovitch, D.P., Steadman, E.N., Zygarlicke, C.J. and Benson, S.A. (1992) Application of SEM techniques to the characterization of coal and coal ash products. In H.L.C. Meuzelaar (Eds.),

Advances in coal spectroscopy. Plenum Press. Larsen, O.H., Laursen, K. and Frandsen, F. (1996) Danish collaborative project on ash deposition in PF-fired boilers. In L.L. Baxter and R. DeSollar (Eds.), Applications of advanced technology to ash-related problems in boilers. Plenum Press. Laursen, K. (1997) Characterization of minerals in coal and interpretations of ash formation and deposition in pulverized coal tired boilers. Ph.D. Thesis. Geological Survey of Denmark and Greenland. Report 1997/65. ISBN 87-7881-022-7. Laursen, K., Frandsen, F. and Larsen, O.H. (1996) Slagging and fouling propensity: full-scale tests at two power stations in western Denmark. In L.L. Baxter and R. DeSollar (Eds.), Applications of advanced tech-

nology to ash-related problems in boilers. Plenum Press. Laursen, K., Frandsen, F. and Larsen, O.H. (1997) Ash deposition trials at three power stations in Denmark. Paper presented at the Engineering Foundation Conference in Kona, November, 1997. Lee, R.J., Huggins, F.E. and Huffman, G.P. (1978) Correlated Mössbauer-SEM studies of coal mineralogy. Scanning Electron Microscopy, 1 , 561–568. Lee, D.J. and Whitehead, M.E. (1982) Microanalysis of scales and deposits formed on corroding furnace tubes in coal-fired boilers. In D.B. Meadowcroft and M.I. Maning (Eds.), Corrosion resistant material for coal

conversion systems. Applied Science Publishers. London and New York. Moza, A.K. and Austin, L.G. (1981) Studies on slag deposits formation in pulverized coal combusters. 1 . Results on the wetting and adherence of synthetic coal ash drops on steel. FUEL, 60 ( 1 1 ) , 1057–1064. Osborn, G.A. (1992) Review of sulfur and chlorine retention in coal-fired boiler deposits. FUEL, 71 (2), 131–142.

Raask, E. (1985) Mineral impurities in coal combustion. Behavior, problems, and remedial measures. Hemisphere publishing corporation. Springer Verlag. Richards, G.H., Harb, J.N. and Baxter, L.L. (1996) Investigations of mechanisms for formation of deposits for

two Powder River Basin coals. In L.L. Baxter and R. DeSollar (Eds.), Applications of advanced technology to ash-related problems in boilers. Plenum Press. Skorupska, N.M. and Carpenter, A.M. (1993) Computer controlled scanning electron microscopy of minerals

in coal. IEA Coal Research. Perspectives, 1–21. Straszheim, W.E., Yousling, J.G., Younkin, K.A. and Markuszewski, R. (1988) Mineralogical characterization

216

K. Laursen and F. J. Frandsen

of lower rank coals by SEM-based automated image analysis and energy-dispersive X-ray spectrometry. FUEL, 67, 1042–1047. Wain, S.E., Livingston, W.R., Sanyal, A. and Williamson, J. (1992) Thermal and mechanical properties of boiler slags of relevance to sootblowing. In Proceedings of the Engineering Foundation conference on Inorganic transformations and ash deposition during combustion. Palm Coast, Florida. 10–14 March, 1991. New York. American Society of Mechanical Engineers.

Walsh, P.M., Sayre, A.N., Loehden, D.O., Monroe, L.S., Beer, J.M. and Sarofim, A.F. (1990) Deposition of bituminous coal ash on an isolated heat exchanger tube: effects of coal properties on deposition growth. Prog. Energy Combust. Sci., 16, 327–346. Zygarlicke, C.J. and Steadman, E.N. (1990) Advanced SEM techniques to characterize coal minerals. Scanning Microscopy, 4 (3), 579–590.

DETERMINATION OF AMORPHOUS MATERIAL IN PEAT ASH BY X-RAY DIFFRACTION Minna S. Tiainen, Juha S. Ryynänen, Juha T. Rantala, H. Tapio Patrikainen, and Risto S. Laitinen University of Oulu, Department of Chemistry

Linnanmaa FIN-90571 Oulu, Finland

1. INTRODUCTION The combustion of peat in power plant boilers has increased in recent years in Finland. While boilers involving pulverized fuel are still common, the utilization of flu-

idized bed boilers is rapidly increasing. FCB boilers are best suited for fuel with a low energy value because the increased efficiency is beneficial for the community heat

distribution. Peat has a low energy value and high water content and therefore resembles low rank lignitic coal [Moilanen et al., 1993]. The sulfur content of Finnish peat, however, is very low. The inorganic material inherent in peat is typical to that of the plants in the

bog [Spedding, 1988] and therefore the slagging tendency of peat ash can be expected to be dependent on the bog from which the peat is originating. Generally the operation of power plants utilising peat is unproblematic, but in some cases severe slagging may occur leading to a plant shutdown thus causing significant economical losses to the entire chain of energy production. Therefore the slagging tendency of peat ash needs to be predicted from the peat fuel prior to its combustion. The slagging of peat ash is connected with the high iron content leading to the formation of low melting point iron aluminosilicates [Heikkinen et al., 1997]. The partial melting of ash particles might lead to agglomerate formation in the fluidized bed. While SEM-EDS connected with an automated image analysis provides a convenient method to investigate the formation of the coating on the bed particle as well as the nature of the adhesive material binding the bed particles [Virtanen et al., 1997], The ease of agglomerate formation and slagging can also be tested by inspecting the compression strength of ash that can be considered as a measure for the degree of sintering of the ash particles [Hupa et al., 1989]. Since the melting and sintering—and ultimately slagging—all involve the formation of amorphous material in ash, it is important to devise an independent method to determine its content. The X-ray powder diffraction (XRD) technique is commonly used to identify the crystalline phases in solid samples. However, it also offers the option to determine the Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

217

218

M. S. Tiainen et al.

content of the amorphous material, since the presence of non-crystalline phases creates a broad hump in the diffraction pattern, called the amorphous halo. The position and the width of the halo indicate the distribution of interatomic distances in the structure. The area under the halo depends on the amount of amorphous material in the sample and can therefore be used in its semiquantitative determination [Nakamura et al., 1989]. In this work we report a systematic study of the formation of amorphous material as a function of peat ash composition and its thermal history. The work is mainly carried out for synthetic oxide mixtures, but comparison is made to actual samples of peat ash obtained in standard laboratory conditions, pilot reactors, and in the power plant boilers.

2. EXPERIMENTAL A Siemens D5000 diffractometer with a goniometer was used for recording of the powder diffraction diagrams using radiation and 40mA) and a zirconium filter. Step size was 0.02° and the counting time was 0.3 seconds for each step. Diffraction diagrams were recorded at a range 5–60°.

The amorphous material induces a broad hump in the diffraction diagram, called a halo. The area of the halo is related to the content of the amorphous material in the sample.

The calculation was made by recording the background at three values of the angle: 52°, 57° and 59°. The integration over the whole range of 5–60° yielded the total area over the background (see Fig. 1). By use of profile fitting the area of peaks could be subtracted thus yielding the area of the halo.

The calibration standards were made by mixing silica gel and synthetic ash at different proportions (0, 10, 30, and 50 w-% of silica gel). They were ground in a

mortar for one minute and homogenized manually in an agate mortar. The powder diffraction diagrams of the calibration standards are shown in Fig. 2 and the calibration curve relating the area of the amorphous halo to the content of silica gel in the standard is shown in Fig. 3. Synthetic ash was made by mixing suitable metal oxides and sulfates in order to mimic the composition of the actual peat ash. A test series involving nine different samples of synthetic ash was carried out by subjecting the samples to heat treatment at different temperatures (300, 500, 600, 700, 800, 900, 1,000, 1,100, 1,200°C) for 1.5h. In addition, the tenth sample was included in the test series without heat treatment. After cooling to room temperature all samples were ground and homogenized as described for the calibration standards.

Determination of Amorphous Material in Peat Ash by X-Ray Diffraction

219

The slagging Finnish peat was ashed using the ASTM (D 3174–89) standard ash procedure [ASTM, 1989a]. The composition of the peat ash was determined by mixing

the ash with (ASTM, D 3682–97) [ASTM, 1989b] and analyzed with a Philips PU-7000 ICP-AES and SpectraSpan IIIB DCP-AES spectrometers. The slagging peat ash contains over 75% of which is thought to be the main reason for its slagging properties [Heikkinen et al., 1997]. A Jeol JSM 6400 scanning electron microscope, Link X-ray spectrometer, and a Link ISIS-image processing program were used to study compositional distribution for the discrete particles in the samples. The acceleration voltage was and the current The sample distance was 15mm and the magnification 130. The sample treatment and the automated image analysis is described elsewhere [Virtanen et al., 1997].

220

M. S. Tiainen et al.

3. RESULTS

3.1. Calibration Direct calibration seems to be possible for the ash samples. The silica gel was selected as an appropriate candidate tor the amorphous material in the standards because the majority of the amorphous material formed in ash during combustion can be thought to be different silicates. Silica gel is known to form a substantial amorphous halo at the range 10–15° (see Fig. 2). Silica gel was mixed with the synthetic ash without heat treatment. It can be seen from Fig. 2 that the area of the halo increases as the content of silica gel increases. It can also be seen that the intensities of the diffraction peaks due to crystalline material decreases as the content of silica gel increases. This is exemplified by the main peak due to quartz at 12°. The calibration curve indicating the halo area as a function of the silica gel content is shown in Fig. 3. The calibration curve is approximately linear with a correlation coefficient of 0.955. This facilitates a semiquantitative determination of the content of the amorphous material in actual ash samples. The formation of the halo is also effected by grinding of the solid material. Long grinding times seem to induce the formation of microcrystalline particle in the powder that may explain the growth of the halo in the samples [Altree-Williams et al., 1981]. It is therefore important to standardise the grinding procedures as fully as possible.

3.2. Peat Ash The actual peat ash that has been selected for this study is known to be problematic in pf-boilers. It easily forms deposits during the combustion. This is believed to be due to the high iron oxide content in the ash [Heikkinen et al., 1997]. The composition of this peat ash is presented in Table 1. The X-ray diffraction diagram for the peat ash is shown in Fig. 4. Two amorphous halos can be seen. One halo located at 20 range 10–13° is caused by silicates and the other that is located at 15–20°, is probably caused by iron silicate phases. The quartz peak at 12° is clearly visible in the diagram. The other possible quartz peaks overlap with those due to iron-containing compounds. The quasi-ternary diagram [Virtanen et al., 1997] for the standard laboratory ash of this peat is shown in Fig. 5. It can be inferred from this figure that iron can form a wide range of aluminosilicates in ash that may explain its melting at low temperatures

Determination of Amorphous Material in Peat Ash by X-Ray Diffraction

221

[Hutchings et al., 1995]. Whereas it is well-known that the composition of the inherent inorganic material in peat is typical to that of the plants from which the peat has been formed [Spedding, 1988], it is also probable that external mineral grains (like quartz) may be introduced into the peat during the production processes.

3.3. Synthetic Ash In this work the synthetic ash was used to model the thermal behaviour of the real peat ash that has been described above. The composition of synthetic ash was made as near to this peat ash as possible. The diffraction patterns of synthetic ash mixtures treated

at different temperatures seem to be very similar (see Fig. 6). This can be explained by the dominating presence of iron oxide in the synthetic ash. The large iron oxide content may well disguise the changes in diffraction patterns. The total area under the two halos

observed at ranges 7–13° and 15–20° is dependent on the temperature during the heat treatment (see Fig. 7). It is obvious that the content of amorphous material in synthetic ash depends on the temperature. The sample that has not been heat treated does not contain amorphous material. Upon heating the synthetic ash at 300°C for 1.5h the content of amorphous material is increased to about 8%. A sample that has been heated at 800–900°C shows

approximately 17% amorphous material. At higher temperature the content of amorphous material decreases. At 1,200°C only 5% of the material seems to be amorphous in the synthetic ash mixture. Possibly new phases are formed at high temperatures that

crystallize during the cooling of the sample to room temperature. It was not possible to identify these phases by XRD because of the dominance of iron oxide. SEM-EDS was

used to elucidate these new phases formed during the heat treatment processes.

222

M. S. Tiainen et al.

The composition of the synthetic ash has also been inspected by SEM-EDS. The quasi-ternary diagrams indicating the compositional distribution of discrete ash particles are shown in Fig. 8. It can be seen that with no heat treatment [Fig. 8(a)] the main components were expectedly pure aluminum oxide, quartz and iron oxide. Upon heating the material at 800°C the formation of aluminosilicates can be inferred [Fig. 8(b)]. At higher temperatures (1,200°C) a second aluminosilicate phase is formed [Fig. 8(c)]. The formation of this second silicate phase may explain the increase of crystallinity in the ash that was deduced by XRD (see Fig. 7). It is interesting to note that the synthetic ash, when heated at 1,200°C, and stan-

dard peat ash exhibited virtually identical contents of amorphous material (see Fig. 7). Consequently, their diffraction patterns are virtually equal (see Fig. 9). However, the SEM-EDS results indicate that the composition of the particles formed during standard ashing of actual peat covers a significantly wider range than those from synthetic ash as

clearly demonstrated by their quasi-ternary diagrams shown in Fig. 5 and Fig. 8(c). The standard peat ash has a more complicated mineral composition than the idealized synthetic ash. The inspection of the compositional distribution of the particles in terms of five elements provides only a rough qualitative approximation. The melting behaviour, however, that is seen as the formation of the amorphous halos may depend on the main phases present in ash. It seems that both the standard peat ash and synthetic ash behave similarly upon heating, and therefore this simple system of synthetic ash may be used to predict the factors affecting the slagging properties of real peat ash.

Determination of Amorphous Material in Peat Ash by X-Ray Diffraction

223

4. CONCLUSIONS X-ray powder diffraction has turned out to be a potential technique to inspect the

slagging properties of ash. In this work we have reported the semi-quantitative determination of the content of amorphous material both in peat ash that is known to be slagging, as well as in synthetic ash the composition of which has been adjusted to mimic

the actual peat ash. It was established that the content of amorphous material in the synthetic ash depended on the temperature the material was heated to. A sample that had not been heated contained virtually no amorphous material. However, already at 300°C about 8% of amorphous material was found in the sample. The content of amorphous material reached a maximum value of 17% at 800°C. Upon further heating to 1,200 °C the content of amorphous material decreased to about 5%. This is probably due to the crystallisation of new phases upon cooling of the sample to room temperature prior to the recording of the XRD diagram. The synthetic ash was found to be a good model for the actual peat ash. Diffraction patterns of the standard peat ash and the synthetic ash that had been heated to

224

M. S. Tiainen et al.

1,200°C were virtually identical. Due to the high content of iron oxide phases in ash XRD could not be used to identify other phases formed during the heat treatments. The quasi-ternary diagrams that were obtained from the SEM-EDS analysis of about 1,000 ash particles provided additional information that can be used to discuss the composition of ash.

5. ACKNOWLEDGEMENTS Financial support from Academy of Finland and Liekki-2 Research Program is gratefully acknowledged.

REFERENCES Altree-Williams, S., Byrnes, J. G. and Jordan, B. (1981). “Amorphous Surface and Quantitative X-ray Powder Diffractometry.” Analyst, 106, 69–75. Nakamura, T., Samcshima, K., Okunaga, K., Sugiura, Y. and Sato, J. (1989). “Determination of Amorphous Phase in Quartz Powder by X-Ray Powder Diffractometry.” Powder Diffraction, 4(1), 9–13. Heikkinen, R., Laitinen, R. S. and Patrikainen, T., Tiainen, M. and Virtanen, M. (In press). “Slagging Tendency of Peal Ash.” Fuel Processing Technology.

Hutchings, I. S., West, S. S., Williamson, J. (1995). “An Assessment of Coal-Ash Slagging Propensity Using an Entrained Flow Reactor.” In L. L. Baxter and R. DeSollar (Eds.), Applications of Advanced Technology to Ash-Related Problems in Boilers, New York, Engineering Foundation.

Hupa, M., Skrifvars, B.-J. and Moilanen, A. (1989). “Measuring the Sintering Tendency of Ash by a Laboratory Method.” J. Inst. Energy, 131–137. Moilanen, A. (1993). “Studies of Peat Properties for Fluidized-Bed gasification.” VTT Publications 149, Espoo, 69. Spedding, P. J. (1988). “Peat” Fuel, 67, 883–900. Standard Test Method for Ash in the Sample of Coal and Coke From Coal, Annual book of ASTM standards,

05.05 (1989a), D 3174–89, 302. Standard Test Method for Ash in the Sample of Coal and Coke From Coal, Annual book of ASTM standards,

05.05 (1989b), D 1857–87, 222. Virtanen, M., Heikkinen, R., Patrikainen, T., Laitinen, R. S., Skrifvars, B.-J. and Hupa, M. (1997). “A Novel Application of CCSEM for Studying Agglomeration in Fluidized Bed Combustion”, Engineering Foundation Coference on The Impact of Mineral Impurities in Solid Fuel Combustion, November 2–7, 1997, Kona, Hawaii.

SYSTEM ACCURACY FOR CCSEM ANALYSIS OF MINERALS IN COAL R. P. Gupta, L. Yan, E. M. Kennedy, T. F. Wall1, M. Masson, and K. Kerrison 2 1

CRC for Black Coal Utilisation Department of Chemical Engineering The University of Newcastle 2 Pacific Power Advanced Technology Centre The University of Newcastle

1. INTRODUCTION The advanced techniques such as CCSEM or QEM*SEM (Skorupska and Carpenter, 1993, and Gottileb et al., 1991) are now able to provide the detailed analysis related to the minerals present in coal. The CCSEM technique uses an automated scanning electron microscope (SEM) and is programmed to scan pre-selected areas of a polished sample to capture the back scattered emission (BSE) images. The mineral particles are automatically detected by an increase in the BSE signal above a pre-set value of signal, termed as threshold intensity, corresponding to that of the coal matter. The electron

micro-beam detects the centre of the mineral grain by an iterative bisection of chords. The area of each mineral grain is also determined.

An energy dispersive X-ray (EDX) spectrum is acquired (0–20 keV) for five seconds from the centre of each particle detected. X-ray intensity data and size and shape parameters for a statistically significant number of particles are then collected at different

magnifications. The relative intensity of elements is related to a mineral type according to some heuristic rules (Steadman et al., 1991). CCSEM and QEM*SEM are the successful automated versions of this technique. The CCSEM technique determines the particular mineral types from the elemental analysis provided by the SEM, its size from measured dimensions and can establish if the mineral grain is included within a coal particle or if it is excluded from coal. This latter analysis requires mounting the coal sample in wax to differentiate the X-ray signal from coal matter and the mounting medium. An analysis of many thousands of mineral grains is then assembled into the final analysis in terms of mineral types, their size distribution and included/excluded Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

225

226

R. P. Gupta et al.

nature. This paper details a study in which two pulverised coals were analysed by three CCSEM systems: the ATC system at Pacific Power, Australia, the BYU system at the Chemical Engineering Department of Brigham Young University in Provo, Utah and MTI system at Microbeam Technologies Incorporated in Grand Forks, North Dakota. The coals were mounted prior to analysis in each laboratory in order to compare sample preparation techniques, and the same prepared sample (stub) was also analysed at ATC and BYU in order to compare the analysis system alone excluding the sample preparation differences. The major aspects leading to differences in analysis results are identified as the sample preparation technique and the number of mineral grains analysed. The ATC technique mixes crushed coal and wax into a pressed pellet, thereby guaranteeing a representative distribution, while the other laboratories disperse coal in molten wax where differential settling effects are possible. If a pulverised coal contains coarse minerals, then a greater number of grains comprising large size minerals will need to be analysed. The estimation of the progressive average of the analysis as the number of mineral grains increase may be used to establish the optimum number. This will depend on the coal characteristics and the application.

2. DIFFERENCES IN CCSEM TECHNIQUES The study has revealed that there are differences in sample preparation and operating conditions of the SEM used by different laboratories. This comparison is detailed in Table 1. Yang and Baxter (1992) have discussed in detail the effect of sample preparation and operating conditions on the results obtained from CCSEM. There are differences in the elemental analysis due to application of ZAF correction when the intensity spectrum is converted to elemental composition. The ZAF correction accounts for the changes in the relative intensity due its atomic number (Z), absorption (A) of X-rays by the sample and the fluorescence (F). A ZAF correction procedure (Holt, 1974) is usually applied to improve the overall elemental composition analysis. BYU applies the ZAF correction to their data, however, the other systems do not apply this correction. The ZAF correction is based on a calibration using standard mineral samples of known

System Accuracy for CCSEM Analysis of Minerals in Coal

227

composition. The correction also accounts for the machine dependent parameters. For Si and Al species, the correction effect is small For the minor components, the correction is significant. The correction procedure should be developed by ATC. The number of mineral grains analysed by ATC was substantially less than that examined by the other laboratories. There were differences in the heuristic rules for the conversion of the elemental analysis to the mineral phases. BYU used a different set of mineral phases and different heuristic rules (Yu et al., 1993). The threshold intensities, defining mineral matter, coal matter and wax, are established by an operator during each analysis. Therefore, there are inherent problems in determining the total mineral content in coal due to assigning of the two threshold intensities required. ATC had problems in setting a threshold intensity to differentiate between coal matter and wax, and consequently, in the estimation of the total coal area. The total mineral matter content (as a percentage of coal) can not, therefore, be determined accurately. It should be noted that MTI determines their total coal area on the basis of the proportion of coal and wax. Setting up of the threshold intensity between the coal matter and the mineral matter appears to be, comparatively, simpler. Although it might be possible to establish a standard technique of sample preparation, and standard operating conditions for the SEM and a set value for threshold intensity, no such standards exist. Comparison of analysis from different laboratories for the same sample is, therefore, the best way of evaluating this technique.

3. RESULTS AND DISCUSSIONS Comparisons are made for the CCSEM elemental analysis presented as equivalent oxides to compare with the XRF analysis of ash, the major minerals, the size distributions of particular minerals and the total minerals as well as coal performance indices. Galbreath et al. (1996) have also investigated the differences in the CCSEM techniques of six different laboratories. There are some inherent differences among these laboratories due to significant differences in operation of respective systems. The present paper also investigates the effect of these differences in CCSEM analyses on the performance indices derived from these analyses.

3.1 Equivalent Oxide Composition The elemental composition can be converted to the equivalent oxide composition and is compared with the standard ash analysis in Table 2, Figs. 1 and 2. It must be noted here that the XRF analysis accounts for all the inorganic material present in the coal,

228

R. P. Gupta et al.

whereas the CCSEM technique cannot detect any inorganic matter that is organically

associated or in grains smaller than one micron in size. The oxide composition determined from CCSEM analysis is close to that determined by XRF except for aluminum and calcium oxides. The CCSEM data from BYU also gave similar values for . It is not been possible to explain the low estimation of alumina by ATC.

System Accuracy for CCSEM Analysis of Minerals in Coal

229

3.2 Comparison of Major Minerals The mineral grains are classified into a number of mineral types based on the elemental composition and heuristic rules. The heuristic rules used by ATC are the same as

those used by MTI (Steadman, 1991). The heuristic rules used by BYU are slightly different (Yu et al., 1993). The mineral compositions for major mineral phases thus determined are compared in Figs 3 and 4 for Coal A and Coal B respectively. Hereafter, BYU(P) is the analysis by BYU on the same coal sample (stub) as that used by ATC. It can be seen that the BYU(P) results of mineral contents are closer to the ATC results than the MTI and BYU results, except that there is a 9% deviation of clays content between the BYU(P) and ATC data. This suggests that the different sample preparation techniques add to the differences in the analyses. Similar differences in major minerals from different laboratories have been observed by Galbeath et al. (1996).

3.3 Comparison of Size Information The Particle Size Distribution (PSD) of mineral matter from ATC is compared here with that determined from BYU. The comparison could not be made with the results from MTI, as that laboratory had crushed the samples before analysis. The differences are quite significant for stubs prepared by different laboratories. However, approximate results are obtained when the same stubs are analysed by ATC and BYU(P). This suggests the sample preparation technique used by BYU is different. The particle size distributions (PSD) from these laboratories are based on 2Dimensional images from SEM and need stereological correction to obtain the true PSD. A stereological correction for the particle size distribution of minerals PSDs was also developed to transform data from two dimensional information (as measured in CCSEM) to three dimensional results, which is required for estimation of particle volumes. The

230

R. P. Gupta et al.

correction does influence the absolute PSD values, making the PSD curve shift to coarser grain sizes. However, it was not found to influence the ranking of coals according to the indices proposed here. The correction has not been used for any of the above results, as other laboratories do not apply this correction. However, the correction is expected to be very important in determining the mineral-mineral and mineral-coal associations. Figures 5 and 6, and Table 3 show the comparison of mineral PSD determined at ATC and BYU.

4. ACCURACY OF RESULTS The accuracy of results is a function of number of particles analysed and the uniformity of coal distribution in the stub being examined. The uniformity of distribution

is improved by refining the sample preparation technique. The effect of number of grains analysed on the accuracy of results is discussed in the following sub-section. The

System Accuracy for CCSEM Analysis of Minerals in Coal

231

232

R. P. Gupta et al.

usefulness of the CCSEM data lies in deriving advanced indices for the performance of coal. The performance indices derived from the CCSEM data from different laboratories for two coals are also compared in this section.

4.1 Number of Mineral Grains The fraction of mineral grains less than size determined by ATC for both the coals matches that from BYU(P). However, there are significant differences in the coarse size fractions for Coal B. One large grain of can result in a difference of about 2% by mass. This can influence the mineral percentage or composition significantly. Table 4 presents the number of grains larger than from the ATC system. It is observed that 1–2 mineral grains (0.1–0.2% by number) of siderite or pyrite larger than 50µm, analyzed in the total area, comprise about 3–4% of these minerals by mass. Rolling averages of individual components of mineral grains shown in Figs 7 and 8 suggest the analysis of additional mineral grains for better estimates of compositions of coal A is needed. Analyzing 400 mineral grains for coal A at 50X magnification is not adequate. For coal B, it appears, an adequate number of mineral grains may have been examined. However, if a coal performance index is based on the presence of large pyrite grains, then additional grains would need to be analysed.

System Accuracy for CCSEM Analysis of Minerals in Coal

233

4.2 Coal Performance Indices There have been several attempts in developing advanced indices for ash deposition and abrasion based on CCSEM data (Gupta and Wall, 1995, Kalmanovich, 1992, Wigley and Williamson, 1995, and Zygarlicke et al., 1992). First order estimates of indices for ash deposition, and are defined as the mass percentage of particles with basic oxides greater than particular levels (40% and 80% respectively). , the mass percentage of equivalent ash particles having viscosity less than particular value (l00 Pa.s or 1,000 poise at 1,250°C) is also defined as one of the ash related index. The first order estimates for indices of abrasion, and are derived from the mass of quartz particles greater than a certain size (10 µm and 20 µm, respectively), expressed as a percent-

age of total minerals. The indices for Coal A and Coal B are compared in Figs. 9 and 10. The indices determined from the same plug, related to ash deposition, are similar. However, the indices from ATC and BYU(P) for abrasion differ for Coal B. This may again be related to the number of mineral grains analyzed by ATC. Fine quartz particles below in size may easily be entrained in PFBC and cause erosion of turbine blades. Another index, is, therefore, defined as mass fraction of quartz particles less than in size (expressed as a percentage of total minerals) for coal erosion performance in PFBC. The indices derived from the two laboratories are seen to be similar. The first order estimates for the indices can be improved with an analysis of the included/excluded nature of the mineral grains. For example, included quartz is reactive and would either react with other mineral grains present in the same coal particle or would soften in a highly reducing environment. On the other hand, excluded quartz particles would not react in this way due to the oxidising environment in the combustion stream. Thus, a second-order estimate for abrasion index would be based on the mass fraction of excluded quartz particles bigger than a critical size. The comparisons for included and excluded minerals could not be made in this study. MTI and BYU did not determine the included/excluded mineral distribution. A semi-automatic and interactive computer program for identifying included/excluded

234

R. P. Gupta et al.

minerals from CCSEM data files and images has been developed in this project. This technique has been validated by comparison with results from the Energy & Environment Research Center (EERC) at BYU. A fully automated system being developed at ATC awaits proving.

System Accuracy for CCSEM Analysis of Minerals in Coal

235

5. CONCLUSIONS It can be concluded from this international cooperative study that the main factor affecting the accuracy of CCSEM analysis is the coal sample preparation technique. The differences of the CCSEM systems including their operating conditions only play a minor

role to account for the discrepancy of CCSEM measurements, as different laboratories gave similar CCSEM results on the same coal sample stub. The performance indices related to ash deposition determined from the CCSEM analyses from BYU and ATC were found to be similar. This study indicates that the examined number of mineral grains which is necessary for a correct analysis will depend on the character of the minerals in a coal, in that the presence of a few large grains can influence the composition reported on a mass basis. The need for this information depends on the application, for abrasion issues it will clearly be important. An analysis system may also be configured to preferentially detect such grains. The reporting of results as the mass of particular minerals related to the mass of coal matter, rather than related to the total minerals, would also give results which are satisfactory for the fine minerals. The analysis procedure can therefore be optimized once the application of the analysis is specified.

6. ACKNOWLEDGEMENTS The authors are grateful to John Harb and Peter Slater of Brigham Young University and Steve Benson of Microbeam Technology Inc. The authors also acknowledge the financial support for the project from Cooperative Research Centre for Black Coal Utilisation.

7. REFERENCES Dehoff, R.T. and Rhines, F.N. (1968), Quantitative Microscopy, McGraw-Hill, New York. Holt, D.B. et al. (1974), Quantitative Scanning Electron Microscopy, Academic Press.

Galbreath, K. et al. (1996), “Collaborative Study of Quantitative Coal Mineral Analysis Using ComputerControlled Scanning Electron Microscopy”, Fuel, Vol. 75 No. 4. Gottlieb, P. et al. (1991), The Characterisation of Mineral Matter in Coal and Fly Ash, Eng. Foundation Conf., Palm Coast., FL., USA. Gupta, R.P. and Wall, T.F. et al. (1995), Inorganic Transformations and Ash Deposition During Combustion,

Eng. Foundation Conf., New Hampshire. Kalmanovitch, D.P. (1991), ibid. Skorupska, N.M. and Carpenter A.M., Computer Controlled Electron Microscopy of Minerals in Coal, IEA

Report (1993). Steadman, E.N. et al. (1991), Inorganic Transformations and Ash Deposition During Combustion, Eng. Foun-

dation Conf., Palm Coast. Williamson, J. and Wigley, F. (1995), ibid. Yang, N.Y. and Baxter, L.L. (1991), The Characterisation of Mineral Matter in Coal and Fly Ash, Eng. Foun-

dation Conf., Palm Coast., FL., USA. Yu, et al. (1993), The Impact of Ash Deposition on Coal Fired Plants, Engg. Found. Conf., Solihull. Zygarlicke, et al. (1991), Inorganic Transformations and Ash Deposition During Combustion, Eng. Foundation

Conf., Palm Coast.

This page intentionally left blank.

THE MICROSTRUCTURE AND MINERAL CONTENT OF PULVERISED COAL CHARS

F. Wigley and J. Williamson Imperial College Department of Materials London SW7 2BP, UK

1. INTRODUCTION The nitrogen release and burnout behaviour of a suite of eight world-wide coals has been studied in a UK collaborative research programme. As part of this programme,

samples of pyrolysis char were prepared in a drop-tube furnace at International Combustion Ltd and analysed by CCSEM at Imperial College. The eight coals used were Asfordby, Betts Lane, Hunter Valley, Kaltim Prima, Koornfontein, La Jagua, Pittsburgh and Thoresby. The objective of this CCSEM study was to characterise the particles from the suite of coal char samples, in respect of the size, porosity, microstructure and ash content of individual char particles. The aim of this work was to provide a description of the char samples that would assist in a comparison between the behaviour of the eight coals in the drop-tube furnace, and would provide basic data for the modelling of char combustion reactions.

2. ANALYTICAL TECHNIQUE Each coal sample was mounted in epoxy resin and prepared as a polished crosssection perpendicular to the settling direction, to eliminate the effects of density segregation. Digital back-scattered electron images were collected at magnification at forty random points across the sample cross-section (e.g. Figs. 1–4). Clusters of pixels with intensities corresponding to the average atomic number range of the carbonaceous component of char were located. In addition, clusters of pixels with intensities corresponding to the average atomic number range of ash derived from coal mineral matter were located. The carbon and ash pixel clusters were combined, to identify the char particle cross-sections. Each char particle cross-section was analysed for size, shape

and composition (carbon, ash and included pores). In addition to the parameters that Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwcr Academic / Plenum Publishers, New York, 1999.

237

238

F. Wigley and J. Williamson

have been measured for each cross-section, an estimate of particle wall thickness was calculated. Although some attempt has been made to automatically rejoin breaks in the crosssections of porous particles, the extent of this process has been limited by the desire not to connect adjacent cross-sections into a single particle. No correction was made for the difference between the apparent size, in cross-section, and the true dimensions of the mineral occurrences. No account has been taken of pores that were not fully enclosed. To minimise the effect of noise within the image, particle cross-sections smaller than in size were ignored. Between 1,200 and 5,000 char particles were analysed for each sample.

The Microstructure and Mineral Content of Pulverised Coal Chars

239

3. RESULTS

3.1. Particle Size There are a number of parameters that can be used to measure the sizes of particles in cross-section. For this analysis, a commonly used approach has been adopted, in which size has been defined as the diameter of a circle with the same area as the carbon, ash and enclosed pores in the particle cross-section. The average particle sizes for the eight samples are listed in Table 1, and the particle size distributions for each sample are shown in Fig. 5. The eight char samples can be grouped by average particle size in the following way: La Jagua, Asfordby, Kaltim Prima Thoresby, Hunter Valley, Koornfontein Pittsburgh, Betts Lane.

240

F. Wigley and J. Williamson

The char particle size distributions were all similar. Figure 5 shows a typical size distribution, in which there are very few particles that are smaller than the lower size limit of the char analysis procedure However, the shape of the size distributions at the

upper end of the size range indicates that the very largest particles in the char samples have not been analysed.

3.2. Particle Porosity Particle porosity has been measured as a pore fraction, using the ratio between the area of the enclosed pores and the total area of the particle cross-section (carbon + ash + enclosed pores). The eight char samples can be ranked and grouped by average particle porosity (Table 2—Pore fraction) in the following way: La Jagua, Asfordby, Kaltim Prima Koornfontein, Hunter Valley, Thoresby Betts Lane, Pittsburgh. Char porosity results from the interaction between plastic carbonaceous material and evolving gasses; included mineral matter probably plays little part in this process. The “pore/carbon ratio”, calculated as the ratio between the enclosed pore area and the carbon area for a particle, has been calculated in order to eliminate the effect of the ash component from the estimation of porosity. Average pore/carbon ratios (Table 2) rank the char samples in the same way as average porosity (above), but this ranking, which is

The Microstructure and Mineral Content of Pulverised Coal Chars

241

independent of the widely varying ash contents of the chars, shows a clearer distinction between the three char groups. From Table 3, it can be seen that char particles below about diameter were fully solid for all eight samples. For char particles larger than about the area fraction of enclosed pores increased with increasing particle size. The epoxy resin used to mount the char samples contains pores, which may occasionally appear as solid circular char particle cross-sections. In addition, the inability of the mounting medium to penetrate large enclosed pores may result in these pores being identified as carbon, rather than pore. As a result, the measured porosity of the larger char particles Table 3) is less precise than for smaller char particles, and probably underestimates the true porosity.

3.3. Particle Ash Content Particle ash content has been measured using the ratio between the area of the inherent ash and the total area of the particle cross-section (carbon + ash + enclosed pores). The average ash contents vary between 5 and 19% (Table 2), and correlate well with the proximate ash contents of the chars with the exception of the Kaltim Prima sample. Because the ash content has been determined on an area basis, rather than by mass, the measured ash contents are lower than the actual values. The way in which ash content varies with particle size is not consistent between samples (Table 4).

242

F. Wigley and J. Williamson

3.4. Particle Composition Each char particle cross-section has been analysed in terms of three components: carbon, ash and enclosed pores. In every char sample, the particles have a wide range of compositions; there are no clear distinctions between porous and non-porous p or between particles with and without ash. Many cross-sections contained carbon, ash and enclosed pores.

3.5. Particle Microstructure Char particle microstructures are conventionally divided into “solid”, “network” and “cenosphere” categories. Both network and cenosphere char particles are porous; cenospheres contain one large pore, while network chars contain several pores. Further microstructural categories have been proposed for particles that contain a significant ash component, for porous particles with thick and thin walls, and for particles that contain features of more than one “ideal” microstructure, as reviewed by Cloke and Lester [1994]. These frameworks for classifying char particle cross-sections are based on optical microscopy, which is unable to reliably identify ash inclusions and usually classifies them as pores. In addition, char classification is frequently performed manually (rather than automatically), with a resultant emphasis on larger char particles that display more “ideal” microstructures. The char classification is frequently described, rather than defined. The analytical procedure used for the char samples described in this paper is readily able to distinguish carbon, ash and enclosed pores, and has measured the same parameters for a large number of randomly-chosen char particle cross-sections of widely varying size and shape. However, this analytical procedure is currently unable to re-create the microstructure of large fractured particles in the same way as the subjective analysis of a skilled microscopist. Given the continuous distribution of particle compositions described above, a simple, numerically based classification scheme for the microstructure of char particle cross-sections has been devised. The particles analysed in this study have been classified as “solid” if they have a pore fraction below one tenth, and as “porous” if the have a greater pore fraction. Between a quarter and one half of the char particles in each sample were solid (Table 5), and the fraction of solid particles increased in the following way: Pittsburgh Betts Lane, Koornfontein, Hunter Valley, Thoresby Asfordby, La Jagua, Kaltim Prima.

The Microstructure and Mineral Content of Pulverised Coal Chars

243

Almost all the char particles smaller than 8µ m are solid (Table 6), and the fraction of solid particles decreased with increasing particle size. Between 15% and 35% of the larger char particles were solid, but these values were not consistent for the two size bands (Table 6) and probably reflect the analytical errors described above. A value for “wall thickness” has been measured for each particle analysed, based on the average distance between each pixel in the cross-section of the particle (combined carbon and ash) and the nearest pixel on the internal or external perimeter of the particle cross-section. The porous particles have been classified as “thin-walled” if their thickness is less than one fifth of their diameter, or as “thick-walled” if they have a greater thickness. The proportion of thick-walled porous particles in each char sample was relatively constant at about 10% (Table 5), although the fraction for the Pittsburgh sample was higher. The proportion of thin-walled porous particles in each char sample showed the opposite variation to the proportion of solid particles, and ranks char samples in the reverse of the ranking shown above. The fraction of porous thin-walled particles in each size band is listed in Table 7. There are very few thin-walled porous particles below in size, and the proportion of this particle type increased rapidly with increasing particle size. For the larger particles, the analytical errors described above made measurements of thickness slightly less reliable.

F. Wigley and J. Williamson

244

4. DISCUSSION From the results described above, it is clear that none of the char samples could be distinguished by the presence or absence of particles of a particular size, porosity or microstructural type. However, there is a high degree of consistency between the ranking and grouping of the eight coal char samples, based on average particle size, average particle porosity, average pore/carbon ratio, average fraction of solid particles, and average fraction of thin-walled porous particles (Table 8). Comparisons between samples based on the properties of individual particle size bands frequently produced the same ranking and grouping. The ranking of the eight char samples by char microstructure did not correlate with the proximate analyses of either the chars or their parent coals (Table 9). Comparison

with data from other experiments within the collaborative project indicates that the char particle properties do not correlate with either initial rates of volatile loss or final levels of burnout during combustion trials of these eight coals.

5. CONCLUSIONS Based on CCSEM analysis of eight char samples prepared from coals of different rank and ash contents, the following conclusions can be drawn: 1. The carbon, ash and pore components of char particle cross-sections can be distinguished in back-scattered electron images; image processing of these images has been used to provided quantitative descriptions of char samples.

The Microstructure and Mineral Content of Pulverised Coal Chars

245

2. There was no significant difference between the eight char samples in terms of the presence or absence of char particles of a particular size, porosity or microstructure. 3. There was a consistent ranking of the char samples with respect to size, porosity, pore/carbon ratio, fraction of solid particles, and fraction of thin-walled porous particles. This ranking separated the coal char samples into three groups: La Jagua, Asfordby, Kaltim Prima; Koornfontein, Hunter Valley, Thoresby; Betts Lane, Pittsburgh. For all eight samples, char particles smaller than about contained very little porosity. The level of porosity increased rapidly in the larger particles, with the most porous particles having walls that were thin with respect to their size.

6. ACKNOWLEDGEMENT The authors gratefully acknowledge the funding and samples provided by Dr A. Thompson of International Combustion Ltd.

7. REFERENCE Cloke, M. and Lester, E. (1994). “Characterisation of coals for combustion using petrographic analysis: a review.” Fuel, 73, 315–320.

This page intentionally left blank.

FIRESIDE CONSIDERATIONS WHEN COFIRING BIOMASS WITH COAL IN PC BOILERS# Allen L. Robinson, Larry L. Baxter, Gian Sclippa¶, Helle Junker § , Karl E. Widell†1, Dave C. Dayton, Deirdre Belle-Oudry 2 , Mark Freeman, Gary Walbert‡, and Philip Goldberg3 1

Combustion Research Facility Sandia National Laboratories Livermore, CA 94551-0969 2 National Renewable Energy Laboratory Golden, CO 80401-3393 3 Federal Energy Technology Center Pittsburgh, PA 15236-0940

INTRODUCTION This paper discusses fireside issues associated with cofiring biomass and coal in pulverized-coal-fired (pc-fired) boilers. The primary motivation for such cofiring is effective reduction of emissions, as sustainably managed biomass feedstocks are essentially neutral. Sustainable management in this context means that biomass resources are consumed at the same rate as they are produced with no adverse environmental effects. Under such management, generated during biomass combustion is reincorporated in plants via photosynthesis, helping to close the carbon cycle and resulting in virtually no atmospheric accumulation. In many cases, cofiring offers several additional environmental benefits, including: (1) reduced production of criteria pollutants such as oxides of sulfur and nitrogen (2) reduced open-field burns and relief from the environmental costs associated with them; and (3) decreased quantities of landfill or other waste materials. The realization of these benefits requires optimization of power systems to make best use of the blended fuels. The US has considerable experience in operating biomass-fired power stations, #

Presented at the Engineering Foundation Conference on the Impact of Mineral Impuriteis in Solid Fuel Combustion, Nov. 2–7, 1997, Keauhou Beach Hotel, Kona, Hawaii ¶ Onsite Engineering, Livermore, CA § lsamprojekt/Aalborg University, Denmark † Professor, Mechanical Engineering, Aalborg University, Denmark ‡ Parsons Power Group, Inc., Library, PA Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

247

248

A. L. Robinson et al.

ranging from reliable operation of wood-based systems to troublesome operation of systems using agricultural residues (Baxter, Miles, Jr., Jenkins, Milne, Dayton, et al., 1997a). Overall efficiencies and availabilities of biomass power plants are low compared to traditional coal-fired plants. For example, net plant conversion efficiencies of biomass systems are roughly half that of coal (16–20% for biomass compared with 33–38% for coal, HHV basis), and availability of biomass stations is considerably lower than that of coal (Baxter, Miles, Thomas R. Miles, Jenkins, Dayton, Milne, et al., 1996). At the same time, the demand for renewable energy is growing, with biomass offering some of the greatest potential for significant market penetration.

Biomass-coal cofiring takes advantage of the best aspects of both coal-based and biomass-based systems while addressing and other environmental and economic issues. Power production from coal exceeds that from biomass by at least two orders of

magnitude. Therefore, even a relatively minor incorporation of biomass as a cofiring fuel with coal would significantly increase the potential unsubsidized market for renewable energy (biomass). Biomass cofired with coal achieves an effective reduction in emissions proportional to the thermal input from biomass. The efficiency of power production from biomass is nearly doubled in a cofired system compared to use of the same fuel in stand-alone biomass facilities. Biomass cofiring can reduce emissions of criteria pollutants compared to use that produced from pure coal systems. In addition, capital and operating costs decrease as a fraction of total power cost in a cofired system compared to existing biomass facilities.

This document presents selected results from a multi-year, multi-laboratory, interagency effort to evaluate the use of biomass as a cofiring fuel during power production from coal-fired power plants. The essential objectives of this project are to assess the potential fireside impacts of firing biomass in coal-based systems and to demonstrate the potential emission reductions and other environmental benefits from cofiring. The scope of the project is limited to fireside issues. Other issues of major importance, such as ash disposal, fuel handling and preparation, regulations and other institutional issues, and resource assessment/economic analyses of cofiring options at specific plants are not included except as they impact the fireside behavior. In this document, we use results of one of our most recent tests using switchgrass (SG) and Pittsburgh (P8) coal to illustrate cofiring combustion behavior. In several cases, we introduce results from our previous experiments (Robinson, Junker, and Baxter, 1997b) to help place the SG/P8 results in context.

TECHNICAL APPROACH

Experimental Facilities This investigation utilizes the experimental facilities and expertise of three institutes. The molecular beam mass spectrometer (MBMS) (Dayton, French, & Milne, 1995) used at the National Renewable Energy Laboratory (NREL) to characterize the composition of fuel off gases provides fundamental chemical data as a function of extent of combustion. The Captive Particle Imaging (CPI) system at Sandia National Laboratories (SNL) is a small flow reactor used to monitor the combustion history of an individual particle over its lifetime (Hurt and Davis, 1994). Sandia’s Multifuel Combustor (MFC) is a small (0.03–0.1 MMBtu/hr, depending on fuel type), pilot-scale, down-fired, turbulent flow facility that is capable of firing a wide variety of fuels and of reproducing the temperature and gas composition histories experienced by particles in commercial-scale pc systems systems (Baxter, 1992). The Combustion and Environmen-

Fireside Considerations when Cofiring Biomass with Coal in PC Boilers

249

tal Research Facility (CERF) at the Federal Energy Technology Center (FETC) is a larger-scale pilot facility (0.5 MMBtu/hr) that allows commercial-scale engineering issues to be addressed (Freeman, Chitester, James, Ekmann, & Walbert, 1997). The combination of the facilities, together with observations from commercial-scale trials, allows more definitive data interpretation than any of the facilities could accomplish alone.

Fuel Selection and Characterization This project has examined a wide range of commercially viable coal and biomass fuels. This document focuses on the combustion behavior of a Wisconsin switchgrass

(SG) and a high-volatile A bituminous Pittsburgh coal (P8), properties of which are indicated in Table 1. Results from the previous tests are used to place the P8/SG results in context. Several replicate analyses were performed on both SG and P8. The coal properties exhibited little variation from analysis to anlaysis, but the biomass properties exhibited much greater variation. Such heterogeneity is a common characteristic of biomass. The Pittsburgh coal was fired as a utility grind (70% through 200mesh) fuel. The switchgrass sample was prepared using a tub grinder followed by a micropulverizer. resulting in a final product that passes through a 1 mm screen. Unlike equant coal particles, many SG particles exhibit aspect ratios of 3 or higher; consequently, the largest particles were approximately cylinders. Specifically, 98% of the sample was less than 20mesh (0.84mm) and 55% was less than 100 mesh (0.15 mm). The biomass and coal were blended prior to injection in all of the experiments reported here. All laboratories conducted comparative experiments with pure coal. SNL and NREL also conducted comparative experiments with pure biomass.

RESULTS AND DISCUSSION

Ash Deposition Mechanisms of ash deposition include inertial and eddy impaction, thermophoresis, condensation, and chemical reaction (Baxter, 1993). The contribution of each of these

A. L. Robinson et al.

250

mechanisms depends on local chemistry, aerodynamics, and operating conditions, but alkali materials influence both the rates of deposition and the properties of the deposits much more than their concentration sometimes suggests. The mechanisms of alkali

release, deposition, and reaction are at least qualitatively understood. Alkali chlorides represent the most volatile of the common forms of alkali and are expected to be the most easily released from the particle. Chlorine is volatile in essentially all of its forms

in fuels and typically is released from the fuel early in the combustion process. Alkali, on the other hand, is dominantly released from the fuel during the high-temperature, char oxidation stage. Once vaporized, the alkali condenses on heat transfer surfaces, where it often contributes to fouling, slagging, and corrosion. Ash deposition in pc boilers is commonly most severe on the first bank of superheater/reheater tubes in the convection pass but can also affect the furnace and backpass regions significantly. Furthermore, reflective deposits that may not present a maintenance problem directly can alter heat absorption patterns in the boiler and raise furnace exit gas temperatures. All of these issues have been noted in previous experiments with a variety of coals, biomass fuels, and blends of

coal and biomass. Investigation and interpretation of the P8/SG ash deposition data are made in the context of these conceptual mechanisms. Figure 1 illustrates MBMS data indicating the extent of alkali vaporization during combustion of P8, SG, and blends of 5, 15 and 25% SG. Blend ratios are based on the high heating values of the fuels. Time-resolved results from the MBMS are integrated during the char oxidation phase for several fuel blends to produce a semi-quantitative

indication of the vapor concentration, as shown in Figure 1 for NO, HCl, K, and The detection limit is approximately 10ppm, which corresponds to an ion signal of about 1 in the units of the figure. The error bars represent one standard deviation as determined

from three replicate experiments. The most significant conclusion from the figure is that the K and HCl signals are both below detection limits, indicating essentially no vapor-

ization of K or evolution of Cl. This contrasts sharply with similar experiments conducted on higher-chlorine biomass fuels. The relatively low chlorine concentrations and high sulfur concentrations of these fuels inhibit the formation of high concentrations of alkali-containing vapors; KCl is the alkali species with the highest vapor pressure. Other than potassium- and chlorinecontaining vapor, the amount of material released from these fuels scales with the concentration of the material in the fuel within the precision of the measurement. The values

indicated for HCl and K are near the detection limits of MBMS in all cases and their trends with respect to fuel type should not be overinterpreted.

The implication of these MBMS results for ash deposition (and corrosion) is that

Fireside Considerations when Cofiring Biomass with Coal in PC Boilers

251

alkali vaporization will play a fairly minor role during the combustion of switchgrass and Pittsburgh #8 coal, which is confirmed by experiments in both the MFC and the CERF. Ash deposits are collected in the MFC by inserting an air-cooled stainless steel probe into the test section. The outside diameter of the sample probe is 1.6cm, to match the Stokes number of tubes used in heat exchangers found in commercial power plants. Two embedded type-K thermocouples monitor probe surface temperature, which is controlled by varying the cooling air flowrate. A probe temperature of 500 °C was selected based on typical operating temperatures of the convective pass in power generation systems. Deposits are collected for a 1-hour period. After each experiment, the deposition probe

is removed from the combustor, the deposits are weighed and photographed, and a removable section of the probe with an undisturbed sample of the deposit is potted in epoxy for SEM analysis. To characterize the ash deposition rate, we calculate the particle capture efficiency, defined as the ratio of the mass of ash deposited on the probe to the mass of ash in the swept area of the probe. The mass of ash in the swept area of the probe depends on crosssectional area of the combustor occupied by the probe and the ash flow. At constant firing rate, the particle collection efficiency is effectively the ash deposition rate normalized the ash content of the fuel and the size of the probe.

Measured particle capture efficiencies indicate that Pittsburgh #8 fuel collects with a much higher particle capture efficiency than switchgrass. Considering the high potassium content of the switchgrass sample (Table 1), this is a potentially surprising result. Alkali metals such as potassium generally cause fouling problems. However, the switchgrass sample examined in this study has a very low fuel chlorine content (Table 1) which, as shown by the MBMS results, significantly inhibits the vaporization of alklai. Similar experiments indicating more significant levels of alkali establish the capacity of the equipment to measure alkali if it is present (Belle-Oudry and Dayton, 1997; Dayton and BelleOudry, 1997). In the MFC, ash deposition characteristics can be characterized for arbitrary blend ratios, ranging from pure biomass to pure coal. An instructive analysis is to compare the observed characteristics of the blends with an anticipated behavior, the latter being a linear interpolation of the measured characteristics of the two pure fuels. Figure 2 presents such a comparison for a variety of fuels tested in the MFC (Robinson, et al., 1997b), including the switchgrass/P8 blend. Each point in the figure requires three measurements:

252

A. L. Robinson et al.

the behavior for the two pure fuels and the blend. When plotted on a parity diagram, points above the parity line indicate the interpolated values are too high and points below indicate it is too low. As illustrated, measured capture efficiencies are commonly less than the interpolated values, indicating that ash accumulation on surfaces proceeds at rates

slower than anticipated based on the behavior of pure fuels. This behavior we attribute

to the combination of alkali from the biomass and sulfur from the coal forming solid alkali sulfates on surfaces. These sulfates are less prone to capture particles than molten salts such as chlorides and hydroxides. However, there is insufficient sulfur in many biomass fuels to completely convert the alkalis to sulfates. The effects of cofiring SG and P8 on heat transfer rate and deposit removability are illustrated by data from the CERF (Table 2). These data indicate the changes in heat transfer rate in convection pass fouling probes and simulated slag panels/waterwalls as measured by the percent of the clean-surface value. Results are reported based on a 4hour sootblower cycle. The deposit cleanability is measured on a relative scale of 0–3, based on the peak impact pressures required to remove the deposits using an incremental sootblowing procedure. This cleanability indicates the ease with which deposits are removed from the surface, with cleanabilities of 2 or less regarded as manageable by conventional sootblowing. The data indicate that the switchgrass blend reduces heat transfer rates slightly more and requires slightly more sootblower pressure to clean from the surface than deposits formed from Pittsburgh #8 coal alone. The heat transer rate during the P8/SG blend tests always returns to its clean surface value after sootblowing at pressures regarded typical of coal-fired boilers. Also, the differences in heat transfer rate are not very large and in all cases the cleanability is within established limits of manageability. The data from the MFC generally indicate that there is a potential for severe ash deposition problems when cofiring some biomass fuels, especially herbaceous fuels, with coal. However, this particular sample of switchgrass exhibited only minor ash-related problematic behavior in both the CERF or the MFC. These conclusions are all based on relatively short deposition tests at modest deposit temperatures. Reactions that lead to sintered and unmanageable deposits include reactions of alkali with silica to form alkali silicates (Baxter et al., 1996). There is ample alkali and silica in these fuels for such reactions to occur. Such reactions are generally slow, perhaps too slow to significantly affect the measured results over the time of the tests. Also, increases in deposit temperature would increase the reaction rate significantly

Fireside Considerations when Cofiring Biomass with Coal in PC Boilers

253

and could lead to considerably different results. Finally, many agricultural practices could lead to significantly higher concentrations of potassium and chlorine in the fuel, resulting in significantly different behavior.

emission is a complicated issue made potentially more complicated when cofiring biomass and coal. The MFC has a capability for measuring while maintaining gas, wall, and flame temperatures essentially constant. The results reported here indicate concentrations in the absence of any reduction technology such as fuel/air staging, reburn, lowburners, etc. Experiments over a wide range of oxidizing stoichiometric ratios conducted in the MFC indicate that concentrations at typical coal-combustion conditions (3% in flue gas) for pure biomass combustion are about 35% lower for the blend than for the pure coal (Fig. 3). At lower exit oxygen concentrations, emissions from biomass are lower still relative to coal, with a maximum of about 45% less from biomass at exit concentrations of 0.5%. However, at higher levels (6% and above), concentrations produced by pure biomass exceed those of the coal. This behavior is believed to be related to combustion of the residual biomass char, which requires higher oxygen content or longer residence time than that of coal due to its size. Qualitatively, the produced during combustion of the blend approximates a linear combination of the measured values for the blend behavior. Quantitatively, an interpolation of the behavior of the pure fuels generally overpredicts the observed behavior by –2 to 10 percent (average of 5 percent, see right ordinate of Fig. 3). Therefore, the blend produces slightly less in these experiments than the interpolation suggests. The absolute concentrations and its dependence on local stoichiometry, temperature, and extent of char combustion is under further investigation. These results are similar to our previous results (interpolation = measured value ±

254

A. L. Robinson et al.

10%) (Robinson, Junker, and Baxter, 1997a; Robinson, et al., 1997b) and those from commercial experience (Aerts and Raglunc, 1997), all of which suggest that cofiring coal with low nitrogen-content biomass can significantly reduce emission under pc conditions. The relatively high production from SG at exit oxygen contents greater than 6%

despite SG’s low fuel nitrogen content compared to P8 illustrates that conversion efficiencies of fuel nitrogen to strongly depend on operating conditions and fuel type. Past experience with both coal and biomass suggests that concentrations are only modestly correlated with fuel N content. Several observations should be considered before extrapolating these results to new conditions. Highly fertilized plants commonly have very high fuel nitrogen contents and can produce concentrations higher than those produced from coal, even under pc conditions, as is observed in MFC results using high-nitrogen biomass. Cofiring biomass with coal can cause flame instabilities, change flame stoichiometry, and produce other

changes that may significantly alter

emissions. Such behavior was observed in the

CERF, where detached flames generated high concentrations. Also, few commercial coal-fired facilities are capable of firing pure biomass for a point of reference. For these reasons, emissions from coal-biomass blends are often difficult to anticipate.

Burnout Biomass is more difficult to comminute than coal and, consequently, is typically fired at large sizes compared to pulverized coal. Furthermore, biomass fuel commonly exhibits high aspect ratios compared to the commonly equant coal particles. The consequences of these size and shape considerations are that biomass fuels most commonly oxidize under diffusion control (because of their size), often at rates that scale with minimum size to the first power (because of the aspect ratios). We have attempted to identify the maximum size biomass particle that can be used without excessive loss of

fuel or increase in residual ash carbon content of either bottom or fly ash. These results depend on fuel moisture content, particle properties (size, density, shape, volatile yield, ash content, etc.), boiler design, and boiler operating conditions. Figure 4 illustrates typical modeled results of particle burning times under the

assumptions of a boiler that measures 40m (131 ft) from bottom to furnace exit, with biomass injected at a height of 9.1 m (30ft) from the bottom. The boiler is assumed to

be operating at full load, with an average furnace gas velocity of about 20m/s (65ft/s).

Fireside Considerations when Cofiring Biomass with Coal in PC Boilers

255

The biomass is assumed to yield 80% of its mass through devolatilization. The cylindri-

cally shaped char is assumed to have a specific gravity of 0.2 and the ash of 2.5. The fuel is assumed to contain 7% ash on a dry basis. Moisture vaporization and char reaction are assumed to be heat transfer and mass transfer limited, respectively. Devolatilization is assumed to occur rapidly relative to both moisture vaporization and char oxidation. All of these assumptions and values are reasonable, but they can vary depending on the fuel, boiler design, and operating conditions. Figure 4 illustrates both the available residence time and combustion time for complete burnout for particles as a function of size. Particle sizes at which the combustion time is less than the residence time are predicted to completely burn. Particles with initial diameters less than about 5mm are predicted to exit the top of the furnace with residence times increasing with size from 1.5s at small sizes. A particle of about 5mm in diameter has a terminal velocity approximately equal to the assumed boiler gas velocity under these conditions. At larger sizes, the particles flow opposite the direction of the gases and exit the bottom of the boiler. Particles larger than about 5.5mm are predicted to require longer to burn than is available in the boiler and are predicted to exit the boiler through the bottom ash. The prediction also indicates that particles between about 1.5 and 3.5 mm are also predicted to leave the top of the boiler with a small amount of residual carbon. The linear relationship between particle size and burnout time indicated in Fig. 4 assumes a cylindrical particle shape and is different than coal behavior, the latter typically exhibiting a burnout time proportional to initial sphere diameter squared when burning under diffusion control. Experiments in Sandia’s CPI facility indicate that the linear relationship between burnout time and size is valid. A statistical analysis indicates that the assumed linear relationship is a significantly better model than a or relationship, but only marginally better. Several practical conclusions can be drawn from Fig. 4. First, the overall residence time of biomass fuel particles can be much longer than that of pulverized coal because of the non-negligible terminal velocities of such particles. The pulverized coal residence times in the boiler modeled for Fig. 4 is about 1.5s. Since many pilot- and research-scale facilities are down fired, the biomass particles have shorter rather than longer residence times, and erroneous conclusions with respect to burnout could easily be drawn based on experiments in such reactors if proper account for this effect is not made. Second,

biomass particles larger than about 1/4 inch in their minor dimension are likely to be found as partly burned char in the furnace bottom ash. While this conclusion strictly applies only to the simulation summarized in Figure 4, most perturbations on the simulation in terms of changes in operating conditions or fuel properties would decrease or have little effect on the top size at which biomass can be burned. Some smaller particles are also predicted to be found incompletely burned in the fly ash, that is, those with initial diameters ranging from about 1.5 to 2.8mm. Third, biomass should not be injected in or under the bottom row of burners, as most of it would relatively rapidly arrive in the bottom ash unless it is extremely small. Similarly, it should not be injected very high on the furnace wall or large fractions of it will enter the convection pass incompletely burned out.

Chlorine and Corrosion One proposed mechanism for chlorine-related corrosion in boilers suggests that chlorine plays an important role in concentrating alkali-metal containing salts on heat transfer surfaces. The conceptual mechanisms involve the combination of chlorine with

256

A. L. Robinson et al.

alkali to form alkali chlorides. Alkali chlorides are the most volatile forms of alkali and

the most stable form of chlorine under a wide range of oxidizing conditions relevant to combustion. Condensation of these alkali vapors selectively deposits both chlorine and alkali on the surface. At surface temperatures, the most stable form of alkali is sulfates. The chlorides react to form sulfates, releasing chlorine-containing gases (typically HCl under moist, oxidizing conditions). However, if there is insufficient sulfur to completely react with the chlorides, the chlorides remain on the surface, potentially leading to rapid corrosion (Baxter and Nielsen, 1997). This scenario suggests that there should be a relationship between deposit chlorine content and the potential for forming sulfates from alkali chlorides. Experimental indication of this relationship is indicated in Fig. 5, where deposit chlorine concentration is plotted versus the ratio of fuel sulfur to twice the maximum fuel alkali content for a variety of fuels tested in the MFC (Baxter, Robinson, Buckley, Shaddix, Lunden, & Hardesty, 1997b). This scaling parameter is chosen because, at values greater than unity, there is sufficient sulfur to convert all of the alkali chloride to sulfate. As is seen, the chlorine content of the ash deposit is very low when this scaling parameter exceeds unity, but increases to high values when the parameter is less than unity. The

highest two chlorine concentrations in the figure represent deposits formed from highchlorine biomass, but the next highest represent blends of coals with high-chlorine biomass. These data illustrate how the combination of biomass and coal properties can influence the potential for corrosion in boilers. The high-sulfur, low-chlorine, and moderate-alkali contents of the SG/P8 blend investigated here place it in the low-depositchlorine region of the graph. The absence of chlorine in the deposits does not necessarily imply that corrosion is not an issue, as it still leaves corrosion by sulfates as a potential problem. In particular, alkali tri-sulfates are known to be aggressive on heat transfer surfaces. The potential for formation of sulfates on surfaces is generally greater for biomass-coal blends than for pure coal because of the often high alkali content of the biomass. Corrosion of surfaces by alkali sulfates can usually be managed by maintaining low surface temperatures, preventing occurrences of locally reducing conditions, and proper sootblowing.

Fireside Considerations when Cofiring Biomass with Coal in PC Boilers

257

CONCLUSIONS Biomass-coal cofiring is an important future technology for power generation that introduces complexities into the fireside behaviors of such fuels. The potential consequences for ash deposition, formation, carbon burnout, and corrosion are all sensitive to biomass and coal fuel properties. The overall result is that judicious combinations of fuels for cofiring and boiler operation can avoid major problems and may reduce pollutant emissions from pc boilers. However, each of these issues has the potential of developing into a significant problem if they are not specifically addressed in the

cofiring plan. The switchgrass and Pittsburgh #8 coals investigated here produced an ash accumulation rate greater for the coal than for the switchgrass in both absolute terms and when normalized by the amount of total ash. Ash deposit formation is higher for the blend than for switchgrass alone, but not as much higher as is predicted by interpolating between the behaviors of pure switchgrass and pure Pittsburg #8 coal. This is consistent with previous observations for a variety of biomass-coal blends. We postulate that sulfur from the coal and the alkali from the switchgrass combine to form sulfates on heat transfer surfaces to which particles are less prone to adhere than chlorides or hydroxides. Deposits from the switchgrass-Pittsburgh #8 coal blend form a higher barriers to heat transfer and are somewhat more difficult to remove than those from the coal alone, but they remained in a manageable range within the duration of these tests. The inorganic composition of the fuels suggests that longer-term behavior and behavior at higher temperatures may result in more severe ash management problems. Many proposed energy crop proposals involve aggressive agricultural practices with regard to fertilization and harvesting and may have significantly higher chlorine and alkali contents than our sample of switchgrass. This, too, would lead to less manageable and more rapidly accumulating deposits. Qualitatively, formation for blends interpolates between that of the pure biomass and coal if combustion conditions are otherwise unchanged. Quantitatively, the interpolated values are about 5% higher than the measured values. The trends of formation with overall stoichiometry are fairly complex, with switchgrass producing lower than coal at about 6% oxygen content in the gases or less but higher at higher exit concentrations. In most cases of practical interest, it may not be possible to keep combustion conditions constant as biomass is introduced and establishing the amount of produced from pure biomass combustion may not be possible. In general, biomass fuels with low nitrogen contents have been observed to decrease emissions significantly under most pulverized-coal-relevant conditions. Most biomass fuel particles are large and have high aspect ratios compared to pulverized coal particles. Burnout time for such particles scales approximately with the smallest dimension. Under typical, full-load, utility-boiler conditions, particles as large as 5.5mm are predicted to burnout before exiting the furnace. This result depends on

operating conditions, point of injection, boiler design, and fuel properties. Larger particles are predicted to exit the furnace bottom while smaller particles exit with the fly ash. The size at which the transition is made is 5mm in the particular simulation here and also depends on operating conditions, etc. Deposit chlorine content is anticipated to be minimal in many biomass-coal cofiring scenarios due largely to the relatively low fuel chlorine content and the high overall sulfur content relative to alkali. This reflects the tendency of alkali chlorides to react to

form sulfates if there is available sulfur. The absence of chlorine in deposits indicates that

258

A. L. Robinson et al.

corrosion may not be extremely severe. However, sulfates, in particular alkali tri-sulfates, can be very corrosive and biomass coal combustion will enhance the formation of sulfates on heat transfer surfaces. Highly fertilized or aggressively managed energy crops may have greater potential for increased corrosion due to increased alkali and chlorine contents. All of these results find support in commercial-scale data but there are not sufficient long-term commercial-scale tests to provide definitive evidence of many of these trends. These laboratory results will be used to help design commercial-scale tests and indicate sampling and analysis needed in future tests to provide better supporting information or indicate weaknesses in our current understanding.

ACKNOWLEDGMENTS Funding for this project was provided by DOE's Office of Fossil Energy through the Advanced Research and Technology Development (FETC & Sandia) and Office of Energy Efficiency and Renewable Energy through the Biomass Power Program (NREL and Sandia).

REFERENCES Aerts, D., and Raglunc, K. Co-firing Switchgrass and Coal in a 50MW Pulverized Coal Utility Boiler (Final Report No. University of Wisconsin-Madison. (1997). Baxter, L. L., Combustion and Flame, 90, 174–184. (1992). Baxter, L. L., Biomass and Bioenergy, 4(2), 85–102. (1993). Baxter, L. L., Miles, T. R., Jr., T. R. M., Jenkins, B. M., Milne, T., Dayton, D., Bryers, R. W., and Oden. L. L. The Behavior of Inorganic Material in Biomass-Fired Power Boilers: An Overview of the Alkali

Deposits Project:. In A. V. Bridgwater & D. G. B. Boocock (Eds.), Developments in Thermocheimcal Biomass Conversion (pp. 1424–1444). London: Blackie Academic & Professional. (1997a).

Baxter, L. L., Miles, T. R., Thomas R. Miles, J., Jenkins, B. M., Dayton, D., Milne, T., Bryers, R. W., and Oden, L. L. The Behavior or Inorganic Material in Biomass-Fired Power Boilers—Field and Laboratory Experiences: Volume II of Alkali Deposits Found in Biomass Power Plants No. SAND96-8225, NREL/TP433-8142). Sandia National Laboratories; National Renewable Energy Laboratory. (1996). Baxter, L. L., and Nielsen, H. P. The Effects of Fuel-bound Chlorine and Alkali on Corrosion Initiation. In 214th American Chemical Society National Meeting, (pp. 1089–1095). Las Vegas, NV: (1997). Baxter, L. L., Robinson, A. L., Buckley, S. G., Shaddix, C. R., Lunden, M., and Hardesty, D. R. Task 4. Development of Guidelines for Cofiring Biomass and Pulverized Coal (Quarterly Progress Report No. Sandia National Laboratories. (1997b).

Belle-Oudry, D. A., and Dayton, D. C. Analysis of Combustion Products from the Cofiring of Coal with Biomass Fuels. In 214th American Chemical Society National Meeting, (pp. 10-96-1100). Las Vegas, NV: (1997).

Dayton, D. C., and Belle-Oudry, D. A. Bench-Scale Biomass/Coal Cofiring Studies. In Engineering Foundation Conference on the Impact of Mineral Impurities on Solid Fuel Combustion, Kona, HI: (1997).

Dayton, D. C., French, R. J., and Milne, T. A., Energy and Fuels, ( to appear ). (1995). Freeman, M. C., Chitester, D. C., James, R. A., Ekmann, J. M., and Walbert, G. F. Results of Pilot-Scale Biomass Co-Firing for PC. Combustors. In USDOEIFETC Advanced Coal-Based Power Systems and Environmental Control ’97 Conference, Pittsburgh, PA: (1997). H u r t , R. H., and Davis, K. A. Near-Extinction and Final Burnout in Coal Combustion. In 25th Symposium (International) on Combustion Pittsburgh, PA: The Combustion Institute. (1994).

Robinson, A., Junker, H ., and Baxter, L. Pollutant Formation, Ash Deposition, and Fly Ash Properties When Cofiring Biomass and Coal. In Engineering Foundation Conference on the Economic and Environmental Aspects of Coal Utilization, Santa Barbara, CA: (1997a). Robinson, A. C., Junker, H., and Baxter, L. L. Ash Deposition and Pollutant Formation when Corfiring Biomass with Coal in PC Boilers. In EPRI Coal Quality Conference, Kansas City, MO: (1997b).

SUMMARY OF RECENT RESULTS OBTAINED FROM USING THE CONTROLLED FLUIDISED BED AGGLOMERATION METHOD Marcus Öhman and Anders Nordin Energy Technology Center, Department of Inorganic Chemistry, University of Umeå, S-901 87 Umeå, Sweden

1. INTRODUCTION The most promising energy conversion technologies for solid fuels, and biomass in particular, are based on fluidized bed combustion (FBC) or gasification (FBG). These

processes enable higher electrical and total efficiency as well as greater fuel flexibility. Due

to the relatively low temperatures in FBC and FBG, sulfur emissions and the extent of deposit formation can also be kept to a minimum. However, bed agglomeration could be a potential problem which can decrease both the heat transfer in the bed and the fluidization quality, resulting in poor conversion efficiencies and loss of control of bed operational parameters. In the most severe cases, bed agglomeration can lead to total defluidization, resulting in unscheduled plant shut down. The state-of-the-art concern-

ing methods to determine fluidized bed agglomeration tendencies were recently compiled by Öhman (1997).

1.1. Review of Previous Work It is generally agreed that agglomeration may proceed through several different sintering mechanisms [Skrifvars 1994, Manzoori 1994; Öhman 1997]; viscous flow sintering of liquid silicates; reactive liquid sintering of molten salt systems; chemical reaction sintering by formation of new compounds; solid state sintering; and vaporization preceded by re-condensation. It can further be assumed that, although they all may contribute to the initial mechanisms of the ash transformations, the melting behavior of the ash or coating material probably is decisive for the final agglomeration in the turbulent fluidized beds.

1.1.1. Prediction of Ash Melting Behavior. Several techniques are currently used in the fuel industry to predict ash fusion temperatures from elemental composition. The most Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

259

260

M. Öhman and A. Nordin

common is the use of phase diagrams to determine liquid temperatures that parallel the fusion temperatures of the formed ash [Huggins et al. 1981; Hastie and Bonell 1985]. The accuracy of this technique declines with an increase of primary elements in the ash and wood ashes, for example, contain up to seven major elements. However, phase diagrams are very useful as a first screening method for a better understanding of the general ash melting behavior. The ideal technique would be to utilize chemical equilibrium model calculations, incorporating all intermediate stoichiometric phases, non-ideal solid and liquid solutions required for the system of interest. Several investigations using equilibrium calculations have been performed during the last few years

and many of these have included non-ideal solid and liquid solutions, see Nordin et al. [1997] for a review. In addition, Blander et al. [1997] recently performed a round robin of different programs and databases. The melting behavior can presently be accurately predicted for the salt system the most important binary and some ternary silicate systems but several ternary and higher ordersilicate systems still remain to be evaluated/optimized. Further, data for the interactions between the salt system and and silicates are needed for a general equilibrium model of the melting behaviors. New evaluations of solution data are continuously made available in the literature, and through for example the FACT database.

There are also many empirical slagging and fouling indexes proposed, that are based on ash composition [Nichols and Reid 1940; Sage and Mcllroy 1960; Reid and Cohen 1944]. However, these indexes have been shown to be of very little general value for different fuel types. Empirical models using multiple regression techniques have also been used as predictors of ash fusion temperature [Wintergarten and Rhodes 1975; Sonderal and Ellman 1975; Vorres 1979]. Where successful, the empirical models have been limited to coals from a certain field.

1.1.2. Available Laboratory Methods. The standard ash fusion tests (AFT; ISO 5401981; BS 1016.15-1970; GOST 2057-1982; SABS 932; ASTM D1857, 1987; DIN 517 30, 1984; AS 1038.15, 1987; GE 219-74) are the most commonly used laboratory methods and also the only standardized technique to predict the behaviors of ashes in different processes. These tests are based on following the external shape (deformation, shrinkage and flow) of a pyramidal or cylindrical pellet of ash during heating in a laboratory furnace in either an oxidizing or a reducing atmosphere. Of particular concern is the estimation of the initial deformation temperature (IDT). The standard AFT have been extensively criticized in the literature. One general criticism is the relevance of the ash sample which is subjected to the test. The ashing temperature used to generate the ash sample, is much lower and the history and atmosphere of the ash is quite different than that experienced in combustion situations. Another criticism is that changes in the shape of the sample during heating, due to phase transitions and chemical reaction, can also be interpreted as initial melting. Further, strict control and observance of the test conditions are necessary to obtain reproducible results. Both repeatability and reproducibility have recently been shown to be poor [Wall et al. 1995; Coin et al. 1995]. In one investigation the reproducibility has been demonstrated to be as poor as 140°C [Slegeir and Singletary 1988]. Gerald et al. [1981] have revealed that significant melting of the ashes occurs far below (200–400 K) the IDT. In addition, Huffman & Huggins [1983] and Huggins et al. [1981] showed that most ashes are completely melted at temperatures far below the IDT and that the progression from the IDT to the fluid temperature can be accomplished by holding the IDT constant for a period

Summary of Recent Results Obtained from Using the Controlled Fluidised Bed Agglomeration Method

261

of 45 minutes. They suggested that this behavior was due to decreasing of the slag viscosity rather than increasing the amount of melt. Compression strength tests of heat-treated cylindrical pellets have been used by several authors, to study the ash melting behaviors of different types of coals [Barnhart and Williams 1956; Conn and Jones 1984; Nowok et al. 1990] and biomass fuels [Skrifvars et al. 1995]. The sample, which can be ashes collected from industrial plants, laboratory ashes, or a mixture of chemical compounds, is first crushed and screened. Then the sample is pelletized to a cylindrical pellet and heated in a controlled gas atmosphere. After cooling, the pellets are crushed in standard compression-testing equipment, and the crushing strength is taken as a measure of the sintering degree. The method has proven to give results with a relatively high reproducibility [Skrifvars 1994], but no thorough evaluation of the accuracy concerning bed agglomeration seemed to be available. Smith [1956] used a dilatometric shrinkage technique to study sintering characteristics of pulverized fuel ash. This technique is based on shrinkage measurements of an ash sample, i.e. if an ash sample shrinks, it indicates sintering. An intercept on the temperature axis is taken to define the initial sinter point. Similar techniques have been used by several authors [Raask 1979; Smith 1956; Manzoori 1990; Coin et al. 1995; Wall et al. 1995] and the method was shown to be superior to the ash fusion tests. However, with some coal ashes, results have been obtained where the shrinkage measurements showed no change although a significant degree of sintering had taken place [Raask 1985]. Raask therefore suggested that the dilatometry method should be supplemented with continuous measurements of the conductance. Before sintering, an ash would show low thermal- and electrical conductance because of the lack of particle to particle contact. As the cross sectional area of sinter bonds between the particles grows, the conductance path (both thermal and electrical) is increased. Simultaneous measurements of thermal conductance and dilatometric shrinkage for detection of the onset of sintering of coal ashes have been used by several authors [Cumming et al. 1985; Conn and Austin 1984]. A similar method based on the electrical conductance of ash has also been proposed [Sanyal and Metha 1993; Sanyal and Cumming 1981; Gibson and Livingston 1991]. A disadvantage of these methods is that satisfactory contact between the ash and the electrodes is hard to achieve and maintain [Wall et al. 1989]. An interesting complement would be the use of capacitive measurements for potential information about initial melt formation, as recently proposed by Nordin and Leven [1997]. To determine the effect of temperature on ash agglomeration, a simple method was developed and used by Stallmann and Neavel [1980], Samples of coal less than 100 mesh were ashed at low temperature. The ash produced was screened to ensure that its top size remained below 100 mesh. These samples were then heated at different temperatures in a platinum crucible. After cooling, the samples were screened to determine the weight percent retained on a 100 mesh screen. The effect of the straw ash characteristics on agglomeration of silica sand [Ghaly et al. 1994] and alumina sand [Ghaly et al. 1993] at various temperatures was investigated using a high temperature muffle furnace. SEM/EDS was then used for the analysis. Padban et al. [1995] used a similar approach to study the effect of two biomass ashes on agglomeration of silica sand.

The viscosity is a critical property when assessing the sintering, slagging and fouling tendencies of different inorganic materials including fuel ash. Previous experimental work have shown that there is a critical viscosity for adhesion of ash particles [Srinivasachar et al. 1988; Wibberly and Wall 1982; Srinivasachar et al. 1990]. At temperatures and velocities typically found in coal fired boilers, this critical viscosity has been shown to be

262

between

M. Öhman and A. Nordin

and

[Senior and Srinivasachar 1995]. The measurable viscosity of

glassy materials can extend over a wide range, more than 10 orders of magnitude [Raask 1985]. There is no single method capable of measuring the viscosity of fuel ash slags over the entire range. Most laboratory viscosity measurements at high temperatures are carried out using a rotating crucible viscometer. The viscosity of less fluid slags in the range of to can be measured by a rod penetration viscometer. Several empirical methods have also been proposed for the estimation of the viscosity’s for coal ashes [Nowok 1994; Urbain and Boiret 1990). In a thermogravimetric analysis (TGA), the mass of a sample is continuously recorded as a function of temperature or time. The information provided by thermogravimetric methods is limited because a temperature variation must bring about a change in mass of the analyte. Thus, thermogravimetric methods are mainly limited to decomposition and oxidation reactions and to such physical processes as vaporization, sublimation, and desorption. Differential thermal analysis (DTA) is a technique in which the difference in temperature between a substance and a reference material is measured as a function of temperature while the substance and reference material are subjected to a controlled temperature program. Usually, the temperature program involves heating the

sample and reference material in such a way that the temperature of the sample increases linearly with time. The difference between the sample temperature and the reference temperature is then monitored and plotted versus sample temperature to give a differential

thermogram. DTA has been used to study thermal behavior of inorganic materials such as silicates, ferrites, clays, oxides, ceramics, catalysts and glasses. Information about different processes such as fusion, dehydration, oxidation, reduction, adsorption and solid state reactions is provided. DTA can preferably be used to accurately obtain eutectic,

solidus and liquidus temperatures in different systems [c.f. Vassilev et al. 1995]. High temperature microscopy (HTM) may be used in combination with DTA to obtain the liquidus temperature of a fuel ash. In this method the solidification of a sample is followed from melt to solid phase. Hereby, the initial crystallization temperature may be obtained with a high accuracy. 1.1.3. Previous Bench Scale Agglomeration Studies. Several detailed agglomeration studies on coal in bench scale have been conducted, under both combustion and gasification conditions. The effect of temperature on the extent of agglomeration has been studied by Sandstrom et al. [1979] and Goblirsch et al. [1983]. An increase in temperature was shown to increase both the agglomeration tendencies and the size of the agglomerates. The different effects of bed materials and additives on agglomeration tendencies was also studied. The effect of limestone addition on agglomeration, was determined

both in combustion [Benson et al. 1982; Dawson and Brown 1992] and gasification [Kline et al. 1990; West et al. 1993]. Depending on the ash composition in the fuel, the limestone addition could both increase and decrease the bed agglomeration tendencies. In addition, experiments with gabbro, alumina sand and dolomite as bed material have been performed in combustion [Goblirsh et al. 1980]. These experiments showed that gabbro was the most effective additive regarding bed agglomeration. Parameters such as fluidization velocity [Atakul and Ekinci 1989; Basu and Sarkar 1983; Sandstrom et al. 1979], bed particle size, bed height [Atakul and Ekinci 1989], bed particle surface and contact area [Siegell 1976; Huang 1985] have also affected the agglomeration/defluidization tendencies. In most of the above bench scale studies, at least one unknown variable was introduced by the burning particles in the bed. This could lead to a incomplete mapping and poor understanding of the relations between the studied process parameters and the

Summary of Recent Results Obtained from Using the Controlled Fluidised Bed Agglomeration Method

263

agglomeration tendencies. In addition, the operational parameters in many of these studies have been chosen outside of practical ranges in full scale plants. Only a few studies on agglomeration in biomass combustion [Salour et al. 1993] and gasification [Ergundler and Ghaly 1993; Bruce and Bitoft 1988, Soltes et al. 1982; Le Pori et al. 1980] have been performed. In contrast to the results obtained from studies with coal, Ergudenler and Ghaly [1993] showed that the operational velocity had no effect on the agglomeration tendency when wheat was used as fuel. In a biomass study [Bruce and Bitoft 1988], none of the studied operational parameters had an effect on the bed agglomeration. In addition, results have shown that the agglomeration tendencies could be greatly affected by fuel mixing. By blending wood into rice straw, Salour et al. [1993] showed that the bed agglomeration could be controlled and prevented. The conclusion form reviewing the literature was that in spite of a relatively frequent reporting, a precise and quantitative knowledge of bed agglomeration processes has not yet been presented. In addition, no reliable and realistic methods were found to be available to determine bed agglomeration tendencies of different fuels, fuel combinations or fuels with additives. As none of the existing laboratory methods comprises the ash transformations in a fluidized bed, a more relevant method would be to use a fluidized bed for actual and controlled bed agglomeration studies. Several attempts were made prior to 1993 to use bench or pilot scale fluidized bed reactors for this purpose but they all suffered from either bed temperature inhomogeneity, caused by the burning particles in the bed, or inadequate evaluation and documentation. Although the publications by Gluckman et al. [1976], West et al. [1993] and Basu and Sarkar [1993] reported commendable work in the right direction, no bench scale studies had previously been performed in such way that the fuel specific bed agglomeration temperature could be determined in a realistically and accurately optimal way. Since the presentation of the controlled fluidized bed agglomeration (CFBA) method [Nordin et al. 1995] at the last EF ash conference, our continued work have resulted in a licentiate thesis [Öhman 1997] including four papers, two subsequent submitted papers and some still not published results, all so far with limited availability. The objectives of the present report was therefore to review all these CFBA studies performed and present our gathered conclusions.

2. THE CONTROLLED FLUIDIZED BED AGGLOMERATION METHOD

2.1. Summary of the CFBA Results Obtained 2.1.1. Particle Temperature Studies in Fluidized Bed Combustion. As discussed in previous sections, the most important individual parameter influencing the bed agglomeration is the actual process temperature. It is well known that the temperature of the burning particles significantly exceeds the bed temperature. It is also clear that the increased temperature of the burning particles is crucial for the ash transformations and bed agglomeration process in fluidized beds. The effects of different bed- and particle variables on the particle temperature have been extensively studied during the last 30 years. However, to our knowledge, no simultaneous evaluation of all the studied variables was previously performed. The objectives of the work presented by Öhman and Nordin [1996] were therefore; to review the literature of existing results from particle temperature studies in

fluidized beds; and based on literature data to determine an empirical PLS (partial least

264

M. Öhman and A. Nordin

squares projections to latent structures) model describing the effects of the different variables on the particle temperature.

The results from the literature survey and from the determined PLS-model, showed that; i) the difference between the burning particle temperature and the bed temperature may be considerable (40-600 K); ii) many different fuel and process variables have a large influence on the temperature difference; and iii) the total oxygen concentration, bed temperature, fluidization velocity and bed particle diameter are the most influential variables.

Thus, it was clear that a modified bench scale method had to be developed without the burning particles, if a fuel specific agglomeration temperature was to be accurately determined. 2.1.2. The Controlled Fluidized Bed Agglomeration Method. A bench scale fluidized bed combustor was therefore constructed, enabling realistic and highly controlled bed agglomeration tests with a homogeneous bed temperature (Fig. 1). The bench scale reactor is made of stainless steel (SS 2,343), being 2m high, 100mm and 200mm in diameter in the bed and freeboard sections, respectively. A perforated stainless steel distributor plate with 1% open area and a total of 90 holes is used. The maximum temperature of the equipment is 1,020°C. To allow for a constant increase of the bed temperature, with a homogeneous temperature profile, much effort was focused on the

bed section of the reactor as well as the controlling system. To keep the walls at the same temperature as the bed, the reactor is equipped with controlled electrical wall heating elements. An air-pre-heater allowing primary air temperatures up to 1,050°C was also constructed to control the bed temperature. Forced convection is utilized in a cyclonelike stainless steel cylinder equipped with Kanthal electrical wall heating elements. All temperatures are manipulated by Eurotherm temperature controllers and the maximum temperature deviation within the bed was determined to be less than The reactor is utilized in one of two operational modes. During the first normal combustion operational mode, a relatively accurate simulation of a full-scale process is accomplished, and the bed is loaded with ash with appropriate characteristics. The subsequent mode of operation is based on controlled increase of the bed temperature by applying external heat to the primary air and to the bed section walls. In addition, temperature homogeneity is secured by switching from normal fuel feeding to a propane precombustor. Bed temperatures at four locations in the bed are measured by shielded type S thermocouples and differential bed pressures at four other positions are determined by differential pressure transducers. The initial agglomeration temperature is determined by on- or off-line principal component analysis of the small variations in measured bed temperatures and differential pressures, preceding the definite defluidization. Samples of ash

and bed material for evaluation of agglomeration mechanisms may be collected throughout the operation. The method is more closely described in the publication by Nordin et al. [1995], which also presents a measure for prevention of agglomeration of biomass fuels. Several repeated combustion tests with two biomass fuels alone (Lucerne and olive flesh), all resulted in agglomeration and defluidization of the bed within less than 30 minutes. By controlled agglomeration experiments, the initial cohesion temperatures for the two fuels were determined to be as low as 670 °C and 940 °C, respectively. However, by cocombustion with coal, the initial agglomeration temperatures increased to 950 °C and more than 1,050°C, respectively. When co-fired with coal during ten hour extended runs, no agglomeration was observed for either of the two fuel mixtures. Samples of bed materials, collected throughout the experimental runs, as well as the produced agglomerated

Summary of Recent Results Obtained from Using the Controlled Fluidised Bed Agglomeration Method

265

266

M. Öhman and A. Nordin

beds, were analyzed using SEM/EDS and X-ray diffraction. The results showed that loss of fluidization resulted from formation of molten phases coating the bed materials; a salt

melt in the case of Lucerne and a silicate melt in the case of the olive fuel. By fuel mixing, the in-bed ash composition is altered, conferring higher melting temperatures, and thereby agglomeration and defluidization can be prevented. 2.1.3. Quantification of Fluidized Bed Agglomeration Tendencies—Sensitivity Analysis. The objectives of the work described by Öhman and Nordin [1997] were to; i) determine the inaccuracy and reproducibility of the new method; and ii) determine potential effects of all the process related variables on the determined agglomeration temperature.

An extensive sensitivity analysis was performed according to a statistical experimental design to evaluate the effects of eight different process analytical variables on the determined agglomeration temperature of a biomass fuel. The results showed that amount of bed material, heating rate, fluidization velocity and air to fuel ratio during both ashing and heating did not influence the determined agglomeration temperature. Only ash to bed material ratio, the ashing temperature and the bed material size had significant effects on the agglomeration temperature, but still the effects were relatively small. The agglomeration temperature of the fuel could be determined to 899°C with a repeatability of

(STD). Based on the results, the inaccuracy was determined to be

(STD), considering the normal variations in all operating variables. A corresponding, still

unpublished, study using another type of fuel (wheat straw) confirmed the above results. The only significant operating parameter was found to be the ashing temperature.

To further evaluate the accuracy of the method, another study has been initiated, where “synthetic ashes” with known melting behaviors are used for comparison. So far, different compositions (50–99% of the binary system was prepared by heat treatment and crushing and then fed to the bed. The rich part of the system has an eutectic temperature of 690°C [Bergman et al. 1995] and the liquidus temperature increases almost linearly up to 1,076°C for the pure The results from

the CFBA method although used at somewhat lower temperatures than normal, were found to be in good agreement with the more accurately determined eutectic temperatures. Further, the composition did not seem to influence the temperature of onset of agglomeration, indicating that only small amounts of melts are needed for salt systems. The continued sub-project will include also silicate systems to qualitatively determine also the effect of the higher viscosity melts.

2.1.4. Mechanisms of Bed Agglomeration in Combustion of Biomass Fuels. In a recent unpublished collaborative study with Abo Akademi University, the mechanisms of the chemical processes during bed agglomeration of ten different biomass reference fuels

[Nordin 1993] were determined by extensive SEM/EDS analysis of bed material sampled continuously throughout the agglomeration process as described above. The biomass fuels were chosen by principal component analysis (PCA) to represent all biomass fuels, with respect to variations in content of ash forming elements. The results from the study show that the agglomeration mechanism of most types of biomass fuels are governed by the formation and stickiness of high Ca- and K-silicate melts (sintering by viscous flow) but also high elemental inhomogeneity, even within ash from specific fuels. S and Cl was in general found not to participate in the final agglomeration mechanism. This was further supported by the careful study of ash formation during waste sludge and bark incineration performed by Latva-Somppi et al. [1997]. The significant vaporization and transport of K, Na, Cl and S was determined by impactor sampling and elemental analy-

Summary of Recent Results Obtained from Using the Controlled Fluidised Bed Agglomeration Method

267

sis both for full scale FBC and corresponding runs in the bench scale reactor. The vaporization in the bench scale unit was also found to increase linearly by increased bed temperature. In addition, no vaporization was found during external heating phase of CFBA runs. 2.1.5. Predicting Agglomeration Tendencies—A Comparison of Three Techniques. A comparison between three different techniques to predict bed agglomeration tendencies during FBC was performed by Skrifvars et al. [1997]. The standard ASTM ash fusion test, a compression strength based sintering test and actual controlled bed agglomeration in a bench scale FBC reactor were used to determine the critical temperature of ten biomass reference fuels (Bark, Bagasse, Cane Trash, Lucerne, Olive Flesh, Peat, RDF, Reed Canary Grass, Wheat Straw, Wood Residues). The fuels, once again, were chosen

based on a PCA of a compilation of about 300 different samples of Nordic biomass fuels. The results from the comparison showed significant differences in the critical temperatures obtained, depending on which technique was used. The limited applicability of the ASTM standard ash fusion test was clearly illustrated as the resulting initial deformation temperatures were found to be 100–900 °C higher than those from the compression strength and bench scale FBC tests. The general tendency, as determined by the compression strength tests, seemed to agree with the results from the actual agglomeration tests, although a large spread was obtained. For five of the fuels, the compression strength test gave sintering temperatures some 20–40 °C lower than those determined as the bed agglomeration temperature with the bench scale FBC reactor. For Reed Canary grass, a significantly lower sintering temperature was obtained and for three fuels significantly higher values were obtained. The results indicated that the reason for this disagreement could be that the critical temperature previously has been taken at the onset of sintering, i.e. at the intercept between the zero pressure line and the line of increased sintering strength. If a higher specific strength is used instead, much better agreement between the two methods was obtained. In addition, Natarajan et al. [1997] used the CFBA method to determine the initial agglomeration temperature, during FBC and FBG in both SiO2 and CaO, for rice husk, bagasse, cane trash and olive flesh. The results were further compared with the results from corresponding ASTM fusion tests. The results showed a significantly higher initial deformation temperature than the actual agglomeration temperature for all fuels considered. For rice husk, the discrepancy was found to be as high as 600°C, due to its high Si rigid cage-like ash structure. The use of lime instead of quartz increased the agglomeration temperature somewhat for combustion conditions. Cane trash and olive flesh resulted in agglomeration of the quartz bed at somewhat lower temperatures during FBG than FBC, while a small increase in agglomeration temperature was obtained for rice husk, going from FBC to FBG.

3. CONCLUSION AND FUTURE WORK The conclusions from the different studies are:

• the new method seems to be accurate in determining fuel specific agglomeration tendencies, with high reproducibility and most importantly closely related to actual full scale behavior. • for fuel analysis, the laboratory methods may suffer significantly from the origin

M. Öhman and A. Nordin

268

of the analyzed samples, and the standard ash fusion tests have been shown to produce erroneous results due to several different reasons. • the identification of low temperature liquids as the reason for the agglomeration can be used to propose measures for prevention, either by sorption of fluxing elements such as K and Na or alteration of the ash composition conferring higher melting temperatures. This can simply be accomplished by co-combustion with coal. • the CFBA method can preferably be used both for FBC and FBG conditions.

In addition to the study of different synthetic ashes for evaluation purposes, our future work will include more of the corresponding studies in gasification environments and with other bed materials. Careful studies of the effect of potential additives for prevention of bed agglomeration, and some related fundamental studies of alkali behavior in fluidized beds have also been initiated.

4. ACKNOWLEDGEMENTS The authors thank Dr Bengt-Johan Skrifvars, Dr Rainer Backman and Professor

Mikko Hupa, Åbo Akademi University for the stimulating multi-year collaboration, as well as our shorter term visiting scientists. The financial support from the Swedish

National Board for Industrial and Technical Development (NUTEK) is gratefully acknowledged.

5. REFERENCES Atakul, H., and Ekinci, E. (1989). “Agglomeration of Turkish lignites in fluidized-bed combustion.” J. Inst. of Energy, March, 56.

Barnhart, D. H., and Williams, P. C. (1956). “The sintering test-An index to ash fouling tendency.” Trans. ASME, 78, 1229. Basu, P., and Sarkar, A. (1983). “Agglomeration of coal ash in fluidized beds.” Fuel, 62, 924. Benson, S. A., Karner, F. R., Goblirsch, G. M., and Brekke, D. W. (1982). “Bed agglomerates formed by atmospheric fluidized-bed combustion of a North Dacota Lignite.” Proc. of the 183rd Nat. ACSM, Div. Fuel Chem.,27, 174. Bergman, A. G., Kislova, A. 1., and Posypiako, V. I. (1954). “System Obshch. Khim., 24, 1722. Blander, M., Milne, T., Dayton, D., Backman, R., Blake, D., Kuhnel, V, Linak, W, Mann, M., Nordin, A., and Ljung, A. (1997). “Equilibrium chemistry of the combustion of biomass: a round robin set of calculations using available computer programs and data bases.” Proc. Eng. Found. Ash Conf. Kona Hawaii. Bruce, K., and Bitowft, B. S. (1988). “A generic study of the sintering aspects of biomass in a fluid-bed gasifier.” Energy Biomass Wastes, 11, 5 1 1 . Coin, C., Kahraman, H., and Peifenstein, A. P. (1995). “An improved ash fusion test.” Applications of Advanced Technology to Ash-Related Problems in Boilers, Ed. Baxter, L. Desollar, R.. 187–200. Conn, R. E., and Austin, L. G. (1984). “Studies of sintering of coal ash relevant to pulverized coal utility boilers.” Fuel, 63, 1664. Conn, R. E., and Jones, M. L. (1984). Eng. Found. Conf., Copper Mountain, Colorado. Cumming, I. W., Joyce, W. I., and Kyle, J. H. (1985). “Advanced techniques for the assessment of slagging and fouling propensity in pulverized coal fired power plant.” J. Inst. Energy., 58. Dawson, M. R., and Brown, R. C. (1992). “Bed material cohesion and loss of fluidization during fluidized bed combustion of midwestern coal.” Fuel, 71, 585.

Ergudenler, A., and Ghaly, A. E. (1993). “Agglomeration of silica sand in a fluidized bed gasifier operating on wheat straw.” Biomass and Bioenergy, 4, 135.

Summary of Recent Results Obtained from Using the Controlled Fluidised Bed Agglomeration Method

269

Gerald, P. H., Huggins, F. E., and Dunmyre, G. R. (1981). “Investigation of the high-temperature behavior of

coal ash in reducing and oxidising atmospheres.” Fuel, 60, 585. Ghaly, A. E., Ergüdenler, A., and Laufer, E. (1993). “Agglomeration characteristics of alumina and sand-straw ash mixtures at elevated temperatures.” Biomass and Bioenergy, 5, 467. Ghaly, A. E., Ergüdenler, A., and Laufer, E. (1994). “Study of agglomeration characteristics of silica sandstraw ash mixtures using scanning electronic microscopy and energy dispersion x-ray techniques.” Biore-

source Technology, 48, 127. Gibson, J. R., and Livingston, W. R. (1991). “The sintering and fusion of bituminous coal ashes.” Eng. Found. Conf. Inorganic Transformations and Ash Deposition During Coal Combustion, 425–447 Palm Coast Florida. Gluckman, M. J., Yerushalmi, J., and Squires, A. M. (1976). “Defluidizalion characteristics of sticky or agglomerating beds.” Fluidization Technology, 2, 395.

Goblirsch, G. M., Benson, S. A., Karner, F. R., Rindt. D. K., and Hajicek, D. R. (1983). “AFBC bed material performance with low-rank coals.” Proc. of the 12th biennial lignite symp., May 18–19, Grand Forks,

DOE/FE/60181-5. Goblirsch, G., Vander Molen, R. H., Wilson, K., and Hajicek, D. (1980). “Atmospheric Fluidized bed combustion testing of North Dakota Lignite”. Proc. of the 6th Int. Conf. Fluidized Bed Combustion, 2, 850. Hastie, J. W., and Bonell, W. (1985). “ A predictive phase equilibrium model for multi component oxide mixtures.” High Temp. Sci., 19, 275. Huang, C. H. (1985). “Fundamentals of agglomeration in a fluidized bed.” Thesis, Illinois Inst. Techn. Huffman, G. P., and Huggins, F. E. (1983). “Investigation of partial ash melting by phase analysis of quenched samples”. In Fouling and slagging resulting from impurities in combustion gases, Engineering Foundation, 259–279, New York. Huggins, F. E., Deborah, A. K., and Gerald, P. H. (1981). “Correlation between ash fusion temperatures and ternary equilibrium phase diagrams.” Fuel, 60, 577. Kline, S. D., Mason, D. M., Carty, R. H., and Babu, S. F. (1990). “The effect of limestone on ash behavior in

fluidized-bed gasification of coal”. Proc. of utilization of high sulphur coals III, Elsevier Science., 687. Latva-Somppi, J., Kauppinen, E. I., Kurkela, J.,Öhman, M., Nordin, A., and Johanson, B. (1997) “Ultrafine

ash particle formation during waste sludge incineration in fluidized bed reactors.” Proc. of AAAR97. Le Pori, W. A., Anthony, R. G., Lalk, T. R., and Craig, J. D. (1980). “Fluidized bed combustion and gasification of biomass.” Agricultural Energy, 2, 330. Manzoori, A. R. (1990). “Role of inorganic matter in agglomeration and defluidization during the circulating fluidized bed combustor.“ Thesis, University of Adelaide. Manzoori, A. R., and Agarwal, P. K. (1994). “Agglomeration and defluidization under simulated circulating fluidized bed combustion conditions.” Fuel, 73, 563. Nicholls, P., and Reid, W. T. (1940). “Viscosity of coal ash slags.” Trans. ASME 62 (1), 141. Natarajan, E., Öhman, M., Gabra, M., Nordin, A., Liliedahl, T. and Rao, A. N. (1998) “Experimental deter-

mination of bed agglomeration tendencies of some common agricultural residues in fluidized bed combustion and gasification.” Biomass and Bioenergy, 15, 163–169. Nordin, A., Dayton, D., French, R., and Milne, T. (1997). Literature review of previous work on alkali metals in combustion systems. Report to be published. Nordin, A. (1994). “Chemical elemental characteristics of biomass fuels.” Biomass and Bioenergy, 6, 339. Nordin, A, and Leve*pln, P. (1997). “Ash related problems in biomass fired boilers.” Thermal Engineering Research Foundation, Report no. 607. ( I n Swedish)

Nordin, A., Öhman, M., Skrifvars, B., J., and Hupa, M. (1995). “Agglomeration and defluidization in FBC of biomass fuels—mechanisms and measures for prevention.” Applications of Advanced Technology to AshRelated Problems in Boilers, Ed. Baxter, L. Desollar, R., 353–366. Nowok, J. W., Benson, S. A., Jones, M. L., and Kalmanovitch, D. P (1990). Fuel, 69, 1020. Padban, N., K i u r u , S., and Hallgren, A. L. (1995). “Bed material agglomeration in PFB biomass gasification.” ACSM., Div. Fuel. Chem , 40(3), 743.

Raask, E. (1979). “Sintering characteristics of coal ashes by simultaneously dilatometry-electrical conductance measurements.” J. Thermal Anal., 16, 91. Raask, E. (1985). Mineral Impurities in coal combustion—Behavior, problems, and remedial measures, Hemisphere Press, New York. Reid, W. T., and Cohen, P.(1944). “The flow characteristics of coal ash slags in the solidification range.” Trans.

ASME 66, 83. Sage, W. L., and Mcllroy, J. B. (I960). “Relationship of coal-ash viscosity to chemical composition.” Trans. ASME 82(2), 145.

270

M. Öhman and A. Nordin

Salour, D., Jenkins, B. M., Vafei, M., and Kayhaian, M. (1993). “Control of in-bed agglomeration by fuel blending in a pilot scale straw and wood fueled AFBC.” Biomass and Bioenergy, 4, 117. Sandstrom, W. A., Vora, M. K., and Rehmat, A. (1979). “Recent developments in high-temperature fluidization at the ash-agglomeration pilot plant.” AICHE 72nd Annual meeting. Sanyal, A., and Cumming, I. W. (1981). “An electrical resistivity method for detecting the onset of fusion in coal ash.” US Eng. Found. Conf. Slagging and Fouling from Combustion Gases, 329. Sanyal, A., and Metha, A. K. (199.3). “Development of an electrical resistance method based on ash fusion test.” Eng. Found. Conf. Impact of Ash Deposition Coal Fired Plants, Sollihul, England. Senior, C. L., and Srinivasacher, S. (1995). “Viscosity of ash particles in combustion systems for prediction of particle sticking.” Energy & Fuels, 9, 277. Siegell, J. H. (1976). "Defluidization phenomena in fluidized beds of sticky particles at high temperatures.” Thesis

the City University of New York. Skrifvars, B. J. (1994). “Sintering tendency of different fuel ashes in combustion and gasification conditions.” Thesis, Åbo Akademi University. Skrifvars, B. J., Hupa, M., Moilanen, A., and Lundqvist, R. (1995). “Characterization of biomass ashes.” In Applications of Advanced technology to ash-related problems in boilers. Ed. Baxter, L.L., Waterville valley. Skrifvars, B. J., Hupa, M., Öhman, M., and Nordin, A. Accepted for publications in Energy & Fuels Slegeir, W. A., and Singletary, J. H. (1988). “How reliable are correlations between coal ash chemistry and ash fusibilities.” Proc. Int. Conf. Santa Barbara, California. Smith, E. J. D. (1956). “The sintering of fly ash.” J. Inst. Fuel, 29, 253. Soltes, B. J., Lepori, W. A., and Pollock, T. C. (1982). “Fluidized-Bed Energy Technology for Biomass Conversion.” Biotechnology and Bioengineering Symp. No. 12, 15. Sondreal, E.A., and Ellman, R. C. (1975). “Fusibility of ash from lignite and its correlation with ash composition.” U.S. Bureau of Mines Rept., GRFREC/RI-75-1, Pittsburgh. Srinivasachar, S., Helbe, J. J., Katz, C. B., and Boni, A. A. (1988). “Transformations and stickines of minerals during pulverized coal combustion.” Proc. Eng. Found. Conf. Miner. Matter Ash Deposition Coal, 210. Srinivasacher, S., Helbe, J. J., and Boni, A. A. (1990). Twenty-Third Symposium (International) on Combustion; The Combustion Institute, 1305, Pittsburgh. Stallman, J. J., and Neavel, R. C . (1980). “Technique to measure the temperature of agglomeration of coal ash.” Fuel, 59, 584. Urbain, G., and Boiret, M. (1990). “Viscosities of liquid silicates.” Ironmaking and Steelmaking, 17, 255. Wall, T. F., Gupta, R. P., Polychronadis, P., Ellis, G. C., Ledger, R. C., and Lindner, E. R., (1989). “The strength,

sintering, electrical conductance and chemical character of coal ash deposits.” NERDDC Project No. 1181—Final report, Vol. I , Summary Report. Wall, T. F, Creelman, R. A., Gupta, R. P., Gupta, S., Coin, C., and Lowe, A. (1995). “Coal ash fusion temperatures: new characterization techniques and associations with phase equilibria.” Applications of Advanced Technology to Ash-Related Problems in Boilers, Ed. Baxter, L. Desollar, R., 541–556.

West, S. S., Williamson, J., and Laughlin, M. K. (1993). “Mineral interactions during fluidized bed gasification of coals.” Proc. of the Eng. Foundation Conf., “The impact of ash deposition on coal fired plants”, Solihull, Birmingham.

Wibberley, L. J., and Wall, T. F. (1982). “Alkali ash reactions and deposit formation in pulverized-coal-firedboilers.” Fuel, 61, 93. Winegartner, E. C., and Rhodes, B. T. (1975). “An empirical study of the relation of chemical properties to ash fusion temperatures.” J. Eng. Power, 97, 395. Vassilev, S., V, Kitano, K., Takeda, S., and Tsurue, T. (1995). “Influence of mineral and chemical composition of coal ashes on their fusibility.” Fuel Processing Technology, 45, 27. Vorres, K. W. J. (1979). “Effect of composition on melting behavior of coal ash.” Eng. Power, 101, 497. Öhman, M. (1997) “A new method to quantify fluidized bed agglomeration in the combustion of biomass fuels.” Licentiate Thesis, Umeå University. Öhman, M., and Nordin, A. (1996). Review of particle temperature studies in fluidized bed combustion.” Proc. Of Nordic Sem. Thermochem. Conv. of Solid Fuels, Trondheim Norway. Öhman, M., and Nordin, A. (1997). “A new method for quantification of fluidized bed agglomeration tendencies

a sensitivity analysis.” Energy & Fuels, 12, 90–94.

DEPOSITION AND CORROSION IN STRAW- AND COAL-STRAW CO-FIRED UTILITY BOILERS Danish Experiences

Flemming J. Frandsen, Hanne P. Nielsen, Peter A. Jensen, Lone A. Hansen, Hans Livbjerg, and Kirn Dam-Johansen1, Peter F. B. Hansen (1) and Karin H. Andersen (2)2, Henning S. Sørensen 3 Ole H. Larsen4, Bo Sander and Niels Henriksen 5 , and Peter Simonsen6 1

Department of Chemical Engineering, Technical University of Denmark Building 229, DK-2800 Lyngby, Denmark Phone: +45 45 25 28 83, Fax: +45 45 88 22 58,

2

E-mail: ff/hpn/paj/lah/hl/[email protected]

Midtkraft I/S Power Company, Studstrup Power Station, DK-8541 Skødstrup, Denmark. Phone: +45 86 99 17 00, Fax: +45 86 99 37 20 E-mail: [email protected],[email protected] (1) Currently with Rockwool International A/S, DK-2640 Hedehusene, Denmark. (2) Currently with SunChemical A/S, DK-4600 Køge, Denmark. 3 Geological Survey of Denmark and Greenland, Thoravej 8, DK-2400 Copenhagen NV, Denmark. Phone: +45 38 14 20 00, Fax: +45 33 63 39 89 E-mail: [email protected] Currently with Danfoss A/S, DK-6430 Nordborg, Denmark. 4 Faelleskemikerne, I/S Fynsvaesrket, Havnegade 120, DK-5000 Odense C, Denmark. Phone: +45 65 90 44 44, Fax: +45 65 90 38 12 5 Faelleskemikerne, ElsamProjekt A/S, Kraftværksvej 53, DK-7000 Fredericia, Denmark Phone: +45 79 23 33 33, Fax: +45 75 56 44 77 E-mail: bos/[email protected] 6 Elkraft A.m.b.a., Lautruphøj 5–7, DK-2750 Ballerup, Denmark Phone: +45 44 66 00 22, Fax: +45 44 65 61 04

Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

271

272

F. J. Frandsen et al.

INTRODUCTION Ash-forming elements, ie. Al, Ca, Fe, K, Mg, Na, and Si, occur in fossil or biofuels as internal or external mineral grains, simple salts such as or KCl or associated with the organic matrix of the fuel. Coals utilized in Danish power stations contain 5–15% (w/w) ash, usually rich in Si, Al, Fe, and/or Ca, while Danish straw has 2–7% (w/w) ash, usually rich in Si, K, Ca, and Cl (Sander (1997)). In pc-firing, approximately 1% (w/w) of the inorganic metals is vaporized, while the rest occurs as ash droplets (Flagan and Friedlander (1978)). Depending on the gas/particle temperature and local stoichiometry during coal particle heat-up, devolatilization and char burnout, these mineral inclusions will undergo phase transformations and approach each other to form residual fly ash, ie. fly ash particles with (Fig. 1). The vaporized metal species may undergo several transformations: nucleation, subsequent coagulation, scavenging, heterogeneous condensation and/or interactions with mineral inclusions in the burning char or residual fly ash particles. These transformations depend on the total specific surface area of the residual fly ash particles, the rate of cooling of the flue gas, the local stoichiometry, and the mixing in the gas phase. Local supersaturation with respect to certain chemical species such as and may lead to the formation of submicron ash particles by homogeneous nucleation (Flagan and Friedlander

(1978); Christensen (1995)) (Fig. 1). Vapors and fly ash particles may be deposited on heat transfer surfaces in the boiler through a number of mechanisms, e.g. inertial impaction, thermophoresis, and diffusion. Ash deposits may cause several operational problems, e.g. changes in the heat uptake of the boiler, corrosion of heat transfer metal surfaces and/or in extreme cases plugging of the convective pass of the boiler. This may cause unscheduled outages of the boiler with significant financial loss as a consequence.

DANISH EXPERIENCES WITH STRAW-UTILIZATION Through the years, Danish utilities have gained significant knowledge about how to minimize/avoid ash deposition problems in utility boilers firing worldwide high-volatile

Deposition and Corrosion in Straw- and Coal-Straw Co-Fired Utility Boilers

273

bituminous coals (Laursen (1997)). Recently, the Danish Government has decided on a 20% reduction in the carbon dioxide emissions before year 2005 with reference to 1988. Biomass is considered (CO2)-neutral due to its short time of regeneration compared to fossil fuels. Thus, the Danish power producers are enjoined to burn 1.0 Mtons of straw,

0.2Mtons of wood chips and 0.2 Mtons of straw/wood chips (free choice) every year beyond year 2004. In addition, the Danish Government has recently decided not to allow new power stations based on coal thermal conversion to be build. As a consequence, over the past few years, a number of full-scale investigations related to power generation from straw combustion has been carried out at Danish power stations. This paper provides a review of Danish experiences with deposit formation and corrosion in utility boilers fired with straw or co-fired with straw and coal. The boilers include various types of grate-fired boilers designed for straw combustion (Larsen (1996); Stenholm et al. (1996); Michelsen et al. (1996); Jensen et al. (1997)), a CFB-boiler cofiring coal and straw with up to 50% straw on an energy base (Henriksen and Hansen (1995); Henriksen et al. (1995); Hansen et al. (1996a)) and PF-boilers co-firing coal and straw with up to 30% straw on an energy base (Hansen (1994); Larsen and Inselmann (1994); Henriksen et al. (1995); Hansen et al. (1996b); Andersen et al. (1996,1997)). Further details on ash chemistry aspects of straw utilization will be provided in a subsequent number of publications (Andersen (1998), Nielsen (1998) and Frandsen et al. (1998)).

THE STRAW-FIRED SLAGELSE AND HASLEV GRATE-BOILERS Stenholm et al. (1996) and Jensen et al. (1997) have investigated the combustion of twelve well-defined batches of different straws (wheat, barley, rape) at two CHP boilers in Slagelse and Haslev, Denmark. In the Slagelse CHP, bales of straw are shredded before entering the combustion chamber through a screw-feeder. The straw burns out on a sloping grate. Boiler data are provided in Table 1. Deposition measurements were carried out at two locations using air-cooled probes, in the upper half of the furnace and in the third pass, the latter probe

274

F. J. Frandsen et al.

being located between the primary and the secondary superheater. The mean gas temperature during the wheat and barley experiments was 873 °C for the furnace location and 647 °C for the superheater location (Fig. 2). In the Haslev CHP, bales of straw are fed to the boiler, which is equipped with four cigar burners implying that the straw bales burn on the front side, from one end to the other as the bales are fed into the combustion chamber. The final burnout takes place on a sloping grate. Boiler data are provided in Table 1. Deposition measurements were carried out at two locations using air-cooled probes: in the top of the furnace and at the entrance to the third pass, just in front of the superheaters. The mean gas temperatures at the two locations were approximately 835 °C and 650°C, respectively (Fig. 2). Each experiment at the Slagelse and Haslev CHPs lasted approximately eight hours, and the parameters measured included: local gas temperatures, exit flue gas composition and aerosol particles in the flue gas. In addition, detailed analyses of the straw, fly and bottom ashes and deposit samples were conducted (Jensen et al. (1997)). For wheat and barley straw, the extent of deposit formation on the inserted probes could be correlated to the content of potassium in the straw fuel. In all experiments firing wheat and barley straw, a faster build-up of deposit was seen at the furnace probes compared to the superheater probes. Although they must be interpreted with great care, the measured total deposition fluxes from Slagelse and Haslev CHPs are provided in Table 2 (Jensen et al. (1997)). From Environmental Scanning Electron Microscopy (ESEM) combined with Energy Dispersive X-Ray (EDX) analysis, the inner and outer layer of the deposits were found to be alike, both consisting primarily of K, S and Cl which correlates with the aerosol findings in the experiments (Christensen (1995)). Detailed Computer Controlled

Deposition and Corrosion in Straw- and Coal-Straw Co-Fired Utility Boilers

275

Scanning Electron Microscopy (CCSEM) analyses of fly and bottom ashes and deposits provided from Slagelse and Haslev CHPs are by Hansen et al. (1997) and Sørensen (1997). The aerosol measurements have revealed high concentrations of submicron aerosols, during wheat-firing, and during barleyfiring. In the rape experiment at Slagelse CHP, a submicron aerosol concentration of more than was measured (Christensen (1995), Christensen et al. (1997)). The aerosols consisted almost solely of K, Cl, and S.

THE STRAW-FIRED RUDKØBING GRATE-BOILER Combined corrosion and deposition studies were carried out at the wheat strawfired Rudkøbing CHP boiler (Michelsen et al. (1996), Larsen and Henriksen (1996), Henriksen and Larsen (1996), Michelsen et al. (1998)). The boiler is a grate-fired unit with both a stationary and a moving grate. Boiler data are shown in Table 1. The experiments were performed mainly in order to evaluate the effect of increasing the steam temperatures on deposit formation and corrosion. The existing superheater tubes experienced only negligible corrosion at a steam temperature of 450 °C, but the corrosion probe experiments have shown that at a steam temperature above 520 °C, a highly temperature-dependent corrosion takes place, showing severe internal corrosion. The internal corrosion is believed to be caused by selective Cl corrosion, where gaseous chlorine reacts with primarily chromium and iron at the metal/scale interface to form volatile metal chlorides which diffuse out through the oxide layer forming a loose non-protective oxide layer and leaves behind a degrated metal phase enriched in nickel (Nielsen (1998)). The deposits collected on an air-cooled probe were larger, darker and more dense with no direct signs of molten phases on the upstream side, and powdery on the downstream side, of the probe. The deposits were rich in K and Cl (K and Cl makes up about 40–80% (w/w)) and to a smaller extent in Si, Ca and S. No significant difference in deposit composition was found as a function of the metal temperature (460 °C and 550 °C) or with increasing sampling periods (2, 4, 14–16h). SEM-analyses have revealed that the upstream deposits mainly consisted of fly ash particles and large amounts of condensed KG. Close to the metal surface, a layer of almost pure KC1 material was found, which was very dense and approximately 20–30µm thick on the (550°C)-probes. The corresponding layer was less dense and only 5–l0µ m thick on the (450°C)-probes. Threads of iron through the molten layer were seen on the (550°C)-probe, but not on the (450 °C)probe. The differences in structure of the inner layer of condensed material could be one of the reasons for the difference in the corrosion behaviour observed at the different temperatures (Nielsen (1998), Michelsen et al. (1998)).

THE STRAW-FIRED KYNDBY PF-BOILER A 2-week test run co-firing straw and oil at the Kyndby Power Station, Unit 11, was performed in September 1995. Boiler data are provided in Table 3. During the test period, the boiler was fired with oil (10–30% on an energy base) and pelletised straw,

which was pulverised in the coal mills (Michelsen (1996)). Deposit measurements were performed at the furnace outlet, and in three positions in the convective pass, at flue gas temperatures between 1,050°C and 600 °C, using air-

276

F. J. Frandsen et al.

and water-air-cooled probes. The probe metal temperature was 530 °C, and sampling periods of 3–5 and approximately 24 hours were used. The deposits collected were light grey, with most deposit forming on the upstream side of the probes. The up-stream deposits from the convective pass had an inner uniform layer with islands of deposited material on top. This was also the appearence of the mature deposits collected from boiler tubes. The deposits collected at the furnace outlet were more loose in appearence and contained more unburned material than the deposits collected in the convective pass. Deposits were rich in K, Si, Ca, S, and Cl, with a maximum content of KCl and K2SO4 of 27–60% (w/w) of the deposit. SEM-analyses revealed, that the deposits consisted of large amounts of condensed KCl and K2 SO4 as well as fly ash particles. No mixing between the salts were observed (Michelsen (1996)).

THE COAL-STRAW CO-FIRED GRENÅ CFB-BOILER Co-firing of biomass and coal have been performed with reasonable succes at the Grenaa CHP, a circulating fluid bed combustor with a maximum biomass share of 60% (energy base). Limestone is added for in-situ capture of S and quartz sand and fuel ash constituted the bed material. Boiler data for the Grenaa CFB are provided in Table 3. During the initial 8 months of operation, the combustor load never has exceeded 80% of full load and no problems with formation of fouling deposits were reported. However, increasing the load to 100%, the cyclone temperatures have increased from app. 850 °C to 900 °C–1,000°C causing serious fouling in the cyclones and on the superheaters in the convective pass (Hansen et al. (1996)). After 1½ year of operation, the final

Deposition and Corrosion in Straw- and Coal-Straw Co-Fired Utility Boilers

277

superheater was severely damaged by selective Cl corrosion and had to be replaced. The subsequent decrease of the furnace temperature by adding extra heat transfer surfaces, changing to a low-S coal, and replacing the limestone by a better quality, has reduced fouling significantly and corrosion to a somewhat lesser extent (Hansen et al. (1996a,b)).

A measuring campaign was carried out, co-firing Danish wheat straw with two different coal types (Hansen (1997a,b)). Deposit measurements were made in front of the screen tubes (just outside the cyclone) using water-air-cooled probes. The mean gas temperature at that position was 875°C–890°C and the probe metal temperature was kept at 525 °C. A significant increase in the content of water-soluble K and Cl was seen when comparing upstream probe deposits with the fly ash (from co-firing of both coals). This indicates that a major mechanism for build-up of upstream deposits (on probes) could be the condensation of KCl and/or thermoforetic transport of KCl aerosols (Hansen (1997a,b)). The effect of coal type was significant: co-firing coal A gave a probe deposit with a CCSEM-based porosity of 30%, while the corresponding coal B deposit was almost solid (CCSEM based porosity: 2.5%). Mature deposits were collected from five different positions in the boiler and analysed. It was observed that the deposits were comprised of numerous distinct layers with K present as K2SO4 or K-Al-Si-species (Hansen (1997a,b)). Deposits removed from

the final superheater tubes situated in the convective pass contained mainly K (30–35% and and some Si, Ca, and Al. Chlorine accounted for less that 0.5%, (w/w) (Hansen (1997a,b)).

Thus, comparing probe deposits with mature in-boiler superheater deposits have shown that a deposit sampling period of 3 hours may not be sufficient to obtain a complete understanding of deposit formation mechanisms in a coal-straw co-fired CFB-boiler (Hansen (1997b)). Test tubes built into the final superheater and corrosion probes inserted into the convective pass for 400–2,800 hours and operated at ultra super critical conditions, showed corrosion rates 10–25 times higher during co-firing of coal and straw as compared to 100% coal combustion. Selective Cl corrosion was found to be the predominant corrosion form (Henriksen et al. (1995)), possibly indicating a significant degree of solid phase sulfation of KC1, according to reaction (1), which will release the HC1 necessary

for the observed selective Cl corrosion to occur:

THE COAL-STRAW CO-FIRED AMAGER PF-BOILER In October 1994, a coal-straw co-combustion test was performed at the Amager Power Station, Unit 3. Part of the test programme was deposition inspections inside the furnace (Pedersen et al. (1995)). The unit is boxer-fired with oil or pulverised

coal, producing 778 tons steam per hour at 250 bar and 545 °C. During the experiments, the straw was added as pellets, and grounded with the coal in the coal mills. The coal used through most of the test period was a Canadian high-S coal. The in-boiler deposit inspections revealed, that the boiler was relatively clean after a test period of coal-firing, whereas the amount of deposit around the burners had increased after a week of coal-straw co-firing (10, 20% straw on an energy basis).

278

F. J. Frandsen et al.

Fig. 3 show pictures of the pendant tertiary superheater placed in the top of the furnace, after a period of coal-firing and one week of coal-straw co-firing, respectively.

THE COAL-STRAW CO-FIRED VESTKRAFT PF-BOILER A full-scale coal-straw co-firing test was performed at the Vestkraft Power Station, Unit 1, from October 1993 to March 1994. Boiler data are provided in Table 4. The unit is wall-fired with three burner levels, each containing four burners. Two of the mid-level burners were replaced with straw burners during the test period, see Fig. 4. Significantly lower temperatures were measured in the straw burners (800°C–900°C) compared to the coal burners (1,200°C–1,300°C). A part of the straw, introduced to the furnace in pieces of 10–12cm, was not burned, but hit the rear wall and fell directly into the bottom ash hopper. The test programme included corrosion tests at 10% straw share (energy base). At a position with a gas temperature of 1,140°C, the results from the corrosion tests showed

Deposition and Corrosion in Straw- and Coal-Straw Co-Fired Utility Boilers

279

no difference in the appearence of the corrosion test material samples at any of the metal temperatures used (560°C–630°C), when comparing coal-firing to coal-straw co-firing (Larsen and Inselmann (1994); Henriksen et al. (1995)). Most of the potassium in the fly ash was found to be non-soluble in water, and thus assumed to be present as K-Alsilicates. The chlorine is assumed to be released as HCl and Cl 2 in concentrations lower than those needed for selective Cl corrosion to occur (Inselmann and Larsen (1995)). An investigation at the Vestkraft power station by Baldacci et al. (1994) have revealed low deposition rates, <12mg/m2/s, independent of the straw-fraction fired (0, 10, 20% straw on an energy base) and sampling periods (20min.–2 hours). Based on SEM analyses, coal-firing produced sticky deposits, while coal-straw co-firing (10, 20% straw on an energy basis) gave non-sticky deposits, probably due to a lower gas temperature in the top of the furnace during co-firing (Baldacci et al. (1994)).

THE COAL-STRAW CO-FIRED MIDTKRAFT PF-BOILER Based on the experiences from the Vestkraft Power Station, a 2-year test programme of coal-straw co-firing was carried out at the Midtkraft Power Station, Unit 1. Boiler data are provided in Table 4. The unit is wall-fired with three burner rows, each of four burners, see Fig.4. The middle burner row has been rebuilt for coal-straw co-firing. Straw pieces with a max. length of 5cm, was introduced through the centre of the burner, forming a core, surrounded by a swirling annulus coal-flame. The test programme was initiated in 1996, and comprised in-situ sampling of deposits, fly ash particles, flue gas temperatures and composition, gas phase alkali metal compounds and aerosol measurements, as well as mass balance closures and examination of the fly ash with regard to

utilization in concrete and cement production. Included in the test programme were also long-term corrosion tests at 10 and 20% straw share (energy base), the effect of co-firing on de-NO x catalysts and boiler performance, including an evaluation of the straw han-

280

F. J. Frandsen et al.

dling, transport and combustion. Experiments have been performed with 0, 10 and 20% (energy base) straw share, and at 50, 75 and 100% boiler load (Andersen et al. (1996, 1997); Andersen (1998); Hansen et al. (1996)). Deposits were collected at probes with metal temperatures from 400°C to 620°C, in five positions ranging from the top of the furnace through the convective pass to the economiser section, see Fig. 4. The upstream deposits collected at flue gas temperatures from 1,300°C to 750°C during coal-straw co-firing were in all cases larger and more tenacious than the deposits collected during coal combustion (Andersen et al. (1997); Andersen (1998)). No visual effect of probe metal temperature on the upstream deposits was observed. All deposits collected at flue gas temperatures below 750°C were powdery on the up- as well as the downstream side of the probes (Andersen et al. (1997); Andersen (1998)). Increased deposit formation was seen when the load was increased from part load (50, 75%) to full load. In the long-term corrosion test at 10% straw share (energy base), the corrosion was generally found to be low, and no significant increases compared to coal-firing was observed (Larsen (1997)). The aerosol measurements revealed shown low submicron particle loadings, 30–110 compared to the measurements in the straw-fired CHPs, ie. The submicron particles from coal-straw co-firing consist of small spherical fly ash particles and even smaller, dendritic clusters, each consisting of numerous very small primary particles. The morphology and chemical composition of the submicron particles were unaffected by the straw share except for a significant enrichment in water soluble which was observed only when straw is co-fired (Nielsen et al. (1996)). The chlorine content in the particles was negligible in contrast to the high content of chlorine in the submicron aerosol particles from straw-fired CHPs (Christensen (1995); Christensen et al. (1997)).

SUMMARY AND CONCLUSION Straw-fired grate boilers produce fly ash and deposits with a very high content of K and Cl (40–80% (w/w)). Ash deposition and corrosion may constitute a significant problem in these boilers, particularly if the steam temperature is increased above 520 °C (Michelsen et al. (1996), Stenholm et al. (1996), Jensen et al. (1997)). During co-firing of straw and coal there seems to be a significant capture of potassium released from the straw, in the coal ash. This phenomenon has been observed both in the Grenaa CFB-boiler (Hansen (1997a,b)) and in the Vest- and Midtkraft pf-fired boilers. Due to the capture of K in the coal ash, only low concentrations of were observed in the fly ash and deposits from these plants. CCSEM compositional data for bottom and fly ashes and probe deposits from the Slagelse and Haslev CHPs, the Grenaa CFB and the Midtkraft Power Station, Unit 1, may be illustrated in a triangular diagram of: 1) quartz, Al silicates, and illite, 2) K + Ca silicate and 3) KCl (Hansen et al. (1997)). The three end-members in the diagram are chosen to represent: 1) relatively refractory quartz and Al silicates, 2) more easily fusible potassium and calcium dominated silicates (formed eg. by reaction between evaporated potassium and Ca silicates), and 3) low melting KCl. Figure 5 summarizes all the compositional data in such a diagram. The top triangle provides composition of fly ashes (FA) and probe deposits (D) and the bottom triangle contain the same information for bottom ashes (BA). In Fig. 5, it is seen that the fly ash and probe deposits in the straw-fired grate-boilers

Deposition and Corrosion in Straw- and Coal-Straw Co-Fired Utility Boilers

281

are rich in KC1, while in the pf-boilers the fly ash and probe deposits are rich in alumi-

nosilicates (often mixed with small amounts of Ca and K). Thus, KCl may constitute a deposition and/or corrosion problem in the grate boilers, but not in the pf-boilers when

utilizing up to 20% straw on an energy base. Recent CCSEM-analyses of probe deposits from the coal-straw co-fired Midtkraft Power Station, Unit 1, has revealed in the deposits from the convective pass, indicating condensation of during cooling of the flue gas. According to Andersen et al. (1997), ash deposition and/or corrosion will not be the major problems in coal-straw co-firing in pf-boilers. Instead, other problems

like fly ash quality aspects and poisoning of Selective Catalytic Reduction (SCR ) catalysts for reduction, may be more pronounced. The Grenaa CFB-boiler has two distinct areas marking fly ash and probe deposit composition in Fig. 5. Only little Cl is present in the fly ash, but Cl is highly enriched in

the (upstream) probe deposits. However, less than 0.5% (w/w) Cl is present in mature superheater deposits from the CFB-boiler, indicating a transformation of KC1 to with time, in the mature (boiler) deposit. This may cause a release of HCl(g) close to the metal surface and could thereby explain the 10–25 times increase in the selective Cl corrosion rate observed when co-firing coal and straw as compared to coal combustion (Henriksen et al. (1995)).

ACKNOWLEDGEMENTS This work has been financially supported by the CHEC Research Programme, ELKRAFT, ELSAM, the Danish Energy Research Programme and the Midtkraft Studstrup Demoprogramme.

REFERENCES Andersen, K. H., Hansen, P. F. B., Wieck-Hansen, K., Frandsen, F. J., Dam-Johanscn, K. (1996) “Co-Firing Coal and Straw in a 150 MWe Utility Boiler: Deposition Propensities”, Proc. 9th European Bioenergy

Conf., Copenhagen, Denmark, June 24–27. Andersen, K. H., Frandsen, F. J., Hansen, P. F. B., Dam-Johansen, K. (1997) “Full-Scale Deposition Trials at a 150 MWe PF-Boiler Co-Firing Coal and Straw: Summary of Results”. Proc. Eng. Found.

282

F. J. Krandsen et al.

Conf. “Impact of Mineral Impurities in Solid Fuel Combustion”, Kona, Hawaii, November 2–7, 1997. Andersen, K. H. (1998) “Deposit Formation During Coal-Straw Co-Combustion in a Utility PF-Boiler”, Ph.D.-Thesis, Department of Chemical Engineering, Technical University of Denmark. Baldacci, A., Bianchi, A., De Robertis, U., Paolicchi, A., Rognini, M., Tani, R. (1994) “Fireside Fouling during Coal/Straw Co-Firing in a Pulverized Coal Fired Plant at Esbjerg (DK )”, APAS-Report, ENEL-S.p.A., Thermal Research Center, Pisa, Italy (CONFIDENTIAL). CATM (1994) National Center for Excellence on Air Toxics; 1st Annual (Kickoff) Meeting, Energy and Environmental Research Center, University of North Dakota, Grand Forks, ND, May 24–25, 1994. Christensen, K. A. (1995) “The Formation of Submicron Particles from the Combustion of Straw”, Ph.D.Thesis, Dept. Chem. Eng, Techn. Univ. of Denmark, ISBN-87-90142-04-7.

Christensen, K. A., Stenholm, M. and Livbjerg, H. (1997) "The Formation of Submicron Aerosol Particles, HCl and SO2 in Straw-Fired Boilers”. To appear in J. Aerosol Sci. Flagan, R. C., Friedlander, S. K. (1978) “Particle Formation in Pulverized Coal Combustion—A Review”, D. T. Shaw (Ed.), Recent Developments in Aerosol Science, John Wiley & Sons, New York, USA,

1978. Frandsen, F. J., Nielsen, H. P., Hansen, L. A., Hansen, P. F. B., Andersen, K. H., Sørensen, H. S. (1998) “Ash Chemistry Aspects of Straw and Coal-Straw Co-Firing in Utility Boilers”, Manuscript to be presented at the 15th Ann. I n t . Pittsburgh Coal Conference, GreenTree Marriott Hotel, Pittsburgh, PA, USA, September 14–18, 1998. Hansen, K. E. (1994) “Co-Firing Straw in a PC-Boiler—I/S Vestkraft Unit 1 ” , Final Report for EU APAS Clean Coal Technology Project, COAL-CT92-0002, No. 24 (In Danish). Hansen, P. F. B., Lin, W., Dam-Johansen, K., Henriksen, N. (1996a) “Can Superheater Corrosion during Co-Combustion of Straw and Coal in a CFB-Boiler be reduced ?”, Proc. 5th Int. Conf. on Circulating Fluidized Beds, Beijing, China. Hansen, P. F. B., Andersen, K. H., Wieck-Hansen, K., Overgaard, P., Rasmussen, I., Frandsen, F. J., Hansen, L. A., Dam-Johansen, K. (1996b) “Co-Firing Straw and Coal in a 150M We Utility Boiler: In-Situ Measurements”, Proc. Eng. Found. Conf. “Biomass Usage for Utility and Industrial Power”, Snowbird,

Utah, April 28–May 3. Hansen, P. F. B. (1997a) “Deposit Formation in a Coal/Biomass Fired CFB Under Load Variations”, Internal Report. I/S Midtkraft, Elsam R&D Project No. 348. Hansen, P. F. B. (I997b) “Deposit Formation in the Convective Path of a Danish 80 MW th CFB -Boiler CoFiring Straw and Coal for Power Generation”, Proc. Eng. Found. Conf. ‘Impact of Mineral Impurities

in Solid Fuel Combustion’, Kona, Hawaii, November 2–7, 1997. Hansen, L. A., Frandsen, F. J., Sørensen, H. S., Rosenberg, P., Hjuler, K., Dam-Johansen, K. (1997) “Ash Fusion and Deposit Formation at Straw Fired Boilers”, Proc. Eng. Found. Conf. ‘Impact of Mineral Impurities in Solid Fuel Combustion’, Kona, Hawaii, November 2–7, 1997.

Henriksen, N., Hansen, P. F. B. (1995) “CFB-Material Test at Grenaa CHP Plant—Final Report of Phase 1 and 2”, ELSAMPROJEKT Note No. EP94/799a. Henriksen, N., Larsen, O. H., Blum, R., Inselmann, S. (1995) “High-Temperature Corrosion when Co-Firing Coal and Straw in Pulverized Coal Boilers and Circulating Fluidized Bed Boilers”, Proc. VGB Conf.

‘Corrosion and Corrosion Protectioning in Power Plant Technology’. Henriksen, N., Larsen, O. H. (1996) “Fouling and Corrosion in Straw and Coal-Straw Fired USC Plants”, Proc. 9th Eur. Bioenergy Conf. on Biomass for Energy and the Environment, Copenhagen, Denmark, June 24–27, 1996. Inselmann, S., Larsen, O. H. (1995) “The Corrosivity of Coal Types on Superheaters”, Fælleskemiker Report No. 4.19 (In Danish). Jensen, P. A., Stenholm, M., Hald, P. (1997) “Deposition Investigation in Straw-Fired Boilers”, Energy and Fuels, 11(5), 1048–1055. Larsen, O. H., Inselmann, S. (1994) “Superheater Corrosion during Co-Firing of Straw and Coal—Investiga-

tions at Vestkraft Unit 1”, ELSAM R&D Project Internal Report. Larsen, O. H. (1996) “Investigation on High Temperature Corrosion during Straw Combustion at Super High Steam Data”, ELSAM R&D Project 268. Larsen, O. H., Henriksen, N. (1996) “Ash Deposition and High Temperature Corrosion at Combustion of Aggressive Fuels”, Proc. Int. Conf. on Power Plant Chemical Technology, Kolding, Denmark, September 4–6, 1996. Larsen, O. H. (1997) “ MKS1 Demoprogramme: Superheater Corrosion at 10% Straw Share”, Fælleskemiker Report No. 4.06 (In Danish). Laursen, K. (1997) “Characterization of Minerals in Coal and Interpretations of Ash Formation and Depo-

Deposition and Corrosion in Straw- and Coal-Straw Co-Fired Utility Boilers

283

sition in Pulverized Coal Fired Boilers”, Geological Survey of Denmark and Greenland, Ph.D.-Thesis, ISBN-87-7871-022-7. Michelsen, H. P. (1996) “Results of Deposit Measurements during Straw Firing at the Kyndby Power Station”, CHEC Report No. 9604, Dept. Chem. Eng, Techn. Univ. of Denmark. Michelsen, H. P., Larsen, O. H., Frandsen, F. J., Dam-Johansen, K. (1996) “Deposition and High Temperature Corrosion in a 10 MW Straw Fired Boiler”, Proc. Eng. Found. Conf. "Biomass Usage for Utility and Industrial Power", Snowbird, Utah, April 28–May 3. Michelsen, H. P., Frandsen, F. J., Dam-Johansen, K., Larsen, O. H. (1998) “Deposition and High Temperature Corrosion in a 10 MW Straw Fired Boiler”, Fuel Processing and Technology, 54 , 95–108. Nielsen, H. P. (1998) “The Influence of Alkali Metals and Chlorine on Deposition and High-Temperature Corrosion in Biomass-Fired Boilers”, Ph.D.-Thesis, Department of Chemical Engineering, Technical University of Denmark. Nielsen, L. B., Pedersen, C, Røkke, M., Livbjerg, H. (1996) “Aerosol Measurements at MKSI—Final Report”, The Aerosol Laboratory, Dept. Chem. Eng. Techn. Univ. of Denmark (CONFIDENTIAL). Pedersen, L. S., Michelsen, H. P., Hansen, L. A., Kiil, S. (1995) “Full-Scale Co-Firing of Straw and Coal”, CHEC Report No. 9503, Dept. Chem. Eng, Techn. Univ. of Denmark. Sander, B. (1997) “Properties of Danish Biofuels and the Requirements for Power Production”, Biomass and Bioenergy, 12(3), 177–183. Stenholm, M., Jensen, P. A., Hald, P. (1996) “The Fuel and Firing Characteristics of Biomass—Combustion Trials”, Final Report, EFP Project No. 1323/93-0015 (In Danish). Sørensen, H. S. (1997) “Computer Controlled Scanning Electron Microscopy (CCSEM) Analysis of Straw Ash”, Proc. Eng. Found. Conf. 'Impact of Mineral Impurities in Solid Fuel Combustion', Kona, Hawaii, November 2–7, 1997.

This page intentionally left blank.

DEVELOPMENT OF BLAST-FURNACE GAS FIRING BURNER FOR COFIRING BOILERS WITH PULVERIZED COAL Takashi Kiga, Takehiko Ito, Motoya Nakamura and Shinji Watanabe Combustion Engineering Dept. Ishikawajima-Harima Heavy Industries Co., Ltd. 3-2-16, Toyosu, Koto-ku, Tokyo 135 JAPAN

1. INTRODUCTION To utilize surplus gases such as blast-furnace gas (BFG), Linz Donawitz gas (LDG) and coke-oven gas (COG) as fuels in power plants is one of the effective energy-saving measures in iron and steal companies. For the purpose, a 145 M We, wall firing boiler was designed and constructed by Ishikawajima-Harima Heavy Industries Co., Ltd. ( I H I ) to fire three kinds of surplus gas, BFG, LDG and COG other than pulverized coal and heavy oil. Among these fuels, BFG assisted by COG or heavy oil and pulverized coal were main fuels with whichever the full load could be obtained. Commissioning tests have started with pulverized-coal firing, and satisfactory performances have been obtained with the coal alone. However, BFG burners and LDG burners were found to be subjected to a serious ash trouble on their burner mouth and gas nozzle after 2 weeks’ coal firing. A lot of amount of ash was deposited on the burner mouth and it was plugged the gas nozzle partially or entirely as shown in Fig. 1. In addition, as shown in Fig. 2, which shows the status of the gas nozzle of the BFG burner after removing ash, it was deformed. In order to investigate the cause of the trouble and to propose the method for burner modifications, experimental and analytical studies were carried out.

2. SPECIFICATIONS OF THE BOILER The subject boiler was constructed in the iron and steal company as a steam generator of a 145 MWe thermal power plant and it started the commercial operation in 1997. The unit is equipped with flue gas treatment system such as selective catalytic reactor (SCR), flue gas desulphurization (FGD) and electrostatic precipitator (ESP). Figures 3 and 4 show the sectional side view of the boiler and the burner arrangement, respectively. Twenty four (24) burners are arranged in three (3) stages with four (4) Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

285

286

T. Kiga et al.

rows for each opposed wall. Upper two (2) stages of front wall are for BFG, an upper stage of rear wall is for LDG, and lower three (3) stages are for pulverized coal from three (3) roller mills. COG nozzles and heavy fuel oil burner are also installed in the BFG and LDG burners. So they are called as “combination burners”. The IHI NOx Preventing Advanced Combustion Technology (INPACT) [Kiga, 1989] is applied so that a very

low NO x operation can be realized; that is, over fire air ports (two-stage combustion air ports) are settled up in consideration of the residence time in the furnace. The main specifications of this boiler are shown in Table 1. Typical properties of surplus gases are tabulated in Table 2. Since the BFG has a very low heating value of around its flame is stabilized by a pilot flame of COG or heavy fuel oil. On the other hand, properties of coal used in the commissioning test are presented in Table 3. The coal was Australian bituminous coal and it was pulverized by the mills into over 90% through 200 mesh sieve [Tamura, 1996].

3. EXPERIMENTAL Several kinds of half-scale size burner model were manufactured and they were utilized both for flow pattern testing using cold air and for combustion testing.

3.1. Test Facilities These tests were conducted in the 12 MW(thermal) pilot-scale combustion test facilities in IHI’s Aioi Works. The specifications of the test facilities are listed in Table 4 and the general arrangement of its furnace is schematically shown in Fig. 5. The furnace comprises of a multi-burner furnace and a single-burner one. The single burner furnace was used in the test.

Development of Blast-Furnace Gas Firing Burner for Cofiring Boilers with Pulverized Coal

287

3.2. Test Burners The tests were conducted for three (3) kinds of BFG burner model, Type I (original burner model), Type II, and Type III, where Type II was a model of a temporally modified burner and Type III was that based on the final modification design. The photographs of these burners tested are shown in Fig. 6.

288

T. Kiga et al.

The BFG is provided from an annulus nozzle, and multi-spud type COG nozzles are arranged in the inside of the inner barrel of the BFG nozzle. The heavy fuel oil burner is

mounted in the center of the burner and a swirler is attached near an atomizer for flame stabilizing. The primary air which is divided from the combustion air at the inlet of the burner is supplied in the center of the burner, and it is minimized when the BFG is not provided. The secondary air is supplied from multiple rectangular nozzles arranged in the BFG nozzle for Type I and Type II, and it is supplied from around the BFG nozzle for Type III. Type II and Type III models are equipped with multiple openings on the outer barrel of the BFG nozzle to lead secondary air into BFG nozzle. In addition, Type II model has a baffleplate at the outlet of the secondary air nozzle so as to avoid the reverse flow toward the secondary air nozzle. Type III model has an baffleplate at downstream of the openings to change the direction of the air flow from the openings into a suitable way to avoid the reverse flow toward BFG nozzle when the burner is out of service.

Development of Blast-Furnace Gas Firing Burner for Cofiring Boilers with Pulverized Coal

289

Two (2) types of burner throat were prepared; one was made by styrene foam used in the flow pattern testing in order to measure the velocity in the burner by removing it‚ and the other was made by steal and refractory used in the combustion testing.

3.3. Test Conditions The ash troubles were occurred during pulverized-coal firing when the BFG burners were out of service. The flow pattern tests were therefore carried out in the condition that

290

T. Kiga et al.

the air flow rate was settled so that the air velocity through the burner throat was simulated for actual conditions when the BFG burner was out of service. In the tests‚ the dis-

tributions of the air velocity were measured by a hot-wire anemometer at the section of the edge of burner tile and throat. On the other hand‚ in the combustion tests‚ an oil stabilizer was operated to simulate the condition in which the metal temperature measured on the BFG nozzle took the highest value in the actual boiler. The light oil was used in the tests.

4. STRESS AND DEFORMATION ANALYSIS The thermal stress and deformation analysis was performed for the real size burners of Type I (original) in order to examine the reason of the deformation of the gas nozzle

of the BFG burner. The analysis was carried out using the finite element program‚ MSC/NASTRAN code. Analytical conditions were given based on the metal temperature measured in the actual operation.

5. RESULTS AND DISCUSSIONS Figure 7 illustrates distributions of air velocity obtained from the flow pattern tests. A reverse flow coming from the outside into the inside on both the secondary air nozzle

and the BFG nozzle is found for the original burner (Type I). Such a reverse flow conveys

Development of Blast-Furnace Gas Firing Burner for Coliring Boilers with Pulverized Coal

291

hot flue gas and fly ash onto the burner throat and into the nozzles‚ and will cause ash depositions on them. On the other hand‚ Type II (temporally modified burner) and Type III (finally modified burner) have positive air velocity at the secondary air nozzle and BFG nozzles and near the burner throat. Especially for Type III burner‚ the strong cleanup and cooling effects on the burner throat are expected owning to the secondary air flow along the throat. The secondary air flow through the BFG nozzle also make the BFG nozzle and the burner tile clean and cool. The modifications of flow pattern enables

292

T. Kiga et al.

to prevent these burners from ash troubles. Furthermore‚ Type III burner forms so-called a strong internal circulation flow‚ which will make the flame stable‚ although the test conditions are to simulate the status of burner’s out-of-service. The flames observed in the combustion tests are shown in Fig. 8. As is expected from the flow pattern test results‚ good flame stability is obtained about Type III burner. As for the metal temperature of the BFG nozzle‚ it was confirmed that that of modified burners (Type II and Type III ) was clearly lower than that of original burner as shown in Fig. 9.

Development of Blast-Furnace Gas Firing Burner for Cofiring Boilers with Pulverized Coal

293

The result of the thermal deformation analysis is shown in Fig. 10. It is obviously found that the deformation pattern obtained in the analysis quite resembles to that observed in the actual boiler as previously shown in Fig. 2. From the analytical results‚

the rigid structure of the secondary air nozzle and the BFG nozzle regardless of thermal deformation are the cause of such damage. In the modification at the actual boiler‚ therefore‚ the swirl plates for secondary air and BFG were not fixed and the BFG nozzle was slit to prevent it from deformation by stretch‚ in addition to making the nozzle cool as mentioned above. Based on the above-mentioned studies‚ all of gas burners including LDG burners in the subject boiler were modified into new type ones. After the modification‚ the plant has been operated without such troubles.

CONCLUSIONS Against a serious ash trouble on BFG burners and LDG burners‚ experimental studies and thermal stress analysis were carried out and the followings were obtained. • Ash deposits on gas nozzles and burner mouth were caused by the reverse flow conveying hot flue gas and fly ash toward the gas nozzles and along the burner mouth. • The deformation of the BFG and secondary air nozzles was caused by its rigid and cooling-less structure.

The studies offered ash-free BFG and LDG burners for cofiring with pulverized coal and the subject boiler has been free from such troubles. Since to utilize surplus gases in power plants is one of the effective energy-saving and measures in iron and steal companies‚ we would like to proceed to spread our technologies widely.

294

T. Kiga et al.

Development of Blast-Furnace Gas Firing Burner for Cofiring Boilers with Pulverized Coal

295

ACKNOWLEDGEMENTS We would like to acknowledge the assistance and suggestions of Mr. Yutaka Takamatsu of Volcano Co.‚ Ltd.‚ a manufacturer of the combination burners‚ in the tests and modifications.

REFERENCES Kiga‚ T‚ Miyamae‚ S.‚ Makino‚ K.‚ and Suzuki‚ K. (1989). “Low NOx Combustion Technologies for CoalFiring Boilers and Its Application Results to Commercial Boilers.” AIChE 1989 Summer National Metting‚ Philadelphia‚ PA. Tumura‚ M.‚ Kiga‚ T.‚ Endo‚ Y.‚ Murakami‚ H.‚ and Tanaka‚ T. (1996). “Practical Studies on Combustion Technologies of Micro-Pulverized Coal.” IHI Engineering Review‚ 29(3)‚ 81–86.

This page intentionally left blank.

CHANGES IN SLAGGING BEHAVIOUR WITH COMPOSITION FOR BLENDED COALS Nicholas J. Manton1‚ and Jim Williamson1‚ and Gerry S. Riley2 1

Department of Materials Imperial College of Science Technology and Medicine London SW7 2BP‚ UK 2 National Power Plc Windmill Hill Business Park Swindon SN5 6BP‚ UK

1. INTRODUCTION Coal blending has traditionally been used to provide fuels of a consistent quality‚ to reduce fuel costs and to improve the combustion behaviour of low volatile coals. In addition‚ blending may be used to enhance fuel flexibility‚ to extend the range of acceptable coal feeds and to control the sulphur content of the fuel to meet emission regulations. UK experience in the use of blended fuels for utility use is limited‚ although increasing amounts of overseas coals are now imported by the power generators. In 1994 British Coal was sold to private industry and the contracts which were in place at the time will expire early in 1998. Intensive bargaining between the coal users and the coal suppliers is now in progress to maintain a UK coal mining industry‚ which currently supplies approximately Tpa. This scenario has to be set against a background of an increasing use of natural gas for power generation and a market price for world traded coals which can be significantly less than indigenous coals. While the amount of imported coal used by utilities could be limited by the problems of transporting the coal from the ports to the power stations‚ the amount of imported coal could rise significantly above the amount currently used. Many imported coals would lie outside the normal range of boiler specifications; thus‚ blending with indigenous coals would be a prerequisite for efficient boiler operation. This study has been undertaken to provide a method of determining the change in slagging propensity when UK and overseas coals are blended for use in pulverised coalfired boilers. Impact of Mineral Impurities in Solid Fuel Combustion‚ edited by Gupta et al. Kluwer Academic / Plenum Publishers‚ New York‚ 1999.

297

298

N. J. Manton et al.

1.1. Properties of Blended Coals Certain properties of a coal blend can be calculated from the weighted average of the measured values of the single coals‚ as reviewed by Carpenter [1995]. These properties include:

• moisture content • volatile matter • ash content • fixed carbon • hydrogen‚ sulphur‚ nitrogen‚ oxygen‚ chlorine and proportions of macerals These properties can then be considered as additive properties of the individual coals. Coal properties which are non-additive include:

• swelling index • Hardgrove grindability • ash fusion temperatures

• slagging propensity

Thus the slagging propensity of a blend requires careful evaluation to avoid costly boiler operational problems.

1.2. Assessment of the Slagging Propensity of a Coal Fuel technologists have long sought reliable methods for predicting the slagging propensity of a coal ash‚ as described by Williamson and Wigley [1994]. Conventional methods of assessment based on ash fusion tests‚ viscosity of coal ash melts and empirical indices calculated from the chemical composition of the ash were reviewed by Couch [1994]. These techniques all fail to reflect the complex mineral matter transformations and reactions which occur at the combustion temperatures. In addition‚ the short residence times for ash particles within the radiant zone of the boiler (1–2 sec) frequently do not allow the decomposition of larger mineral particles to reach equilibrium. Wigley and Williamson [1996] reported the use of advanced microstructural analysis‚ combined with chemical analysis of individual ash particles obtained using computer controlled scanning electron microscopes (CCSEM). This demonstrated the diverse chemical nature of ash particles and the lack of chemical homogeneity. Thus‚ to obtain reliable data‚ fuel technologists have developed combustion test facilities which closely simulate the conditions in large pulverised coal boilers. The results obtained with such facilities still require

validation with full scale boiler trials‚ as reported by at a previous Engineering Foundation Conference by Gibb [1996]. The high cost of these trials‚ and the difficulty of obtaining an objective assessment of the slagging propensity of a coal ash stimulated the design and construction of a entrained flow reactor (EFR) at Imperial College to provide a

laboratory facility to study ash formation and deposition phenomena‚ as reported by Hutchins et al. [1996]. This facility has provided a means of preparing coal ash slags and deposits from both single and blended fuels‚ with preliminary findings being reported by Manton et al. [1996].

The EFR was designed to closely simulate the conditions which pf and ash parti-

Changes in Slagging Behaviour with Composition for Blended Coals

299

cles experience in large utility boilers. The reactor consists of a vertical multi-zoned furnace‚ approximately 5m in length‚ with an internal diameter of 100mm. Four independently controlled furnaces heat the reactor‚ providing a temperature gradient from 1‚650°C at the top to 1‚200°C at the bottom. A series of sample ports between each furnace allows ash and char samples to be withdrawn from the combustion atmosphere. Ash deposition probes may also be inserted into these ports. Pulverised coal is introduced at the top of the furnace at a rate of the chosen feed rate depending on the ash content of the coal or blend under investigation. Gas flow rates of approximately 70 (STP)‚ equivalent to at 1‚650°C‚ entrain all coal and mineral particles with a density of or less and a particle size of Under these conditions‚ particle residence times from the top to bottom of the reactor are approximately three seconds. The reactor is shown schematically in Fig. l(a)‚ with the position of the sample ports. Gas temperatures at the top sample port are approximately 1‚400°C and 1‚250°C at the second port. Both air-cooled metal deposition probes and uncooled ceramic probes have been used to collect ash deposits. Deposits collected at 1‚400°C are generally of a highly fused nature and give little indication of the likely slagging propensity of the coal. However‚ when deposits are collected at 1‚250°C on an uncooled ceramic probe‚ they range from a thin dusty covering of lightly sintered ash particles to well-bonded and coherent deposits‚ the type of the deposit depending on the nature and proportions of the mineral matter present in the coal. Deposits collected under these conditions closely resemble those which form on the platten superheaters of large utility boilers and are therefore suitable for characterisation.

2. EXPERIMENTAL TECHNIQUES The ceramic deposition probe used in this study was of simple construction‚ consisting of a mullite tube or coupon‚ 17mm OD and 12.5mm ID‚ and 150mm in length‚ see Fig. l(b). The coupon is held in place over a longer (approximately 1 m) mullite tube of 12mm OD‚ with a small amount of a high alumina cement mixed with a sodium silicate solution. The same cement was used to block the open end of the 12mm tube‚ thus preventing gas leakage into or out of the reactor. An aluminium holder with an “O” ring seal is used to hold the probe in position in the sample port.

2.1. Production of Deposits Each of the coals was ground separately to pf grade‚ i.e. 70% was less than and blends were prepared in a laboratory mixer. The blended coals were dried at 100 °C for an hour or more to remove adsorbed moisture which can cause problems with the coal feed system to the EFR. Over a l–2hr period‚ the deposits varied in thickness from 1–5mm depending on the amount of mineral matter in the coal and the feed rate.

2.2. Characterisation of Deposits and Determination of Slagging Propensity A small sample of each deposit was first removed for X-ray powder diffraction analysis to establish the crystalline phases present‚ while the remainder of the deposit was left undisturbed on the ceramic coupon. The deposit was then coated in a low viscosity epoxy resin‚ which penetrated the open porosity of the sample‚ but when set gave

300

N. J. Manton et al.

Changes in Slagging Behaviour with Composition for Blended Coals

301

stability to even the most fragile deposits. Sections through the deposit were cut with a diamond saw, perpendicular to the surface, to give samples approximately 5mm in length. These samples were then set in resin blocks so that a cross-section through the deposit from the interface with the ceramic to the surface of the deposit could be examined. The samples were ground and polished to a diamond finish, and coated with carbon for SEM examination. At low magnification a series of backscattered images (BSI) were first obtained to give an overall view of the microstructure of the deposit. Typical microstructures are shown in Fig. 2. More detailed examination was then made of selected areas at higher magnifications, typically Quantitative chemical and image analysis data was then acquired using a Tracor Noran low element detector and Tracor Noran Voyager III software. Typically, a chemical analysis was acquired using a matrix of analysis points, giving 256 separate analyses for each area examined. Quantitative EDS analyses were obtained from this data using Proza corrections. The porosity of the microstructure, which gave a strong indication of the amount of viscous flow sintering which had occurred, could be obtained from the grey scale digital image which differentiated between the ash and the resin (i.e. the filled pores) or from the fraction of analysis points reporting oxide analyses, where the criteria used for this were that analyses with could reasonably be assumed to have originated from an ash particle. The chemical and microstructural data was normally collected from several areas within each section and from samples taken at intervals along the coupon length. Typically, the number of analysis points was 1,500 or so, thus giving sufficient data for reliable statistical variations to be established. An assessment of the slagging propensity of the blended coal ash was made by a consideration of the analysis and the chemical variations shown by the ash particles and the bonding phase(s). The four main oxides which make up the ash composition are and CaO and the chemical data may be displayed by normalising each ash composition to one of two ternary systems, i.e. and Only those analysis points where or exceeded 70wt% were taken, since the normalisation of compositions totaling less than this could lead to substantial errors. The data obtained may then be plotted on a ternary diagram, see for example Figs. 3 or 5. Each plot shows not only the range of chemical compositions shown by the structure, but also the degree of chemical inhomogeneity in

302

N. J. Manton et al.

the sample. This indicated the degree of interaction between the decomposed mineral residues‚ i.e. the clays (kaolinite and illites)‚ pyrite‚ calcite‚ quartz etc. An assessment of the slagging propensity was made by calculating the proportion of the analyses where the content lay between 10 and 50 wt%‚ or the CaO concentration was between 5 and 40 wt%. Within these compositional limits the aluminosilicate melts would be considered to contain sufficient fluxing oxides to reduce the viscosity of the slag to a value at which the sintering of the ash particles by a viscous flow process would proceed at a measurable rate.

4. RESULTS AND DISCUSSION The three UK coals (UK 1‚ 2 & 3) have been blended with either American coals (AM 1&2) or South African coals (SA 1 & 2). The ash contents and ash compositions of each coal are shown in Table 1. The UK coal ashes are typical of the range of and CaO contents to be found in boiler ashes in UK utility coals. UK 1 has the highest and CaO contents‚ and this is reflected in the B/A ratio. This coal would be considered a moderate to high slagging coal‚ whereas the other coals with much lower B/A ratios would not be expected to cause any real problems.

4.1. CCSEM Mineral Matter Characterisation A CCSEM analysis of the UK 1 pf indicated a coal with much free pyrite and some

pyrite intimately mixed with kaolin and illitic clays low in

By contrast‚ coals UK

2&3 contained much less pyrite but an illitic clay much richer in

The

CCSEM analysis of the American coal AM 1 is shown in Fig. 3. This coal contains relatively little free pyrite and only a small amount of pyrite mixed with the clays; the mineral matter consisting largely of quartz‚ kaolin and illites‚ some of which are intimately mixed. The South African coals are low in Fe‚ contain little free quartz‚ and show mostly kaolinite as opposed to illitic clays. Ca‚ Mg and Fe are found as calcite‚ dolomite and

siderite. Fig. 4 shows how the Fe and Ca is distributed amongst the various mineral types for UK 1 and AM 1 coals. The relatively high proportion of Fe present as free pyrite (Fe-rich) in the UK 1 is striking‚ whereas for coal AM 1 most of the Fe is as an intimate

Changes in Slagging Behaviour with Composition for Blended Coals

303

pyrite-clay mixture (Fe-Alsil). Almost all the Ca in UK 1 is found as carbonates (nonCFAS)‚ whereas for AM 1 the Ca is divided between carbonates‚ carbonate-clay mixtures (Ca-alsil) and illitic clays (Alsil). The differences in the distributions of the Fe and Ca for these two coals may play a crucial role in the initial stages of ash formation. The results of the CCSEM analysis of UK 1 are shown in Fig. 5. The differences compared to coal AM 1 (Fig. 3) could not be more marked. UK 1 shows significantly less association between the pyrite and the clays‚ with virtually no free quartz‚ and much less quartzclay inter-mixing.

3.2. Change in Slagging Propensity of UK Coals on Blending The results on blending the American coal AM 1 with UK 1 are shown in Fig. 6. The variation in slagging propensity‚ as determined from a number of cross-sections from

304

N. J. Manton et al.

each deposit‚ are shown in the Figure by the length of the error bars. Cross-sections closest to the centre of the deposit were in general slightly higher in than sections towards the end of the coupon. This is undoubtedly due to the heavier pyrite particles showing less inclination to be deflected around the ceramic probe by the gas stream than the lighter aluminosilicate ash particles. Thus‚ within the limits of error of the technique‚

Changes in Slagging Behaviour with Composition for Blended Coals

305

the slagging propensity of the UK 1—AM 1 blends would appear to be a linear function of the blend composition. The change in slagging propensity of UK 3 when blended with the South African coal SA 1 is shown in Fig. 7 by the solid line. An initial slight decrease in slagging propensity on adding SA 1 is followed by a sharp increase. The slagging propensity of pure SA 1 lies below the curve for the blends‚ suggesting that a small amount of UK 3 when blended with SA 1 leads to a rapid increase in the slagging propensity of this coal. The dotted curves in Fig. 7 show the proportion of points from either the or the systems which would be deemed to be of sufficiently low viscosity to allow viscous flow sintering. The increase in slagging propensity is thus seen to be entirely due to the Ca-aluminosilicate compositions‚ which for the pure SA 1 had insufficient time for all the CaO to flux and enter the melt; thus these compositions tended to be rich in CaO and to lie outside the arbitrary limits chosen to assess the slagging propensity.

Coal UK 1 and the American coal AM 2 were observed to have similar slagging

306

N. J. Manton et al.

propensities‚ see Fig. 8. Although the pyrite content of UK 1 is much higher than that of AM 2‚ much of the pyrite in UK 1 is present as free mineral matter and is therefore slow to react with the decomposing clays forming the aluminosilicate melts during the initial stages of ash and deposit formation. However‚ when these two coals were blended‚

pronounced deviations from linearity with respect to blend composition were observed. Similarly‚ the effect of blending the South African coal SA 2 with UK 2 showed

Changes in Slagging Behaviour with Composition for Blended Coals

307

positive deviations from linearity‚ as shown in Fig. 9‚ although the magnitude of the nonlinearity is less than with the other blends. CCSEM examination of the fly ash particles from a blended coal has revealed a similar chemical distribution to that from a mechanical mixture of the constituent ashes. Thus the origin of the observed non-linear behaviour must follow the deposition of the ash particles. At this stage the composition of individual ash particles is crucial in determining the rate at which undissolved iron oxide particles will dissolve in the melt. The

308

N. J. Manton et al.

evidence to date is that CaO species readily combine with decomposing clays‚ since the residues of high CaO minerals are seldom observed in fly ashes. As the basicity of an aluminosilicate melt increases‚ the viscosity decreases and its power to enhance the rate at which the iron oxides will dissolve increases. An understanding of the complex nature of the interactions between ash particles at high temperatures remains one of the key issues to improved slagging predictions‚ be it from a single coal or a blend.

4. CONCLUSIONS Blends of UK coals with either American or South African coals show non-linear behaviour‚ with enhanced slagging propensity frequently observed with relatively small additions of an overseas coal to a UK coal. The most pronounced increases were observed when a UK coal‚ high in was blended with a coal high in CaO. The ability of a Ca-aluminosilicate melt to increase the rate of dissolution of free pyrite from a UK coal would appear to be a significant feature in accounting for the observed phenomena.

5. ACKNOWLEDGEMENTS The authors gratefully acknowledge the financial support provided by National Power plc and the UK Engineering and Physical Sciences Research Council (EPSRC) and the help and guidance given by Mr F. Wigley with the CCSEM analysis.

6. REFERENCES Carpenter‚ A. (1995)‚ “Coal Blending tor Power Stations”‚ IEA Coal Research London‚ IEACR/81. Couch‚ G. (1994)‚ “Understanding Slagging and Fouling in PF Combustion”‚ IEA Coal Research London‚ 1EACR/72. Gibh‚ W.H. (1996)‚ “The U K collaborative research programme on slagging in pulverised coal-fired boilers: summary of findings”‚ in “Applications of Advanced Technology to Ash-Related Problems in Boilers”‚ Eds. L. Baxter and R. DeSolar‚ pp 41–65‚ Plenum‚ New York. Mulchings‚ I.S.‚ West‚ S.S. and Williamson‚ J. (1996) “An assessment of coal-ash slagging propensity using an entrained flow reactor” in “Applications of Advanced Technology to Ash-Related Problems in Boilers”‚ pp 201–222‚ Plenum‚ New York. Manton‚ N.J.‚ Riley‚ G.S. and Williamson‚ J. (1996)‚ “A laboratory assessment of the slagging propensity of blended coals”‚ 212th American Chemical Society National Meeting‚ Fuels Division Vol. 41‚ No 3‚ pp 1113–1117. Wigley‚ F. and Williamson‚ J. (1996)‚ “Modelling fly ash generation for UK power station coals” in “Applications of Advanced Technology to Ash-Related Problems in Boilers”‚ Eds. L. Baxter and R. De Solar‚ pp 593–603‚ Plenum‚ New York. pp. 41–65‚ Plenum‚ New York. Williamson‚ J. and Wigley‚ F. (1994)‚ “Mineral Impurities in Coal-Fired Plants”‚ published by Taylor and Francis‚ Washington.

ROLE OF INORGANICS DURING FLUIDISEDBED COMBUSTION OF LOW-RANK COALS

Hari Babu Vuthaluru and Dong-ke Zhang

CRC for New Technologies for Power Generation from Low-Rank Coal C/- Department of Chemical Engineering‚ The University of Adelaide‚ Australia 5005

1. INTRODUCTION Inorganic matter in coal continues to be one of the major sources of operational problems during coal combustion. The role of inorganic matter during coal combustion is governed by its nature and occurrence in the coal matrix [Raask‚ 1985]. In pulverised coal combustion processes‚ slagging‚ fouling and corrosion are the main problems which

arise from the coal mineral matter. These problems can be even more severe when firing low-rank coals with high alkaline and transition metal ions. In contrast‚ fluidised bed combustion processes operating at much lower tempearures ( 800–900 °C) offer better prospects for reducing ash-related problems associated with the firing of low-rank coals. However‚ low-rank coals do not present trouble-free operation in fluidised beds. A major problem is presented by particle agglomeration‚ bed defluidization and ash deposition on heat exchanger surfaces. The inorganic constituents in low-rank coals (mainly sodium‚ calcium and organically bound sulphur) transform to low-melting eutectics at fluid bed operating temperatures [Manzoori and Agarwal‚ 1993; Hodges and Richards‚ 1989; Manzoori and Agarwal‚ 1992; Souto et al.‚ 1996]. These eutectics are transferred to the surface of inert bed particles leading to agglomeration and defluidization of the bed. This phenomenon is dependant to a great extent on the operating temperature of the fluid bed combustor. The present work is aimed at understanding the effects of bed temperature on ash characteristics during fluid bed combustion of an Australian and an Indonesian low-rank coal. Data is presented to demonstrate the effect of bed temperature on the distribution of inorganic matter into variuos streams such as bed ash and cyclone ash. Comparisons of the characteristics of ash deposition on bed material and bed defluidization for Australian and Indonesian low-rank coals are presented. In addition‚

the chemical nature of the ash deposits is discussed‚ along with the chemical interactions that might have occurred during the combustion experiments to cause ash-related problems. Impact of Mineral Impurities in Solid Fuel Combustion‚ edited by Gupta at al. Kluwer Academic / Plenum Publishers‚ New York‚ 1999.

309

310

H. B. Vuthaluru and D.-k. Zhang

2. EXPERIMENTAL Experiments were carried out in a spouted bed combustor as described by Man-

zoori and Agarwal (1993). It consisted of a cylindrical furnace‚ 77mm i.d.‚ with a conical distributor (Fig. 1). The combustion zone was confined to the lower part of the furnace‚ where the bed particles and char recirculates internally. Combustion air was introduced at the base of the furnace after passing through an electric heater. The air flow was controlled to create an average fluidisation velocity in the combustion zone of A 200 g batch of bed material (silica sand) was fed into the fluidisation vessel from an opening at the top of the reactor. Air dried coal‚ ground and sieved to 1.0–3.35 mm‚ was fed into the combustion zone via a water-cooled screw feeder. Low-rank coal samples of South Australia (Coal A) and Indonesia (Coal B) were used in the present work. The

chemical analyses of coal samples are given in Table 1. At the end of a combustion run‚ the ash coated bed material was discharged into the ash can. Ash coated sand particles were collected on defluidization with Coal A runs whereas with Coal B the samples were collected at a given time (typically 6 h duration). Ash coated bed material and fly ash samples retrieved from the combustor and cyclone were weighed. Samples of ash coated bed material and cyclone ash collected at the end of each run were analysed using several analytical techniques which included: wet chemical analysis‚ electron microprobe (EMP)‚ scanning electron microscopy (SEM) and X-ray diffraction (XRD). Chemical analysis was done to determine the composition of bed material coating and cyclone ash. Elemental analyses were performed on polished cross-sections of bed material grains using SEM equipped an energy dispersive X-ray detection system (EDAX). Ash coating thickness was estimated by examining the polished cross-section of bed particles with an electron microprobe (EMP) operated in a back scattered-electron imaging mode. XRD was also used to identify the several mineral phases present in various ash

samples.

Role of Inorganics During Fluidised-Bed Combustion of Low-Rank Coals

311

3. RESULTS AND DISCUSSION

3.1. Ash Coating The ash samples showed an ash coating on the surface of the bed particles. Comparison of the samples from the two coals revealed that only a very small amount of Coal B ash deposited on bed material particles. The cause for the difference in ash coating was attributed to the small amount of low melting point species in Coal B ash as compared to Coal A (in particular‚ 0.25% in Coal B as compared to 15% in Coal A).

3.2. Composition of Ash Coating The composition of the ash coating on the bed material was calculated from the chemical analysis on a silica-free basis in order to eliminate the contribution of the quartz bed particles. Fig. 2 shows a comparison of the average composition of the bed material coating‚ cyclone ash and raw coal ash for the tested coals at a bed temperature of 800 °C. The results show that the coating on the quartz particles was enriched in Na and S compounds especially with Coal A runs. Electron microprobe spot analyses (Fig. 3) and X-ray maps of polished cross-sections of coated sand grains showed the presence of a highly sulphated ash in the coating‚ confirming previous observations [Manzoori and Agarwal‚ 1992]. On account of the lower interfacial tension between the ash and silica bed particles compared to that between the ash and char surface‚ previous researchers [Manzoori and Agarwal‚ 1993; Manzoori and Agarwal‚ 1992] suggested that the ash transfers

to the surface of bed particles by collision between sand grains and the burning char particles. It may be noted that in comparison to the composition of feed coal ash‚ a large proportion of Ca and Mg compounds was retained in the bed material coating and cyclone ash. For Coal B‚ however‚ no sulphur was seen in the ash coating (from electron microprobe spot analyses) and a significant proportion of Al has escaped into the cyclone.

312

H. B. Vuthaluru and D.-k. Zhang

i

Figure 4 shows the effect of temperature on the distribution of inorganics in bed material coating and cyclone ash for the tested coals. The effect of bed temperature was more pronounced on Coal A than Coal B. The proportions of inorganic constituents (namely Fe‚ Ca‚ Mg and Al) in bed ash and cyclone ash showed an increase with increasing bed temperatures for Coal A runs. At a bed temperature of 850 °C‚ the levels of Ca‚ Mg and Al decreased in both ash coating and cyclone ash streams. For Coal B‚ similar trends (for all the inorganic constituents in the ash coating except Fe) were observed with increasing bed temperatures. At higher bed temperatures‚ the level of Al decreased in the ash coating. However‚ the bed temperature showed little influence on the inorganic constituents in the cyclone ash. The results indicated that Coal A is likely to cause more ash-related problems than Coal B at typical fluid bed temperatures. Owing to increased levels of sodium in fly ash with coal A (Fig. 5)‚ deposition of Na-rich compounds on heat transfer surfaces [Souto et al.‚ 1996; Dawson and Brown‚ 1992] can be significant‚ which presents another potential problem‚ fouling and high-temperature corrosion‚ in fluidised bed combustion of lowrank coals‚ in addition to agglomeration and defluidization.

3.3. Ash Coating Thickness Figure 6 shows the effect of temperature on coating thickness on the bed particles for the tested coals. It should be noted that the ash coated samples for Coal A were collected on defluidization whereas with Coal B the samples were collected for the same operating time (typically 6 h duration). The coating layer thickness decreased with increasing bed temperatures for Coal A runs. The increased proportion of low-melting point compounds (in the ash layer) in this coal decreases the critical thickness (ash coating with sufficient amounts of sticky ash material‚ Manzoori and Agarwal‚ 1994) required for agglomeration and/or defluidization. Hence the bed defluidized after 0.5hrs of operation at 850 °C‚ with only a relatively

thin coating of ash. For Coal B‚ however‚ the coating layer increased with increasing bed temperatures. This may be due to decreased proportions of low-melting species in the ash coating [Manzoori and Agarwal‚ 1994] in Coal B. The results suggest that the combustion of Coal A in fluid bed combustors for long operating periods would not be feasible as the coal is likely to defluidize after few hours of operation (even with only a thin ash coating). On the other hand‚ utilisation of Coal B is less likely to cause such prob-

Role of Inorganics During Fluidised-Bed Combustion of Low-Rank Coals

313

314

H. B. Vuthaluru and D.-k. Zhang

Role of Inorganics During Fluidised-Bed Combustion of Low-Rank Coals

315

lems when employed in fluid bed combustors mainly due to the lower proportions of liquid phases in the ash coating.

3.4. Ash Characteristics 3.4.1. Ash Coated Sand Grains. XRD analysis of sand grains obtained from Coal B showed the presence of several minerals (quartz‚ maghemite and hematite) in various proportions (Table 2). In addition‚ several minerals were present in trace quantities. Analyses of sand grains obtained from Coal A also revealed quartz as the dominant phase in all the samples. However‚ all the other minerals were observed in minor and trace proportions. With increasing bed temperatures‚ the distribution of mineral phases in various proportions changed significantly for Coal A runs. This effect was not so pronounced with Coal B runs except at bed temperatures of 700 °C. In addition to quartz‚ maghemite and hematite were seen as the dominant phases in the ash coating at 700 °C. Thenardite was not present in the ash samples of Coal B at all temperatures. This indicates that the higher ash build-up with Coal A may have been due to the presence of sulfur bearing mineral phases (anhydrite and thenardite) in the ash coating. 3.4.2. Fly ash samples. XRD analysis of fly ash samples from Coal B showed that the most abundant minerals in fly ash are quartz‚ maghemite and hematite (Table 3). Other minerals present in trace quantities include mullite‚ anhydrite‚ hematite and nosean. However‚ fly ash samples from Coal A indicated the presence of periclase and maghemite in addition to quartz‚ with periclase becoming more dominant phase with increasing bed temperatures.

4. CONCLUSIONS 1. Significant differences were seen in the characteristics of ash coating layers on

bed material used in fluid bed combustion of two coals tested. The amount of

316

H. B. Vuthaluru and D.-k. Zhang

inorganic matter in these coals were found to influence the nature of low-melting eutectics and their proportion in the ash coating. These in turn determined the agglomerating propensity of bed material.

2. Bed temperature was found to have a significant influence on the distribution of inorganic constituents into various streams viz. bed ash and cyclone ash. The distribution was found to depend on the proportions of inorganic constituents in the coal samples. Especially with Coal A‚ the bed temperature was found to influence the proportion of sodium in bed ash and cyclone ash. This was

expected to result in the formation sodium-rich deposits on heat exchanging

Role of Inorganics During Fluidised-Bed Combustion of Low-Rank Coals

317

surfaces with this coal‚ causing additional ash-related problems during fluidised

bed combustion. 3. The ash coating on bed material from Coal A showed the presence of low melting point compounds‚ and at higher bed temperatures‚ the proportions of low melting compounds increased‚ which led to decreased critical coating thickness at higher temperatures. Hence the bed defluidised after a short period of combustion of Coal A. The ash coating from Coal B‚ however‚ showed the presence of solid material‚ leading to reduced agglomerating propensity of the bed particles. This is expected to prolong the combustion operation with such coals with less ash-related problems.

ACKNOWLEDGEMENTS The authors gratefully acknowledge the financial and other support received for this research from the Cooperative Research Centre‚ New Technologies for Power Generation from Low-Rank Coal‚ which is established and supported under the Australian Government’s Cooperative Research Centres program.

REFERENCES Dawson‚ M.R. and Brown‚ R.C. (1992). “Bed material cohesion and loss of fluidization during fluidized bed combustion of midwestern coal.” Fuel‚ 71 585–592. Hodges‚ N.J. and Richards‚ D.G. (1989).“The fate of chlorine‚ sulphur‚ sodium‚ potassium‚ calcium and magnesium during the fluidized bed combustion of coal.” Fuel‚ 68 440–445.

H. B. Vuthaluru and D.-k. Zhang

318

Manzoori‚ A.R. and Agarwal‚ P.K.. (1992). “The fate of organically bound inorganic elements and sodium chloride during fluidized bed combustion of high sodium‚ high sulphur low rank coals.” Fuel‚ 71

513–522. Manzoori‚ A.R. and Agarwal‚ P.K. (1993). “The role of inorganic matter in coal in the formation of agglomerates in circulating fluid bed combustors.” Fuel‚ 72 (7) 1069–1075. Manzoori‚ A.R. and Agarwal‚ P.K. (1994). “Agglomeration and defluidization under simulated circulating fluidized-bed combustion conditions.” Fuel‚ 73 (4) 563–568. Raask‚ E. (1985). Mineral Impurities in Coal Combustion. Washington: Hemisphere Publishing Corporation.

Souto‚ M.A.‚ Rodriguez‚ J.C.‚ Conde-Pumpido‚ R.‚ Guitian‚ F.‚ Gonzalez‚ J.F. and Perez‚ J. (1996).“Formation of solid deposits in the gas circuit of a pressurized fluidized bed combustion plant.” Fuel‚ 75 (6) 675–680.

ROLE OF ADDITIVES IN CONTROLLING AGGLOMERATION AND DEFLUIDIZATION DURING FLUIDIZED BED COMBUSTION OF HIGH-SODIUM‚ HIGH-SULPHUR LOW-RANK COAL

Temi M. Linjewile* and Alan R. Manzoori** CRC New Technologies for Power Generation from Low-Rank Coal Thebarton Commerce and Research Precinct Adelaide‚ SA 5031‚ AUSTRALIA

1. INTRODUCTION South Australian lignites contain high levels of sodium and sulphur which combine during combustion to form sodium sulphate. Due to the low melting temperature of sodium sulphate and its formation of low-temperature eutectics with other ash species‚ the use of these coals in fluidized bed combustors causes fouling of cyclone and heat exchange surfaces‚ and deposition of ash on bed material particles leading to agglomeration and defluidization. The problem of growth in the size of bed material particles during fluidized bed combustion while burning Lochiel coal‚ a South Australian lignite has been reported by Manzoori [1990]. The presence of sodium‚ calcium and organically bound sulphur in this coal caused the formation of mostly sulphates of sodium and calcium. These formed low-melting eutectics which caused the deposition of ash on the

surface of inert bed particles. The molten phase present in the coating of bed particles facilitated sintering and formation of agglomerates eventually leading to defluidization of the bed. Dawson and Brown [1992] also reported problems of bed material ash deposition‚ agglomeration and defluidization occurring during circulating fluidized bed combustion of a high-sulphur Iowa coal. In pulverized coal fired plants‚ by comparison‚ problems of slagging and fouling spurred development of effective countermeasures for ash deposition involving mechanPresent address: * Department of Chemical and Petroleum Engineering‚ University of Wyoming P.O. Box 3295‚ Laramie‚ WY 82071-3295‚ USA‚ [email protected] ** CSIRO Minerals‚ Box 312‚ Clayton Sth Vic 3169‚ AUSTRALIA [email protected]

Impact of Mineral Impurities in Solid Fuel Combustion‚ edited by Gupta et al. Kluwer Academic / Plenum Publishers‚ New York‚ 1999.

319

320

T. M. Linjewile and A. R. Manzoori

ical procedures such as soot blowing with air‚ steam or water [Gorbaty et al.‚ 1983]‚ treatment of coal for addition or removal of specified elements in the coal structure [Vuthaluru et al.‚ 1990] and the use of additives [Raask‚ 1985; Barrett‚ 1989]. The introduction of additives‚ either with the coal or separately‚ offers the opportunity for controlling deposit formation without interruption of operations. As for control of bed material ash deposition‚ agglomeration and defluidization in fluidized bed combustion (FBC) systems‚ the use of additives does not feature prominently‚ presumably because lower operating temperature and the use of higher rank coal ruled out occurrence of bed material deposition and sintering. Clearly the situation changes dramatically when lowrank coals are considered. In recent times‚ developments in advanced coal-based power generation technologies such as pressurized fluidized bed combustion (PFBC)‚ integrated gasification combined cycle (IGCC) and direct coal fired turbine (CFT) present the need for clean-up of the hot flue gases for removal of alkali salts which may cause hot corrosion and ash deposition in gas turbines. This has increased research interest in the use of additives which may be classified into two general categories. The first approach involves the use of a bed of sorbent for removal of gaseous alkali salts [Lee and Johnson‚ 1980; Luthra and Blanc‚ 1984; Uberoi et al.‚ 1990]. The second approach entails in-situ use of additives for removal of alkali compounds [Logan et al.‚ 1990; Spiro et al.‚ 1990]. In each case kaolin

and alumina were used. Although additives have been shown to be effective in pulverized coal fired systems for deposit control and in advanced coal utilization systems for alkali vapour removal‚

mechanisms behind their effectiveness are not fully understood. This study is a report of the control of bed-material ash deposition—a primary step to agglomeration and defluidization—by the use of additives in a simulated circulating fluidized bed combustion (CFBC) unit.

2. EXPERIMENTAL The fluidization vessel‚ Fig. 1‚ consisted of a 77mm ID stainless steel tube with a conical section at its base connected to a 25mm ID gas inlet and solids discharge tube.

The vessel was fitted with an air preheater‚ coal screw feeder‚ additives feeder and a cyclone separator for collection of fines. The reactor is also fitted with thermocouples and differential pressure probes along its height. These were interfaced to a chart recorder for continuous monitoring. Below the conical section of the reactor a sealed bed material receptacle was provided for collection of spent bed particles. The bed material consisted of silica sand particles in the size range 0.85–1 .0mm. Coal samples obtained from the Lochiel coal deposit of South Australia were air-dried‚ crushed and sieved to a size range of 1.4–3.25 mm. Typical analysis of low mineral Lochiel coal is given in Table 1.

Role of Additives in Controlling Agglomeration and Defluidization

321

The additives investigated included dolomite‚ CW clay—a kaolinite and sillimanite-rich

clay‚ DV clay—a kaolinite and quartz-rich clay and gibbsite—a hydrated alumina referred to as (AH). The additives were screened prior to utilization to obtain particles of size less than 0.125mm. The chemical and mineralogical composition of the additives is presented in Table 2.

322

T. M. Linjewile and A. R. Manzoori

Fluidization air‚ at 140 lit/min at room temperature‚ from a compressor was allowed to flow through a honeycomb-design air preheater where its temperature was raised to 873 K prior to entry into the reactor. A l00g batch of bed material was fed into the fluidization vessel from an opening at the top end of the reactor. A known amount of coal was loaded into the coal feed hopper and the coal was fed continuously into the combustion zone by means of a water-cooled screw feeder at an average rate of 240g/h. Upon sighting of the first spark via the viewing mirror‚ the additives screw feeder was also started. An additive was introduced into the combustion zone by preheater air blowing past the discharge end of the additive screw feeder. For the purpose of comparison‚ some experiments were conducted without the introduction of additives‚ referred to as the no additive run. During combustion‚ the bed temperature was maintained at 1‚073K by adjusting the preheated air temperature. At this temperature‚ the average fluidization velocity in the combustion zone was 8 m/s. The fast fluidization velocity and the internal recirculation of the bed material particles provided conditions similar to CFBC systems. After three hours of combustion‚ the bed particles were discharged into an ash can at the bottom of the fluidization vessel. Coated bed particles retrieved from the combustor were weighed and some were subjected to wet chemical analysis‚ examination with a scanning electron microscope (SEM) equipped with an energy dispersive X-ray analyzer (EDAX)‚

electron microprobe analysis (EMPA) and X-ray diffraction analysis.

3. RESULTS AND DISCUSSION

3.1. Ash Deposition on Bed Material Particles The bed material recovered at the end of a combustion run without additives was weighed and the size distribution determined. The samples showed an ash coating buildup on the surface of the bed particles as shown in Fig. 2. Some agglomerates were also formed. In some combustion runs of long duration the formation of agglomerates led to defluidization of the bed. During combustion‚ inorganic matter introduced by the coal appears as an ash coating on bed material and fly ash in the cyclone. The partition of coal inorganic matter between the ash coating and cyclone ash was determined from the chemical analysis of the coal‚ its feed rate and the quantities of ash collected on bed

Role of Additives in Controlling Agglomeration and Defluidization

323

material and in the cyclone separator. It was found that 21–32 percent of the inorganic matter in the coal feed deposited on the surface of sand bed particles.

The composition of the inorganic constituents in the bed material coating was calculated on a silica—free basis in order to eliminate the contribution of the quartz bed particles. Figure 3 shows a comparison of the average composition of the bed material coating (bottom ash)‚ cyclone ash and coal ASTM ash. The results show that the coating on quartz particles is enriched in Na and S compounds. On account of the lower interfacial tension between the ash and silica bed particles compared to that between the ash and char surface‚ Manzoori [1990] suggested that the ash transfers to the surface of bed particles by collision between sand grains and the burning char particles. It may be noted that in comparison to the composition of feed coal ash‚ a large proportion of Ca and Mg compounds was not retained by the bed particles and escaped into the cyclone. Electron microprobe spot analyses and X-ray maps of polished cross-sections of coated sand grains showed the presence of a highly sulphated ash in the coating confirming observations previously made by Manzoori and Agarwal [1993]. XRD analysis of the coated sand grains obtained from no additive run presented in Table 3 shows the

324

prevalence of anhydrite‚ thenardite and

T. M. Linjewile and A. R. Manzoori

-sulphate. Thenardite and

-sulphate

are low-melting temperature compounds capable of forming low-temperature eutectics

with other ash compounds and hence are considered to be the basis of ash binder in the coatings. The influence of AH‚ CW clay‚ DV clay and dolomite on ash deposition presented as percent increase in mass of the bed material with time is shown in Fig. 4. The results show that dolomite and DV clay increase the rate of deposition of ash on bed particles‚ whereas AH and CW clay significantly reduce the formation of ash deposit on bed particles. The results presented in Figure 4 can be explained as follows: In the case of dolomite‚ the quantity of ash increased due to incorporation of dolomite in the ash but the stickiness of bed particles reduced due to dilution of melt. For DV clay deposition rate increased due to increased amount of melt (silicate) perhaps while on the char surface and the incorporation of clay particles in the ash. The molten phase solidifies at bed temperature and the ash coating on bed particles will have less melt (cement) and more aggregate — a scenario which minimizes formation of agglomerates. With AH no agglomeration occurred because deposition was prevented. CW clay fixed the sodium species present in ash into solid phases and hence less deposition and less tendency for agglomeration. Although dolomite is preferred for in-bed sulphur capture‚ it has been found to enhance the build-up on sand particles and proved to be the least effective of the additives investigated from the point of view of ash build up. These observations raise several questions regarding the nature of additive/ash interactions in the vigorously stirred environment of the CFBC. It is not discernible from the results whether the additive/ash interaction has occurred on the char surface or after deposition of ash on the bed material surface. It is also unclear whether the interaction of additives with ash is entirely due to chemical processes or a combination of physical and chemical processes. These questions are important in understanding additive/ash interactions and may assist in the selection and formulation of additives for a specific application problem.

3.2. Aluminium Hydrate Additive Combustion with aluminium hydrate in the CFBC prevented the formation of a significant ash coating on bed particles. The bed particles examined under scanning elec-

Role of Additives in Controlling Agglomeration and Defluidization

325

tron microscope (SEM) showed uneven ash coating as shown in Fig. 5. The coating thickness ranged from 3 to On higher magnification (not shown here) the morphology of the coating showed individual particles embedded in a molten matrix. An EDAX analysis of the coating revealed that aluminium was the most abundant element in the coating. Other major elemental species identified were sodium‚ magnesium‚ calcium and sulfur. These observations were confirmed by spot analysis at various locations of the ash coating with an electron microprobe analyzer (EMPA). Results for some arbitrarily selected regions of the coating are presented in Table 4. Observation of Table 4 shows that for AH run all the major ash species (Na‚ Mg‚ Ca‚ and S) are present in the coating. Chemical analysis of the bed particles‚ given in Fig. 6‚ shows a significant increase

in the content of aluminium in the ash coating with the resultant decrease in the concentration of ash forming constituents. These results also show that the proportion of sulphur in the coating is higher than that found in the clay run ash coatings. It may be noted that the results of XRD analysis‚ given in Table 3‚ verify that (thenardite) is still present in the ash coating. The mineralogical analyses do not provide any evidence to suggest that sodium in the coal has reacted with the aluminium hydrate to form sodium aluminate (an amorphous compound).

326

T. M. Linjewile and A. R. Manzoori

Sondreal et al. [1978] found no evidence of reaction between an aluminiumbased additive and sodium to form high-melting products that would tie up sodium in a harmless form during p.f. combustion of a Western U.S. lignite at 1‚366K. Lindner et al. [1988] also found that during p.f. combustion of Bowmans coal (South Australian lignite) at 1‚523K‚ AH was ineffective in fixing sodium to non-fouling aluminium compounds

when compared to kaolin. Clearly‚ more evidence is needed to determine the extremely successful role that aluminium hydrate plays in preventing ash deposition on bed particles. It is‚ however‚ suggested that the behaviour of AH additive is attributable to

the fact that heat-treated gibbsite transforms to activated alumina which has an internal surface area [Gerhartz‚ 1985] of 88m 2/g at the bed temperature considered. The porous alumina adsorbs the molten constituents of the ash into its structure upon collision with the ash possibly on the char surface where the molten constituents of the ash are less viscous.

3.3. Clay Additives 3.3.1. CW Clay. Combustion with CW clay resulted in the formation of a moderate deposit around the bed material particles. The coating formed was friable and broke off easily from the sand surface. However‚ it had sufficient integrity to remain mostly bonded on sand particles in the vigorously stirred combustion environment except for a few shell fragments which were collected in the bed material receptacle at the end of a run. SEM micrograph of a polished cross-section of a bed sand grain from a combustion run with CW clay additive‚ Fig. 7‚ showed a thick‚ nearly even coating ranging in thickness from 35 to 80µm. The coating appears to be porous‚ an indication of a sintering process. An EDAX analysis of the coating showed aluminium‚ silica and sodium to be the dominant species present in the coating. Further analysis of the coating with EMPA‚ Table 4‚ confirmed that the coating is rich in Al‚ Si and Na. Some trace quantities of calcium are also present. The analysis shows a conspicuous absence of sulfur in the coating. The chemical analysis of bed material‚ given in Fig. 6‚ shows a high proportion of clay-forming aluminium together with the inorganic constituents in the ash such as sodium and calcium sulphates. The higher proportion of sodium to sulphur suggests that sodium is present as compounds other than sulphates. The XRD analysis‚ given in Table

3‚ identified the presence of feldspar as a minor phase in the coating. This evidence‚ nevertheless‚ is inconclusive since trace amounts of feldspar are also present in the fresh additive. However‚ given the presence of kaolinite in the original clay additive‚ it is suggested

Role of Additives in Controlling Agglomeration and Defluidization

327

that kaolin has reacted with sodium in coal forming high melting temperature sodium aluminosilicates. This may occur via the following reaction pathways:

In addition‚ Wall et al. [1975] suggested that the sodium aluminosilicate products may combine to form eutectic mixtures melting at 1‚340K. Such reaction could only occur on the char surface where temperature is high enough for the reaction to be thermodynamically favoured. This is confirmed by observations from previous work on the effect of additives during combustion of South Australian lignites under pulverized fuel combustion conditions [Lindner et al.‚ 1988]. The emphasis in that study was placed on the effect of additives in controlling deposit formation on boiler tubes. In the high temperature environment of pulverized fuel firing‚ it was found that clay additives proved to be most effective by fixing gaseous sodium as sodium aluminosilicate‚ suggesting an additive mechanism of a chemical nature.

3.3.2. DV Clay. The addition of DV clay resulted in large deposits of strongly bonded ash on bed particles. Figure 8 shows an SEM micrograph of a polished crosssection of a bed sand grain from a combustion run with DV clay additive. The micrograph shows a layered coating of between 290 and 380 µm thickness. The coating appears to be porous and may have been formed by sintered particles. A qualitative analysis of the coating by EDAX showed that the coating is dominated by Al‚ Si‚ Na and Ca. This was proved by quantitative analysis with EMPA as shown in Table 4. The results are similar to those from the CW clay run in that there is an obvious lack of sulfur in the

coating. Chemical analysis of the ash coating‚ Fig. 6‚ shows relatively high concentration of sodium and very low levels of sulphur. Given the presence of quartz in DV clay‚ it is suggested that a portion of sodium has reacted with it forming low-melting sodium silicates. These sticky sodium silicates‚ possibly formed by reaction routes of Eq. (1) and (2)‚ could have been responsible for increasing the deposition of ash on bed particles. The presence

of kaolinite in DV clay may have resulted in the following reaction:

328

T. M. Linjewile and A. R. Manzoori

Raask [1986] observed that reaction (5) can take place at temperatures as low as 1‚085K. As in the case for CW clay‚ such reactions are favoured at temperature similar to that found at the char surface. The results of XRD analysis for DV clay run‚ Table 3‚ show the formation of a dominant amorphous phase and trace nepheline. The amorphous phase could be attributed to glassy sodium silicate which may have formed as a result of reaction between quartz present in clay and sodium in coal. Clays are known to react with sodium to form sodium aluminosilicates [Lindner et al.‚ 1988; Hodges and Richards‚ 1989; Lowe et al.‚ 1993]. However‚ the effect of clay properties and combustor operating parameters are not well established. It is apparent from the discussions on AH and clay additives that alumina in combination with silica is more potent in fixing sodium than alumina alone. Although DV clay is likely to convert sodium sulphate to higher melting compounds‚ it does not form a friable deposit and hence it is not as effective as CW clay in controlling bed material deposit formation. It may be noted‚ however‚ that insignificant quantities of agglomerates were collected in DV clay run but none in CW clay run.

3.4. Dolomite Additive Combustion of Lochiel coal in the presence of dolomite resulted in the formation of a thick coating on bed material particles. SEM micrograph of a polished cross-section of a bed material particle from a combustion run with dolomite additive‚ Fig. 9‚ showed

a coating thickness of between 170 and 330µm. The morphology of the coating reveals a combination of molten and crystalline matrix. EDAX analysis of the coating showed the prevalence of Ca‚ S‚ Mg and Na. Spot analysis of the coating with EMPA at various locations‚ Table 4‚ confirmed the presence of Ca‚ S‚ Mg and some regions rich in Na possibly as The presence of in the coating may facilitate formation of eutectic mixtures with other salts as discussed later. Chemical analysis of the bed material after combustion with the various additives is shown in Fig. 6 where in the dolomite run‚ the analysis shows a significant increase in the contents of Ca and Mg‚ indicating the incorporation of dolomite in the ash coating. Also significant is the higher proportion of sulphur in the coating in comparison to the other additives‚ indicating that the ash coating contains sulphated dolomite. XRD analysis confirmed that the ash coating in the dolomite run‚ Table 3‚ consists of anhydrite

Role of Additives in Controlling Agglomeration and Defluidization

329

thenardite and some of the species formed as a result of transformations of the inorganic matter in coal. There is no evidence to suggest that dolomite has in fact reacted with the ash. The presence of these salts may form sticky eutectic mixtures with low-melting points. It is suggested, therefore, that sulfated dolomite diluted the “cementing” material in the ash (sodium sulphate eutectics) resulting to a less sticky ash coating compared to undiluted ash and hence it had less tendency to agglomerate. Some agglomerates were formed during the dolomite run but were too few to initiate defluidization.

3.5. Effect of Additive Dosage The effect of additive feed rate on bed material ash deposition for the clays and AH additive is shown in Fig. 10 where the lines represent best-fit curves to the data. DV clay data shows that the build-up rate on bed material particles increases with feed rate. Results for aluminium hydrate additive show a sharp decline in the rate of build-up with increase in additive feed rate. Results for CW clay show a gradual decline in build-up rate with increase in additive feed rate. Increasing the DV clay feed rate increases the quantity of aggregate resulting in an increase in the rate of deposition. In the case of those

T. M. Linjewile and A. R. Manzoori

330

additives removing the melt (cement) from the system, i.e. AH and CW clay, the more the additive the more of the cement was taken out and hence the rate of deposition decreased despite the fact that more aggregate was available. The dosages reported here reflect on the characteristics of a once-through system where additive/ash interaction depends on successful collisions between additive particles and the ash melt on char particles. In a commercial CFB plant where a significant amount of additive is recycled the ratio of additive to coal is expected to be significantly lower.

3.6. Proposed Additive Mechanism The sequence of events during combustion with additives may thus be summarized as follows: When an additive is introduced into the combustor it undergoes transformation and subsequently the additive particles collide with char particles during which the additive becomes incorporated in the molten ash on the char surface. In the case of AH and dolomite incorporation is a physical dilution. AH however soaks the molten species and hence reduces the available melt whereas dolomite simply dilutes the melt without removing it. In the case of the clays additive/ash reactions occur after incorporation. For DV clay the nature of the melt may change but the increase in the rate of build-up is attributed to the dilution effect. In the case of CW clay reactions between the additive and sodium species in ash produce a solid phase and hence reduce the amount of melt available to cause stickiness. The physical shape of sillimanite (needle-like grains) in this clay may also contribute in reducing the stickiness by covering the sticky ash.

CONCLUSION Four additives have been tested to determine their effectiveness in controlling ash deposition, agglomeration and defluidization under circulating fluidized bed combustor conditions. Gibbsite was found to be most effective. The evidence obtained from the various analytical techniques suggests that depending on the type of additive used, interference in the process of ash deposition may be due to either a chemical reaction or physical dilution. These interactions apparently occur at the surface of burning char particles. Although previous investigations have identified clays to be effective in sodium capture, care must be taken in selection of clay additives since the mineralogical composition of the clays appears to play a major role in the extent of deposit formation control. Gibbsite and the clays have demonstrated potential for in-situ sodium capture which may facilitate use of low-rank coals in advanced FBC-based power generation plants.

ACKNOWLEDGEMENTS The authors gratefully acknowledge the financial and other support received for this research from the Cooperative Research Centre for New Technologies for Power Generation from Low-Rank Coal, which is established and supported under the Australian Government’s Cooperative Research Centres program. The support given by ETSA Corporation and the Department of Chemical Engineering, University of Adelaide is also acknowledged. The assistance provided by John Terlet and Hugh Rosser of the Centre for Electron Microscopy and Microstructure Analysis (CEMMSA) in SEM and EMPA

Role of Additives in Controlling Agglomeration and Defluidization

331

analyses is greatly appreciated. TML would like to thank A. Kosminski for helpful discussions; D. Linard of ETSA Corporation for assistance in preparation of samples for electron microprobe and XRD analysis; Ashley Hull for assisting with the digitization of the SEM micrographs and Dr P. K. Agarwal and the National Science Foundation— Grant number OSR-9550477 for extending facilities for preparation of this paper at the

University of Wyoming.

REFERENCES Barrett, R.E. (1989). “Slagging and Fouling Due to Impurities in Combustion Gases.” Engineering Foundation, New York, 593–672. Dawson, M.R. and Brown, R.C. (1992). “Bed Material Cohesion and Loss of Fluidization During Fluidized Bed Combustion of Midwestern Coal.” Fuel, 71, 585–592. Gerhartz, W. (1985). Ullmann’s Encyclopedia of Industrial Chemistry. VCH Weinheim, Federal Republic of Germany. Vol A1, 561–562.

Gorbaty, M.L., Larsen, J.W. and Wender, I. (1983). Coal Science. Academic Press, New York, Vol 2, 2–64. Hodges, N.J. and Richards, D.G. (1989). “The Fate of Chlorine, Sulphur, Sodium, Potassium, Calcium and

Magnesium During the Fluidized Bed Combustion of Coal.” Fuel, 68, 440–445. Lee, S.H.D. and Johnson, I. (1980). “Removal of Gaseous Alkali Metal Compounds from Hot Flue Gas by Particulate Sorbents.” J. Engng Power , 102, 397–402. Lindner, E.R., Kosminski, A., Taylor, C. and Williams, R.G. (1988). “Effects of Additives on Fouling Behaviour Characteristics of South Australian Brown Coals.”, Australian Coal Science Conference, University of Adelaide 16–18 May 1988, B2:l 1.1–B2:11.14. Logan, R.G., Richards, G.A., Meyer, C.T. and Anderson, R.J. (1990). “A Study of Techniques for Reducing Ash Deposition in Coal-Fired Gas Turbines.” Prog. Energy Combust. Sci. 16, 221–233. Lowe, A.J., McCaffrey, D.J.A. and Richards, D.G. (1993). “An Investigation into the Effectiveness of Fireside

Fuel Additives.” Fuel Processing Tech, 36, 47–53. Luthra, K.L. and LeBlanc, O.H. (1984). “Adsorption of NaCl and KCl on Chem., 88, 1896–1901.

at 800–900°C.” J. Phys.

Manzoori, A.R. (1990). Role of the Inorganic Matter in Agglomeration and Defluidisation During the CFBC

of Low-Rank Coals. PhD Thesis, The University of Adelaide. Manzoori, A.R. and Agarwal, P.K. (1993). “The Role of Inorganic Matter in Coal in the Formation of Agglomerates in Circulating Fluid Bed Combustors.” Fuel, 72, 1069. Raask, E. (1985). Mineral Impurities in Coal Combustion. Hemisphere Publishing Corporation. 293–312. Raask, E. (1986). “Flame Vitrification and Sintering Characteristics of Silicate Ash.” ACS Symposium Series,

301, 138–155. Sondreal, E.A., Gronhovd, G.H., Tufte, PH. and Beckering, W. (1978). “Ash Fouling Studies of Low-Rank Western U.S. Coals.” In Ash Deposits and Corrosion Due to Impurities in Combustion Gases. R.W.Bryers (Ed.), Washington, Hemisphere Publishing Corporation, 85–111.

Spiro, C.L., Chen, C.C., Kimura, S.G., Lavigne, R.G. and Schields, P.W. (1990). “Deposit Remediation in CoalFired Gas Turbines through the Use of Additives.” Prog. Energy Combust. Sci. 1990, 16, 213–220. Uberoi, M., Punjak, W.A. and Shadman F. (1990). “The Kinetics and Mechanism of Alkali Removal from

Flue Gases by Solid Sorbents.” Prog. Energy Combust. Sci., 16, 2 0 5 – 211. Vuthaluru, H.B., Gupta, R.P., Wall, T.F. and Domazetis, G. (1990). “Minimising Ash Fouling from High

Sodium Brown Coals.” Volume 4: Laboratory Experiments to Examine the Fouling Characteristics of Loy Yang Coal. SECV Report No ND/90/058, (1990), Herman Research Laboratories, Melbourne Victoria. Wall, C.J., Graves, J.T. and Elliot, J.R. (1975). “How to Burn Salty Sludges.” Chem Engng., 77–82.

This page intentionally left blank.

THE AGGLOMERATION IN THE FLUIDIZED BED BOILER DURING THE CO-COMBUSTION OF BIOMASS WITH PEAT

Ritva E. A. Heikkinen, Mika E. Virtanen, H. Tapio Patrikainen, and Risto S. Laitinen University of Oulu, Department of Chemistry Linnanmaa, FIN-90570 Oulu, Finland

1. INTRODUCTION The use of peat for energy production in Finland dates back to early 1970’s [Asplund, 1996]. In recent years the use of biomass (wood chip, bark, sawdust, agricultural and municipal waste, etc.) as fuel has become increasingly popular because of the

environmental benefits associated with these fuels. Presently, the peat combustion is responsible for over 6%, and the biomass combustion for over 15% of the energy production in Finland [Energy Statistics, 1996]. The co-combustion of biomass with peat in fluidized bed boilers starts to be rather common in population centers. Boilers of this type are best suited for fuel with a low energy value because the increased efficiency is beneficial for the community heat distribution. While this biomass fuel contains little or no sulfur, it may contain high amounts of alkali and alkaline-earth metals. As Finnish peat often contains large amounts of iron either as limonite or in the organic matrix, the partial melting of the ash may take place during the fluidized bed co-combustion. This results in the agglomeration and eventual defluidization of the bed material, thus causing severe problems in boiler functionality leading to the boiler shutdown and significant economical losses. Bed agglomeration and defluidization can, however, also be prevented by intelligent

fuel mixing [Nordin et al., 1995]. When coal was co-combusted with Lucerne or Olive flesh in a pilot scale FBC, no signs of agglomeration were observed whereas combustion tests of biomass alone resulted in the rapid agglomeration and defluidization of the bed. The chemical equilibrium calculations and the SEM-EDS analysis indicated that the agglomeration was caused by the formation of sticky or partially molten coating of ash on the bed particles. The coating consisted of salts in the case of Lucerne and silicates in the case of Olive fuel. Dawson et al. [Dawson et al., 1992] have reported that the formation of an aluminosilicate cohesive material that wetted the surface of bed particles by viscous flow is the Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

333

334

R. E. A. Heikkinen et al.

primary cause for the bonding of bed material. The presence of iron in the bonding material indicated that it was present in the melting reactions and may have acted as a flux. The formation of sticky bed material coatings and agglomerates during the fluidized bed combustion of coal has been investigated by SEM-EDS [Manzoori et al., 1993]. The results suggested the formation of a highly sulfated ash coating on the surface of bed particles. In the case of agglomerates, it was evident that they were formed as a result of sintering of the ash coating on bed particles. Therefore it was suggested that agglomeration occurs as a result of random collision between bed particles coated with ash. We have applied SEM-EDS, combined with Automatic Image Analysis (AIA) in the prediction of the slagging properties of peat ash during the pulverized fuel combustion [Heikkinen et al., 1997]. This work has been extended for the evaluation of the factors affecting the agglomeration tendency of fluidized bed particles. Special attention was directed to the analysis of bed material coatings. We have explored samples collected from an actual power plant boiler during an extensive test series. Both fly ash samples and coated bed material samples, as well as agglomerated bed samples were investigated.

2. EXPERIMENTAL 2.1. Boiler Tests The Toppila power plant of Oulu Energy Board has two peat-fired units. The older boiler Toppila II is a circulating fluidized bed boiler and has a total capacity of 310 MW and the new Toppila I CFB boiler has a capacity of 260MW. The total peat consumption of both units is about 3.000GWh/year with a full load, and the total consumption of biomass is about 300GWh/year [Kantola, 1997]. An extensive series of tests was carried out in the Toppila II boiler. The peat fuel used in the tests was a mixture of different peat types from several bogs in Northern Finland. The biomass co-combusted with peat was a mixture of saw dust and wood chip. The combustion of this peat-wood mixture was expected to be unproblematic. During the tests the fuel samples were taken from the inlet line. The fly ash samples were collected in the electrostatic precipitator and the bed material samples were collected from the bottom outlet line. The agglomerated bed samples were not collected during the test, but were obtained earlier when there were disturbances in the normal operation of the boiler. In addition to the agglomerates the bed material samples were also collected and investigated.

2.2. Scanning Electron Microscopy All SEM samples were mounted with epoxy resin (Struers Epofix, 15:2) into a

brass-ring. The mounted samples were ground with 200, 800 and 1,000 mesh SiC powder in glycerol or Butane-1,2-diol and polished using the 3 and 1 µm diamond paste. The samples were coated with a thin carbon layer to eliminate electrostatic effects. The SEM-analysis was carried out using a Jeol 6,400 scanning electron microscope equipped with a Link ISIS energy dispersive X-ray microanalyser that was used for the determination of sodium, magnesium, aluminum, silicon, sulfur, phosphorus, potassium, calcium, titanium, and iron. The acceleration voltage was 15keV and the current 120*

The sample distance was 15mm and the magnification 250× or 400×. During two hours of instrument time about 1,000 particles were analyzed for every sample from

approximately ten different image fields. The image analysis was performed with an IMQuant software contained in the Link ISIS.

The Agglomeration in the Fluidized Bed Boiler During the Co-combustion of Biomass with Peat

335

Data collected from SEM-EDS is visualized by utilizing appropriate quasi-ternary distribution diagrams [Virtanen et al., 1997]. The composite corners of these diagrams are defined in terms of the total relative content of selected elements, which can be chosen freely within the ten elements. The vertical axis in the diagram indicates the relative number of particles with a given composition. The particle is counted in the diagram if the total content of all the appropriate elements exceeds 80%. The number of classified particles is indicated in each diagram shown below.

2.3. The Image Processing The purpose of image processing is to exclude the bed particles (sand originated from peat bogs) from the image and to direct the EDS-analysis to the coatings and adhesive ingredients in the agglomerates. An example of the image processing is given in Fig. 1. During the first step of processing the back scattered electron image (BEI) collected from the scanning electron microscope is smoothened by strengthening of contrast and by removing roughness. The second step of processing is based on a gray scale shading. Bed material particles are excluded from the image and only coatings or adhesive ingredients remain for the final step. Several binary operations are performed to disconnect, open, erode, etc. the image in order to create special areas, where the ZAF corrected point analysis of ten elements (Na, Mg, Al, Si, P, S, K, Ca, Ti and Fe) are directed.

3. RESULTS The conditions of operation and the composition of fuel mixtures during the test series are specified in Table 1. The SEM-EDS data of the analyzed bed material coatings, visualized as quasi-ternary diagrams, are seen in Fig. 2. In Fig. 2 the total content of silicon and aluminum is chosen for one corner, the total content of alkali and alkaline-earth metals for the other, and the content of iron for the third corner. The vertical axis indicates the relative number of artificial particles

336

R. E. A. Heikkinen et al.

The Agglomeration in the Fluidized Bed Boiler During the Co-combustion of Biomass with Peat

337

with a given composition. Typically, over 90% of observed particles are classified with this method. The sample 1 in Fig. 2 is bed material. Coating, that was expected to be nonagglomerative, was formed during the combustion of peat. Boiler power was at 100% of capacity. It can easily been seen from the diagram that the composition of coating in the

bed material is mainly aluminosilicates of iron, as well as of alkali and alkaline-earth metals. Sample 2 has been collected at a lower level of boiler capacity. The particle distribution is more dispersed. During the co-combustion of peat and saw dust (samples 3 and 4), the distribution is shifted towards alkali and alkaline-earth metal silicates. When the fuel mixture consisted of 55% of wood chips, the content of alkali metal aluminosilicates increases and that of iron decreases (samples 5 and 6). Despite the differences in the fuel composition and power levels, the operation was unproblematic throughout the test series. Fig. 3 shows the compositional distribution of two samples (7 and 8) collected during the operation failure of the boiler. The samples 7a–7c are taken from the bed

338

R. E. A. Heikkinen et al.

particle coatings and the samples 8a–8c from the adhesive ingredient in the agglomerate. The differences in the compositional distribution are easily seen. Peat was used as fuel at the time of operation failure through agglomeration. The particle distribution in 7a is similar to that in the samples 1–4 (see Fig. 2). The composition of the agglomerate (8a) differs from that of the coating (7c). The adhesive ingredient mainly consists of alkali and alkaline-earth silicates or aluminosilicates. A closer investigation of particle distribution in samples 7b and 8b by separating silicon and aluminum in different corners of the quasi-ternary diagrams, leads to a conclusion that both the bed material coatings and the adhesive ingredient of agglomerate consists of aluminosilicates, rather than silicates. A small amount of pure quartz, a remainder from the image processing can also be observed. While iron plays an important role in the formation of the bed material

coatings, 7c, it is less prominent in the case of agglomerate. This is due to the melting of bed material, which increases the amounts of alkali and alkaline-earth metals in the adhesive ingredient (8c). The role of sulfur and the formation of sulfates, is shown in Fig. 4. The formation

The Agglomeration in the Fluidized Bed Boiler During the Co-combustion of Biomass with Peat

339

of sodium and calcium sulfates is obvious, when the amount of wood was raised to 55% of the fuel mixture. Sulfates were not found from the bed material coatings (c and d), but were observed in fly ash (a and b). While no sulfate formation was observed in the case of the adhesive ingredient of the agglomerate there are signs that sulfur plays a minor role in aluminosilicates (e).

4. CONCLUSIONS The formation of bed material coatings during the co-combustion of peat and biomass is caused by iron, calcium, aluminum and silicon. No signs of sodium or calcium

sulfates were observed in bed material samples. Sulfates were observed in fly ash samples, when the amount of wood was 55% of the fuel mixture. Thus the bed material agglomeration during peat and biomass co-combustion is due to the partial melting of aluminosilicates, rather than the formation of low melting salts. Iron is found from the bed material coatings and may act as a flux in the melting processes of the silicates. When the agglomeration progresses, the coated sand particles are molten on the surface as seen from the increased amounts of potassium, sodium and calcium. The role of iron is not so significant in the adhesive material.

5. ACKNOWLEDGEMENTS Financial support from the Ministry of Trade and Industry (LIEKKI-2 Research

Program), is gratefully acknowledged.

REFERENCES Asplund, D. (1996). “Energy Use of Peat. “In H. Vasander (Ed.), Peatlands In Finland. Helsinki: Finnish Peatland Society. Dawson. M.R., and Brown, R.C. (1992). “Bed Material Cohesion and Loss of Fluidization During Fluidized Bed Combustion of Midwestern Coal”. FUEL, 71, 585–592. Energy Statistics of Finland (1996). Energy Review 211997. Helsinki, Statistics Finland. Heikkinen, R., Laitinen, R.S., Patrikainen, T., Tiainen, M., and Virtanen, M. ( I n press). “Slagging Tendency of Peat Ash”. Fuel Processing Technology.

Kantola, R. (1997). Oulu Energy Board, Personal Communication. Manzoori, A.R., and Agarwal, P.K. (1993). “The Role of Inorganic Matter In Coal.” FUEL, 72, 1069–1075. Nordin, A., Öhman M., Skrifvars, B.-J., and Hupa, M. (1995). “Agglomeration and Defluidization in FBC of Biomass Fuels—Mechanisms and Measures for Prevention”. In L. Baxter and R. DeSollar (Eds.), Application of Advanced Technology to Ash-Related Problems in Boilers. New York, Engineering Foundation.

Virtanen, M., .Heikkinen, R., Patrikainen, T., Laitinen, R.S., Skrifvars, B.-J., and Hupa, M. (1997). “A Novel Application of CCSEM for Studying Agglomeration in Fluidized Bed Combustion”. Engineering Foundation Conference on The Impact of Mineral Impurities in Solid Fuel Combustion, November 2–7, 1997,

Kona, Hawaii.

This page intentionally left blank.

ASH FUSION AND DEPOSIT FORMATION AT STRAW FIRED BOILERS Lone A. Hansen 1 , Flemming J. Frandsen 1 , Henning S. Sørensen2, Per Rosenberg2, Klaus Hjuler 3 , and Kim Dam-Johansen 1 1

Department of Chemical Engineering Technical University of Denmark DK-2800 Lyngby, Denmark 2 Geological Survey of Denmark and Greenland (GEUS) Thoravej 8, DK-2400 Kbh. NV, Denmark 3 dk-TEKNIK Energy & Environment Gladsaxe Moellevej 15, DK-2860 Soeborg, Denmark

1. INTRODUCTION Much attention has been drawn towards the burning of biomass for power production in recent years. The Danish Government has agreed to a 20% reduction in emissions from power production by the year 2005 with reference to year 1988. Since biomass is considered neutral, the substitution of coal with biomass as fuel for power production is one of the initiatives taken to meet this goal. In Denmark, straw is the most frequently used biomass type, since surpluses of straw frequently appear. The use of straw for heat production in small scale furnaces, e.g. at individual farms, has been practiced for a number of years. However, power generation from biomass is a fairly new task. Straw fired boilers have generally experienced serious problems with slagging and fouling [Miles et al., 1995; Michelsen et al., 1996; Stenholm et al., 1996; Jensen et al., 1997] primarily due to the large content of troublesome deposit initiating elements such as potassium and chlorine in the straw. The straw contains sufficient quantities of volatile fluxing alkali to significantly lower the fusion temperature of the ash, so that it partly melts during combustion and increases the tendency of the ash to stick to heat transfer surfaces. Alternatively, the potassium and chlorine species may vaporize and subsequently condense on boiler tubes and refractory surfaces, creating sticky surfaces that accelerate deposit build-up and/or work as “glue”, thereby increasing deposit strength. This paper presents results from two new techniques for characterisation of biomass ash fusion and relates the obtained results to compositional CCSEM data and measurements of deposit formation rates as found in two biomass fired boilers. The aim of this work has been to gain a more detailed understanding of the chemistry and fusion of fly and bottom ashes in biomass fired boilers and the mechanisms that lead to deposit formation. Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwcr Academic / Plenum Publishers, New York, 1999.

341

342

L. A. Hansen et al.

2. HASLEV AND SLAGELSE COMBINED HEAT AND POWER PLANTS Deposition measurements have been carried out at two straw fired boilers, the 23 Haslev Combined Heat and Power Plant (CHP) equipped with four cigar burners and the stoker fired Slagelse Combined Heat and Power Plant. Details about the measurements are reported by Stenholm et al. [1996] and Jensen et al. [1997]. The boilers and deposit sampling positions are shown schematically in Fig. 1. At the Haslev CHP, furnace temperatures of 720–820 °C were measured, and deposit samples were collected at two locations: in the top of the furnace and at the entrance to the third pass, just before the superheaters. At the Slagelse CHP, furnace temperatures of 745–910°C were prevailing, and deposit measurements were carried out in the middle of the furnace and in the third pass, the probe being located between the primary and the secondary superheater (Fig. 1). During the test period at Slagelse CHP, a number of different straws were burned in order to investigate the influence of fuel composition on the deposit formation rate. The chemical composition of these five types of straw and the single straw burned at Haslev CHP are listed in Tables 1 and 2. As seen in the tables the numbers showing the largest deviation are the ash contents (varying between 3.5 and 7.1%), and the contents of and Cl (0.06–0.86) in the ashes. In the SLA8 test, rape straw was burned, which explains the markedly different ash content and chemistry compared to the other tests, in which wheat and barley straw was used. Samples of fly ash, bottom ash and deposits were collected during the test period, and the chemical composition of these are shown in Tables 3a and 3b. Since also the association of elements in the fly and bottom ashes is of crucial importance for the ash fusion, the chemical species present in the ashes were identified by means of Computer Controlled Scanning Electron Microscopy, CCSEM. The mineral categories relevant for biomass ashes are developed and described by Sørensen [1997]. From the data, it was seen that fly ashes were characterised by a high content of KC1, a minor content (below 5% (w/w)) of other potassium and calcium salts (chlorides and sulphates), and a relatively low content of potassium and calcium silicates. The bottom ashes were characterised by a very high content of silicon containing species, including quartz, potassium silicates, calcium silicates and small quantities of alumina silicates. Most of

Ash Fusion and Deposit Formation at Straw Fired Boilers

343

the deposits examined were found to contain large quantities of KC1 and varying quantities of potassium and calcium silicates, the ratio between silicates and KC1 varying for different deposit locations in the boiler. The compositional differences between the various ashes and deposits will be discussed later.

3. DETERMINATION OF ASH SOFTENING The fusion of ashes is generally of great importance for the ability of these ashes to form problematic deposits on heat transfer surfaces in boilers [Moza and Austin, 1981; Srinivasachar et al., 1990; Walsh et al., 1990; Benson et al., 1993; Richards et al., 1993]. In this study, the fusion of ashes and deposits collected was investigated in order to find a relation between ash fusibility and deposit formation observed. The fusion of ashes and deposits was quantified experimentally using two different approaches: High Temperature Light Microscopy and Simultaneous Thermal Analysis.

3.1. Experimental Techniques Using High Temperature Light Microscopy, HTLM, the quantification of melt in ash is based on the change in light transparency of the ash as it is melted. Shortly described the method works as follows: A sample of ash is placed in the high temperature light microscope. During heat up, the area covered by the sample is quantified at short intervals using advanced image analysis. As part of the ash is melted, this part becomes transparent, thereby not contributing to the determined area any more. In this way, the area at any given temperature, T, divided by the area at the initial temperature, is used as an estimate of the solid fraction of the ash at the temperature

T. The method is described in detail by Hjuler [1997].

344

L. A. Hansen et al.

Using Simultaneous Thermal Analysis, STA, for fusion determination implies continuous measurement of sample weight (Thermogravimetric Analysis, TGA) and sample temperature (Differential Scanning Calorimetry, DSC) during heat up of the ash. The weight measurement reveals any mass changes taking place in the sample. By comparing the sample temperature to the temperature of an inert reference material, any heat producing or heat consuming processes (chemical or physical) occurring in the sample is detected, and the energy involved subsequently quantified. On the resultant STA curves, melting is detected as an endothermic process involving no change in mass. The conversion of STA curves to melting curves (showing the fraction of melt in the ash as a function of temperature) can be carried out as follows: Initially, the energies related to other processes than melting (typically evaporation) have to be subtracted from the DSC curve to obtain a DSC curve showing only melting energies. The area of a given melting peak in the resultant DSC curve, corresponds to an absolute quantity of energy used for melting. The position of the peak (onset and peak temperature) gives an indication of the identity of the melting substance(-s), which means that a reasonable estimation of the relevant melting enthalpy can be given. Based on these two numbers, the mass of material melted in the given temperature interval can be calculated. By relating this melted mass to the total mass of ash analysed, the mass fraction of ash melted in the given temperature interval is obtained. If the identification of discrete melting peaks is not possible, the quantification of melt formation is carried out based on a direct comparison of areas below the DSC curve, i.e. energies used for melting in different temper-

Ash Fusion and Deposit Formation at Straw Fired Boilers

345

ature ranges. A complete description of the method is given by Hansen [1997] and Hansen et al. [1997].

3.2. Comparison of Techniques To illustrate the connection between the two techniques applied for ash fusion quantification, results from respectively HTLM and STA are presented in Fig. 2, showing

fusion of a fly ash (from exp. 3 at Slagelse CHP), a bottom ash (from exp. 8 at Slagelse) and a deposit (SH outer deposit from Haslev). For the fly ash and the deposit the figure shows that the two techniques do not agree on melting onset: the HTLM detects the first

346

L . A. Hansen et al.

melting at temperatures 50–150 °C lower than the STA. The initial melt formation detected by the HTLM method is believed, though, to be caused by physical rearrangement of the ash, and not necessarily melt formation. At the temperature, where the STA detects the first melting to occur, the HTLM detects 10–15% melt. These are typical numbers/figures. Generally, the methods agree nicely at the temperatures for which significant melting occurs: both the slopes of the curves and the amounts of melt obtained agree. At higher temperatures the STA technique tend to detect slightly higher fractions of melt than the HTLM technique, which could be due to some of the ash species (e.g. some high melting silicates) forming melts of dark colour, which may not be light transparent. For the bottom ash, the agreement on the melting onset is better, and the quantity of melt detected by the HTLM at the onset temperature determined by the STA is only 0–5%. For half of the bottom ashes investigated, the two techniques agree well at higher temperatures (as for the ash presented), whereas for the others, the STA detects significantly higher degrees of melt than the HTLM. As explained above this may be due to some (silicate) species not being light transparent even when melted. Another feature illustrated is the occasional decrease in melt fraction observed with the HTLM method. This often appears after significant melt formation, and is due to physical rearrangement of the solid ash in the liquid phase.

In conclusion, the two techniques are found to agree qualitatively, and for about half of the ashes, results from the two methods do not deviate more than 10% (absolute) at temperatures between 650°C and 1,100°C. Taking into account experimental uncertainty and the difference in the principles applied, the two techniques are judged to agree reasonably well.

4. RESULTS In this section, the ash fusion is evaluated based on the chemical composition and the origin of the ashes. In order to limit the number of fusion curves, only curves produced by the STA method will be shown, and comments will be given on any different

trends found from the HTLM method.

4.1. Straw, Ashes and Deposits from Haslev The CCSEM compositional data for the ashes is illustrated more clearly in a triangular diagram of 1) quartz and alumina silicates, 2) K and Ca silicates and 3) KC1 (Fig. 3). The three end-members in the diagram were chosen to represent: 1) relatively refractory quartz and alumina silicates, 2) more easily fusible potassium and calcium

dominated silicates (probably formed by reaction of evaporated potassium and calcium with straw derived silica or quartz grains), and 3) low-melting KC1 formed during the combustion process [Sørensen, 1997]. Added up the three end-members account for 63 to 99% of the samples presented, the sum of which is indicated beside each point. The compositional differences between the bottom ash, the fly ash and the deposits are evident in the triangular diagram (Fig. 3), with the deposits located close to the KC1 apex, fly ash occupying an intermediate position and bottom ash located close to the “silicate line”. It is seen that the compositions of the two outer superheater deposits are very similar, as might be expected since both are bulk sintered deposits collected at each end of the same deposit probe. Similarly, the two outer furnace deposits are quite alike. The

Ash Fusion and Deposit Formation at Straw Fired Boilers

347

inner deposit from the furnace probe is depleted in K and Ca silicates compared to the two outer parts, indicating that the initial furnace deposit is formed primarily from condensation of KC1, whereas for the bulk sintered deposit part, which contains a larger fraction of silicates, impaction of fly ash particles has played a more dominant role. For the bulk sintered part of the super heater deposit, the KC1 content is higher than for the bulk sintered furnace deposit, indicating that impaction of silicate rich fly ash particles is a more important deposit forming mechanism in the furnace compared to the super

heater area. An alternative explanation for the higher KC1 content in the outer super heater deposit may be that the lower temperatures in the convective pass may have caused some KC1 to have condensated heterogcneously on the fly ash that impacts. The fusion of fly ash, bottom ash and deposits collected at Haslev CHP is shown in Fig. 4 as determined by the STA method. It is seen that for all ashes, the melting occurs in two temperature ranges: 1) between 630 °C and 750° and 2) above 1,000°C. The low temperature range is believed to be caused by melting of an eutectic mixture of potassium and calcium salts (chlorides and sulphates, the exact composition of which is not known), while the high temperature range is believed to include melting of the various silicates. The fusion curves clearly reflect the compositional differences between the ashes: the bulk sintered super heater deposit and the initial furnace deposit containing the largest quantities of KC1, form larger quantities of melt in the low temperature range than the bulk sintered furnace deposit. Likewise, the fly ash shows lower melt fractions

348

L. A. Hansen et al.

than the deposits, but higher melt fractions than the bottom ash. which does not melt at all in the low temperature range. This means that the melting behaviour of the ashes correlate with the position in the ternary diagram, i.e. moving (in composition) away from the KC1 apex results in smaller melt fractions obtained in the low temperature range.

4.2. Straw, Ashes and Deposits from Slagelse 3 For fly ash, bottom ash and the three deposits from the Slagelse 3 experiment, the chemical composition as determined by CCSEM is illustrated in a triangular diagram

(Fig. 5). Also here is seen a dominance of KCl in the deposits compared to the bottom ash, as was the case for the Haslev samples. However, the fly ash from Slagelse 3 does not extinguish itself from the deposits in respect of KCl content, as did the Haslev fly ash. It is seen that the compositions of outer furnace and superheater deposits are quite alike, as these are two sets of bulk sintered deposits, from the first and second half of the superheater and the furnace deposit probe, respectively. For the deposits collected in the furnace, the initial layer is seen to contain a moderate quantity of KCl, whereas the bulk sintered layer contains only a small fraction (between 2.5 and 18.1%). For the deposit collected in the superheater area, both inner and outer layers contain large quantities of

KCl, and it actually seems that the KCl content of one of the bulk sintered deposit layers is a little higher than that of the initial layer. This is quite surprising, but is probably a consequence of the high KCl content of the impacting fly ash particles. These findings thus support the trends seen for the Haslev samples: The content of silicates is higher in the bulk sintered layer than in the initial layer of any deposit, meaning that condensation and thermophoresis of KCl is of highest importance for formation of the initial deposit layer, whereas impaction of fly ash particles is more important for formation of the bulk sintered deposit layer. The fusion curves for fly ash, bottom ash and deposits collected at Slagelse 3 is shown in Fig. 6 as determined by the STA method. Again, the melting is found primarily to occur in a low-temperature range between 600 °C and 750° due to melting of eutectic salt mixtures and a high-temperature range above 900 °C covering melting of silicates. These fusion curves also clearly reflect the compositional differences between the ashes:

Ash Fusion and Deposit Formation at Straw Fired Boilers

349

The bottom ash that contains no KC1 shows no melting in the low temperature range, whereas the bulk sintered furnace deposit containing 3% KC1 forms a small amount of melt below 900°C; the inner furnace deposit, which contains moderate quantities of KC1 shows an intermediate fraction of melt at 800 °C, and the fly ash and both superheater deposit layers, which have the largest KC1 contents consequently show the largest melt fractions below 800 °C. As for the Haslev samples, the position in the triangular diagram thus illustrates the relative fusibility of the ashes.

4.3. Slagelse Fly and Bottom Ashes The CCSEM compositional data for the Slagelse fly and bottom ashes are illustrated in a triangular diagram (Fig. 7) identical to the ones shown for the Haslev and Slagelse 3 experiments. The difference in mineral composition between bottom ashes and fly ashes is obvious: the bottom ashes are dominated by K and Ca silicates, quartz and alumina silicates, whereas the fly ashes are dominated by KC1 and contain varying fractions of silicates. KC1 and silicate often appear as composite grains which are found mainly in two different forms: they are either formed by condensation of KC1 on silicate particles, or they represent KC1 and silicate particles being located so closely that they

350

L. A. Hansen et al.

cannot be distinguished by the SEM. Observations in the SEM backscatter images underline that the presence of condensed KC1 on silicate particles is a common feature in the biomass fly ashes. In the Slagelse 8 experiment rape is used as fuel instead of wheat or barley as in

the other experiments. The difference in ash composition is especially evident for the fly ash, which contains large amounts of and Ca rich categories in accordance with the Ca and S rich character of rape relative to wheat and barley. This is the reason that only 11% of the fly ash composition is represented in the diagram. For the remaining ashes, the strong compositional difference between bottom ashes and fly ashes is evident: the bottom ashes being located at or close to the silicate line, whereas the fly ashes stretch from the KC1 apex towards a composition of approximately 85% “quartz and alumino silicates” and 15% “K and Ca silicates”. This ratio (of “quartz and alumino silicates” to “K and Ca silicates”) is higher for the fly ashes than for any of the bottom ashes. Fusion curves for the five fly ashes as determined by STA are shown in Fig. 8. It is seen that all fly ashes show distinct melt formation in the temperature range from 600 to 720 °C, except for the fly ash from experiment 8. The fly ashes from experiment 3 and 6 reach complete melting first (just below 1,200°C). The fly ashes from experiment 7 and 1 are almost, but not completely melted at 1,250°C, whereas the fly ash from experiment 8 showed significantly lower fusion at 1,250 °C. Fusion curves for the five bottom ashes are shown in Fig. 9 as determined by HTLM. All bottom ashes start melting at around 800°C. From 900°C to 1,000°C the melting process is very slow, and above 1,000°C significant melting occurs in bottom ashes from experiment 3, 6 and 8. Bottom ashes from experiment 1 and 7 do not show any sign of additional melting above 1,000°C. To test the usefulness of the diagram in Fig. 7 as a basis for classification into more or less fusible ashes, melt fractions extracted from the melting curves are assigned to the data points. The numbers in square frames beside the data points represent melt fractions at 880 °C. A consistent trend is evident for the bottom ashes towards lower values of “melt fraction at 880°C” with an increasing content of “quartz and alumino silicates”. This is in accordance with the more refractory nature of quartz and most alumino silicates compared to K and Ca silicates. For the fly ashes a trend is seen of lower melt fractions for a given temperature (data labels show melt fractions at 800 °C, at which temperature the salts are assumed to be melted) with an increase in silicate content relative to KC1 content. This is in accordance with the low melting point of KC1 compared

Ash Fusion and Deposit Formation at Straw Fired Boilers

351

to most silicates. In conclusion, it is found that the proposed triangular diagram may be useful to classify the relative fusibility of biomass ashes, and the figure furthermore illustrates the logical relation between the chemical composition of the ashes and their fusion characteristics.

4.4. Comparison to Standard AFT For all fly and bottom ashes, the fusion was additionally characterised by the standard ash fusion test (ISO540, 1981). This test implies heating up of a cube of ash and noting the temperatures corresponding to 1) rounding of the cube corners, at the Initial Deformation Temperature, IDT, 2) the shape change from cubic to hemispherical, at the

Hemispherical Temperature, HT, and 3) flowing of the ash, at the Fluid Temperature, FT. Table 4 shows the melting onset (given by difference, to the IDT) as determined by the STA and the fractions of melt present in the various ashes at the IDT, HT and

FT as determined by HTLM and STA, respectively. It is seen that for the fly ashes, the melt formation starts at temperatures well below the IDT, ranging from 298°C below for the fly ash from Haslev to 50°C below for the fly ash from experiment 3 at Slagelse. For

the bottom ashes, melt formation is detected between 0 and 95 °C below the IDT, excepting the bottom ash from experiment 1 at Slagelse, for which melt is detected 79°C above the IDT. For the fly ashes, substantial quantities of melt have been formed at the IDT,

averages of 45.8% and 52.1% melt are found by the HTLM and the STA, respectively. However, the melt fractions detected for the bottom ashes at the IDT is much lower, respectively 15.5% and 3.6% on average. The reason for the STA to find no melt at the IDT for two silicate rich bottom ashes might be that the rounding of the corners has been due to sintering of the ash instead of melting. At the Hemispherical Temperatures melt fractions are generally between 35 and 70% melt, the bottom ashes generally showing smaller melt fractions than the fly ashes, and the STA generally detecting slightly higher melt fractions than the HTLM. At the fluid temperature melt fractions measured generally lie between 60 and 80%, the averages found to 74.5 and 73.2% melt. These findings support previous criticism [Huffman et al., 1981; Huggins et al., 1981; Coin et al., 1996; Wall et al., 1996] stating that the IDT does not indicate the melting

352

L. A. Hansen et al.

onset. The IDT has been known to indicate the temperature, at which ash particles in an

operating furnace have cooled sufficiently to have only a slight tendency to stick together [Bott, 1991], but with melt fractions of 30 to 60% the stickiness of the ash may be somewhat higher than only “slight”, meaning that keeping furnace exit gas temperatures below the IDT does not guarantee avoidance of deposit problems in the boiler convective pass.

5. SIMPLE DEPOSIT MODELING The quantity of melt in an fly ash is known to increase its tendency for adhering to surfaces after impaction [Srinivasachar et al., 1990; Walsh et al., 1990; Benson et al., 1993; Richards et al., 1993]. In this study simple calculations have been made to estimate the deposit formation rate as it would occur if only condensation and inertial impaction of the collected fly ashes were important. This means that thermophoresis has not been taken into consideration. In the model, the build up of deposit is assumed to be initiated by condensation of gaseous KC1. After a layer of 20µm thickness has formed, the condensation process is assumed to end, and the deposition is continued by impaction of ash particles.

The ash deposit formation rate, calculated as [Frandsen, 1997]:

caused by inertial impaction has been

where is the bulk gas velocity, is the mass concentration of fly ash particles in the flue gas, is the weight fraction of fly ash particles that impacts the tube, and is the fraction of those particles impacting the tube that sticks to it. In the equation, the product represents the total flux of fly ash particles in the gas. For calculating the deposition flux, the total flux is corrected for the facts that not all particles impact the tube (multiplication by and not all impacting particles adheres to the tube The mass concentration of fly ash particles

Ash Fusion and Deposit Formation at Straw Fired Boilers

353

in the flue gas is estimated based on the mass flow of fly ash compared to the mass flow of fuel ash introduced into the furnace, and the prevailing air excess number. The velocity of the gases have been measured to 3.1–3.6 m/s in the furnace and 9.1–10.3 m/s in the convective pass. Estimation of the impaction coefficient, may be carried out based on the Stokes number, Stk, for potential flow around a tubular cylinder [Wessel and Righi, 1988], using the following relationship [Israel and Rosner, 1983]:

where e = Stk – a, and the coefficients applied are those given by Israel and Rosner [1983]. The sticking coefficient of the fly ash particles has been estimated based on the presented STA ash fusion curves. The sticking coefficient was estimated to simply equal the fraction of melt in the fly ash at the measured gas temperature, after having corrected the melt fraction for the fraction of melted ash that has been evaporated at the gas temperature. The sticking coefficient of the tube is estimated to zero, since melt only occurs at temperatures above 610 °C, and the deposit surface structure did not show any sign of melting. Based on these figures the ash formation rates presented in Fig. 10 were found. As seen in the figure, the model overestimates the deposition rate for all experiments. The ratio between calculated rates and measured rates varies between 1.4 and 11.2 except for one extreme case, where the model overestimates the deposition rate by a factor of 47.2. The average ratio is 10.3 including all experiments, and 6.6 excluding experiment SLA3, SH. Regarding the deposition measurements, it is known that deposit typically detaches the probe, both during experiment, but particularly during probe withdrawal. This means that the mass of “loose” deposit measured is probably lower than the “actual” value. On the other hand there is no doubt, that the model in addition overestimates the deposition rates. Several reasons for this can be given: — the applied model only accounts for deposit build up, that is, no deposit removal mechanisms are taken into account. In reality, erosion of the deposit due to the

354

L. A. Hansen et al.

impaction of dry fly ash particles may be severe. This causes unrealistically high deposition rates. — taking the melt fraction directly as the sticking probability is of course a rough simplification. On its route through the convective pass of the boiler, the fly ash particles may be enriched in condensible alkali species. Since the condensible species lower the melting point, probably the fly ash collected in the particulate removal device and thus the one that has been used as basis for calculations, show larger melt fractions at low temperatures than the ash present in the flue gas passing the deposit probes. This causes that the sticking probability used is too high. — the rate of condensation in the convective pass is probably overestimated, since the concentration of gaseous KCl has been estimated as if all sampled aerosols were present in the gas phase. Taking the prevailing temperatures into account, a large fraction of the KCl has probably already homogeneously condensed before the gas enters the superheater probe location. The KCl is therefore probably present as aerosols which will be transported to the probe surfaces by thermophoresis instead of gas diffusion. The rate of condensation may not equal the rate of thermophoresis, and the rate of condensation may therefore be overestimated.

— finally, the modeling of the condensation phenomena is greatly simplified. A layer of equal thickness present for all experiment conditions is not credible, and since the condensation process plays a large role in purely straw fired boilers, this part of the model should be refined to obtain an overall model in

better agreement with the measured data. In Fig. 10, the measured deposition rates are shown as a function of the calculated values. There is a clear correlation between measured and calculated deposition rates, when excluding two data points, the one from SLA8,SH (206, 120) and the one from SLA3,SH (783, 17). The reason for these data points to distinguish themselves from the others is not quite clear yet. For the rest of the data points, the obtained linear correlation means that even though the model does not quantitatively agree with the measurements, there is a clear qualitative agreement, indicating that the model may be useful for ranking the deposit formation propensities of one fly ash compared to an other from the same kind of boiler. Taking the great simplicity of the model into account, this is not a bad result.

6. CONCLUSIONS The fusion of fly ashes, bottom ashes and deposits collected in the furnace and the convective pass of two straw fired boilers have been quantified. CCSEM of the ashes revealed that the fly ashes and deposits contained high quantities of salts, especially KCl, and varying quantities of potassium and calcium silicates, whereas the bottom ashes consisted of a mixture of quartz, potassium, calcium and alumina silicates. The compositional differences of the ashes were clearly reflected in the fusion, as the salt containing fly ashes and deposits showed significant melting in the temperature range from 600 °C to 750 °C, whereas the melt formation in the bottom ashes primarily occurred at temperatures between 1,000°C and 1,200°C. Furthermore, it was generally found that increasing contents of salts in the ashes implied a larger fraction of the ash fusion to occur below 800 °C.

Ash Fusion and Deposit Formation at Straw Fired Boilers

355

Comparison between results from the standard ash fusion test and the newly developed techniques revealed that melt formation was initiated at temperatures much lower than the IDT, the temperature difference being in average 142 °C below for fly ashes and 30 °C below for bottom ashes. Melt quantities of approximately 40–50% were formed in the fly ashes at the initial deformation temperature, which means that probably the tendency for the ashes to form deposits can not be neglected at the IDT. For the bottom ashes the melt fraction at the IDT was lower, around 0–15%. The agreement between the standard ash fusion test and the new techniques is thus worse for salt rich compared to silicate rich ashes. The chemical composition of the deposits as determined by CCSEM indicated that the formation of the initial part of the deposits was dominated by condensation of KCl, whereas inertial impaction of fly ash particles played a higher role in formation of the bulk sintered layer on top. Furthermore inertial impaction was indicated to play a higher role for the deposits in the furnace compared to deposits in the convective pass. A very simple model set up to estimate the deposition rates was found not to agree quantitatively with the deposit formation rates measured, since deposit rates were overestimated by a factor of 1.4 to 11.2. Never the less, a linear correlation between calculated and measured deposition rates was found, meaning that the simple model was capable of ranking deposition rates from the different fly ashes.

ACKNOWLEDGEMENTS This work has been financially supported mainly by the Danish Energy Research Programme (EFP), and partly by the Combustion and Harmful Emission Control (CHEC) research programme, ELSAM (the Jutland-Funen Electricity Consortium) and

ELKRAFT. Peter Arendt Jensen and Michael Stenholm are acknowledged for giving the project participants access to the collected ash samples.

REFERENCES Benson, S.A., Jones, M.L., and Harb, J.N. (1993). “Ash Formation and Deposition”. In L..D. Smoot (Eds.), Fundamentals of Coal Combustion for Clean and Efficient Use. New York: Elsevier

Bott, (1991). “The Assessment of Fouling and Slagging Propensity in Combustion Systems”. In S.A. Benson (Eds.), Inorganic Transformations and Ash Deposition during Combustion. New York: ASME Coin, C.D.A., Kahraman, H., and Peifenstein, A.P. (1996). “An Improved Ash Fusion Test”. In L.L. Baxter and R. DeSollar (Eds.), Proceedings of the Engineering Foundation Conference. New York and London: Plenum Press Couch, G. (1994). Understanding slagging and fouling in pf combustion. IEA Coal Research, Report No. IEACR/72 Frandsen, F.J. (1997). Estimation of Ash Deposition Fluxes in Utility Boilers. Internal Report, Department of Chemical Engineering, Technical University of Denmark Hansen, L.A. (1997). “Melting and Sintering of Ashes”. Ph.D. Thesis, Department of Chemical Engineering, Technical University of Denmark Hansen, L.A., Frandsen, F.J., and Dam-Johansen, K. (1997). “Ash Fusion Quantification by Means of Thermal Analysis”. These conference proceedings Hjuler, K. (1997). “Ash Fusibility Detection Using Image Analysis”. These conference proceedings Huffman, G.P., Huggins, F.E., and Dunmyre, G.R. (1981). “Investigation of the High-Temperature Behaviour of Coal Ash in Reducing and Oxidizing Atmospheres”. Fuel (60) 585 Huggins, F.E., Kosmack, D.A., and Huffman, G.P (1981). “Correlation between Ash-Fusion Temperatures and Ternary Equilibrium Phase Diagrams”. Fuel (60) 577–584 ISO 540 (1981). “Determination of Fusibility of Ash”. International Standard Organisation

356

L. A. Hansen et al.

Israel, R. and Rosner, D.E. (1983). “Use of a Generalized Stokes Number to Determine the Aerodynamic Capture Efficiency of Non-Stokesian Particles from a Compressible Gas Flow”. Aerosol Science and Technology (2) 45–51 Jensen, PA., Stenholm, M., and Hald, P. (1997). “Deposition Investigation in Straw Fired Boilers”. Energy and Fuels 11 (5) 1048–1055 Michelsen, H.P., Larsen, O.H., Frandsen, F, and Dam-Johansen, K. (1996). “Deposition and High Temperature Corrosion in a 10 MW Straw Fired Boiler”. Fuel Processing Technology (54) 95–108 Miles, T.R., Miles, T.R.Jr, Baxter, L.L., Bryers, R.W., Jenkins, B.M., Oden, L.L. (1994). “Alkali Deposits Found in Biomass Power Plants. A Preliminary Investigation of Their Extent and Nature”. National Renewable Energy Laboratory, 1617 Cole Boulevard, Golden, Colorado Moza, A.K. and Austin, L.G. (1981). “Studies on Slag Deposit Formation in Pulverized Coal Combustors. 1. Results on the wetting and adherence of synthetic coal ash drops on steel”. Fuel (60) 1057–1064 Richards, G.H., Slater, P.N., and Harb, J.N. (1993). “Simulation of Ash Deposit Growth in a Pulverised Coal-

Fired Pilot Scale Reactor”. Energy & Fuels (7) 774–781 Srinivasachar, S., Helble, J.J., Katz, C.B., and Boni, A.A. (1990). “Transformations and Stickiness of minerals during pulverised coal combustion”. In R.W. Bryers and K. Vorres (Eds.), Proc. of the Engineering Foundation Conference on Mineral Matter and Ash Deposition from Coal. New York: Engineering

Foundation Stenholm, M., Jensen, P.A., and Hald, P. (1996). “Fuel and Firing Characteristics of Biomass—Combustion Trials”. An EFP-93 Project (in Danish), Journal No. 1323/93-0015, The Danish Energy Research Programme Sørensen, H.S. (1997). “Computer Controlled Scanning Electron Microscopy (CCSEM) of Straw Ash”. These

conference proceedings Wall, T.F., Creelman, R.A., Gupta, R.P., Coin, C.. and Lowe, A. (1996). “Coal Ash Fusion Temperatures: New Characterisation Techniques and Associations with Phase Equilibria”. In L.L. Baxter and R. DeSollar (Eds.), Proceedings of the Engineering Foundation Conference on Applications of Advanced Technology to Ash-Related Problems in Boilers. New York: Plenum Press Walsh, P.M., Sayre, A.N., Loehden, D.O., Monroe, L.S., Beér, J.M., and Sarofim, A.F. (1990). “Deposition of Bituminous Coal Ash on an Isolated Heat Exchanger Tube: Effects of Coal Properties on Deposit Growth”. Prog. Energy Comb. Sci. ( 1 6 ) 327–346 Wessel, R.A. and Righi, J. (1988). “Generalized Correlations for Inertial Impaction of Particles on a Circular Cylinder”. Aerosol Science and Technology (9) 29–60

INFLUENCE OF METAL SURFACE TEMPERATURE AND COAL QUALITY ON ASH DEPOSITION IN PC-FIRED BOILERS

Karin Laursen, Flemming J. Frandsen, and Ole Hede Larsen Geological Survey of Denmark and Greenland Thoravej 8, 2400 Copenhagen NW, Denmark Phone: +45 3814 2000 Fax: +45 3814 2050 Technical University of Denmark Building 229, 2700 Lyngby, Denmark

Phone: +45 4525 2883 Fax: +45 4588 2258 Faelleskemikerne, Funen Power Station Havnegade 120, 5000 Odense C, Denmark Phone: +45 6590 4444 Fax: +45 6590 3812

1. INTRODUCTION During coal combustion, the inorganic constituents are transformed into fly ash particles that may deposit on the heat transfer surfaces of a boiler. Flame-side deposits cause many problems that lead to operational problems such as reduction in the heat transfer and possible reduced output of the boiler. Additionally, some types of deposits increase the corrosion of the heat transfer surfaces. Under normal operational conditions ash deposits are removed regularly from the boiler either “naturally” by shedding or

mechanically by soot-blowers. However, types or quantities of ash deposits may develop that cannot be removed with these simple methods and in severe cases, boilers may be damaged by the collapse of large blocks of ash deposits and it may be necessary to temporarily shut down the boiler for cleaning and repair. In order to increase the general knowledge of ash deposition in boilers, a three year collaborative project on “Mineral transformations and ash deposition in pulverized coal fired systems” was initiated in Denmark in 1993 (Larsen et al., 1996). The participants in the project included: 1) ELSAM (project coordinator and manager of full-scale combustion trials); 2) the Geological Survey of Denmark and Greenland (coal, ash, and deposit analyses); and 3) the Combustion and Harmful Emission Control (CHEC) research program at the Technical University of Denmark (modeling of ash deposition propensities). The project focused on existing boilers (steam parameters up to 250 bar/540 °C), the new ultra super critical (USC) boilers (steam parameters 290 bar/580 °C) under construction in western Denmark, and the next generation pf-boilers (325 bar/620 °C– Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

357

358

K. Laursen et al.

700°C). Many of the existing boilers and all the new boilers are equipped with lowburners. In general, the purposes of these full-scale trials were to gain fundamental knowledge of ash deposit formation and elucidate factors controlling the growth and consolidation of deposits. A major specific objective of the trials was to evaluate the influence of

increasing steam temperatures on the morphology and chemistry of ash deposits, but also to evaluate the influence of load and general operation of the plant on the ash deposition.

2. EXPERIMENTAL A total of six full-scale trials were conducted: one at the Ensted, three at the Funen and two at the Vendsyssel power stations (Laursen, 1997). The Ensted Power Station, Unit 3, is a 600 pc-fired boiler with 36 lowcombination burners situated in 3 levels in a front and rear wall configuration. Unit 7 at the Funen Power Station is a 350 pc-fired unit with 16 tangential arranged lowburners situated in 4 levels. The Vendsyssel Power Station, Unit 2, is a 300 pc-fired unit with 16 retrofitted lowburners situated in 4 levels in a front and rear wall configuration. Boiler profiles for the

three boilers are illustrated in Fig. 1 and Table 1 shows major plant and operational data for the boilers. All the trials included in this project were carried out in collaboration with other EESAM activities (i.e., coal test burns and research and development projects). The coals burned during the tests were mainly selected because of their general combustion behavior rather than for their ash deposition propensities, however, the coal types covered the spectrum and a major fraction (on an energy basis) of the coals consumed in power stations in Denmark. The coal burned during the trial at the Ensted Power Station was an Indonesian

sub-bituminous coal (Coal A). The trials at the Funen Power Station were conducted during combustion of three bituminous coals: 1) a Colombian coal (Coal B); 2) a South African coal (Coal C); and 3) a Polish coal (Coal D). The trials at the Vendsyssel Power Station were conducted during combustion of two coal types / blends: 1) 67% Coal D— 33% high-S coal blend (Coal E) from the USA; and 2) 100% Coal E. The sub-bituminous Coal A had a higher moisture content (15.5%) than the five bituminous coals (9.6–12.6%)(Table 2). Except from Coal E, all the coals were low-sulfur coals. Coal A and Coal E had a slightly higher content of volatile matter compared with the other coals. Coal A had a low ash content (6.5%), whereas the other coals had medium ash contents. The heating value for Coal E (28.2MJ/kg) was slightly higher than for the other five coal types (25.16–25.88 MJ/kg). The contents of the two main components in the ash, and were similar for all coals except Coal C; a higher amount of and, Coal B; a higher amount of (Table 3). The and contents indicated that Coal C contained more Alsilicates or Al-silicates with higher contents compared with the other coals, while Coal B had a higher quartz content. These mineral contents agreed well with results of CCSEM analyses of the coals (Laursen, 1997). The content varied significantly for the coals, from very low in Coal C (3.15%) to high in Coal E (16.6%). Coal A contained a relatively higher content of and a lower content of compared to the other coals which indicated that the iron in Coal A is not only present as pyrite but also as iron-oxides or iron-carbonates (e.g., siderite

Influence of Metal Surface Temperature and Coal Quaulity on Ash Deposition in PC-Fired Boilers

and ankerite

359

Manual SEM-EDX investigations of

Coal A revealed that both siderite and ankerite were common in this coal. This Indonesian coal apparently had mineral characteristics similar to Australian coals, with siderite and ankerite as the most common iron containing minerals, and pyrite only occurring as a minor phase (Couch, 1994). Based on the ash composition all coals except from Coal E are classified as low slagging coal. Coal E considered to be a medium slagging coal (Laursen, 1997). Analyses of ash samples from the electrostatic precipitator and the furnace bottom are shown in Table 4. As expected, the bottom ash was enriched with Fe2O3, especially for the iron-rich coal, Coal E. In general, all the bottom ashes were enriched in SiO2 (except for Coal C) and CaO, and depleted in (except for Coal A) compared to the fly ashes. The trials at Ensted and Vendsyssel power stations lasted 3 and 2 days, respectively, during which full-load was maintained. Trials at the Funen Power Station lasted three

360

K. Laursen et al.

days for each of the three coal types. The operation of the plant and the excess air was constant on a given day for each coal type, but varied from Day 1 to Day 3 of the test as follows: Day 1, (1.17% excess air, normal coal load); Day 2, (1.15% excess air, normal coal load); and Day 3, (1.17% excess air, and higher coal load plus lower excess air on the lowest burner level).

Two types of probes were used for collecting deposits from the boilers: an air cooled probe for exposure in the convective pass and a water-air cooled probe for exposure in areas of the boiler subjected to radiation (Laursen, 1997). The locations of the probes are indicated on the boiler profiles in Fig. 1. During the trials at Ensted

and Vendsyssel power stations, the probe metal temperatures (furnace: 450–650 °C, convective pass: 500–700 °C) and exposure time (2–32 h.) were varied in order to

Influence of Metal Surface Temperature and Coal Quaulity on Ash Deposition in PC-Fired Boilers

361

362

K. Laursen et al.

evaluate the effects of the two parameters on the morphology, texture and chemistry of the deposits. During the trials at the Funen Power Station the effects of boiler operation were evaluated and the metal temperature for all the probes was held at 570 °C. Additionally, at Funen Power Station, deposits were also collected on an un-cooled ceramic probe inserted adjacent to Probe 2. The deposits from the un-cooled probes simulated a deposit that was not influenced by cooling from the tubes and indicated the appearance and chemistry of a thick deposit in the consolidation phase. In addition, during the trials at the Vendsyssel Power Station, an in-furnace video recording system was used for recording ash deposition on the probe from the convective pass and in the furnace in general. The probes were generally covered with a loose powder deposit on the downstream side (shelter side) and to a lesser extend on the sides. These loose powder deposits were often blown off the probes when the probes were retracted from the boiler due to air pressure from the probe and the pressure from the boiler. In addition, some of the powder was lost during dismantling the test tube. Thus, no test element contained a complete, intact powder deposit that indicated the magnitude of the deposition rate on the downstream side. Hardly any deposits were visible on the upstream (windward side) of the test elements from the Funen and Ensted power stations, except for a thin, black scale and

some small irregularities (islands). The upstream side of the test elements from the

Vendsyssel Power Station were normally covered with a hard-bonded, rough deposit. Video recordings at the Vendsyssel Power Station taken shortly before the probe was removed showed that most of the deposits on the upstream side of the probe were lost during retraction. These deposits were thick (approximately 3–4cm) especially during combustion of Coal E. Thus, the deposits collected on the probes only represented the hard-bonded deposits. However, they were a significant part of a deposit since they represented the parts that were difficult or impossible to remove with soot-blowers, whereas the loose deposits were easily removable. Based on scanning electron microscope analyses, the deposits collected on the probes during the full-scale trials could be classified into five main textural types: 1) a porous deposit; 2) a powder deposit; 3) an iron-rich deposit; 4) a semi-fused slag; and 5) a fused slag (Laursen, 1997; Laursen et al., 1997). Table 5 summarizes the deposit types collected on the cooled metal and un-cooled ceramic probes during combustion of the six coals. These five textural types have both textural and chemical characteristics and have been used for interpretation of the mechanisms controlling the deposition of each of the deposit types (Laursen, 1997; Laursen et al., 1997).

363

Influence of Metal Surface Temperature and Coal Quaulity on Ash Deposition in PC-Fired Boilers

3. INFLUENCE OF OPERATION AND METAL TEMPERATURES ON PROBE DEPOSITS In general, the influence of coal type, load, probe temperature and excess air on the thickness of the iron-rich deposits collected on the probes at the Ensted and Funen power stations were limited. As previously discussed, most of the powder deposits were lost as the probes were retracted from the boilers. Thus, was not possible to provide quantitative information on the influence of coal types, probe temperatures, and load on the deposition rates for powder deposits. The influence of the probe temperature on the thickness of the porous deposits was also minor. No variation in the chemical composition of the deposits due to changes in the metal surface temperatures or boiler operation were observed for the powder deposits and the iron-rich deposits. However, a systematic variation in the chemical composition of the porous deposits due to increasing metal surface temperatures was found. Bulk chemical compositions calculated from the SEMPC analyses show that an increase in the probe temperature led to a systematic increase in the ratio (Table 6). This increase in the ratio is the only chemical changes observed on the probes caused by an increase in the probe temperature. The increase in the ratio was also apparent on the ternary diagrams for probes with different metal temperature. However, the increase was even more noticeable on histograms illustrating the mass fraction of particles with a specified iron content (Fig. 2). These histograms were constructed by determining the total iron content in strips cut from the ternary diagrams along lines parallel to the

binary.

The change in the chemical composition of the deposits towards a higher content indicated that the change from preferential deposition (i.e., porous deposit) to bulk deposition (i.e., semi-fused slag and fused slag) was faster when the metal temperature was increased. The faster change towards bulk deposition suggested that the deposition rate would increase with higher metal temperatures, thus leading to accelerated deposition problems.

364

K. Laursen et al.

Video recordings indicated that porous deposits usually shed easily. The question is whether this change in the chemical composition of the deposit changed the

strength of the bonding between the initial deposit (i.e., porous deposits) and the bulk deposit (i.e., semi-fused slag and fused slag). In that case, increasing ash deposition was likely to occur. However, if the strength between the two deposit types was

unchanged, the ash deposit formation was likely to remain the same, as on the low temperature (500°C) metal surfaces. Further research is needed to clarify the influence of metal temperature on the strength of the bonding between the initial deposits and the bulk ash deposits.

Influence of Metal Surface Temperature and Coal Quaulity on Ash Deposition in PC-Fired Boilers

365

4. SUMMARY AND CONCLUSIONS The six full-scale trials carried out at the Funen, Ensted and Vendsyssel power stations provided valuable information on the types and chemistry of deposits formed on heat transfer surfaces in pulverized coal-fired boilers. Based on the appearance (texture) of the deposits collected on probes during the trials, five distinct textural deposit types have been defined: 1) porous deposit; 2) powder deposit; 3) iron-rich deposit; 4) semifused slag; and 5) fused slag. These five textural deposit types have characteristic chemical compositions and their formation is controlled by various mechanisms. In-furnace video recordings of deposition probes at the Vendsyssel Power Station revealed that the majority of the collected deposits were lost during retraction of the probes and even more was lost during dismantling of the test elements: Thus, it has not been possible to collect an undisturbed, intact deposit. Deposition rates could therefore only be evaluated for hard bonded deposits. However, deposition rates for the hardbonded deposits and dependence of the rates on boiler operation and probe temperatures must also be viewed cautiously. The loose bonded deposits were of minor importance for the operation of boilers compared to the hard bonded deposits because the loose bonded deposits were easily removed by soot-blowers. The thickness of the iron-rich deposits collected at the Ensted Power Station increased slightly with increasing probe temperature, whereas the same type of deposits collected at the Funen Power Station showed no consistent changes in the deposition rate with coal load and excess air changes. Porous deposits collected at the entrance of the convective pass at the Vendsyssel Power Station increased with increasing probe temperature, whereas deposits from the

furnace showed ambiguous results. Deposits collected at both positions showed a systematically increase in the ratio with increasing probe temperature.

This increase indicates that increasing probe temperature may result in a faster build-up to bulk-deposition and thus increased slagging tendencies. In general, the trials indicated that combustion of a coal that leads to minor ash deposition problems in an existing boiler may create major problems in the same boiler if the steam temperature and thereby the metal surface temperatures is increased. However, combustion of a non-problematic coal in an existing boiler may not lead to any problems related to increasing steam temperature. Thus, these trials indicated that increasing steam temperature may restrict the types of coals that can be burned without problems in such boilers, but further research is needed to quantify the

influence.

ACKNOWLEDGMENTS This project was funded by ELSAM (The Jutland-Funen Power Consortium, Denmark) and the Danish Research Academy. The three power stations: Ensted, Funen and Nordjylland-Vendsyssel are acknowledged for allowing and for helpful support during the full-scale trials. The Energy and Environmental Research Center, University of North Dakota, is acknowledged for support with development of SEM-EDX techniques at GEUS. This work has been carried out partly within the Combustion and Harmful Emission Control (CHEC) Research Program at the Department of Chemical Engineering, Technical University of Denmark. The CHEC Research Program is cofunded by Elsam, Elkraft, the Danish Technical Research Council, the Danish and the

366

K. Laursen et al.

Nordic Energy Research Programs and the European Union Joule and Thermie Research Programs.

REFERENCES 1 . Larsen, O.H., Laursen, K. and Frandsen, F. (1996) Danish collaborative project on ash deposition in PF-fired boilers. In L.L. Baxter and R. DeSollar (Eds.), Applications of advanced technology to ash-related problems in boilers. Plenum Press. 2. Laursen, K. (1997) Characterization of minerals in coal and interpretations of ash formation and deposition in pulverized coal fired boilers. Ph.D.Thesis. Geological Survey of Denmark and Greenland. Report

1997/65. ISBN 87-7881-022-7. 3. Couch, G. (1994) Understanding slagging and fouling during pf combustion. IEA Coal Research Report No. IEACR/72, London. 4. Laursen, K. and Frandsen, F.J. (1997) Classification system for ash deposits based on SEM analyses. Paper presented at the Engineering Foundation conference in Kona, November, 1997.

FULL SCALE DEPOSITION TRIALS AT 150 MWE PF-BOILER CO-FIRING COAL AND STRAW Summary of Results

Karin H. Andersen1, Flemming J. Frandsen2, Peter F. B. Hansen1, and Kim Dam-Johansen2 1

MIDTKRAFT Energy Company, Studstrup Power Station DK-8541 Skoedstrup, Denmark Phone: +45 86 99 17 00, Fax: +45 86 99 37 20 E-mail: [email protected] or [email protected] 2 Dept. of Chemical Engineering, Technical University of Denmark DK-2800 Lyngby, Denmark Phone: +45 45 25 28 83, Fax: +45 45 88 22 58

E-mail: [email protected] or [email protected]

1. INTRODUCTION The Danish Government has committed the Danish power companies to burn between 1 and 1.2 million tonnes of straw per year by the year 2000. This is part of the national act to reduce the emissions to 80% of the 1988 level, in year 2005. As part of the ELSAM (The Jutland-Funen Electricity Consortium) biomass strategy, a conventional 150 PC-fired boiler at the Danish energy company I/S Midtkraft has been converted for coal-straw co-combustion. A two-year demonstration programme was initiated in January 1996, and will be concluded in 1998. The main purpose of the programme is to evaluate the economical and technical feasibility of this concept compared to other technologies. The demonstration programme includes experiments to demonstrate the plant operational data and operating costs when co-firing coal and straw, as well as experiments to clarify the fuel and process chemistry. This includes analysis of fuels, fly ash and flue gas, deposition propensities, and high temperature (HT) corrosion. Also a number of in-situ gas, temperature and aerosol measurements were performed to establish the chemical environment in the boiler during the deposition and corrosion experiments. For further details on the demonstration programme as a whole is referred to Hansen et al. [1996] and Overgaard and Hansen [1997]. Straw-fired boilers have generally experienced serious problems due to slagging and Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

367

368

K. H. Andersen et al.

fouling [Grunge, 1989; Miles et al., 1995; Henriksen and Hansen, 1995; Baxter et al., 1996; Stenholm et al., 1996], and also high temperature (HT) corrosion problems have been extensive, with the steam temperatures usually kept relatively low (<500 °C) to limit corrosion damages [Michelsen et al., 1996; Nielsen, 1996]. However, the co-firing of straw with coal, at low straw shares, in some cases seems to have a beneficial effect on the corrosion and deposition problems often associated with the combustion of straw alone [Larsen and Inselmann, 1994; Henriksen et al., 1995; Robinson et al., 1997]. The potential risks of slagging/fouling of the boiler, and HT corrosion of the superheaters when introducing straw to the boiler, have led to comprehensive investigations of both subjects in the demonstration programme. An overview of the Danish experiences with combustion of straw alone or in conjunction with coal is given by Frandsen et al. [1997], and ash fusion and deposit formation in straw-fired boilers is examined by Hansen et al. [1997].

2. BACKGROUND The ash forming elements occur in solid fuels mainly as internal or external mineral grains, as simple salts, e.g. NaCl and KC1, or associated with the organic matrix of the fuel. In coal, a large fraction of the inorganics is present as minerals, whereas in most biomasses the major part of the inorganics is present as simple salts or organically associated. Approximately 0.5–wt% [Flagan and Friedlander, 1978] of the inorganics in coal vaporise during combuation, whereas between 30 and 75wt% of the inorganics in straw

[Livingston, 1991; Baxter, 1993; Dayton and Milne, 1995; Olanders and Steenari, 1995] will be vaporised at 1200°C. During combustion, the inorganics in the fuel are transformed into ash through a complex combination of chemical and physical processes. The resulting ash can be divided into two fractions: The residual ash which mainly consist of the transformed minerals and the sub-micron ash which consist of the inorganic species vaporised during combustion [Benson et al., 1993; and others]. The minerals can undergo fragmentation, shedding and coalescence during combustion. Similarly, the vaporised inorganic species may take part in several reactions during combustion. These include heterogeneous condensation on heat transfer surfaces or fly ash particles, or homogeneous nucleation due to a local supersaturation of salts, e.g. or KC1 [Flagan and Friedlander, 1978; Christensen, 1995], thereby forming a condensation aerosol which subsequently can coagulate to larger particles. The resulting vapor and fly ash particles may deposit by a number of mechanisms on heat transfer surfaces in the boiler. The most important transport mechanisms involved in ash deposition are inertial impaction, thermophoresis and diffusion, but also chemical reactions such as sulphation can play a major role [Baxter, 1993; Benson et al., 1993], The general effect of ash deposits is to change the heat uptake pattern in the boiler, which may lead to an increase in flue gas temperature until the deposits are removed, either by gravitational or thermal shedding or mechanically. In case build-up of ash deposits can not be controlled by soot blowing (or other mechanical cleaning), they can constitute a serious operational problem, and may in extreme cases lead to plugging of the convective pass or can seriously damage the bottom ash hopper if large (several tonnes) pieces of deposits fall down. This may lead to unscheduled outages of the boiler and thus serious economic loss.

Full Scale Deposition Trials at 150MW, PF-Boiler CO-Firing Coal and Straw

369

The actual extent of ash related problems in utility boilers depend upon the quantity and type of inorgancs present in the fuel, the geometry of the combustion system and the combustion conditions.

3. EXPERIMENTAL FACILITY The demonstration programme was conducted at the Studstrup Power Station, Unit 1 (MKS1), which was first commisioned in 1968. The boiler is a 150MW e wall-fired unit

with 12 burners arranged in three burner rows. The four burners in the middle burner row have been converted for coal-straw co-combustion with up to 50% straw share (energy base). The major plant data are given in Fig. 1 and Table 1.

4. EXPERIMENTAL The investigation of deposit formation is based primarily on deposits collected on temperature controlled probes. Simultaneous in-situ measurements of temperature and gas composition, as well as sampling of fly ash particles and aerosols were carried out. The deposition and in-situ experiments were carried out in five positions throughout the

boiler at a range of operational conditions. Aerosol measurements were conducted at a position in between the two sections of the economiser. In order to establish the reliability of the experimental results, mass balance closures for major components were performed at all the operational conditions used. Thus, the test programme offers a unique possibility of evaluating deposit formation in a fairly well-characterised full scale combustion system. Deposition trials, in-situ experiments and mass balance closures were performed at

370

K. H. Andersen et al.

several operational conditions, including variation of load (50, 75, 100%) and straw share (0, 10, 20%, energy basis), with a South American (coal 1) and a US (coal 2) high-volatile bituminous coal. See Table 2. Deposits were collected in positions 1 through 5 with coal 1, and in positions 2 and 3 with coal 2. In each experiment, a minimum of 12 hours was allowed for stabilisation before deposition trials were performed. Coal 2 has a lower ash content than coal 1, the content of S, Fe and alkali is higher and the content of Si lower for the coal 2 ash compared to coal 1 ash. The straw burned during the test periods was primarily Danish wheat from the 1995 harvest. The compositions of the fuel ashes are given in Table 3. The slagging and fouling indices [Winegartner, 1974] indicate, that coal 2 can be considered medium slagging and fouling whereas coal 1 can be considered low slagging and fouling The main reasons for this categorisation are given by the increase in Fe and S and the decrease in Si for the ash from coal 2 compared to coal 1.

5. DEPOSIT SAMPLING Two types of probes were used for sampling of deposits. Water-air cooled probes were used in the hot positions 1–3, and air cooled probes in the cooler positions 4 and

5. Probe metal temperatures were measured with 3 thermocouples placed at 12, 4 and 8 o’clock in a ring situated before the first test element. The probe temperature on the upstream side was kept at a set point temperature value by automatical control of the

Full Scale Deposition Trials at 1SOMWe PK-Boiler CO-Firing Coal and Straw

371

cooling air flow. All three metal temperatures were continuously registered during exposure in the boiler. The set point metal temperatures used in the experiments were 400 °C in probe positions 4 and 5, 540 °C in probe positions 1–4, 580 °C in probe positions 1–3 and 620 °C in probe positions 1–3. At each metal temperature, deposit samples were collected after 15 minutes, 3 hours and 18 hours exposure in the boiler. In addition, a few deposits were collected after 72 hours exposure.

6. RESULTS OF DEPOSIT ANALYSIS More than 200 deposit samples were collected during the demonstration programme, and the need for a cheap and relatively quick analysis of these samples was evident. Thus, a procedure for visual analysis of the samples was set up, based exclusively upon the physical appearence of a deposit. This procedure has been used for all samples, and a limited amount of samples were selected for bulk chemical analysis and SEM analysis. The samples were selected in order to evaluate changes with all or some of the five key factors examined: Straw share, boiler load, exposure time, probe metal temperature and coal type. Thus, results with increased levels of information was gathered, going from the evaluation of physical structure by visual analysis and chemistry by bulk chemical analysis, to evaluation of structure by SEM-BSE (Back Scattered Electrons) micrographs and chemical information from XRAY mappings and SEM-EDX analysis. In the following, results from each type of analysis are given with main emphasis on the upstream deposits from probe position 2, and the effect of introducing straw as part of the fuel. In addition, an introduction to the visual analysis system will be provided. Probe position 2 is chosen since it is the first position where operational problems could be expected to occur due to fouling deposits. The average flue gas temperatures measured in position 2 at 100% load under combustion of coal 1 was 1,105°C, and 1,195 °C was measured during combustion of coal 2. The main reason for the higher temperature during combustion of coal 2 is operation at a lower excess air during the periods of measurement. Position 1 is closer to the slagging (radiant) part of the boiler, with a measured average temperature of 1,264°C during combustion of coal 1.

6.1. Visual Analysis of Deposits The classes used in the visual analysis are based upon 1) the physical appearence of the collected deposits, and 2) the relative tenacity of each class. Each deposit sample is evaluated at four positions, 90° apart: Upstream, side, downstream and side. The classes are numbered from 1 through 5, where the tenacity increases with increased number. See Figs. 2 and 3. For the well adhered deposits, increased amount of deposit leads to an increase in visual classification. Loose deposits (class 1) are easily removed by sootblowing regardless of thickness, whereas deposits of class 2 and higher are more difficult to remove. The

372

K. H. Andersen et al.

increased amount of deposit from class 3 to 5 leads to a less effective heat uptake. An example of a class 5 deposit (position 1 , 1 8 hours, 100% load, 20% straw share) is given in Fig. 3, with the arrow indicating the direction from the upstream towards the downstream side. The numbering system used allows graphical representation of the results from the visual analysis, which makes the evaluation of the results easier. An example is given in Fig. 4. It is seen from Figure 4, that no effect of straw share and exposure time can be observed in positions 4 and 5, whereas a clear effect is seen in positions 1–3. In positions

1–3, the deposit tenacity (visual classification) is most often seen to increase from 15 minutes exposure to 3 and 18 hours exposure, or not to change at all with increased expo-

sure time. In position 1, at 20% straw share, a decrease in tenacity is seen from 3 to 7

373

hours exposure. This is most probably due to deposit shedding. In general, the deposits collected in positions 1–3 at 20% straw share are more tenacious than the deposits at 0% straw share. The results of the visual analysis were evaluated with regard to gas temperature (probe position), straw share, boiler load, probe metal temperature, and coal type. The conclusions from the analysis are summarised below, first for coal 1 and then also for coal 2. • Powdery, non-tenacious, upstream deposits were found in positions 4 and 5, whereas more tenacious upstream deposits were found in positions 1–3. • The tenacity decreased from position 1 to position 3, as would be expected from the flue gas temperatures and the amount of radiation the deposits have been exposed to. • No distinct trend was seen for the influence of probe metal temperature. • The amount and tenacity of upstream deposits increased with increased straw share, primarily for positions 1–3. See Figure 4 for an example. • The formation of well adhered upstream deposits increased when increasing the load from part load (50, 75%) to full load (100%). • Deposit amount and tenacity increased with increasing exposure time, primarily from 15 minutes to 3 or 18 hours, as would be expected. • More tenacious deposits were formed in positions 2 and 3, when burning 100% coal 1 as compared to 100% coal 2. • At 20% straw share, the deposit amounts and type were similar for coal 1 and 2. However, the deposits collected with coal 2 were more prone to outer sintering and melting on the downstream side, probably due to the higher flue gas temperatures during the experiments with coal 2. • Deposit amounts were alike for coal 2 at 0 and 20% straw share.

374

K. H. Andersen et al.

6.2. Chemical Bulk Analysis of Deposits Deposit samples from positions 1 and 2 were selected for bulk chemical analysis to support the findings in other analysis. The amount of upstream deposit collected in position 3 was too small for chemical analysis. The results are given in Table 4. A comparison of the ash compositions of the fuels (see Table 3) on an energy basis reveals, that the coal ashes are rich in Fe and Al, whereas the straw ash is rich in Ca, K, P and Cl. One could thus expect the effect of Fe and Al to decrease, and the participation of primarily K, Ca, P and Cl to increase when straw is added to the fuel. From a comparison of fly ash composition and the bulk analysis of the deposits, Fe is found to decrease and Ca, K, P and S to increase at 20% straw share. Only very small amounts of Cl are observed in the probe deposits, and high emissions of Cl species were observed during the experiments, indicating that most Cl leaves the boiler as gaseous HC1, which is also expected from a global thermodynamic evaluation. Similar bulk chemical composition was found at two upstream deposits formed at position 2 during almost identical experimental conditions (exp. 3 and 3B), but one year apart. This indicates a relatively good reproduction of deposits during similar operational conditions, and also validates the results found in experiment 3, compared to experiment 3B. The contents of the upstream deposits are seen to be higher during coal combustion compared to co-combustion deposits. This corresponds to the fact, that the coal ash is diluted by the almost Fe-free straw ash during co-combustion, thereby decreasing the amount of Fe introduced to the boiler. However, the change observed for the upstream deposits is approximately three times higher than would be expected from the fuel ashes. The CaO content is found to be significantly higher in the upstream deposits during co-combustion with both coals than during coal combustion, and to be low in the downstream deposits both at 0 and 20% straw share. This indicates that Ca from the straw fuel takes actively part in the formation of upstream deposit.

Full Scale Deposition Trials at 150MW, PF-Boiler CO-Firing Coal and Straw

375

From the composition of the deposits, it is evident that the participation of and in both up- and downstream deposits increases significantly with the addition of straw to both coal types. This is probably part of the explanation for the increase in amount and tenacity for the upstream deposits observed during the co-combustion experiments and confirmed from the visual analysis.

6.3. Deposit Structure Based on evaluation of BSE micrographs from the SEM, a few general points are seen for the upstream deposits from position 2, as well as for the inner part of the

upstream deposits from position 1. A general change in textural appearence is seen for the upstream deposits in positions 1 and 2 when the fuel composition changes from 0 to 20% straw share. In the coal combustion case, the deposit is generally very orderly structured in a dendritic growth, with “fingers” formed from the larger particles, often with an Fe-based skeleton, and “necks” formed by deposition of smaller Al-silicate (clay-derived) particles, which are also deposited between the large particles in the “fingers”. This deposit type was also described by Laursen [1997] for coal ash deposits, and is shown in Fig. 5a. At 20% straw share, the orderly structure of the upstream deposits is broken, and the large particles are deposited in a more random manner between the small particles. In addition, the large particles are generally larger than in the pure coal case, and often contain mainly Ca and Si. This deposit type is illustrated in Figure 5b. The observations of Fe and Ca from the bulk chemical analysis corresponds to the change seen in deposit structure, where the effect of Fe is taken over by primarily Ca in the co-combustion deposits. The Fe-rich particles that primarily constitute the skeleton in the dendritic “fingers”

376

K. H. Andersen et al.

obtained during coal combustion, are generally seen to be deformed upon impact, which indicates that they consist mainly of reduced Fe-compounds with relatively low viscosity. The most probable deposition mechanism is selective inertial impaction of Fe-rich particles. The formation of “fingers” could be initiated by the deposition of a relatively large Fe-rich particle, with the outer part penetrating a thermal boundary layer around the tubes, thereby entering a considerably higher flue gas temperature [Bryers, 1996]. The temperature could even exceed the melting point of the particle or eutectic mixtures formed by fluxing agents on the surface of the particle and thus lead

to a sticky surface for further deposition. Primarily smaller, but also some larger Al-silicate (clay-derived) particles are “caught” in the 3-dimensional Fe-based skeleton, either by eddies or sedimentation in the voids (small particles) or by adherence to the sticky surface of the Fe-rich particles (large particles). Between the Fe-based “fingers”, deposition of small Al-silicate particles take place, probably due to eddy formation between the “fingers”. The change in upstream deposit structure from 0 to 20% straw share, along with the fact that the influence of Fe decreases during co-combustion, indicate that the Ferich particles in co-combustion are more oxidised than during pure coal combustion, and are thus more likely to rebound upon impact [Bool and Helble, 1996]. This can most likely be due to an increase in the

level in the boiler and/or in the residence time of

the particles during co-combustion. Both of these factors increase during co-combustion of coal 1, whereas both the level and residence time are found to decrease during cocombustion of coal 2. Thus, a change in combustion stoichiometry combined with a dilution of the ash due to the addition of straw can explain the change in deposit structure for coal 1, whereas another factor (or factors) must govern the deposit structure for coal 2. One factor, that could play a more significant role for coal 2, is the effect of dilution of the coal ash by the straw ash. The ash content in coal 2 is only two thirds of the ash content in coal 1, when evaluated on the basis of calorific value. Moreover, the content of Ca in the coal 2 ash is higher than in coal 1. These two factors, along with the increased content of Ca and Si caused by the addition of straw to the fuel feedstock, could possi-

bly lead to the observed change in deposit structure. However, this is only part of the explanation, since the dilution of the Fe in the deposits is approximately 3 times the dilution in the fly ash for coal 2. The mode of occurrence of Fe in the coal could also have an effect on the deposit structure. The downstream deposits are of the powdery kind for all the probe positions, where large amounts of small particles deposit on the downstream side of the probe due to eddies. In addition, a small amount of larger particles can be found in the downstream deposits. This deposit type is shown in Figure 5c.

6.4. XRAY Mappings A number of XRAY mappings were performed to provide an overview of the way the elements are associated in the deposits. In Fig. 6, the results of an XRAY mapping of the marked area in Fig. 5b are given. This represents the upstream deposit collected in position 2 during combustion of coal 2 at 20% straw share. The large particles in the mapped area consist mainly of Ca with some Si and Mg, but without any Al. Among the large particles, a number of relatively large Fe-rich particles have impacted, often with deformation as a result, indicating reduced Fe compounds. However, also a few smaller, almost spherical, Fe-rich particles are seen among the small particles, indicating the deposition of small oxidised

Full Scale Deposition Trials at 150MWe PF-Boiler CO-Firing Coal and Straw

377

or completely cooled Fe-rich particles. Otherwise, the small particles are seen to be dominated by Al with some participation of Si and K, probably as K-Al-silicates. Primarily around the Ca containing particles, a corona of K, Ca and S, probably as and and/or an eutectic mixture of the two, is seen. However, small areas seem to contain a K and S based melt as well, indicating formation of sulphate bonds in the deposit. The coronae indicate, that some of the deposited particles were fully or partly covered with K and/or Ca sulphate upon deposition, which may have provided the necessary sticky surface for adhesion. This was also reported by Hurley et al. [1995] in

“Project Calcium”. A similar XRAY mapping of a coal 2 upstream deposit from position 2 (Figure 5a) shows a skeleton of impacted Fe-rich particles in the “finger” structures, which most likely consist of reduced Fe compounds, as discussed previously. Between the Fe-rich particles several smaller and a few larger Al-silicate fly ash particles are observed, as the “necks” and in voids in the “finger” structure shown in Figure 5a. In addition, some small Ca and spherical Fe-rich particles are deposited as part of the “necks”. No significant amount of K participates in the upstream deposit formation, but some S containing coronae are seen around the Ca particles, probably consisting of CaSO4. The same picture as described for coal 2 is seen for upstream deposits of coal 1 in position 2, which mainly consist of large Fe and Ca particles, along with Si, Ca, K (20% straw) or Si, Al, Ca (0% straw) containing particles. The amount of Fe in the deposit decreases from 0% to 20% straw share for both coal 1 and 2. Packed between the large particles are large amounts of smaller particles, which mainly consist of Si, Al and K, with the highest It-content at 20% straw share. Some melt formation is seen for coal 1 at 20%) straw share, where K, Ca and S containing melt is formed between Ca and Fe containing particles. Thin S containing coronas are seen mainly at 20% straw share for both coal 1 and 2, primarily around Ca containing particles. The coronae primarily contains Ca, S and K, with increasing K content as the straw share increases.

378

K. H. Andersen et al.

The amount of the large Ca-containing particles increase during co-combustion with 20% straw share, which is consistent with the observed increase in the CaO content in the bulk analysis. This fact can also explain the decrease in the downstream deposits in the bulk chemical analysis, since the number of particles above diameter in the downstream deposits is low, and Ca is most often found in the large particles. XRAY-maps of downstream deposits from position 2 revealed, that the deposits from both coal types are dominated by small K-Al-silicate particles, with a few additional particles containing Fe, Ca, and occasionally Ca and S, at 20% straw share. For coal 2, melt formation was observed between some of the particles in the downstream deposit, primarily consisting of Ca, K and S, i.e. most probably consisting of and

6.5. SEM-EDX Analysis of Deposits Selected deposit samples were casted in epoxy, polished and analysed by SEMEDX. A system was set up for radial analysis at the four positions evaluated in the visual analysis. One or more lines were drawn through representative parts of the deposit, and an operator determined number of spot analysis were performed along these lines. An evaluation of the SEM-EDX results for selected samples from position 2 was performed using ternary diagrams with regard to and and either , CaO or For all the spot analysis results used in the evaluation, the content of the three end-members in the ternary diagrams exceeded 50%, and in most cases also 80%.

The ternary diagrams for coal 1 at 0 and 20% straw share are shown for position 2 in Fig. 7. For CaO, the effect is seen to be the opposite of with more CaO in the upstream deposits at 20% straw share. This effect supports the bulk chemical analysis results and the more qualitative findings in the XRAY mappings, where a shift from Fe to Ca and Si is seen in the composition of the larger particles in the upstream deposits when straw is added to the fuel. The same trend is seen for the coal 2 deposits. For both up- and downstream deposits, the participation of is seen to increase with increasing straw share, which corresponds to the findings in the bulk chemical analysis, is found to be mainly in connection with but with some present, which corresponds to the presence as K-A1 silicates. The same trends are seen for coal 2 deposits, but here the participation of in the deposit at co-combustion with 20% straw share is smaller than observed for coal 1. However, approximately a doubling of the is seen at 20% straw share compared to 0%. Fly ash based calculations of upstream deposit formation on a dry tube/deposit [Frandsen, 1997] shows, that the increased deposit formation and tenacity during cocombustion is probably primarily due to a more sticky deposit, since the calculations results in the same flux of sticky particles for 0 and 20% straw share.

7. CONCLUSIONS The results from the deposition trials at the MKS1 power station, with special focus on probe position 2 and the effect of straw addition, are summerised in the following:

• A structural change in the upstream deposits was observed for positions 1 and 2 when 20% straw share was introduced to the fuel. • Pure coal ash deposits were found to be dominated by reduced Fe-rich particles,

Full Scale Deposition Trials at 150MW e PF-Boiler CO-Firing Coal and Straw

379

which make up a skeleton in a dendritic “finger” structure. Smaller particles, primarily Al-silicates, are deposited in “necks” between the “fingers” and in voids in the “fingers” themselves. • Deposits from co-combustion with 20% straw share were found to be formed by very large particles, primarily Ca-Si-based, which are more randomly distributed

380

K. H. Andersen et al.

between the small fly ash particles, mainly consisting of Al-silicates. However, Fe-rich particles still participate in the deposit formation, both as reduced Fecompounds and small spherical Fe-rich particles. • During co-combustion at 20% straw share, the participation of K and S increase in both the up- and downstream deposits for both coal 1 and 2. This indicates,

that one of the reasons for increased deposit amount and/or tenacity in the hottest positions during co-combustion is an effect of K and S, possibly due to the formation of a sticky surface primarily on the deposit, but possibly also on some of the fly ash particles. • No content of Cl was observed in the deposits from co-combustion by SEMEDX, and based on the deposits collected, selective chlorine corrosion is thus not expected to constitute a problem in co-combustion of coal up to 20% straw share for the coal types utilised in the tests. Selective chlorine corrosion was not observed in a long term corrosion experiment at 10% straw share. • The deposit amount and tenacity increased in the hottest positions during cocombustion with coal 1, but were alike the pure coal ash deposits for coal 2. However, more research is needed to fully evaluate the effect of coal type. Overall, the deposition during co-combustion with coal 1 and 2 is manageable by increased soot blowing frequency in the furnace and the first pass.

8. ACKNOWLEDGMENTS The demonstration programme is funded by the ELSAM Research and Development programme. Part of the work is conducted by the CHEC Research Programme, which is co-funded by ELSAM, ELKRAFT (the Zealand Electricity Consortium), the Danish Technical Research Council, the Danish and Nordic Energy Research Programmes, and the European Union. The employees at I/S MIDTKRAFT and ELSAMPROJEKT A/S who helped with the practical experiments are greatly acknowledged for their efforts, especially John B. Larsen, Michael Christensen, Tage Rasmussen, Lars Christiansen, Lars Harpøth and Keld Nicolaisen.

9. REFERENCES Baxter, L.L. (1993). “Ash Deposition during Biomass and Coal Combustion: A Mechanistic Approach.” Biomass and Bioenergy, 4 (2), 85–102 Baxter, L.L., Miles, T.R., Miles, Jr., T.R., Jenkins, B.M., Dayton, D.C., Milne, T.A., Bryers, R.W. and Oden, L.L. (1996). Alkali Deposits Found in Biomass Boilers, vol. II: The Behavior of Inorganic Material in Biomass-Fired Power Boilers—Field and Laboratory Experiences, N R E L TP-433-8142, SAND96-8225, vol 11 Benson, S.A., Jones, M.L. and Harb, J.N. (1993). “Ash Formation and Deposition.” In L.D. Smoot (Eds.), Fundamentals of Coal Combustion—for Clean and Efficient Use, Coal Science and Technology 20, 99–373. New York: Elsevier Science Publishers.

Bool III, L.E. and Helble, JJ. (1996). “Iron Oxidation State and its Effect on Ash Particle Stickiness.” In L. Baxter and R. deSollar (Eds.), Applications of Advanced Technology to Ash-Related Problems in Boilers.

New York and London: Plenum Press, 281–292. Bryers, R.W. (1996). “Fireside slagging, Fouling and High Temperature Corrosion of Heat-Transfer Surface due to Impurities in Steam-Raising Fuels.” Progress in Energy and Combustion Science, 22 ( 1 ) , 29–120. Christensen, K.A. (1995). The Formation of Submicron Particles from the Combustion of Straw. Ph.D. Thesis, Department of Chemical Engineering. Technical University of Denmark.

Full Scale Deposition Trials at 150MW e PF-Boiler CO-Firing Coal and Straw

381

Dayton, D.C. and Milne, T.A. (1995). “Mechanisms of Alkali Metal Release During Biomass Combustion.” ACS, Div. of Fuel Chemistry, 40 (3), 758–762. Flagan, R.C. and Friedlander, S.K. (1978). “Particle Formation in Pulverized Coal Combustion—A Review.” In D.T. Shaw (Eds.), Recent Developments in Aerosol Science. New York: John Wiley & Sons. Frandsen, F.J. (1997). Estimation of Ash Deposition Fluxes in Utility Boilers: 1. Inertial Deposition. Reserch Progress Report, CHEC, Department of Chemical Engineering, Technical University of Denmark. Frandsen, F.J., Nielsen, H.P., Jensen, P.A., Hansen, L.A., Livbjerg, H., Dam-Johansen, K.., Hansen, P.F.B.,

Andersen, K.H., S0rensen, H.S., Larsen, O.H., Sander, B., Henriksen, N. and Simonsen, P. (1997). “Deposit and Corrosion Problems in Straw- and Coal-Straw Co-Fired Utility Boilers—Danish Experiences.” Proc. Eng. Found. Conf, Impact of Mineral Impurities in Solid Fuel Combustion, Kona, Hawaii, USA, November 2–7. Grunge, T. (1989). “Material Experiences with Woodchip Combustion.” Materialnyt, 2 (In Danish) Hansen, L.A., Frandsen, F.J., Sørensen, H.S., Rosenberg, P., Hjuler, K. and Dam-Johansen, K. (1997). “Ash Fusion and Deposit Formation at Straw Fired Boilers.” Proc. Eng. Found. Conf., Impact of Mineral Impurities in Solid Fuel Combustion, Kona, Hawaii, USA, November 2–7.

Hansen, P.F.B., Andersen, K.H., Wieck-Hansen, K., Overgaard, P., Rasmussen, I., Frandsen, F.J., Hansen, L.A. and Dam-Johansen, K. (1996). “Co-Firing Straw and Coal in a 150MW e Utility Boiler: In-Situ Measurements.” Proc. Eng. Found. Conf., Biomass Usage for Utility and Industrial Power, Snowbird, Utah, USA, April 28th May 3rd. Henriksen, N. and Hansen, P.F.B. (1995). CFB-Material Test at Grenaa CHP Plant—Final Report of Phase 1 and 2. ELSAMPROJEKT Notat EP94/799a. Henriksen, N., Larsen, O.H:, Blum, R. and Inselmann, S. (1995). “High-Temperature Corrosion when CoFiring Coal and Straw in Pulverised Coal Boilers and Circulating Fluidized Bed Boilers.” In Proceedings of the VGB Conference Corrosion and Corrosion Protection in Power Plant Technology. Hurley, J.P., Benson, S.A., Erickson, T.A., Allan, S.E. and Bieber, J. (1995). Project Calcium, Final Report. DOE/MC/10637-3292, Combustion and Environmental Systems Research. Larsen, O.H. and Inselmann, S. (1994). Superheater Corrosion in Co-Combustion of Coal and Straw. Experiments at Vestkraft. Unit I , ELSAM R&D Project, Internal report from Faelleskemikerne, June 1994 (In

Danish) Laursen, K. (1997). Characterization of Minerals and Interpretations of Ash Formation and Deposition in Pulverized Coal Fired Boilers. Ph.D. Thesis, volume 1 and 2, Geological Survey of Denmark and Greenland, Ministry of Environment and Energy, ISBN 87-7871-022-7 Livingston, W.R. (1991). Straw Ash Characteristics. Babcock Energy Ltd., Babcock Report Number 34/91/08, Contractor Report, Department of Energy, ETSU B 1242. Michelsen, H.P., Larsen, O.H., Frandsen, F. and Dam-Johansen, K. (1996). “Deposition and High Temperature Corrosion in a 10 MW Straw Fired Boiler.” Proc. Eng. Found. Conf., Biomass Usage for Utility and Industrial Power, Snowbird, Utah, USA, April 28th–May 3rd. Nielsen, H.P. (1996). Chlorine-Induced High-Temperature Corrosion of Superheater Tubes—a Literature Survey, Internal report, Department of Chemical Engineering, Technical University of Denmark, CHEC Report No. 9615 Miles, T.R., Miles, T.R. Jr., Baxter, L.L., Bryers, R.W., Jenkins, B.M. and Oden, L.L. (1995). “Alkali Deposits Found ind Biomass Power Plants. A Preliminary Investigation of Their Extent and Nature.” Summary Rep. for National Renewable Energy Laboratory, Subcontract TZ-2-11226-1. Olanders, B. and Steenari, B-M. (1995). “Characterization of Ashes from Wood and Straw.” Biomass and

Bioenergy, 8 (2), 105–115. Overgaard, P. and Hansen, P.B.H. (1997). “Co-Firing Coal and Straw in a 150 MWe Power Boiler.” Conf. Proc.

POWER-GEN EUROPE ‘97. Utrecht: PennWell. Robinson, A., Junker, H. and Baxter, L. (1997.) “Pollutant Formation, Ash Deposition, and Fly Ash Properties when Cofiring Biomass and Coal.” Proc. Eng. Found. Conf., Economic and Environmental Aspects of Coal Utilization, Santa Barbara CA, February 18–21. Stenholm, M., Jensen, PA. and Hald, P. (1996). The Fuel and Firing Characteristics of Biomass—Combustion Trials. EFP-93 Project, Jounal No. 1323/93-0015 (In Danish) Sørensen, H.S. (1997). Preliminary results of CCSEM and SEMPC analysis performed as part of the EFP-95 project, project no. 1323/95-0007 Winegartner, E.C. (1974). Coal Fouling and Slagging Parameters. The Am. Soc. Mech. Eng.

This page intentionally left blank.

SLAGGING TESTS ON THE SUITABILITY OF ALTERNATIVE COALS IN A 325 MWF PC BOILER Timm Heinzel, Jörg Maier, Hartmut Spliethoff, Klaus R. G. Hein 1 , and Werner Cleve 1

Institute for Process Engineering and Power Plant Technology University of Stuttgart Pfaffenwaldring 23, 70569 Stuttgart, Germany 2 Braunschweigische Kohlenbergwerke AG Schöninger Str. 2-3 38350 Helmstedt, Germany

INTRODUCTION Several small lignite seams are located near Helmstedt, Germany, which have been exploited since the last century. At present two power plants, of 325 and 350 capacity, are operated nearby, each directly connected to the excavation system of one seam. After more than twenty years of excavation, the seam which supplies the older 325 MWe plant will be exhausted in 2001. For further operation of this plant, the coal can be blended with or replaced by other coals. The most probable replacement coal originates in a local seam with higher alkali content. A second possible coal type is located

in a different area. The plant was operated with the alternative coals and different blends

in a six-week test series during 1996 in order to evaluate the feasibility and the effects on operation when the coals are changed or blended. Besides evaluating the effect of the

change on general operation, coal feeding and milling system, steam cycle parameters, ash quality and handling etc., mainly done by the plant owner, the goal of the measurements done by the authors was to evaluate the influence of the coal types on slagging, fouling, temperatures and gas concentrations in the boiler.

DESCRIPTION OF THE BOILER The Offleben unit C was commissioned in 1972 and is a conventional tangentially fired “brown coal” boiler. It has twelve burners, in six pairs, on two levels. Six impact pulverizers (high speed fan-beater mills) dry and mill the raw fuel with hot flue gas, each feeding two burners. Five mills and burner pairs are required for full load operation. The cross-section of the boiler is 16.5m ×16.5m, and the furnace is 76m high, followed by the downstream convective zone, sketched in Fig. 1. Steam parameters are 190 bar/530 °C past Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

383

384

T. Heinzel et al.

final superheater and 42.5 bar/535 °C past reheater. For

reduction, the combustion air supply was changed to an air-staging concept in the eighties, and the boiler was equipped with urea injection and a desulfurization plant to meet legislative exhaust gas limits.

EXPERIMENTS During six weeks in autumn 1996, twenty-seven experiments with the three coal types and different blends were carried out at different fuel rates and loads. Due to restrictions of possible burner loads and fuel supply, the blending rates were restricted to 40% vs. 60% and the inverse ratios, combined with output power levels of and the maximum available load . Most of the 27 experiments were carried out over a minimum ten-hour period. All experiments of one coal type or

blend were carried out without switching to a different fuel or blend within the series to achieve comprehensive results for each coal. Figure 2 represents the five coal or blend experimental series, arranged by load, which are referred to as coal 1, coal 2, coal 3, blend 1 and blend 2: — 7 experiments with 100% of coal 1, the regular lignite (Helmstedt) — 5 experiments with blend 1 of 60% coal 1 and 40% coal 2 — 3 experiments with blend 2 of 40% coal 1 and 60% coal 2 — 9 experiments with 100% of coal 3, the alternative seam lignite (Schöningen) — 3 experiments with 100% coal 3, the external area lignite (MIBRAG) Of these experiments, those that had the best comparable conditions at full load

were selected for detailed analysis and comparison. Full load experiments were chosen

Slagging Tests on the Suitability of Alternative Coals in a 325 MWe PC Boiler

385

because the influences of the coals were less pronounced at partial load than at full load. Two experiments at full load were selected for coal 1, and one each for coal 2, coal 3 and the blends. Two experiments with coal 1 were selected because this coal unexpectedly caused serious slagging problems in one of the experiments while not causing trouble in the rest. Hence, one experiment with coal 1 without slagging at full load and another where heavy slagging happened were selected. Measurements of the other experiments

were analyzed in addition to subsidize the results.

MEASUREMENTS AND SAMPLING All relevant process data were recorded using a data acquisition system. Measurements of flue gas temperatures from the furnace wall to the depth of 4.5m were carried out using a suction pyrometer probe. Local gas concentrations of and were measured with a gas analyzer using gas from a suction probe. These measurements were carried out at levels of 12m, 25m, 42m and 62m from the bottom of the boiler (Fig. 3). Samples of coal from the feeder and from ash of wet ash discharge, convective path ash discharge and electrostatic precipitator were taken every two hours and stored separately. At the levels of 12m, 40m and 62m, slagging probes, inserted into the boiler to a depth of approximately four meters, were installed and slag that had deposited on ceramic carriers was sampled at the end of each experiment (Fig. 4). On the levels of 12m, 21 m and 36m, samples were taken at the end of each experiment during normal operation directly from the furnace walls with suitable probes in a distance of approximately 2 m from the ports, from the front side of the first row of final superheaters and from spaces between the superheater tubes. On the level of 62m in front of the entrance to the convective section a deposition sampling probe with exchangeable tubes of regular 13CrMo44 steel was installed. The surface temperature of the tubes was controlled to 620–650 °C by coupled water/air cooling and dual-thermocouple measurements. During each experiment, the furnace wall conditions were regularly inspected by visual checks through more than 50 available ports. The visible condition and peculiarities were recorded and photographed. During the experiments, soot and water blowers were not operated, but the furnace was cleaned carefully before the experiments to

386

T. Heinzel et al.

provide comparable conditions. A continuous acoustic flue gas temperature measurement system was installed with six corresponding ports on the level of 45m with 30 path measurements and a system with two ports and two 2 path measurements was installed on level 62m.

ANALYSES METHODS Coal samples were taken every two hours during each experiment. These were mixed and milled to get representative samples of each experiments. The coal samples

Slagging Tests on the Suitability of Alternative Coals in a 325 MW e PC Boiler

387

were analyzed by standard methods, on proximate and ultimate analyses, on heating values, and elemental composition of ashes. Based on these data, further analysis methods were applied to selected samples. The coal samples from the selected experiments using 100% parent coals were analyzed by petrographic techniques to identify possible changes in the degree of carbonization and structure. Fluctuations in the composition of the coal were identified by analysis of the single samples taken each hour, in particular with regard to ash content and ash composition. CCSEM was carried out to identify the distribution of species in the mineral inclusions. Ashes from wet ash discharge, convective path ash discharge and from the electro-

static precipitator were analyzed for elemental composition. Slag samples, furnace-wall and superheater deposits were analyzed with regard to their bulk chemical composition by standard analysis, to morphology and local composition by SEM-EDX and to mineral species by XRD. Superheater deposits were analyzed by CCSEM. Ash fusion tests were carried out on more than 100 selected samples.

RESULTS

Analysis of the Fuels in the Tests The coals tested were typical German open-pit mining lignites referred to as “brown coals” (Table 1). Water content was about 46% for coal 1 and 2 (regular and alternative seam from the local area) and 50% for coal 3. The mean heating value of the three fuels

and two blends examined in the experiments was between 10.5 and 11.3MJ/kg. Sulfur content was high at values of 2–3%, whereas the nitrogen content was below 0.3%. Compared to coal 1 (regular seam), coal 2 (alternative seam) had a higher ash content resulting in less combustible matter and lower heating value. Coal 3, the externally purchased lignite, was similar the regular coal in combustibles and heating value, because the higher water content was compensated for by a lower ash content.

388

T. Heinzel et al.

The coals differed in petrographic characteristics. It can be noticed that the external coal was similar to the regular coal. The alternative seam had a lower content of Huminite. The chemical composition of the laboratory ashes (produced by high-temperature ashing of the coal samples in a laboratory furnace at 575 °C) is shown in Table 3. The four main components are and in ash. Silicon and aluminum content were highest in the alternative seam. Aluminum content was low in the regular and in the external area coal. The alkali content in the ash of coals 1 and 3 was low. The alkali content was slightly higher in the alternative seam, but still below 2%. This coal

from the alternative seam was initially expected to contain up to 8% alkalis in ash, which was predicted by test boring samples, and therefore a high risk of furnace slagging was expected. No influence of the alkali species on furnace depositions was detected. The base-acid ratio of the ash was lowest for the alternative seam, high for the regular coal and highest for the external area lignite. The fuel compositions of the local seam samples deviated from the mean values against time and single experiments more strongly than the purchased one (Fig. 6a). The

Slagging Tests on the Suitability of Alternative Coals in a 325 MW e PC Boiler

389

local seams were smaller and less homogeneous, and the coal was less intermingled due to being directly conveyed from the excavation system to the plant. This was resulting in an ash content between six and more than 15% in the day’s mean samples of the raw fuel in regular plant operation, shown for fifty days in Fig. 6b. Besides varying in the amount, the ash varied in its composition. Figure 7 shows the variation of the main components of the samples taken every two hours during an experiment with coal 1 (the regular lignite) and which were ashed in a laboratory furnace at 575°C. The ratios of silicon to aluminum and calcium varied strongly. The three coal types, represented by the mean samples of the main experiments, are distinguished by mineral inclusions, too. The CCSEM analysis showed that the main minerals in the regular lignite consisted of almost pure SiO2, whereas the alternative scam contained few SiO2 but high amounts of montmorillonite-class minerals with a Si/Alratio of approximately 4. Both contained smaller amounts of kaolinite-class minerals with a Si/Al-ratio of 2. The external-area coal comprised a broader range of inclusions with larger amounts of kaolinite-class minerals. Inclusions, mainly with iron and sulfur content, had a share of 15–20% in the regular lignite and the external-area coal but were low in the alternative coal, although this coal contained the most sulfur. The distribution of elements in mineral inclusions, which was obtained by the CCSEM analysis, is shown in Fig. 9 versus inclusion size ranges. The element distribution was almost constant in coal 2. In coal 1, aluminum was very low in larger particles where higher silicon contents were found. Detected amounts of calcium are very low in the mineral inclusions, indicating that most Ca was distributed in the organic matrix.

390

T. Heinzel et al.

Influences on the Changes in Furnace Temperatures To achieve reliable data of furnace temperatures in the tests with the new coals, an acoustic gas temperature measurement system was installed on two furnace levels. Suction pyrometer gas-temperature profile measurements were carried out on different

levels. All measurements, in accordance with the regular operation measurements, displayed that, at equal loads, the combustion of the substitute coals resulted in unexpectedly strong changes of furnace temperatures and related heat transfer. Combustion of coal 3 resulted in increased furnace-exit gas temperatures (FEGT) compared with the reference coal 1. Coal 2 from the local alternative seam reduced the FEGT (Figs. 10 and 11). The heat exchange in the furnace changes therefore, too. Combustion of coal 3

transferred less heat into the radiative section of the furnace. Because the total energy flux remained the same until the end of the last convective heat exchanger was reached, more heat had to be transferred into the convective section, where the main parts of superheaters and reheaters are located. This resulted in a higher amount of water injection into the high-pressure (HP) and medium-pressure (MP) section of the steam cycle

Slagging Tests on the Suitability of Alternative Coals in a 325 MW e PC Boiler

391

and a reduction of live steam in the HP section of the turbine and a raise of steam at the MP and LP turbine re-entries. The effects of this were undesirable changes in steam

cycle operation and worse cycle efficiencies. The analysis of the obtained data showed which influences can be responsible for the measured major changes in furnace temperatures and heat transfer under the prevailing conditions of the experiments. The identified main influences were the changes in heat release due to changed radiation of the combustion products, namely by the solid state radiation of ash particles. Besides that, the changes in the combustion course and the heat release of coal particles (combustion temperature, changes in burner operation parameters) and the reduced heat transfer through the water walls due to deposits, which insulate and change the emis-

sivity of the water wall surface, must have resulted in superposed changes. None of the effects could have been the only one because significant deviations remained in possible regressions for single explanations.

Changes in the Radiative Heat Transfer of the Combustion Products. The changed radiative heat transfer was identified as the main reason for the changed furnace temperatures due to the changed values of the solid state radiation surface in the gas, which is basically related to the higher load of ash particles. In the planning of the experiments the importance of the changes of the heat transfer was underestimated and no specific measurement of heat flux, emissivity and heat release in the boiler were carried out. Despite of this lack of direct proof, both calculations and measured correlations of temperatures and ash contents gave rise to the assumption that the changes in the radiative heat transfer of the combustion products must have been the major reason for the temperature changes. In order to compare the possible influence of the ash, calculations for coals and

blends in the experiments were carried out according to the models described in [VDI, FDBR], using as far as possible data from the single experiments. Air ratios and flue gas compositions were similar in the experiments. For comparison of the influence of ash radiation the same surface emissivity and temperature of furnace walls were assumed. The basic calculations prove that the solid state radiation can be influenced strongly by the ash content. The results for a furnace, which is homogeneously dust-loaded, show that the changes in dust emissivity can result in single solid state heat transfers for the coal changing to +48% and –10%, respectively. When taking the interference with the gas into account this results in temperature changes of approx. –105 K and +22 K, respectively. In the real furnace the combustion is not homogeneous and more complex.

392

T. Heinzel et al.

This approach is supported by the measurements, which show a correlation between the ash content of the coals and the furnace temperatures, shown in Fig. 12. When plotting the mean data of all single experiments for all loads, the dependence both on load and on ash content can be combined in a multiple regression wherein the deviations are expected to be caused by superposed influences and short-time variations in fuel properties.

Reduced Water-Wall Heat Transfer Properties. In the radiation part of a boiler, the surface condition of the water-walls influences the heat transfer and the energy yielded from the combustion gas. The visual observations during the experiments brought about that the least depositions occurred in the experiments with the alternative seam coal. Compared with the regular coal, higher amounts of deposits were detected on the front tubes of the superheaters with coal 3, but in the middle and lower furnace parts only few deposits were observed, so the experienced staff did not worry about. All deposits could easily be removed by the regular water and soot blowers (see following chapter). These trends correlate with the measured temperature changes, but they are rather a result of the temperature changes than a reason for it. This was concluded from the evaluation of the different temperature changes at the beginning and during the experiments. Each experiment started after cleaning the boiler in the night at partial load. When the load was increased, furnace temperatures followed immediately to the temperature level typical for the used coal. But no trend in temperature changes could be detected during the experiments (although the measured temperatures fluctuated within 100K),

shown in Fig. 13. Assuming that the cleaning of the furnace has at least measurable effect and that furnace slag forms continuously during a period of time (and not immediately

Slagging Tests on the Suitability of Alternative Coals in a 325 MW e PC Boiler

393

at the beginning of an experiment with no further change), the very low FEGT of coal 2 would be close to the regular FEGT for a clean furnace. A continuous temperature rise would have taken place if growing furnace deposits would have had a major influence on temperature levels and heat transfer in the experiments. With coals 1 and 3 an FEGT as low as with coal 2 was measured in neither case, and the fact that the switching from the regular coal 1 to coal 3 instantly led to rising the needed water injection to a constant level with minor further changes throughout the experiments proved that the furnace deposits had not been a major reason for the changes in temperature levels. This is supported by the regular observations during the experiments, which showed the large majority of the furnace walls being not covered with slags in all experiments even after many hours of operation (Fig. 14, 18–20 in next chapter).

Shifts in Combustion and Heat Release Due to Fuel Characteristics and Operational Parameters. Tables 5–7 present a brief summary of relevant operational parameters and fuel properties. The total air ratio and the individual combustion air supplies were related to

394

T. Heinzel et al.

the needed stoichiometric air supplies and kept almost constant during the experiments. Particle and velocity measurements in the coal-dust ducts revealed that a higher portion of coal and air was supplied to the lower burner level for coal 2, while the supply for coal 1 and 3 was similar. The influence of unburnt substance in the ashes, with shares of 0.05%, 0.035%, and 0.64%, related to the combustible substance for coals 1–3, was negligible.

Coal 3 was milled finer than coals 1 and 2 (d50 approx. 50 microns and 150 microns, respectively). The mean values of the coals do not strongly differ in the basic fuel properties responsible for ignition, combustion behavior and heat release (Table 6). The previously presented petrographic analysis (Table 2) shows a lower degree of transformation into coal for the alternative seam. Due to minor changes in composition and structure of the combustible substance and depending on the water and ash contents, and by taking into account the real air ratios and parameters of the experiments, the adiabatic combustion temperature of the fuel/air mixture in the experiments varied within 60 K.

Slagging Tests on the Suitability of Alternative Coals in a 325 MWe PC Boiler

395

Neither the changed fuel ratios on the burner levels nor the milling fineness matched the measured changes in heat transfer and temperatures, and cannot account for having caused these changes. Although none of the single effects of the measured fuel characteristics, operational parameters and furnace slagging was capable of explaining the higher temperature changes at the furnace exit, these effects are supposed to have interfered with the changes in the radiative heat transfer.

Boiler Slagging As previously mentioned, the slagging of bulk ash caused trouble in regular operation and in the tests whereas the deposits which formed selectively did not pose problems. The growth rate of deposits formed by the impact of a main ash stream containing viscous particles was much higher than the growth rate of selective deposits, and may result in the build-up of massively thick layers within hours. Therefore this type of slagging is the most serious and must be prevented as a matter of prime importance. Other deposition types of different characteristics were detected to a small extent in the furnace, in front of the superheaters and in the convective path, which, apparently, did not cause troubles though. Routine observations of and sampling from tubes and tube spaces in the superheater section of the furnace showed small amounts of typical and easily removable deposits. No indications for unusual deposits were found in the convective path. On the side of the cooled deposition probe facing the boiler, in front of the superheaters, a bonded sticky particle layer slowly formed [Fig. 5], whereas the side leewards was covered with a thin layer of loose, fine, pale-gray ashes rich in calcium. Deposits in Substoichiometric Zones. In the zone below the tertiary and burnout air input, hard crusty deposits on the furnace walls settled slowly (Fig. 14). On top of the crusty and insulating layer, sintered, less bonded bulk ash deposits formed, very similar to the slag samples. The bulk ash did not stick to the clean tubes, most probably due to cooling down to non-stickiness before reaching the surface. This corresponds to the measurements on the deposit probe, where bulk ash did not stick to the tube surfaces which had a temperature of 620–650°C. When sampling these deposits during full boiler operation, the crusty layer of partly oxidized iron-sulfur compounds from zones with air deficiency was taken out without cooling reaching the ambient air where it emitted SO2. The remaining samples were red or black in color and mainly consisted of iron in forms of hematite and magnetite. The presence of FeS in the coal, the emission of SO2 during contact with air and the com-

396

T. Heinzel et al.

position of the samples after removal fit the theory that the samples mainly consisted of partially oxidized iron-sulfur phases which are known to form slags under reducing conditions [Raask, Zelkowski], what can also be shown by equilibrium calculations. This type of layer was found to build up slowly to a maximum thickness of approximately 5mm but not further. Coals 1 and 3 contain higher amounts of pyrite, possibly leading to more extended and faster formation of the crusty slag layer, but in the tests this could not be proven. Besides the undesirable effect of heat resistance of the layer itself and the higher surface temperatures this deposition type was not found to cause trouble, which is in accordance to the operator’s experience. In addition, after the test series with coal 3

(which contained most pyrite) the furnace was shut down for one day because of a problem not related to the measurements. The furnace inspection from inside during the

shut-down also did not reveal strongly bonded layers on the walls. The regular furnace observations during operation clearly show that the water-wall blowers clear the walls even from the crusty deposits in zones with air deficiency.

Deposits on Cooled Probes. On the 620–650°C tube probe in front of the superheater entrance one probe for all experiments of a coal type was inserted to provide longer residence times. Between experiments the probe was taken out, and one area of the probe was photographed, brushed and washed with distilled water, while other sections were not touched. When coal was switched, the probe was taken out and replaced by the probe for the new coal type. Initial layers formed on the tubes but no bonding of bulk ash mass deposits was found. In an initial period, the tube surface oxidized and a ground layer formed. The deposits, which were found on the tubes after this period of approximately 20 hours, varied strongly between front side and back side. On the front, inertial impactation of larger sticky particles begun and a bonded layer formed slowly due to the impact of large sticky particles, most probably from coarsely ground stones, whereas on the lee side, only fine particles deposited. Figure 17 shows the front side of a tube probe after 20 h of operation, with the highest slagging for coal 1. More than 20 superheater deposition samples were analyzed with regard to soluble alkalis. Their contents, however, as a possible measure of condensation fouling on the superheater probes, did not show great differences with respect to the coal type, or depending on different loads. This coincides with the unchanged low amounts of deposits in the superheater spaces and the further convective section, and can be related to the low alkali contents of all coals. Slagging on Ceramic Slag Samples, on Walls in the Burnout Zone and on Superheaters.

The formation of sintered deposits from the bulk ash on furnace walls, superheaters and, in the tests, on the corresponding ceramic slag sampling probes is, for the high growth

Slagging Tests on the Suitability of Alternative Coals in a 325 MWe PC Boiler

397

rate, much more serious than other deposit formations, especially if there is a transformation into slag. For the case of using brown coals in this boiler, the possible formation of slag is the main obstacle, and was the prime objective of observation. The deposits on the ceramic probes, which were inserted on three levels in the furnace (Figs. 3 and 4), corresponded to the bulk ash deposits found on furnace walls, and both showed variations in furnace slagging among the coals. Deposits in the tests with coal 2 and coal 3 each showed uniform deposition characteristics of the single coals. Few scarcely agglomerated ash deposits were found in all experiments with coal 2 (alternative seam) on samples, water-walls and on the front side of superheaters. A balance between erosion, formation and fall of losses due to gravity resulted in partial water-wall covering and easily removable agglomerated deposit patches. On the front side of the superheaters large, heavily sintered deposits formed, which were removable though to a large extent by soot-blowing and in the sampling procedure (Fig. 18). Variations in the amounts and characteristics of deposits were found in different tests with the regular coal 1. In most of the tests with coal 1 slagging on the ceramic samples was comparable to or less than in the tests with coal 3. In few experiments, even at partial load (260MWel and 290MWel), deposits were heavily sintered and formed molten slag on a ceramic sample at full load. In this experiment (test 2), in parallel, a

398

T. Heinzel et al.

I

slag block formed on the ash hopper chute, which eventually had an area of several square meters and was high up to half a meter. Boiler load was reduced to minimum power and the block was thermally shocked by pumping cooling water with a fire engine to burst it. Parts of the block were sampled from the wet ash discharge (Figs. 26 and 27). The chemical composition of a variety of samples taken of ash, deposits and slag by common analysis, CCSEM and SEM-EDX analyses, and the mineral composition of twelve of the samples analyzed by XRD were determined for the six selected experiments. In addition, more than 100 fusion tests were carried out. The whole of the analyses proved that in the burnout zone the composition of the mass deposits was close to the composition of the bulk ash in all experiments. The main elements Si, Al and Ca remained the predominant constituents of the deposits. Compared with the ashes, lower aluminum content and higher amounts of silicon in most slag samples, except in test 1, were found. In Fig. 19, only the ratios of the main components disregarding all the rest are compared, in order to clarify the shifts. In some slag samples, shares of 10% or even more were undetectable by AAS analysis. Only a selection of samples is therefore presented. SEM-EDX and XRD analysis supported the results, showing a glassy, solidified, partially porous matrix with comparable composition, in which spherical particles were

Slagging Tests on the Suitability of Alternative Coals in a 325 MWe PC Boiler

399

bonded and intercalated, exemplarily shown in Fig. 20. The mineralogical composition of the deposits was analyzed for the slag samples on the ceramic slag sampling probes on level 62m, which are very comparable to the superheater deposit samples. In the tests with coal 1 and coal 3 mainly quartz was found. Other minerals were composed of Ca, Si, Al, S and Fe in forms of gehlenite (Ca2Al2SiO7), pseudowollastonite (CaSiO3), gypsum (CaSO4), anorthite (CaAl2Si2O8), and essenite (CaFe3AlSiO6). In contrast, in the slag sample from the test with coal 3, less quartz but higher amounts of anorthite, some gehlenite, and in addition few mullite (Al[AlSiO5]) and iron oxide were detected, whereas no CaSiO3, CaSO4 and CaFe3AlSiO6 were bound in crystalline forms. CCSEM analysis of the samples from the front of the superheaters did not deviate from these results in general but showed that the distribution of elements depends on

particle size. For all coals the same trend was observed: sulfur, calcium and aluminum

were enriched in the particles of 2–16 microns, whereas much less silicon was found in these particles. The mineral classification using a chemical classification system for CCSEM data is shown in Fig. 21. The identification criteria for the main mineral classes [GEUS] found in the samples is shown in Table 8, whereas the ideal composition of detected minerals by XRD is given in Table 9. The high quartz content was detected with both methods for all coals, and the detection of high amounts of Ca-Al-silicates in the slag sample of coal 2 corresponds to the XRD-peak of anorthite. The differences in the detected compositions concerning the gypsum/aluminosilicate contents and the class of Ca/Si-rich ones is expected to be caused by heterogeneous and partially amorphous phases and in the undefined compositions of the model of about 15%. Compared with the mineral inclusions in the coals (Fig. 8), only quartz remains unchanged, but the Capure kaolinite and montmorillonite classes arc not found anymore, whereas other components formed including the originally organically bound calcium from the coal.

400

T. Heinzel et al.

Bulk Ash Deposits Related to Fusion and Furnace Temperatures. In general, the behavior of the bulk ash, which mainly caused trouble, corresponded to the ash fusion temperatures for all coals and blends. Although standard ash fusion tests (AFT) are regarded as not being capable of predicting depositions, it is obvious that the melting of a deposit sample in a laboratory furnace should correspond to the melting behavior of the same deposit in the boiler. Besides the problem of how to define the fusion of a heterogeneous sample in a laboratory furnace, one major problem with AFT is to achieve samples representative for deposits, or at least to know their composition. Fortunately, regarding this specific boiler, the critical mass depositions were identified to be non-selective, and samples from the boiler were available and representative. In consequence, the comparison of bulk ash fusion temperatures with the actual furnace temperatures could

explain the slagging behavior of the coals in a simple way. When the boiler temperature was in the range of or higher than the fusion temperatures detected in the laboratory, heavily sintered deposits or slags occurred in the boiler, whereas very low agglomeration took place when the ash fusion was above the furnace temperatures in the upper furnace.

Figure 23 compares the melting ranges of three deposition samples (namely the slagging probes at levels 42m and 62m and the samples taken directly from superheaters) with the mean flue gas temperature ranges on level 45m and 62m (derived from standard deviations of the acoustic pyrometer measurements of all full load experiments with the specific coal type). The continuous lines represent the shrinkage of the samples in the laboratory furnace (mean values of multiple samples). In the tests with coal 3 (upper chart) the temperatures in the furnace had been in the range of the melting of deposits. The share of sticky particles was high, resulting in the previously described sintered deposit formation. The furnace temperatures in the experiments with coal 2 from the alternative seam (second chart) were much below the fusion range of bulk ash and deposits. For this reason the fly ash could not form sintered deposits. Of course, a few lower melting particles

Slagging Tests on the Suitability of Alternative Coals in a 325 MWe PC Boiler

401

sticked on the surfaces but merged with the predominating non-sticky particles they could not form strong bonding deposits. In the experiments with the two blends (third and fourth chart) the temperatures were higher compared with those utilizing coal 2, and the fusion temperatures of the deposits were shifted according to the fuel mixture. The span between furnace and fusion temperatures was smaller. In consequence, more sticky particles formed. The lower two charts represent the two tests with coal 1, the non-slagging test 1 on the bottom and the slagging test 2 second lowest. Within the same furnace temperature ranges, all samples of test 1 melt at the upper temperature limits, but the samples of slagging test 2 melt at reduced temperatures in the mean range of furnace temperatures, corresponding well to the identified real slagging. The extraordinary slagging of coal 1 in test 2 was caused by the temporary high slagging ash with reduced ash fusion temperatures and strongly variable ash chemistry (see Chapter “Fuel analysis”). These changes in fuel quality are obvious in the changing base/acid-ratio from 0.29 to 0.89 and the changing silica ratio from 0.44 to 0.62 of different ash samples from the fuels sampled in this experiment.

402

T. Heinzel et al.

But there had been an extraordinary operating condition in one of the mills, which influenced the slag block formation on the spot shown in Figs. 24a and b. The mill located nearest to and heading towards the spot where the block formed was operated in an unusual way at maximum fuel supply without adequate mill speed, resulting in a higher fuel load of the transport gas and with too low a secondary air supply at that burner.

CONCLUSIONS The operation of the plant in the tests worked well with all coals and blends, except for one extraordinary case with the regular coal. The operation with the preferred sub-

Slagging Tests on the Suitability of Alternative Coals in a

PC Boiler

403

stitute coal from the alternative seam did not lead to any problems in the boiler. The sulfur content of approximately 3% resulted in a very high load of the desulfurization plant, which is expected to be the main restriction for long-term operation with this coal. The coal from an alternative seam was initially expected to contain 2 to 8% of

alkalis in ash, which was predicted by earlier test-boring samples, hence entailing a high risk of furnace slagging. The coal in the tests did not contain such high amounts of alkalis, so no influence on furnace deposits by alkali species was detected. Furnace temperatures were effected substantially by the type of coal. In comparison to the regular lignite, the third coal tested brought about increased temperatures at the entrance of the convective path, while much lower temperatures at the entrance of the convective path were measured in the tests with the alternative-seam coal. This is put down to the changed radiation of the higher ash amounts in this coal, to the fuel properties, which cause lower combustion temperatures, to operational parameters and, finally, to less deposition resulting from this coal. The strength of furnace slagging differs from coal to coal. The new-seam coal left the cleanest furnace which corresponds to few fragile ash deposits on the probes. Strong sintered deposits occurred on slag probe samples and on the superheaters in the experiments with coal 3 (external area), which still could be cleaned by soot-blowing. In the experiments with the regular coal, variable intensities of slagging were registered. Deviating from the common average medium slagging, one experiment revealed strong slagging occurring simultaneously on the ash hopper chute and on the probes, forcing a load reduction for removal. The main reason for this is the unusually changing quality of coal 1, due to the small and heterogeneous seam, which resulted in temporarily unfavorable ash compositions. Therefore the ash fusion temperatures occasionally were below the

actual temperatures in the boiler and resulted in molten ash.

There was an impact on slagging of a variety of fuel properties and operational parameters. The changed temperature levels in the boiler, for instance, strongly influenced the bulk-ash slag formation. Although this effect was unusually pronounced in this case, the real boiler temperatures must be taken into account in the slagging evaluation in general. Acoustic temperature measurements turned out to be a very helpful tool for the interpretation of depositions, because suction pyrometers cannot provide continuous flue-gas temperature measurements simultaneously at different ports. The growth rate of bulk ash deposits was much higher than the growth rate of other selective deposits, and it can build up massively thick layers within hours. For this reason, this type of slagging is the most serious, and must be prevented as a matter of prime importance. The comparison of furnace temperatures with ash fusion temperatures provides reasonable information about the bulk-ash slagging which cannot be achieved so clearly by the other methods applied.

ACKNOWLEDGMENTS The IVD wishes to gratefully acknowledge the power plant owner BKB for the commission of the measurements, and the management, staff and laboratory of BKB for the excellent co-operation. The authors further wish to acknowledge the subcontractors: Prof. Spliethoff and Klaus Blug from the University of Saarbruecken carried out the acoustic temperature measurements, Mrs. and Mr. Werner from mpa GmbH carried out SEM-EDX and XRD analyses, Henning Soerensen from Geological Survey of Denmark

404

T. Heinzel et al.

and Greenland carried out the CCSEM analyses and W. Pickel from the University of Aachen carried out the petrographic analysis.

REFERENCES Raask, E. (1985). Mineral Impurities in Coal Combustion. Washington: Hemisphere Publishing Corporation. V D I (1984). VDl Wärmeatlas. Düsseldorf: VDI-Verlag.

FDBR (19••). FDBR Handbuch Wärme- und Strömungstechnik. Düsseldorf: FDBR. Laursen, K. (1997). Advanced Scanning Electron Microscope Analysis at GEUS. Kopenhagen: Geological

Survey of Denmark and Greenland, Rapport 1997/1. Zelkowski, J. (1986). Kohleverbrennung, Brennstoff, Physik und Theorie, Technik. Essen: nik, Volume 8 of the series “Kraftwerkstechnik”.

VGB-Kraftwerkstech-

PREDICTING SUPERHEATER DEPOSIT FORMATION IN BOILERS BURNING BIOMASSES Rainer Backman, Mikko Hupa, and Bengt-Johan Skrifvars

Department of Chemical Engineering Åbo Akademi University, Finland

INTRODUCTION Deposits formed on superheater surfaces during combustion of biofuels alone, or mixed together with other fuels, consist mainly of carbonates, sulfates and/or chlorides of calcium, magnesium, potassium and sodium originating from the fuel. In some cases

this kind of deposits can grow to a thickness of several centimeters (Hupa, 1982). If silica rich fuels as coal or peat are co-fired together with the biofuel, various amounts of Si-compounds (silicates, glassy material) can be present in the deposit. The composition of superheater deposits differs considerably from the inorganic composition of the fuel ash (Hupa, 1979 Miles; 1995). Thus, the fouling tendency cannot be studied by testing the fuel ash alone but measurements of deposit compositions in situ are required. Such measurements show also that the composition in a deposit varies with the distance from the metal surface. Sodium and potassium concentrations are higher near the tube surface, whereas calcium concentrations are higher at the outer surface (Hupa, 1979). Concentrations of chlorides are higher at the tube surface, indicating their role in the initial stage of the deposit forming process (Hupa 1982). The transport of inorganic material to the metal surface differs from boiler to boiler and depends, besides on the fuel mix, on many factors, including location of the heat transfer surfaces, gas flow velocity near the surface, gas temperature, heat radiation and heat flow. Bigger, inorganic ash particles strike the surface by impaction and can stick to the surface causing build-up of a deposit. Submicron particles are transported by diffusion or thermophoresis processes. Volatile compounds as alkali chlorides can condense directly from the gas onto the surface. Therefore the composition and quality of deposits on the side of a tube directed towards the fluegas flow (windward side) often differs considerably from those on the leeward side of the tube. The properties of deposits is dependent of the melting behavior of the material in the deposit. Salt mixtures usually have a wide melting range where liquid and solid phases

can coexist. The difference between the temperature where the first melt forms, T0, and Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

405

406

R. Backman et al.

the point where the last solid phase disappears, T100, can be several hundreds of centigrades, depending on composition. Inside this range two characteristic temperatures are of particular interest, the sticky temperature, Ts, and the flow temperature, Tf. We define the sticky temperature as the temperature above which the material stick to a metal surface when it strikes it. This happens when the amount of liquid phase in the material exceeds a certain value. For alkali and earth alkali salt mixture, where the viscosity is not a strong function of the composition, a deposit become sticky when it contains 15% melt (Isaac, 1984). For deposits containing silicates the stickiness criterium is a function of the temperature (Skrifvars 1994). The flow temperature is defined as the temperature where the amount of liquid is big enough for the deposit to flow down along a vertical superheater tube. For alkali and earth alkali salt mixtures this occurs when about 70% of the deposit is molten. Again, for silica rich deposits the flow temperature will be dependent, not only of the amount of melt, but also of the composition of the melt due to the strong relationship between composition and viscosity in such systems. The flow temperature determines how thick a deposit can grow on a vertical superheater tube. When the outer surface of the deposit reaches the flow temperature, no further growth occurs, because the excess material flows down from the tube. We call this thickness the steady state thickness. The time needed for a deposit to reach steady state is dependent on the transport of material to the surface, which is a function of particu-

late load in the flue gases and the size distribution of particles as well local temperature and flow patterns.

A steady-state deposit is schematically shown in Fig. 1. The temperature gradient in the deposit is determined by heat flux through the deposit and the heat conductivity, but is approximately linear between the metal surface temperature and Tf, the outer surface temperature of the deposit. The position of the first melting temperature, T0, in the deposit is indicative for the risk for corrosion of the tubes. If T0 is close to the metal surface, the risk for contact between melt and metal increases, leading often to very rapid general molten phase corrosion. The sticky temperature does not have any particular significance in a steady-state deposit. Its role is in the initial deposit formation mechanism where it determines whether or not a deposit will grow at a certain location.

Predicting Superheater Deposit Formation in Boilers Burning Biomasses

407

The thickness of a superheater deposit can be estimated by the following simplified expression:

where

d = deposit steady state thickness (mm) = heat conductivity of deposit (W/mK) = heat transfer coefficient Tf = deposit flow temperature (°C)

Tm = metal temperature (°C) TK = flue gas temperature (°C). The influence of the flow temperature and the heat transfer coefficient on the steady state thickness of deposits are demonstrated in Fig. 2, where the deposit thickness is plotted as a function of flue gas temperature for different local conditions. This model has been verified for kraft recovery boiler superheater deposits (Backman, 1987) where probe measurements in two kraft recovery boilers showed good agreement with calculated steady state thicknesses.

ESTIMATING THE MELTING BEHAVIOR OF SUPERHEATER DEPOSITS Mixtures of salts melt stepwise in a temperature range where the difference between the temperature of the first appearance of melt and the temperature of complete melting can be several hundred degrees. The melting behavior of a mixture with a known overall composition can be shown by a melting curve, i.e. the amount of melt present as a function of the temperature. To construct such a curve, all the phases that can appear in the system must be known. Also, information about how these phases interact with one another in the melting process must be available. These basic data can only be obtained experimentally. However, the use of fundamental thermodynamic phase theory and modelling of non-ideal mixture phases offer a way to effectively utilize all available experimental data. Also, first approximations can be done of the melting behavior into composition regions where experimental data are not available.

In the following we present a method to estimate the melting curves in salt mixtures with compositions similar to deposits formed on superheaters in boilers firing biomass fuels, that is carbonates, sulfates, sulfides, and chlorides of calcium, magnesium, potassium and sodium. The method is based on a theoretical routine for estimation of melting curves of salt mixtures in the system The method, called MELTEST has been developed for determination of fouling tendencies in kraft recovery boilers (Backman, 1995). In the present work it has been extended to cover also sulfates of calcium and magnesium. The method is based partly on experimental data from the literature, partly on recently published thermochemical evaluations of binary systems. All parameters in the model are thermodynamically consistent, which make extrapolations into higher order systems possible.

408

R. Backman et al.

Predicting Superheater Deposit Formation in Boilers Burning Biomasses

409

The System The melting points for the pure salts in the system range from 771°C for KCl to 1,180°C for The thermodynamic data used in this work are based partly on standard thermodynamic tables (Barin, 1973, 1977), partly on critical evaluations of phases data from the literature (Sangster 1987, Backman 1989). The two-component systems for the salts can be divided into 12 common-cation and 4 common-anion binaries. The phase relationships in the binary systems for carbonates, sulfates and chlorides have been extensively studied experimentally. The phase diagrams are collected in Phase Diagrams for Ceramists (Amer. Ceram. Soc. 1972–1992). The binaries including sulfides are not very well known. Published data are available for (Tegman, 1972), (Andersson, 1982), (Shivgulam 1979), (Babcock, 1988) and (So, 1981). A tentative study by differential thermal analysis of the system has been done recently (Mäkipää, 1998). In Table 1 are listed all possible binaries in the system together with the type (solid solution or eutectic) and the lowest temperature at which melt occurs. All common-anion systems contain solid solutions with minimum azeotropic behavior. The common-cation binaries show eutectic melting behavior except for the carbonate-sulfate systems, where a hexagonal solid solution forms in both the sodium and the potassium system. The system has not been studied experimentally, but it is assumed to be eutectic in analogy with the corresponding sodium system, where experimental data are available. Limited solid solubility has been reported also for the eutectic systems, especially those containing sulfide (Goubeau, 1938; Andersson, 1982; Tegman 1972 So, 1981; Winbo, 1997). However, quantitative data on the solubility ranges are not available in the literature. The binary with the lowest eutectic temperature is where melt appears at temperatures above 600°C. Of the higher order systems six have been partly investigated. Experimentally deter-

410

R. Backman et al.

mined first melting temperatures are in the range 518°C (sulfate-chloride) to 622°C (carbonate-sulfate-chloride). The thermodynamic model we use to estimate the melting curves for the complex salt mixtures is a sublattice ionic model with equivalent site fractions (Pelton 1988). In this model, the multicomponent liquid and the solid solution phases are assumed to constitute of two sublattices, one with metallic cation sites and one with anion sites. It is further assumed that sodium and potassium ions mix randomly on the cation sublattice and that the carbonate, sulfate, sulfide and chloride ions mix randomly on the anion sublattice. Furthermore, the numbers of cation and anion sites are equal to the numbers of cations and anions, i.e. no empty sites, vacancies, are allowed. All binaries for sodium and potassium carbonates, sulfates and chlorides have recently been critically optimized by the sublattice model (Sangster, 1987). The nonideality is devoted both to liquid and solid phases. They were able to reproduce all liquidus curves with an accuracy between and Solidus curves had an accuracy between and compared to the experiments. The binaries and was optimized with the same model (Backman 1989) with an estimated accuracy of Our method to calculate the melting curves for the multicomponent mixtures in this work is based on an extrapolation of the binary behavior into higher order systems and adding two ternary terms, which are adjusted according to experimental data in the

system . In this way the total number of non-ideal parameters in the system is 32. The solid phases assumed to be in chemical equilibrium with the liquid are the solid solutions mentioned earlier.

The System Pure CaSO4 melts at 1,462°C and pure Together they form an eutectic system with a lowest melting point of 900°C. The system contains an intermediate compound, which melts at 900°C. Together with potassium sulfate both calcium and magnesium sulfate form eutectic systems containing double salts, viz. called calcium langbeinite, melts incongruently at 1,011°C. Magnesium langbeinite melts congruently at 930°C. The eutectic temperature in the binary is 875°C and in the binary 746°C. The lowest melting temperatures in the system are around 900°C. Formation of an intermediate compound and solid solutions are possible. In the system an intermediate phase, melts congruently at 820 °C. The eutectic temperature in this system is 660°C. In this work the liquid in the -system has been modeled by a sublattice model similar to the one used for alkali salts. The four metal ions are introduced on a cationic lattice with one single sulfate ion on the anionic lattice. Non-ideal interaction coefficients are introduced for the - system and one interaction coefficient between Solid phases considered are , hexagonal solid solution of and The latter phase is a hypothetical one introduced in order to get the first melting point high enough in the sodium rich part of the system. The lowest calculated melting temperature in the ternary In the quaternary system the lowest melting temperature is 615°C. This model reproduces the calcium rich the binary rather well, within

Predicting Superheater Deposit Formation in Boilers Burning Biomasses

411

from experimental values. On the potassium rich side the deviations are bigger due to difficulties to model the solid solubility of The accuracy for the

other systems are lower due to lack of thermodynamic for the double salts of magnesium and sodium sulfates. Thus, the validity of this model in restricted to compositions where the magnesium and sodium sulfate concentrations do not exceed 20wt-% respectively.

The System In this work the data for the alkali melt and the earth alkali sulfate melt have been combined to one melt containing four cations and four anions. The ionic sublattice model is used to describe the melt. It is assumed that the components are miscible in each other in the whole composition range, i.e. only one melt is formed. All together 32 interaction

parameters are used to describe the non-ideality of the melt. For the earth alkali metals only molten sulfates are taken into considerations. The constitution of the system is given in Table 2. No systematic experimental studies on the melting behavior of salt mixtures containing all the eight components have been done. Thus, extrapolations into higher order systems are uncertain and cannot be verified experimentally. Because formation of earth alkali carbonates, sulfides and chlorides are not considered in the model, it can be used to estimate melting properties of mixtures rich in sulfates. This is usually the case for superheater deposits when biomass fuels are co-fired with oil. In firing systems with little or no sulfur, calcium carbonate is formed. The melting behavior of carbonate melts of calcium, potassium and sodium is not known and is not considered in the present model. Calcium carbonate decomposes into calcium oxide and carbon dioxide at temperatures above some 770°C. Alkali carbonates do not decompose at oxidizing conditions and can exist at temperatures above 1,100°C. Thermodynamic phase stability analysis shows that

412

R. Backman et al.

alkali chlorides are more stable than earth alkali chlorides at temperatures typical in the superheater area. Thus, the assumption done in the present model that all chlorine is bound as alkali chlorides is reasonable, at least at relatively low chloride contents.

EXPERIMENTAL Method Deposit samples were taken from four boilers using air-cooled probes (Hupa 1978). Two boilers were kraft recovery boilers and two were boilers firing wood waste together with oil. Data for the boilers are given in Table 3. In all boilers samples were taken before the superheater section at flue gas temperatures of 850–1,000°C. The surface temperature of the probe at the location of a removable sample ring was 480 °C in the kraft recovery boilers and 525 °C in the wood waste boilers. Sampling times were 3 h in the kraft recovery boilers and 42 h in the wood waste boilers. After exposure in the boiler, the probe was cooled rapidly and the thickness of the deposit was measured. The recovery boiler samples were analyzed for Na, K, Cl, and The wood waste boiler samples were analyzed for Ca, Mg, K, Na, Fe, Al, Mn, Si, and S. The probe ring area was also studied by means of SEM/EDXA.

Results The deposit samples from the two kraft recovery boilers had both reached their steady-state thickness during the 3 hour exposure time. It could clearly be seen that the outer surface of the deposit had been molten. The thickness of the deposit was 6.5mm in Boiler#l and 3.1 mm in Boiler#2. The deposit samples in Boiler#3 and Boiler#4 did not reach a steady-state thickness during the 42 hours exposure. The outer surface of the deposit was rough and contained in both cases mainly solid particles. The thickness of the deposit was about 5mm in Boiler#3 and about 2mm in Boiler#4. From these both boilers also long-term samples (about four weeks) were taken from the superheater tubes. These samples had a rather smooth outer surface, indicating presence of melt. This could be verified by SEM studies. The long-term sample from Boiler#3 was about 25mm thick and that from Boiler#4 about 5mm. The chemical analysis of the probe deposit samples are given in Table 4 recalculated to alkali and earth alkali sulfates, carbonates and chlorides. The part indicated “Inert” contains oxides of Al, Fe, Mn, Si, and the analytical error. The deposits from the kraft recovery boilers are mixtures of alkali carbonate, sulfate, and chloride. Ca and Mg compounds are present in very low concentrations in kraft black liquor. Superheater deposits in kraft recovery boilers sometimes contain

Predicting Superheater Deposit Formation in Boilers Burning Biomasses

413

sulfide (Backman, 1987), caused by local reducing conditions and the presence of elemental carbon in the deposit. In these two samples, no sulfide or elemental carbon was detected. The main difference between deposits from Boiler#1 and Boiler#2 are in the potassium and chlorine content. Neither of the deposits from Boiler#3 and Boiler#4 contains chlorine. The deposit from Boiler#3 has significantly higher content of sodium and magnesium. In both case alkali and earth alkali metals are present as sulfates due to the high sulfur input from the fuel oil.

Melting Behavior The melting curves, i.e. the portion melt as a function of the temperatures have been calculated for the four deposits using the melting model presented above. The compositions given in Table 4 are used as input for the calculations. Calculations were done using the computer program ChemSage (Eriksson, 1990). The results are given in Fig. 3. In the

414

R. Backman et al.

deposits of the kraft recovery boilers (#1 and #2) melt is formed at 560°C. The amount of melt formed at this temperature is 8% for Boiler#l and 2% for Boiler#2. Complete melting occurs at 800°C.

The deposits from Boiler#3 and Boiler#4 has a first melting point at 620°C. In both cases about 20% melt is formed at this temperature. The deposit from Boiler#3 reach 85% melt at 1,000°C and that of Boiler#4 at 1,200°C. The portion not melting is the

part considered as inert. The characteristic temperatures, i.e. the sticky temperature and the flow temperature, were determined from the melting curves. A comparison between these temperatures is given in Fig. 4.

CONCLUSIONS The deposit formation process in the superheaters of boilers burning biomass is strongly related to the melting properties of the depositing material. Knowing the melting range makes it possible to estimate the degree of deposit problems in various parts of the supert heaters. We have developed a melting prediction model which can be applied to calculate the amount of liquid phase in any alkali salt mixture with known composition. This melting prediction can be used as a tool to predict in which locations in the flue gas duct ashes of different compoistion are expected to cause deposit problems. We have introduced the terms sticky temperature, and flow temperature to give simple parameters for the depositing material. Sticky temperature is the temperature below which the salt mixture contains less than 15% molten phase, and flow temperature the temperature above which the share of the liquid phase is largher than 70%. In our approach we assume that the material does not stick on surfaces at flue gas

Predicting Superheater Deposit Formation in Boilers Burning Biomasses

415

temperatures below the sticky temperature of the depositing material. The sticky temperature of the depositing material in Boiler#l was around 665°C and in Boiler#3 and

Boiler#4 around 620°C. In Boiler#2 the sticky temperature was clearly higher, around 770°C. In alkali salt mixtures the sticky temperature can be as low as around 510°C, if the potassium and chloride contents are high. The flow temperature is an interesting parameter, too, because it gives the final, stady state thickness of the deposit. This steady state thickness can be reached if the flue gas temperature is higher than the flow temperature of the depositing material. Boiler#4 shows an example of an alkali ash with a “wide” melting range: a low sticky temperature but still a high flow temperature. This type of an ash has a very difficult behavior in the flue gas channel. In this boiler deposits are formed throughout the superhetares down to flue gas temperatures at some 620°C. Further, there will be no limiting thickness even at the hottest locations, because the flow temperature always exceeds the flue gas temperature in the superheaters. On the other hand, Boiler#2 shows a very “narrow” melting ash. In this boiler the ash is sticky at flue gas temperatures above 770°C, but at higher temperatures the deposit growth will be effectively limited by the flow temperature. Consequently, for this ash the deposit problems will be concentrated in a very narrow flue gas temperature interval.

REFERENCES Andersson, S., (1982), “Studies on Phase Diagrams Na2S-Na2SO4, Na2CO3-Na2S-Na2SO4, Na2CO3Na2SO4-NaOH, and Na2CO3-Na2S-NaOH”, Chemica Scripta, 33, 164–70.

Babcock, K., Winnick, J., (1988) “Solid-Liquid Equilibria in the Reciprocal Ternary System K,Li/S,CO,”, J. Chem. Eng. Data, 33, 96 98.

Backman, R., Hupa, M., Uppstu, E., (1987) “Fouling and corrosion mechanisms in the recovery boiler superheater area”, Tappi J., 70(3), 123–127. Backman, R., (1989), “Sodium and Sulfur Chemistry in Combustion Gases”, Thesis, Report 89–4, Combustion Chemistry Research Group, Åbo Akademi University, Finland. Backman, R., Skrifvars, B-J., Hupa, M., (1995) “Flue gas chemistry in recovery boilers with high levels of chlorine and potassium”, Proc. 1995 International Chemical Recovery Conference, Toronto, A95–A103. Barin, I., Knacke, O., Kubaschewski, O., “Thermochemical properties of inorganic substances”, Springer Verlag 1973, 1977. Cook, L. P., McMurdie, H. F., (editors), “Phase Diagrams for Ceramists, Volume VII”, The American Ceramic Society, Inc. Westerville, Ohio, 1989.

Eriksson, G., Hack, K., (1990), “ChemSage-A Computer Program for the Calculation of Complex Chemical Equilibria”, Trans. Metal. Trans. B21B, 1013–1023. Goubeau, J., Kolb, H., Krall, H. G., (1938), Z. Anorg. Allg. Chemie, 236, 45–56. Hupa, M., Eriksson, B-E., Klingstedt, G., (1978) “A rapid method for investigation of ash deposits in boilers”, Paper and Timber, 60. 719 723. Hupa, M., (1979) “Ascheprobleme bei der Verbrennung von Holzrinde mit öl“ VGB Kraftwerkstechnik, 59, 568–576. Hupa, M., Backman, R., (1982) “Slagging and Fouling during Combined Burning of Bark with Oil, Coal, Gas and Peat”, Proc. 1982 Engineering Foundation Conference about Fouling in Boilers, 419–432.

Isaak, P., Tran, H., Barham, D., Reeve, D.. (1986), “Stickyness of Fireside Deposits in Kraft Recovery Units”, J. Pulp and Paper Science, 12(3), J84–J92. Miles, T. R., (1995), “Alkali Deposits Found in Biomass Power Plants. A Preliminary Investigation of Their

Extent and Nature” Summary Report, NREL. Mäkipää, M., Backman, R., (1998) “Corrosion of floor tubes in reduced kraft smelts: Studies on effects of chlorine and potassium”, Proc. 9th International Symposium on Corrosion in the Pulp & Paper Industry, Montreal. Pelton, A. D., (1988), “A database and sublattice model for molten salts”, Calphad, 12(2), 127–142.

416

R. Backman et al.

Sangster, J., Pelton, A. D., (1987), “Critical Coupled Evaluation of Phase Diagrams and Thermodynamic Properties of Binary and Ternary Alkali Salt Systems”, American Ceramic Society, Westerville, Ohio, 1987. Shivgulam, N., Barham, D., Rapson, H., (1979) “Sodium chloride, potassium: their effects on kraft smelt”, Pulp and Paper Canada 80(9), 89–92. Skrifvars, B-J., Hupa, M., Backman, R., Hiltunen, M.,(1994), “Sintering mechanisms of FBC ashes”. Fuel, 73, 171–176. So, C, W., Barham, D., (1981) “The system K 2 SO4-K 2 S”, J. Thermal Anal. 20, 275–280. Tegman, R., Warnqvist, B., (1972) “On the Phase Diagram Acta Chem. Scand. 26(1), 413–414. Winbo, C., (1997) ”Structural and therochemical studies of double carbonates in the system and solid state emf measurements in the system“, PhD Thesis, Ume Universitet,

Sweden.

DEPOSIT FORMATION IN THE CONVECTIVE PATH OF A DANISH CFB-BOILER CO-FIRING STRAW AND COAL FOR POWER GENERATION Peter F. Binderup Hansen I/S Midtkraft Energy Company, 8541 Skødstrup, Denmark

1. INTRODUCTION The Danish government has committed the national power companies to reduce

the emissions from Danish power stations by 20% based on the level of 1988 before the year 2005. In order to meet this goal Elsam (the Jutland, Funen Energy Consortium) has proposed a biomass program, for combustion of neutral straw, according to which three boiler concepts will be investigated: combustion of 100% straw in a separate stoker fired boiler build as an addition to an already existing PC-boiler, co-combustion of coal and straw in a utility type boiler [Hansen et al., 1996; Andersen et al., 1997, Frandsen et al., 1997] and co-combustion of straw, wood chips and coal in a circulating

fluidized bed boiler. During the combustion process, the inorganic constituents of the fuel, such as Fe, Si. Al, Mg, Ca, K, and Na, are transformed into ash. The ash may be roughly divided into two ash fractions; a residual ash (mainly containing transformed minerals with a particle size ) and sub-micron ash or aerosols (mainly including vaporized minerals with particle smaller than 1 µ m), [Benson et al., 1993]. In PC boilers approximately 1 wt% of a coal ash will vaporize [Flagan and Friedlander, 1978]. In CFB-boilers the combustion temperatures are some 300–600°C lower than in a PC-boiler. Therefore the evaporation of inorganic minerals, ie. the aerosol formation, may be reduced compared to PC-combustion. However, during straw combustion as much as 30–75 wt% of the inorganic constituents in the ash will vaporize [Livingston, 1991, Baxter, 1993; Olanders and Steenari, 1995] mainly due to high concentrations of simple salts such as KC1 in the straw. Combustion of bituminous and sub-bituminous coal in CFB-boilers with in-situ sulfur capture by limestone, generally pose no real problems with respect to superheater deposit formation. A typical bulk chemical analysis of an upstream superheater deposit Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

417

418

P. F. Binderup Hansen

could be: 30wt% 10wt% 18wt% CaO, 20wt% FeO and, 13wt% [Skrifvars et al. [1996]. For a compari-son deposits collected from the final superheater of a CFB-boiler co-firing 50% straw and coal contains 13–16wt% SiO2, 8–10wt% 8–10wt% CaO, 28–32wt% and 30–35wt% Hansen [1997]. Moreover, as much as 90wt% of the K is found to be water soluble, probably in the form of It could therefore appear that the superheater deposit formation in a CFB-boiler co-firing straw and coal will be largely affected or even controlled by aerosol formation of simple Ksalts.

2. BACKGROUND FOR THE INVESTIGATION In January 1992 an CFB-boiler, designed for co-combustion of coal and straw was commissioned by the Danish energy company l/S Midtkraft. The boiler was designed to operate with 50% straw and 50% coal based on thermal input. During the initial 8–9 month of operation, the boiler load never exceeded 80% and the system operated without problems. However, when the load was increased to 100%, the temperatures in the top of the furnace and in the cyclones increased to or even exceeded 1,000°C. The extreme temperatures combined with high concentrations of K

and Cl from the straw let to severe fouling of the cyclones and superheaters which

again necessitated several unscheduled boiler shut-downs to clean the plant. During the summer of 1993, only 18 month after commissioning, the final superheater had to be replaced due to high temperature Cl induced corrosion. Moreover, a high sulfur US coal (2–2.5 wt% S) was replaced by low sulfur coal (<1 wt% S) and a better limestone was selected. This change caused a drop in the limestone consumption by 80–90wt%, whereby a potential fouling agent was significantly reduced. However, deposit formation and high temperature corrosion remained a serious problem. Therefore, an extensive test programme to elucidate on the high temperature corrosion during coal-straw cocombustion was initiated, [Henriksen and Hansen, 1995, Henriksen et al., 1995, Hansen et al., 1996].

From an operater point-of-view, high-temperature corrosion may be dealt with by designing the combustion system so the most vulnerable superheater tubes are rapidly replaced, and include this extra expense in the operating costs. However, with straw combustion severe deposit formation and potential blockage of the convective path may be very difficult to control and unscheduled stand still periods may prove fatal for the boiler economy. Moreover, as high temperature corrosion of the superheaters is often closely related to the composition of the superheater deposits, a research project aimed at disclosing the impact of boiler load, straw share, and coal type on the deposit formation rate and composition was initiated. The result of this work has been described by Hansen [1997]. In this paper, the effect of the coal type on deposit formation in the convection pass are discussed based on probe samples. Subsequently, the probe deposits are compared to mature deposits and the differences evaluated.

3. BOILER AND PROBE DESCRIPTION An CFB-boiler designed for co-combustion of coal and straw was commissioned by I/S Midtkraft in January, 1992 in the Danish city of Grenaa.

Deposit Formation in the Convective Path of a Danish 80 MWTH CFB-Boiler

419

3.1. The CFB-Boiler The boiler is an Ahström (Foster Wheeler) Pyroflow cogeneration plant designed for 0–60% straw and 40–100% coal on an energy base. The maximum steam production is 29 kg/sec, at 92 bar/505°C at the boiler outlet. The furnace is 19.6m high, with a cross section of 4m x 6.2m. The bed height is 3.2m giving a total riser height of 22.8m. As shown in Fig. 1 the boiler is equipped with three superheater sections. Superheater 2 is situated in the furnace 11.7m above the distributer plate followed by superheater 3 and superheater 1 both located in the convective path. Additional evaporator tubes were installed in the

furnace during the second revision period in the summer of 1993 with the goal of keeping the temperatures at the top of the furnace below 900°C.

3.2. Deposit Sampling Position Deposits were collected in a position just upstream from the screen tubes as shown in Fig. 2. Following the screen tubes are the superheaters 3 and 1.

3.3. Probe Design And Probe Handling Air/water cooled probes, developed for long term testing for high temperature corrosion by Elsam and I/S Midtkraft, were used. Two test elements, each 100mm long, are mounted at the tip of the 3 m long deposit probe. The typical sampling periods used were 1 and 3 hours. The metal temperature was fixed at 525°C. After each test, pictures were taken of the test elements. They were subsequently stored in sealed plastic containers for further analysis. Typically the deposit from one test element was to determine the formation rate and the bulk chemical composition whereas the second test element was used for

SEM/EDX-analysis.

420

P. F. Binderup Hansen

4. FUEL COMPOSITION During the study of coal/straw co-combustion two coal types, A and B, and wheat straw from the summer of 1995 was tested. The Danish summer of 1995 was very dry giving straw with high concentrations of K, Cl, and ash (Sander, 1997). The key data are listed in Table 1. The data reveals significant differences between the two coal types in volatile and ash contents. The straw differs from the two coals in volatile content, ash content, and heating value. Furthermore, the straw contains high concentrations of K and Cl. Fuel ash compositions will be presented later.

5. EXPERIMENTAL CONDITIONS During all tests discussed in this paper, 50% straw share (based on thermal input) has been used. The boiler load was varied between 94 and 99% during the co-firing tests with straw and coal A. The load was kept around 90% when co-firing with coal B.

Deposit Formation in the Convective Path of a Danish

CFB-Boiler

421

The flue gas temperatures, in the convection path, just upstream from the sampling position, was in the 875–890°C range during the tests. All SO2 concentrations were measured in the stack. The average SO2 concentration level was in the 210–240 ppmv range (normalized to 4% O2, in dry air) independently of the coal type.

6. RESULTS 6.1. Coal A The surface of a three hour deposit from the convective pass co-firing straw and coal A, is shown in Fig. 3. A close-up picture of the same deposit is presented in Fig. 4. From Fig. 4 it appears that even though the deposit may appear compact for the eye, a highly porous structure with a surface resembling corals or lung tissue is formed. The photo also seems to indicate that new material is predominantly deposited—possibly by thermophoresis—or condensed at the tip of the fine “fingers” (0.5–1.0µm) pointing outwards from the surface of the deposit. Subsequently, the deposit seems to sinter forming a more dense structure towards the metal surface. It is not possible to distinguish fly ash particles in the deposit.

422

P F. Binderup Hansen

If the deposit shown in Figs. 3 and 4 is cut through, as done for SEM-analysis, the deposit reveals a structure of varying thickness and porosity as shown in Fig. 5. The test tube metal is shown as the white reflection at the top of the picture. The deposit reveals a granular surface where the “grains” appears more dense compared to the surrounding deposit. The dense structure seems, sometimes, to stretch, like fingers, from metal tube to the deposit surface. The average porosity of the deposit is around 30% [Hansen, 1997]. In Fig. 6 a close-up picture of the deposit in Fig 5 is shown. The marked areas indicate where elemental analysis by EDX have been performed.

6.1.1. Chemical Composition. In Table 2 the bulk chemical compositions of upstream and downstream deposit are compared to the fly ash as well as fuel ash compositions during co-combustion of straw and coal A. For a comparison the fly ash composition for combustion of 100% coal A is included in the table. A strong increase in K, especially the water soluble K, is seen when comparing the chemical composition of fly ash from co-firing to the upstream deposit. Even compared to the straw ash, the content of water soluble K in the upstream deposit increase. A similar increase in the Cl content can be observed indicating that the upstream deposit could be largely formed by condensation of gaseous KC1 and/or thermoforesis of KC1 aerosols. Under the prevailing reaction conditions K2SO4 aerosols may be formed by

Deposit Formation in the Convective Path of a Danish

CFB-Boiler

423

sulfation of KC1 [Christensen, 1995]. Assuming that the water soluble K is present as KG and K2SO4, a rough estimate shows that it represents about 50% wt of the deposit. Also interesting is the Si/Al ratio which in the upstream deposit is close to 1 and thus differs significantly from the Si/Al ratio in the fly ash as well as the coal ash. The high content of water soluble K in the downstream deposit is most likely related to the particle size. Only fine fly ash particles (<10µ m) are able to settle on the lee side of the probe [Benson et al., 1993] and these particles are strongly enriched in KCl [Hansen, 1997]. Comparing the fly ash from co-combustion of straw and coal A to the fly ash from pure coal A combustion it appears that the shares of Cl and K, especially water soluble K, drops dramatically while the A1 content is much more dominant in the fly ash from pure coal A combustion. Other-wise only minor differences in the two fly ash samples can be noticed. 6.1.2. Upstrean Deposit: Granular And Neck. The two areas (granular and neck) marked with squares in Fig. 6 indicates where elemental analysis by EDX have been conducted. The results are listed in Table 3. The most striking difference between neck and granular is the shift in the Cl and K content and Cl/K ratio. Assuming that all Cl is retained as KCl, about 70% of the K in the granular part of the deposit is present as KCl, whereas less than 15% of the K is present as KCl in the neck part. This seems to indicate that KCl has a preference for certain locally attractive regions in which the dense granulas are formed. Between the dense areas, a more porous deposit structure, depleted in K and Cl and enriched in S, Ca, Si and Al, is formed, possibly by settling of fine particles as described by Laursen [1997].

424

P. F. Binderup Hansen

6.1.3. Downstream Deposits. In Table 4, elemental analyses of the inner and outer layer of a 6 0 µ m thick downstream deposit. The data reveals a strong depletion in Cl as well as K in the outer layer of the deposit. The composition of the inner deposit layer indicate that 90wt% of the K is present as KC1, whereas less than 25wt% of the K is present as KC1 in the outer layer. The different compositions of the inner and outer deposit layer could be explained partly by the temperature gradient across the deposit, partly by chemical reactions. The low KC1 content in the outer deposit layer could indicate that the local temperatures exceed the level where KC1 is thermodynamically stable, still some K transported to the downstream deposit remain as KC1. Based on a simple ion balance this shift in the Ksalts could easily be explained by part of the KC1 in the deposit reacting with SO2 and H2O from the flue gas forming K2SO4 and gaseous HCl.

6.2. Coal B Figure 7 shows a three hour deposits collected during co-combustion of straw and coal B. The very dense deposit structure differs strongly from the deposit shown in Fig. 5. The white deposit just below the metal surface is the initial layer containing predominantly KC1. The close-up picture of the deposit shown in Fig. 8, reveals an almost solid structure with distinct fly ash particles embedded. The average porosity (based on SEM) of the deposit is less than 3%, [Hansen, 1997]. The average porosity of a one hour deposit,

collected under similar conditions, is less than 5%, which is still less than 20% of the porosity found for deposits collected during co-combustion of straw and coal A. The figures 5– 8 indicates that the coal type is a key factor for the physical structure deposits formed when co-firing coal and straw in a CFB-boiler.

Deposit Formation in the Convective Path of a Danish 80MWTH CFB-Boiler

425

6.2.1. Chemical Composition. In Table 5 the chemical composition of the upstream deposit is listed together with the matching fly ash and fuel ash compositions. The data reveals that 34wt% of the deposit is water soluble K compared to the 9 wt% found in the fly ash. Assuming that the water soluble K is present as KCl and K 2SO4 it makes up about of the total upstream deposit. This is quite a lot more than the 50 wt% found in the upstream deposits from co-combustion of straw with coal A.

6.3. CCSEM Analyses Of Fly Ash Samples It is interesting to see how the coal type appears to affect the physical structure of

the probe deposit as described in the preceeding sections. When looking at the bulk chemical analyses of the deposits the main differences are the high contents of S and water soluble K in coal B deposits and the relatively high contents of Si, Al, and Ca in the coal A deposits. When studying the corresponding fly ashes, the high content of water soluble K in the fly ash from coal B is probably the most stricking difference. In Table 6 the results from Computer Controlled Scanning Electron Microscopy

426

P. F. Binderup Hansen

analyses on fly ash samples from the two tests are presented. About 97 wt% of the fly ash is represented by the 14 classes listed in Table 6. The data reveals that the main fly ash components for both coals are K-Al-silicates, K/Ca-silicates (a mixture of K-silicates and Ca-silicates), quarts, unknown silicates, and a large group named unknowns. The unknowns, in general, contains substantial amounts of K, S, and Cl, which helps explain the discrepancy between the KC1 content in Table 6 and the Cl content in the fly ash according to Tables 2 and 5. The unknowns represent mixtures of inorganic minerals and the share of unknowns increases with decreasing particle size. This indicate that the finest ash particles could be formed by nucleation and agglomeration and/or condensation of inorganic salts such as KC1 and K2SO4 possibly combined with some extent of Al- and Ca- silicates. Based on CCSEM-analysis data, the distribution of Cl, S, K, Ca, Al, and Si in the fly ash and bottom ash as a function of particle size (1–2 µrn, 2–4 µm,. .. , 125–250 µm) for the two coal types is presented by Hansen [1997]. A summary of the results for fly ash in the particle range (1 –2 µm) is shown in Table 7. The data are presented in a mole/kg fly ash basis as well as in a n/K molar ratio, where n represent any of the six listed elements and K represents the potassium content.

Deposit Formation in the Convective Path of a Danish 80 MWTH CFB-Boiler

427

The data reveals how the finest fly ash particles from co-combustion of straw with coal B contains more Cl and especially K and S whereas the content Si and Al and to some extent Ca is much lower than found in the fly ash when co-firing straw with coal A. Based on the n/K molar ratio it appears even more clearly how the “inert” ash primarily represented by Al, Si and to some extent Ca, are much more dominant in the fly ash from co-combustion of straw and coal 4.52/1.90) compared to the fly ash from co-combustion of straw with coal B: (1.37/1.92).

6.4. Mature Deposits During the study described by Hansen, [1997] mature deposits from five positions within the combustor, were collected and analyzed. The general observation was that the deposits were compriced of numerous distinct layers. XRD-mapping of a mature superheater deposit indicated that K content in the layers was shifting between predominantly K2SO4 and K-Al-Si. It is believed that the layers are reflecting the shifts in boiler load caused by the daily variations in the public demand for district heating. A bulk chemical analysis of the deposit revealed a chemical composition (based on weight) of about 7wt% Si, 5wt% Al, 12wt% S, 5wt% Ca, less than 0.5–1 wt% Cl and 29wt% K. Almost 90% of the K in the deposit was found to be water soluble [Hansen, 1997]. In Fig, 9, a one week old, 1–1.3 mm thick, superheater deposit is shown. The deposit reveals distinct layers with the most dense layer covering the initial 200–400 µm. The subsequent deposit layers appears to vary significantly in porosity and possibly even in chem-

ical composition. The vertical line marked in the deposit indicates where elemental analyses have carried out. The results are shown in Fig 10. The data have been recalculated to give the local content of the components in mole/kg deposit. The upper part of Fig. 10 reveals a close correlation between the Si and Al content in the deposit with a Si/Al ratio of 1. However, it can not be seen if other elements, ie. K, are involved. From the lower part of Fig. 10 the for-mation of KC1, K2SO4, and CaSO4 may be evaluated. Four profiles are shown where Ca + K represent the kations that may react with S, whereas 2S + Cl represent the anions that may react with K. By depicting 2 moles S + 1 mole Cl in the latter profile, the four profiles will show

when a major part of the S is retained as CaSO4, by largely exceeding the Ca + K profile.

428

P. F. Binderup Hansen

This is clearly the situation at a deposit thickness around 0.8 and 1.0mm. Otherwise the S is predominantly retained as K2SO4 while any Cl is present as KCl. The profiles shows that the initial 0.2mm consist of predominantly K2SO4 which is followed by distinct KCl peaks around 0.2, 0.3, and 0.45mm. After this point, only very little KCl is found in the deposit. In the following 0.2–0.3 mm Si, Al, and Ca seems to dominate with some K2SO4 present. Around 0.8, 1.0, and 1.15mm CaSO4 seems to dominate, whereas the deposit surface seems to be made from almost pure K2SO4. The observed profiles corresponds reasonably well with the back scatter image in Fig. 9.

7. HOW DOES PROBE DEPOSITS REPRESENT MATURE DEPOSITS? The preceeding sections have shown that the one and three hour probe deposits differs quite significantly from the mature deposits in both texture and in chemical

composition. The probe deposits is—as a maximum—formed by two radial layers, ie. an initial layer of pre-dominantly KCl (20–30µm thick) followed by a more or less dense deposit

layer (

thick) formed predominantly by KCl and to a lesser extent by K 2SO4,

CaO/CaSO4, K-Al-silicates and other silicates. If any, then variations of the deposit

Deposit Formation in the Convective Path of a Danish 80MWTH CFB-Boiler

429

texture and chemical composition in the axial direction seems to be controlled by the coal type. This could be named a transition deposit. Mature deposits, represented by the sample shown in Figs. 9 and 10, consists of numerous layers with an inner layer (200–300 µm thick) which seems almost solid and contains predominantly K2SO4 mixed with some Al-silicate, possibly attached to K, and local zones rich in KC1. The inner layer is followed by deposit layers of varying thickness, physical texture, and chemical composition, which seems to shift between dense K2SO4-dominated deposit layers and more porous layers dominated by “inert” fly ash components such as K-Al-silicate, K-silicate, Ca-silicate, and CaO/CaSO4. Whether the dense K2SO4-layers are formed by thermoforesis and/or impaction of K2SO4 aerosols or whether KC1 is initially deposited and subsequently sulfated or a combination of the two, will depend on the deposit surface temperature, the gas temperature and the gas composition. Naturally no more KC1 will be present in the deposit beyond a certain radial position, where KC1 becomes thermodynamically instable. Observations from the 80MWth CFB-boiler reveals that beyond a location in the convective path where the flue gas temperatures is cooled below 760–780°C (ie. below the melting temperature of KC1) the dense and troublesome upstream deposit are not found. Hansen [1997] tested the impact of load variations on the deposit composition but saw little or no effect when reducing the boiler load even though load variations is believed to be the reason for the layers observed in the mature deposits. To summarize, it appears that even though one and three hour probe deposit samples are collected in a full scale combustor under the correct reaction conditions, the deposit samples may not give the complete or even the correct picture of the mature deposits formed when co-firing straw and coal. The explanation is believed to be the gradually increasing deposit surface temperature which follows as the deposit grows combined with a substantial flue gas concentration of SO2. It is not quite clear how long sampling period will be needed to resemble the mature deposits but a guess would be 24–48 or even 72 hours. It should be noted that other boilers burning other fuel mixtures would not necessarily experience the same deviations between short term and mature deposits. It may depend on deposit composition, the local gas temperatures, and deposit growth. However, short term deposit probe samples is still a very strong tool when trying to understand the initial and early deposit formation mechanisms [Baxter et al., 1996], when determining deposit formation rates [Hansen, 1997], and when studying high temperature corrosion [Hansen et al., 1996, 1997; Larsen et al., 1996; Michelsen et al., 1996].

8. CONCLUSIONS One and three hour probe superheater deposits have been collected in a full scale CFB-boiler co-firing coal and straw. The deposit samples are analyzed and the results compared to mature superheater deposits (older than 150 hours) collected from the boiler during boiler shut-down. With a high ash, high Al-containing coal, co-combustion of 50% straw and coal produces an upstream deposit with an overall CCSEM-based porosity of approximately 30%. Compared to the fly ash, the deposit is strongly enriched in KC1. The deposit grows in a coral-like structure, by condensation of gaseous KC1 and/or thermoforesis of KC1 aerosols, at the tip of numerous ultra fine (~0.5µm) “fingers”. Gradually a granular surface forms with grains highly enriched in KC1 (~40 wt%) and a more porous neck structure between the grains, depleted in KC1 (~5wt%) but enriched in Ca, Si, Al, and S. With a low ash, low Al-containing coal, co-combustion of 50% straw and coal pro-

430

P. F. Binderup Hansen

duces an almost solid upstream deposit with an average CCSEM-based porosity of less than 3%. The deposit forms a dense, almost fused, matrix of KC1 and K2SO4 with individual fly ash particles, of up to 10–15µ m in diameter, embedded in the matrix. Compared to the coral-like deposit obtained with the high ash coal, the amount of “inert” ash particles is significantly reduced. As a consequence, the estimated KC1 and K2SO4 content in the upstream deposit increases from about 50% with a high ash coal to about 67% with a low ash coal. Mature deposits collected from the boiler is compriced of numerous layers with an inner—and almost solid—layer containing predominantly K2SO4 with locally KCl-rich zones. The subsequent layers are of varying thickness, physical texture, and chemical composition, shifting between dense K2SO4-dominated layers and more porous layers dominated by “inert” or dry (not melted) fly ash particles rich in K-Al-silicate, K-silicate, Ca-silicate, and CaO/CaSO4. As the deposit grows the surface temperature exceeds the thermodynamic equilibrium for solid KC1 which combined with a high flue gas concentration of SO2 makes K2SO4 more attractive thermodynami-cally than KCl, thus explaining why practically all water soluble K in mature deposits, especially at a certain distance of the cold superheater surface, is present as K2SO4. Downstream deposit from co-firing straw with a high ash coal reveals a significant increase in KC1 and K2SO4 compared to the fly ash it self. Particle settling, including only

particles smaller than 10 µm, is the predominant deposit formation mechanism and these particles are strongly enriched in KC1 and K2SO4. As the deposit grows thicker the local temperatures increases and the content of water soluble K in the downstream deposit shifts from predominantly KC1, near the cool metalprobe surface, towards K2SO4 at the warmer deposit surface. Even though three hour probe deposits are collected in a full scale combustor under true reaction conditions, the deposit may provide an incorrect picture of the mature deposits formed during co-combustion of straw and coal in a CFB-boiler. Depending on the boiler and fuel mixture, this may or may not be a problem. However, it should be considered during coal/biomass co-firing. It is therefore suggested to collect deposit samples exposed for 24–48 or even 72 hours. Although the short term deposit probe samples may not tell the full story of deposit formation it is still a strong tool when trying to understand the early deposit formation mechanisms, when determining deposition rates, and last but not least when studying high temperature corrosion.

ACKNOWLEDGMENTS This study was funded by the ELSAM Research and Developement programme. This auther would like to acknowledge the help and support from the personel at the Grenå power station. Karin Hedebo Andersen is acknowledged for conducting the EDXanalysis on the mature deposit. The many fruitfull discussions with Dr. Bengt Johan Skrifvars and Lone A. Hansen are acknowledged and so are the comments by Dr. Flemming J. Frandsen to the final paper .

REFERENCES Andersen, K.H.. Frandsen, F.J., Hansen, P.F.B., Dam-Johansen, K., (1997). “Full Scale Deposition Trials at l50 M w e PF-Boiler Co-Firing Coal and Straw: Summary of Results.” Proc. Eng. Found. Conf. “Impact of Impurities in Solid Fue1 Combustion”, Kona, Hawaii.

Deposit Formation in the Convective Path of a Danish 80 MWTH CFB-Boiler

431

Baxter L.L., (1993). “Ash Deposition during Biomass and Coal Combustion: A Mechanistic Approach.” Biomass and Bioenergy, 4 (2), pp 85–102. Baxter, L.L., Miles, T.R., Miles, T.R. JR., Jenkins, B.M., Dayton, D.C., Milne, T.A., Bryers, R.W., Oden, L.L.

(1996). “The Behaviour of Inorganic Material in Biomass-Fired Power Boilers Field and Laboratory Experiences.” National Technical Information Service Benson, S.A., Jones, M.L., Harp J.N., (1993). “Ash Formation and Deposition.” in L.D. Smooth (Eds.), Fundamentals of Coal Combustion for Clean and Efficient Use, Coal Science and Technology 20, 299–373.

New York: Elsevier Science Publisher. Christensen, K.A. (1995). “The Formation of Submicron Particles from the Combustion of Straw,” Ph.D. Dissertation. Dept. Chem. Eng. Technical University of Denmark, ISBN-87-90142-04-7. Flagan, R.C., Friedlander, S.K., (1978). “Particle Formation in Pulverized Coal Combustion A Review.” In D.T. Shaw (Eds.), Recent Developement in Aerosol Science. New York: John Wiley & sons.

Frandsen, F.J., Nielsen, H.P., Jensen, P.A., Hansen, E.A., Livbjerg, H., Dam-Johansen, K., Hansen, P.F.B., Andersen, K.A., Sørensen, H.S., Earsen, O.H., Sander, B., Henriksen, N.. Simonsen, P., (1997). Depo-

sition and Corrosion in Straw- and Coal-Straw Fired Utility Boilers- Danish experiences.” Proc. Eng. Found. Conf. “Impact of Impurities in Solid Fuel Combustion”, Kona, Hawaii.. Hansen, P.F.B., Lin, W., Dam-Johansen, K.., Henriksen, N., (1996) Can Superheater Corrosion during CoCombustion of Straw and Coal in a CFB-Boiler be Reduced?“ In Preprints from 5th International Conference on Circulating Fluidized Beds, Beijing, China. Hansen, P.F.B., Ein, W., and Dam-Johansen, K.., (1997). “Chemical Reaction Conditions in a Danish 80 Wm t h CFB-Boiler Co-tiring Straw and Coal.” In F.D.S.Preto (Eds) Prceerings from 14th Fluidized Bed Combustion Conference, pp 287–294. Vancouver, Canada, ASME. Hansen, P.F.B., (1997) “Deposit Formation in a Coal/Biomass Fired CFB under Load Variations.” Internal Report, I/S Midtkraft, ELSAM R&D Project No 348. Henriksen, Hansen, P.F.B., (1995). “CFB-Material Test at Grenaa CHP Plant Final Report of Phase 1 and

2. ELSAMPROJEKT Notat EP94/799a. Henriksen, N., Earsen. O.H., Blum, R., Inselmann, S., (1995). “High Temperature Corrosion when Co-Firing

Coal and Straw in Pulverised Coal Boilers and Circulating Fluidized Bed Boilers.” In Proceedings of VGB Conference Corrosion and Corrosion Protection in Power Plant Technology. Larsen O.H., Henriksen, N., Inselmann, S., Blum, R., (1996). “The Influence of Boiler Design nad Process Conditions on Fouling and Corrosion in Straw and Coal/Straw-Eired Ultra Supercritical Power Plants.” In 9th European Bioenergy Conference, Copenhagen, Denmark. Laursen, K. (1997). “Characteristics of Minerals in Coal and Interpretations of Ash Formation and Deposition in in Pulverized Coal Fired Boilers.” Geological Survey of Denmark and Greenland, Ph.D. dissertation, ISBN-87-7871-022-7. Livingston, W.R. (1991). “Straw Ash Characteristics.” Babcock Energy Ldt., Babcock Report No: 34/91/08, Contractor report. Department of Energy, ETSU B 1242. Michelsen H.P., Larsen, O.H., Frandsen, F.J., Dam-Johansen, K., (1996). “Deposition and High Temperature Corrosion in a 10 MW Straw Fired Boiler.” In Biomass Usage for Utilitv and Industry Power, Snowbird, Utah. Olanders, B., Steenari, B-M., (1995). “Characteristics of Ashes from Wood and Straw.” Biomass and Bioenergy, 8(2), 105–115. Sander, B., (1997). “Properties of Danish Biofuels and the Requirements for Power Productions.” Biomass and

Bioenergy, 12, (3), 177 – 183. Skrifvars, B.J., Backman, R., Hupa, M., Sfiris, G., Åbyhammer, T, (1996). “Ash Behaviour in a CFB-Boiler during Combustion of Coal, Peat, or Wood.” Preprint from the fifth International Conference on Circulating Fluidized Beds, May 28-June 1, Beijing, China.

This page intentionally left blank.

RESEARCH ON THE MELTING POINTS OF SOME CHINESE COAL ASHES

Shen Xianglin, Chen Ying, and Liu Haibin

Thermoenergy Engineering Research Institute Southeast University Nanjing China

1. INTRODUCTION The fusion temperature of coal ash directly influences the slagging condition in coal fired equipment. Slagging can seriously hinder the equipments burning coal from safe

and economical operation. The formation of massive “clinkers” in fuel beds, particularly

if they become attached to a furnace wall or grate surface, disturbs the normal flow and distribution of fuel and combustion air, and may also cause damage to the equipment

[Ely, 1963]. Accumulation of slag on steam generating and superheating surfaces retards heat transfer. It decreases boiler efficiency and capacity. Molten slag may also corrode refractory material and heat transfer surfaces. In general, preventing coal ash slagging is significant in coal combustion, and understanding the variation characteristics of the melting point of coal ash is very helpful in relation to controlling slagging. Coal ash is composed of a variety of mineral matter where the principal elements exist as oxides and sulfates etc. Even though the melting temperatures of the oxides, generally, are high, a series of complex physical and chemical reactions will take place between them when they are heated in a high temperature furnace, so that some low melting point constituents are formed, for example, ferrous iron, calcium, and magnesium contained in coal ash tend to form complex silicates with lower melting temperature than their oxides. The melting point of coal ash is controlled by both its chemical components and the reactions taking place between them. Because the factors that influence the melting temperature of coal ash are complex, it seems that some research results on the melting temperature of coal ash are out of accord with each other. The heating microscope may more precisely measure the melting point. A heating microscope was used to measure the melting temperature of coal ash in this investigation. The relationship between the measured temperature and coal ash chemical components was discussed, and the influences of furnace atmosphere and measuring procedure on the melting temperature were also discussed. Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

433

434

S. Xianglin et al.

2. EXPERIMENT Eight Chinese coals were used as research samples. They are Xiaolongtai lignite, Pingzhuang lignite, Jingxi anthracite, GBW11104a anthracite, Liuzhi lean coal, Hebi lean coal, GBW11109b bituminous, and Hanqiao bituminous, respectively. The results of their proximate analysis and element analysis are listed in Table 1. Coal ash samples were made with reference to the Chinese state standard (GB212–77) on measuring coal ash content. The procedure is: grind coal until the maximum coal particle size is smaller than 0.2mm. Spread it out on a ceramic boat bottom in a 1mm thick layer and place the ceramic boat in an adequately ventilated muffle furnace in which temperature is set at Burn the coal for 1.5 hours and take the residue as the ash sample. The chemical compositions of the ash samples were obtained by means of chemical analysis. They are listed in Table 2. The mineral constituents of Chinese coal ash are mostly silicate, or mica containing potassium and kaolin. Therefore, as can be seen from Table 2, the coal ash is chiefly composed of CaO and in order of their content, with smaller amounts of other oxides [Lou Yindu, 1981]. A heating microscope was used to measure the melting temperature of the coal ash. The maximum heating temperature was 1,660°C in the experiment. The tested ash specimen was made up into a 3 mm high and 3 mm in diameter cylinder. The instrument can show an enlarged image of the ash specimen by 40 times on CRT by means of a video camera system, which is convenient to observe and to measure the changes in the ash specimen. As compared with the traditional procedure that directly observes the change of a triangular pyramid in a pipe furnace, it reduces the observation error. Measurements were carried out under an oxidizing atmosphere and the heating rate was 10°C/min.

Research on the Melting Points of Some Chinese Coal Ashes

435

The typical variation of ash specimen in heating process is shown in Fig. 1. The variation of Xiaolongtai lignite is shown in Fig. la, from which obvious shrinkage, expansion and fusion of ash specimen can be observed. Three characteristic temperatures in the melting process were defined. The initial deformation temperature (IT) is noted when the specimen changes from shrinkage to expansion or its sharp edges first become rounded. The hemisphere temperature (HT) is obtained when the specimen becomes a hemisphere shape, that is H = W/2 (h is the height of specimen, w is the maximum width of specimen). The fluid temperature (FT) is recorded when the ash has become completely fused, and spreads on the base plate with H = W/6. Figure la is a general situation, and a special situation is shown in Fig. lb, which was taken when Hanqiao bituminous was tested. In this case, the specimen has an obvious contraction in the earlier heating process, but after reaching a certain temperature its shape had no more change and it did not completely fuse until 1,660°C. The measured characteristic temperatures of fusion for the eight ash samples are listed in Table 3. Because the measurement of melting point of coal ash depends on standardized procedures, different procedure could result in the difference between their results. Chinese state standard (GB219–74) specifies

436

S. Xianglin et al.

the measuring procedure for the melting point based on the observation of ash pyramid deformation. Three characteristic melting temperature are defined in it, that is: the deformation temperature at which the sharp apex of the specimen first becomes rounded or bent; the soften temperature at which the ash pyramid bends so that its apex touch the base plate, or the shape of the specimen becomes a sphere or hemisphere; the fluid temperature at which the specimen completely fused or spreads on the base plate with its height equal to 1.5mm. According to the mentioned definitions there should be a reasonable comparability between HT and however, as can be seen from Table 4 [Sun yilu, 1987] the HT is larger than

Figure 2 shows another comparison between

HT and The data in the figure are based on 21 Chinese coal ashes. The Fig. 2 also indicate that HT is larger than T2, which may result from differences in the standard used and the associated observation error.

3. DISCUSSION 3.1. Influence of Basic Composition on Melting Point As shown in Table 2, the components of coal ash can be divided into two groups: one is the acidic components including and another is the basic components including and . The ratio of the acidic components to basic components greatly influences the melting point of the coal ash. It is often used to determine the slagging property of coal. The relation between basic components and

Research on the Melting Points of Some Chinese Coal Ashes

437

hemisphere temperature (HT) in an oxidizing atmosphere was firstly discussed here, and is plotted in Fig. 3. The solid points in Fig. 3 were obtained in our experiments, and the actual line was the regressive curve based on the experimental data. For the purpose of comparison some other data [Gong Desheng, 1987; SunYi lu, 1987] obtained under the same conditions are also plotted as hollow points in Fig. 3. As can be seen from Fig. 3 they are in accordance with each other, especially when the basic matter is less then 30%. According to the regressive curve in Fig. 3, it could be found that the HT decreases with an increase in the basic matter when basic matter is less than 40%, but the tendency is

opposite when basic matter is larger than 40%. Therefore, there is a minimum value of

HT at the basic components content of about 40%. The change of the tendency is, chiefly, caused by CaO. Generally, the increase in basic matter could decrease the fusion temperature, but CaO may elevate it if its content is more than a certain value. The situa-

tion is similar when coal ash is in reducing atmosphere. The IT of 10 Chinese coals in a reducing atmosphere were plotted, as solid points, against the basic components content in Fig. 4, and the actual line is the regressive curve based on these data, from which the same tendency as Fig. 3 could be found and there is also a minimum IT when basic components content is about 40%. For the purpose of comparison, a line of dashes is also shown in Fig. 4, which is taken from a reference [Kittle, 1978]. It is a plot of the initial

438

S. Xianglin et al.

ash fusion point against the basic components of coal ash under reducing conditions. The data was taken from numerous sources. As can be seen from Fig. 4, the variation of the actual line and the line of dashes are similar, but their values at a certain basic components content are different, which implies that the initial ash fusion point does not equal to the IT for identical coal ash. Further, the hemisphere temperatures (HT) of the 10 Chinese coals against their basic components are also plotted, as hollow point, in Fig. 4. A reasonable accordance between the HT and the initial fusion temperature can be found in Fig. 4, which means that the so called initial fusion temperature approaches the hemisphere temperature. Melting condition is another important factor affecting the melting point. The fusion temperature of coal ash under mildly reducing condition is lower than that under oxidizing condition. The difference between the hemisphere temperature under an oxidizing condition and that under a mildly reducing condition is plotted against their content in Fig. 5. It can be seen from the figure that the difference increases with the increase in the content of coal ash. This difference can even approach 200°C ~ 300°C when the content of is high enough. It was well known that the iron of coal

ash exists as under oxidizing conditions, and as FeO under reducing conditions. The melting temperature of FeO is lower than that of which could result in the lower melting point of coal ash in reducing condition. But more important factor is that the FeO can react with the oxides of silicon and aluminum of coal ash to form silicate and aluminate, which can more notably reduce the melting point.

3.2. Effect of Acidic Composition on Melting Point The ratio of to in the acidic compositions of coal ash has prominent effect on the melting temperature of coal ash and is another characteristic quantity to judge coal slagging property. Fig. 6 shows the variation of the hemisphere temperatures

of 18 Chinese coal ashes under an oxidizing condition with the changing in the ratios of the silica to aluminum oxide of the coal ashes. The data in the figure were divided into two groups for curve fitting. The first group included all of the data, based on which a regressive curve, line of dashes, was obtained and second group excluded the Xiaolongtai lignite from the first group because of its specially high content of CaO, based on which another regressive curve, actual line, was obtained. Both curves indicate that the

439

Research on (he Melting Points of Some Chinese Coal Ashes

hemisphere temperature decreases with the increase in However, as can be seen from Fig. 6‚ the actual line is more accordant with test data than the line of dashes‚ because the correlation coefficient of first group data is –0.600 and that of second group is -0.726. The high CaO content coal‚ Xiaolongtan lignite‚ makes the correlation between the hemisphere temperature and the worse. The fact indicated that as a judge criterion of coal slagging property the

is appropriate only if the coal ash has

a smaller CaO content. The effect of the on melting point depends on the reactions taking place during heating process. Mineral analysis indicates that the can react with forming mullite when the temperature is about 1‚000°C ~1‚200°C.

The mullite has a higher melting temperature‚ about 1‚810°C [Gao zhihong‚ 1986]‚ The mass ratio of in mullite is 0.39. When heated‚ some are converted to mullite‚ and the surplus free state can react with the basic components of coal ash‚ and form low melting point matter. Therefore‚ the greater the proportion of the surplus free state the lower the melting point. In general‚ the ash which has larger can produce more free state which can result in a lower overall melting temperature.

3.3. Correlation The oxides that have a notable effect on the melting temperature of coal ash are

and the basic components. The aluminum oxide can elevate the melting temperature‚ and the basic components can decrease it generally. The action of silicon oxide is complex‚ which depends on the contents of aluminum oxide and basic components.

They‚ as three independent variables‚ were correlated with the hemisphere temperature in an oxidizing atmosphere‚ based on the experimental data of 15 Chinese coals. The correlation is as follows:

The plural correlation coefficient of the equation (1) is 0.93.

440

S. Xianglin et al.

4. CONCLUSIONS 4.1. The measuring results of melting temperature‚ to some extent‚ depend on the measuring method. The hemisphere temperature (HT) measured with heating microscope is generally higher than its corresponding temperature obtained in the method of pyramid observation. 4.2. The melting temperature of coal ash‚ both under oxidizing and under reducing conditions‚ has a minimum value when its basic components is about 40%. 4.3. The ratio of silicon oxide to aluminum oxide can be used as a criterion of coal slagging property only if the coal ash contains a small amount of calcium oxide. 4.4. Equation (1) can be used to calculate the hemisphere temperature of coal ash in an oxidizing atmosphere.

REFERENCES Ely‚ F. G.‚ Barnhart‚ D. H. (1963).”Coal Ash—Its Effects on Boiler Availability.” In H. H. Lowry (Eds.)‚ Chem-

istry of Coal Utilization‚ New York: John Wiley &Sons. Gao Zhihong‚ Gong desheng (1986).” Variation of Mineral Matter in Heating Process”‚ Thermal Power Generation‚ No. 6‚ 56– 64 Gong Desheng (1987)‚”Melting property of Coal Ash in High Temperature”‚ Thermal Power Generation‚ No. 2‚ 58–64. Kittle P. A.‚ Bennett R. P. (1978).”Chemical Solutions to Problems Encountered in The Gasification of Coal”‚ Combustion‚ 50(4)‚ 18–23. Luo Yindu (1981). Coal Property and Chemical Analysis. Beijing: Coal Industry Press. Sha Yinglu (1991).”Coal Ash Property”‚ In Zhang Baokang (Eds.)‚ Handbook of Industrial Boiler‚ Nanjing: Jiangsu Science and Technology Press.

Sun Yilu‚ (1987). “Experimental Research on The Slagging Property of Dongshen-Shenmu Coal”‚ Thermal Power Generation‚ No. 2‚ 1–12.

COMPUTER CONTROLLED SCANNING ELECTRON MICROSCOPY (CCSEM) ANALYSIS OF STRAW ASH

Henning Sund Sørensen

Geological Survey of Denmark and Greenland Thoravej 8‚ DK-2400 Copenhagen NV Denmark phone: Fax:

e-mail: [email protected]

1. INTRODUCTION One of the main problems involved in straw combustion is related to the behaviour and fate of the ash forming inorganic species contained in the fuel. These elements can end up in fly ash‚ bottom ash or may be incorporated in various types of deposits. The inorganic species in straw either occur disseminated or ionically bound in the organic structure‚ are located in inorganic straw constituents or they are constituents of terrigeneuos dirt that is incorporated in the straw during harvesting and handling. Clearly the fate of the inorganic species during combustion depends on their mode of occurence in the straw. Knowledge of straw ash composition and melting behaviour is important to be able to predict‚ at least to some extent‚ the fate of the inorganic elements in terms of practical operation of boilers. The aim is to be able to predict and possibly avoid formation of troublesome deposits and potential corrosion. Additionally‚ knowledge of straw ash compositions and morphology is useful in terms of possible utilization or disposal of the ash. This paper presents part of a project that aims to acquire knowledge about the formation‚ composition and physical-chemical behaviour of ash from combustion of straw and co-combustion of straw and coal. One of the goals is to develop or modify useful analysis techniques for straw ash characterization. The techniques utilized in the project involves Computer Controlled Scanning Electron Microscopy (CCSEM)‚ High Temperature Light Microscopy (HTLM) (Hjuler‚ 1997) and Simultaneous Thermal Analysis (STA) (Hansen‚ 1997). The CCSEM analysis provides information of the composition and morphology of the ash whereas HTLM and STA provides information about melting temperatures and melting behaviour. The combined data on composition and melting Impact of Mineral Impurities in Solid Fuel Combustion‚ edited by Gupta et al. Kluwer Academic / Plenum Publishers‚ New York‚ 1999.

441

442

H. S. Sørensen

has been utilized to study straw ash and deposits in two Danish straw fired power plants (Hansen et al.‚ 1997). The CCSEM method is well suited to characterize minerals in coal (e.g. Birk‚ 1989; Zygarlicke and Steadman‚ 1990; Laursen‚ 1997a; Jones et al.‚ 1992; Huggins et al.‚ 1980). This is especially the case for bituminous or subbituminous coals where the inorganic elements are preferentially located in distinct minerals wheras in brown coals a significant part of the inorganic elements are situated in the organic structure and are therefore not detected by CCSEM analysis. Additional to coal mineral analysis CCSEM analyses of coal ash supplies valuable data on particle composition and morphology‚ parameters that are important for the behaviour of ash particles in a boiler. Therefore CCSEM data can provide important inputs to modelling of ash deposit formation in coal fired boilers. In straw and other biomass fuels a large proportion of the inorganic species are dispersed in the organic structure and is therefore not detected by CCSEM analysis. Hence‚ CCSEM analysis of biomass fuel is of limited value on its own‚ and should at least be coupled with for example chemical fractionation analysis by progressive leaching. However‚ the CCSEM technique is a strong tool for characterization of biomass ash due to the fact that it provides data on the diversity of ash particles‚ i.e. each individual particle is analyzed with respect to both size and composition. This is in contrast to tra-

ditional bulk chemical analyses which combine all constituents into one batch yielding

an “average” composition. This paper deals with the modifications and development of the CCSEM method to be suited for characterization of straw ash and ash from cocombustion of straw and coal. Additionally‚ the algorithms will be suitable for characterization of deposits formed during straw combustion.

2. ANALYTICAL METHOD Between 0.2g and 1 g of each sample were embedded in epoxy and after hardening sectioned into two pieces parallel to the settling direction. The two pieces were embedded with the new face pointing upwards to avoid density bias due to settling. During the work it was found that many particles settled to touch each other. This induces an adverse effect on the CCSEM analysis. Therefore some samples were mixed with a filler consisting of 10-50 mm diameter acrylic balls to keep ash particles from touching. The blocks were ground and polished using 1/4mm diamond powder for the final step. No water or alcohol were used during preparation since both can readily dissolve various salts present in the sample material. Either Buehler Oil or distilled petroleum were used as lubricant and cooling during grinding and polishing. The polished blocks were coated with a thin carbon layer in a Polaron TB500 coater. The SEM and CCSEM work was performed at The Geological Survey of Denmark and Greenland (GEUS) on a Philips XL 40 Scanning Electron Microscope with a Noran Instruments Voyager II X-ray analysis system attached. The use of CCSEM analysis for characterization of minerals in coal and coal ash at GEUS has recently been described in detail by Laursen (1997a; 1997b) and therefore only the basics of the technique will be mentioned here: i) A number of points is set up randomly to cover a large part of the sample surface. At each point an analysis is performed at three different magnifications and to achieve good resolution in the whole size range.

Computer Controlled Scanning Electron Microscopy (CCSEM) Analysis of Straw Ash

443

ii) A Backscatter Electron (BSE) image with a good separation between inorganic and organic material is acquired (Fig. 1a). iii) An appropiate grey-level threshold value is set to create a binary image with only inorganic particles singled out (Fig. 1b). iv) The binary image is used to control the beam to perform a scan across each

individual particle for five seconds. A particle is here defined as a cluster of connecting pixels in the binary image (Fig. 1c). v) An energy-dispersive X-ray-spectrum and morphological data are acquired during each raster scan yielding compositional and morphological data for each individual particle. It shall be stressed that the CCSEM results are semi-quantitative for several reasons: i) The EDX spectrum is not corrected for ZAF effects (Z: atomic number‚ A: absorbtion and F: fluorescence). ii) Particle size is based on cross sectional area and weight percent is estimated on the assumption that all particles are spherical. This involves an uncertainty for particles with more complicated shapes. iii) Particles with average diameters below 1mm are omitted from the CCSEM analysis because much of the detected X-rays are emitted from a volume beyond the particle boundary‚ potentially leading to erroneous or even meaningless results.

iv) Organically bound or otherwise disseminated elements are not detected by the CCSEM method which only “sees” particles with high BSE reflection‚ i.e. inorganic or inorganic-rich particles. These aspects are‚ however‚ not critical for the CCSEM analysis because the X-ray spectra are primarily used to group particles into mineral categories with rather wide ranges of composition. The strength of a CCSEM analysis lies in information about element distribution between different particle types and different samples and in the fact

444

H. S. Sørensen

that each particle is characterized with respect to both size and composition. However‚ it is emphasized that a CCSEM analysis is not directly comparable to a traditional bulk

chemical analysis. The raw X-ray data are used to group the analyzed particles into a number of catagories‚ commonly called mineral catagories. However‚ the catagories are only defined by composition and bears no information on crystallinity. The data reduction will be described in detail below.

3. SAMPLE MATERIAL The sample material used for the present work consists of fly ash‚ bottom ash and straw chars from full scale experiments and wheat straw and grains from the harvest of 1995. Bottom ash and fly ash were sampled at two danish straw fired boilers‚ the Haslev Combined Heat and Power Plant (CHP) equipped with four cigar burners and the stoker fired Slagelse CHP‚ in due of an earlier biomass characterization project (Stenholm et al. 1996). Fly ash were collected in a bag filter at Haslev CHP and in an electro filter at Slagelse CHP. Bottom ash were collected from material that were automatically scraped off the grate in both cases. Four of the experiments utilized wheat

straw‚ one used barley straw and one rape straw. Wheat chars were obtained from a cocombustion (coal-straw) pf-experiment at Studstrupværket‚ Denmark. The chars were sampled from the burner zone with a water-cooled suction pyrometer. The samples are listed in Table 1.

4. RESULTS

Computer Controlled Scanning Electron Microscopy (CCSEM) Analysis of Straw Ash

445

for modifying the CCSEM analysis procedures and data-reduction. Secondly a number of straw ashes and some fuel samples were analyzed by CCSEM and the raw data were used to develop a new set of data-reduction algorithms suitable for characterization of straw ashes and deposits.

4.1. Initial SEM investigations Selected ash and fuel samples were investigated by SEM to provide a basis for setup of the CCSEM analysis. The results of this study is summarized below. Figure 2 shows BSE-images of wheat straw‚ wheat grains‚ wheat straw char‚ fly ash and bottom ash. Figure 2a shows the elongated structure of this milled straw sample. Inorganic particles (bright) occur either as equant grains or as elongated rims on the straw fragments. Qualitative SEM/EDX analysis indicate that the former are mainly terrigeneous aluminosilicates and quartz that were incorporated during harvesting and handling. The elongated rims consist almost solely of silica and represent amorphous hydrous silica (opal) which is known to be common in cereal straw (Marschner‚ 1995). The presence of these silica-“skeletons” is to a large extent responsible for typical high silica-contents in straw

and straw ashes. It must be noted that any inorganic species that are disseminated in the organic material is not visible on the BSE-image.

Figure 2b shows a BSE-image of wheat straw char‚ sampled from the burner-zone of a pf-fired boiler co-firing coal and wheat straw. The straw chars are readily distinguished from coal chars by showing relict elongated straw-structures and by being relatively large (up to 2mm long) compared to the rounded and smaller coal chars

(<200mm). Many of the straw chars are partly outlined by inorganic rims rich in silica and potassium typically constituting close to 90%. The rims stand out bright in contrast to the organic char material. The ratio of the rims ranges typically between 4 and 6 on a weight basis and under equilibrium conditions 60-80% will be melted at the peritectic temperature of 769°C in the binary system (Morey el al.‚ 1931). Clearly the presence of this low-melting compound is important for the melting behaviour of the inorganic material in the chars and probably in the ash itself. Figure 2c shows a BSE image of a polished embedded sample of mixed straw and coal chars corresponding to the one in Fig. 2b. The Si-K rich rims discussed above are readily observed. Qualitative SEM/EDX analysis show that remaining inorganic constituents of the straw chars consist mainly of small (1–5 mm) particles of KC1 and that are either attached to the Si-K rims or are located in the char interior. In detail these

particles appear to be condensed on the larger char particles. It is very likely that condensation of KG and occurs during sampling‚ as both chars and gas are cooled on their passage through the cooled sampling probe. Nevertheless‚ the presence of condensed KC1 and shows that K‚ Cl and S were present in the gas phase indicating that these species were least to partly evaporated already during this early stage of combustion. This is in agreement with studies of aerosol particles in the plants in question which indicated respectively that 18% K‚ 66% Cl and 43% S (Christensen et al.‚ in press) and 22% K‚ 74% Cl and 52% S (Christensen and Livbjerg‚ 1996) were evaporated from the straw at combustion temperatures. Inorganic particles in wheat grains are clearly different from those observed in the straw (Fig. 2d). They consist predominantly of small K‚ Mg and P-rich‚ spherical particles that are located in clusters close to the grain margins. These particles are phytates‚ which are salts of K‚ Mg and P that function as the main storage sites of these elements in cereal grains (Marschner‚ 1995). Based on these observations it is evident that ashes

446

H. S. Sørensen

Computer Controlled Scanning Electron Microscopy (CCSEM) Analysis of Straw Ash

447

derived from the studied straw and grains will be of contrasting compositions and that they therefore will pose different practical problems in combustion. The grains will produce ashes (and potentially deposits) rich in phosphor whereas straw will lead to ashes

(and potentially deposits) rich in Si‚ K and possibly Cl. Bottom ash from wheat straw combustion consists mainly of relatively large (up to 1 mm) composite Si-rich particles (Fig. 2e). However‚ the bottom ash were milled before analysis so the size distribution does not represent the original conditions. The particles are commonly zoned with grey Si-rich cores surrounded by brighter margins that are enriched in K and to a minor extent Ca (Fig. 2f). Minor elements such as Fe‚ Mg and Na are relatively enriched in the margins as well. The suggestion is that this type of ash particles are formed by reaction between dehydrated fragments of Si-rich inorganic straw components or terrigeneous quartz grains with vaporized species in the gas. As discussed above based on observations of straw chars significant parts of K is likely to evaporate in the early combustion phases. The fly ash samples exhibit a much wider range in particle size than the bottom ash. Fly ash ranges from sub-micron particles that appear as a greyish mass on a BSE image and up to around 150mm large particles with compositions and features similar to those observed in the bottom ash. Rims (Fig. 2g) or aggregates (Fig. 2h) of primarily KC1 and to a minor extent are commonly located on the larger particles Si-rich particles. Most of the Si-rich particles have subordinate contents of K and Ca and many exhibit some zoning with K-enriched margins similar to the features observed i the bottom ash. Since potassium is known to lower the melting point of silicium glass (Morey et al.‚ 1931) it is evident that the stickiness of such fly ash particles and thereby their propensity to adhere to for example heat exchange surfaces is controlled more by composition of the margins rather than bulk- or core compositions.

4.2. CCSEM Analyses The strategy for modifying the CCSEM procedures was to obtain an elaborate set of CCSEM raw data on straw ashes from Slagelse CHP and Haslev CHP and subsequently to utilize them to develop a new data reduction algorithm. In total 12 ashes were analyzed by CCSEM. Initially the raw data were reduced by using coal mineral characterization algorithms developed by Laursen (1997b). Due to the contrasting composition of straw ash and coal ash this resulted in a high number of unclassified particles (35-70%)‚ i.e. particles that do not fit the criteria for any of the remaining catagories. The compositions of the unclassified.particles were subsequently studied in detail‚ and it was found that particles rich in chlorine and sulphur constituted a significant part pointing to the presence of chlorides and sulphates‚ compounds that have been reported present in biomass ashes elsewhere (e.g. Olanders and Steenari‚ 1994; Steenari and Langer‚ 1988;

Bryers‚ 1994). On the basis of semi-quantitive SEM/EDX analysis and corresponding ratios between X-ray intensities a number of new catagories were established in order to characterize the ash particles more completely. After each modification of the algorithms the datasets were re-catagorized and the amount of unclassified particles were taken as a measurement of the success of the algorithm. Modifications of the algorithm included addition of a number of new catagories. A composite catagory of KC1 associated with silicates (KCl + silicate) were included‚ an association that is commonly observed in especially fly ash samples. These composite

particles either represent Si-particles with attached rims or agglomerates of KCl

448

H. S. Sørensen

(Figs. 2g and h) or they result from close association between KG and silicate induced during preparation as discussed above. Such closely adjoined particles are seen as one single particle by CCSEM analysis and the result will be an “average” composition with characteristics from all of the adjoined particles. A catagory of particles rich in K‚ Ca and Si called “K-Ca silicate” were included in the algorithm. Characteristic for this catagory is a very low Al-content in contrast to terrigeneous derived clay-particles. The K-Ca silicates are believed to form by reaction between silica from the straw and vaporized inorganic gas species as discussed above‚ and are therefore considered to be glassy‚ amorphous alkali-containing silica glass particles. However‚ some of the K-Ca silicates may form by reaction between volatilized inorganic species and terrigeneous derived quartz grains‚ but notably not by reaction with aluminosilicates as evidenced by the low Al-content. Additional catagories for and KC1 were included in the algorithm as these phases constitute significant parts of especially the fine fly ash fractions. A range of phosphates: K phosphate‚ Ca phosphate and K-Mg phosphate were included in order to encompass ash particles derived from wheat grain. The unclassified group were subdivided into unclassified phosphates‚ -sulphates‚ -chlorides and -silicates on the basis of their contents of P‚ S‚ Cl and Si‚ respectively‚ based on the assumption that these catagories will respond differently when subjected to

combustion temperatures. Most of the catagories from the coal mineral algorithms were retained to facilitate future comparison between CCSEM analyses of coal and straw ashes and especially for CCSEM analysis of ash from co-combustion of straw and coal. The proportions of catagories typical of coal did not differ outside the estimated analytical uncertainty when going from the initial coal data-reduction to the modified straw algorithms as should be the case if such comparison shall be made.

The modification of the algorithms resulted in a decrease in the proportion of unclassified to between 4% and 25% for bottom ash from wheat and barley straw‚ whereas fly ash contained between 6% and 27% unclassified. In the case of rape

straw unclassified still constitute 60% of the fly ash (SL8FA)‚ underlining that it is problematic to encompass ash from biomass fuels with contrasting compositions into one single data-reduction algorithm. The algorithm presently used is primarily tuned to characterize ash from combustion of wheat and barley straw as well as ashes from cocombustion of these straw types and coal. Table 2 shows examples of criteria used in the modified algorithm.

Computer Controlled Scanning Electron Microscopy (CCSEM) Analysis of Straw Ash

449

4.3. Results from Haslev CHP and Slagelse CHP The CCSEM results for bottom ashes and fly ashes are presented on a catagory basis in Table 3. Some general observations can be made on basis of the ash results. The bottom ashes are typically rich in silica-rich catagories‚ especially K-Ca silicate‚ quartz and Si-rich. Only minor KC1 is present and is only detected in HA1BA. Unclassified constitute less than 5% in the case of wheat and barley combustion and only 4% in the case of rape combustion‚ i.e. the data reduction algoritm are considered to characterize the bottom ashes well. The fly ashes are generally rich in KC1 and KC1 + silicate and has a rather low but variable content of K-Ca silicates. is present in minor quantities except in the rape fly ash (SL8FA) where it constitutes 8%‚ underlining the contrating composition of rape compared to wheat and barley. Unclassified constitute between 4% and 14% in the case of wheat combustion‚ 27% for barley and 60% for rape combustion‚ illustrating that the

data reduction algorithm is suited to characterize fly ash from wheat and barley combustion. However‚ in the case of a fuel of contrasting composition such as rape‚ a high amount of unclassified particles remain. To illustrate the variations in composition and in an attempt to correlate catagory composition with melting behaviour of straw ashes‚ the results were plotted in a triangular diagram of KC1‚ K-Ca silicates and quartz+almino-silicates (Fig. 3). The last group is to be mainly derived from terrigeneous dust and possibly from straw-derived Si-rich material. The three groups constitute more than 90% of the bottom ashes and more than 60% of the fly ashes with the exception of rape fly ash (16%). The basis for creating this type of diagram is that it is expected that the melting of the ash will correlate with position due to the contrasting melting temperatures of the end-members. KCl melts at 770°C whereas K-Ca silicates may melt at temperatures as low as around 740°C if present in proportions close to the eutectic composition (Morey et al.‚ 1931). However‚ most of the K-Ca silicate particles have compositions in the range of 5–15% CaO‚ 70–85% and 10–25% as illustrated in the three dimensional triangular diagram in Fig. 4. These compositions will begin to melt at the eutectic temperature‚ but the major part will‚ under equilibrium conditions‚ not melt until higher temperatures have been reached. The group of “quartz + aluminosilicates” will clearly have

450

H. S. Sørensen

Computer Controlled Scanning Electron Microscopy (CCSEM) Analysis of Straw Ash

451

a wide range of melting temperatures‚ but it is assumed that in general the melting will occur at higher temperatures than for K-Ca silicates. This is especially true of the quartz catagory which on its own will not melt until 1‚713°C (Deer et al.‚ 1966). Separation between bottom ash and fly ash is distinct in the diagram. Bottom ash is situated close to the silicate baseline‚ whereas fly ash range from close to the KCl apex roughly towards a ratio of 85% “quarts + aluminosilicates” and 15% K-Ca silicates. This indicates that the silicate association in the fly ash generally has a lower content of particles formed by reaction between silica and gas alkalies (K-Ca silicates) than what is observed in the bottom ash. The location of S18FA (rape fly ash) is not considered since only 16% of the total is represented in the diagram. A discussion of sample location in the proposed triangular diagram compared to their melting behaviour is discussed by Hansen et al. (1997).

4.4. Repeatability Repeatability of the CCSEM analysis were evaluated by preparing three embeddings of one fly ash (SL7FA) and one bottom ash (SL7BA). In this way the errors induced by both sample preparation and analysis procedures were taken into account. The results of the repeated analysis are presented in Tables 4 and 5. The repeated analyses indicate that there is a substantial uncertainty involved in the CCSEM data on catagory basis. The agreement is considered good for bottom ashes where the problem with touching particles is smaller than in the case of fly ash whereas the fly ash shows poorer repeatability. The problem with the fly ashes is that many particles are connected in aggregates (composite particles) which can change the catagory composition significantly even though the total raw data are unaltered.

452

H. S. Sørensen

For bottom ash there is good agreement between CCSEM bulk data for the three runs (Table 5). However‚ in the case of fly ash there is a rather large variations in and This may be a result of slight variations in the setting of threshold values since much of the masses of submicron material in the fly ash has a rather low BSE reflection. Variations in the threshold setting may cause some of the submicron material to be omitted form the analysis. In conclusion it is noted that this test indicates a satisfying repeatability for bottom ash analysis whereas the fly ash is more problematic. The preparation technique will be further standardized and developed in the future to increase repeatability for straw fly ash.

5. SUMMARY AND CONCLUSIONS The CCSEM method has been modified from coal mineral and ash analysis to be suited for characterization of ash from straw and straw+coal combustion. The CCSEM

results are internally comparable‚ but it is emphasized that direct correlation to traditional bulk chemical analysis is as yet not realistic. The strenght of the CCSEM analysis lies in combined compositional and morphological data for individual particles‚ facilitating advanced modelling of ash behaviour. The internal repeatability of results is satisfying for bottom ash‚ whereas it is worse for fly ash mainly due to the presence of touching particles. The preparation procedure will be further developed and standardized to increase repeatability.

CCSEM analysis of ashes from two straw fired plants showed a marked difference between bottom ashes and fly ashes from wheat and barley combustion. The bottom ash is dominated by silica-rich particles in general and K-Ca silicates in particular. These particles are suggested to be formed primarily by reaction between straw derived Si and K. In contrast the fly ash is dominated by KCl and combined KCl+silicates. K-Ca silicates are also present but in minor amounts than in bottom ash‚ whereas is only present in minor quantities. In conclusion it has been shown that CCSEM analysis is useful for characterization of ashes from straw and straw+coal combustion. The modified algorithms are also useful for characterization of powdery deposits that contain similar compositions as the straw ash.

Computer Controlled Scanning Electron Microscopy (CCSEM) Analysis of Straw Ash

453

6. ACKNOWLEDGEMENTS This project has been financially supported by ELKRAFT‚ ELSAM and the Danish

Energy Research Programme.

7. REFERENCES Bryers‚ R.W. (1994). “Analysis of a suite of biomass samples.” Foster Wheeler Development Corporation‚

FWC/FWDC/TR-94/03. Christensen‚ K.A. and Livbjerg‚ H. (1996). “A Field Study of Submicron Particles from the Combustion of Straw.” Aerosol Science and Technology‚ 25‚ 185–199. Christensen‚ K.A.‚ Stenholm‚ M. and Livbjerg‚ H. (in press). “The Formation of Submicron Aerosol Particles‚ HC1 and in Straw-Fired Systems.” To appear in Journal of Aerosol Science. Deer‚ W.A.‚ Howie‚ R.A. and Zussman‚ J. (1966). “An introduction to the rock-forming minerals.” Longman Group Ltd. Hansen‚ L.A. (1997). “Ash Fusion Quantification by Use of Thermal Analysis.” Proc. Eng. Found. Conf.: Impact of Mineral Impurities in Solid Fuel Combustion‚ Kona‚ Hawaii‚ November 2–7‚ 1997.

Hansen‚ L.A.‚ Frandsen‚ F.J.‚ Dam-Johansen‚ K.‚ Sørensen‚ H.S.‚ Rosenberg‚ P. and Hjuler‚ K. (1997). “Ash Fusion and Deposit Formation at Straw Fired Boilers.” Proc. Eng. Found. Conf.: Impact of Mineral Impurities in Solid Fuel Combustion‚ Kona‚ Hawaii‚ November 2–7‚ 1997. Hjuler‚ K. (1997). “Ash Fusibility Detection Using Image Analysis.” Proc. Eng. Found. Conf.: Impact of Mineral Impurities in Solid Fuel Combustion‚ Kona‚ Hawaii‚ November 2–7‚ 1997. Laursen (1997a). “Characterization of Minerals in Coal and Interpretations of Ash Formation and Deposition in Pulverized Coal Fired Boilers.” Ph.D. Thesis‚ Geological Survey of Denmark and Greenland‚ Report 1997/65.

Laursen‚ K. (1997b). “Advanced Scanning Electron Microscope Analysis at GEUS.” Geological Survey of Denmark and Greenland‚ Report 1997/1. Marschner‚ H. (1995). “Mineral Nutrition of Higher Plants.” Second edition‚ Academic Press. Morey‚ G.W.‚ Kracek‚ F.C. and Bowen‚ N.L.(193I). “The ternary system ” Journal of the Society of Glass Technology‚ Vol.14‚ 149–187. Olanders‚ B. and Steenari‚ B-M. (1995). “Characterization of ashes from wood and straw.” Biomass and Bioenergy‚ Vol.8 (2)‚ 105–115.

Steenari‚ B-M. and Langer‚ V. (1988). “Fasanalys av sintrade och osintrade halmaskor med och utan tillsats av

kaolin respektive dolomit.” Report OOK‚ 88:12. Stenholm‚ M.‚ Jensen‚ P.A. and Hald‚ P. (1996). “Biomasses brændsels-og fyringskarakteristika. Fyringstbrsøg. EFP-93.” 1323/93-0015. (in Danish)

This page intentionally left blank.

PROGNOSIS OF SLAGGING AND FOULING PROPERTIES OF COALS BASED ON WIDELY AVAILABLE DATA AND RESULTS OF ADDITIONAL MEASURAMENTS Alexander N. Alekhnovich‚ Natalja V. Artemjeva‚ Vladimir V. Bogomolov1‚ Vyacheslav I. Shchelokov2‚ and Vasilij G. Petukhov 3 1

Ural Heat Engineering Institute Ural Heat Engineering Laboratory Chelyabinsk‚ Russia

2

3

Podolsk Machine-Building Factory Podolsk‚ Russia Barnaul Boiler Plant Barnaul‚ Russia

INTRODUCTION Depending on the objectives of determining slagging and fouling properties of coals there is a need for experimental data of various levels‚ and various methods are used. The data based on the information of the bulk matter composition have low extrapolate possibilities and yield limited information about the behavior of the mineral matter in the boiler. At the present stage‚ in a number of cases‚ even the reliable ranging of coals according to their slagging properties is not considered sufficient. The task is to mathematically model the transformations of the mineral matter in the boiler and various processes resulting in deposits. With the help of mathematical models‚ in the long run‚ the properties of actual deposits at various boiler surfaces are calculated (e.g. Borio et

al.‚ 1993; Allan et al.‚ 1996). At the same time‚ their use asks for the usage of expensive and unique equipment for defining the initial data. The models are at the stage of implementation in operation and testing. The data concerning the sub-model of consolidation

and growth of deposits wherein the data about the composition of individual particles of fly ashes‚ developed by UralVTI‚ have been reported earlier (Alekhnovich et al.‚ 1996). Further up dating of this model concerning the specification of marginal conditions of Impact of Mineral Impurities in Solid Fuel Combustion‚ edited by Gupta et al. Kluwer Academic / Plenum Publishers‚ New York‚ 1999.

455

456

A. N. Alekhnovich et al.

sticking was done according to the calculated value of the critical temperature or critical viscosity. The future‚ with no doubt‚ belongs to mathematical modeling of deposition processes. However‚ many practical tasks‚ including the designing of boilers‚ are solved by simpler methods. The improvement of these methods‚ as well as mathematical models‚ will be of high priority in future. Comparative analysis and ranging of coals with similar ash type‚ supplied to a power station or coals from one and the same coal-field could be made with relative accuracy based on widely available data on the basis of reference data (chemical composition of the mineral matter‚ element content and technical specifications) or‚ in certain cases

(Hatt‚ 1996)‚ based on still more restricted scope of information. Using the reference data we could solve such problems as the assessment of possible restrictions in changing one coal with another‚ optimization of fuel supply of the power station as well as the ranging of coals with various ash types‚ for which a certain experience of burning exists. To evaluate properly the slagging and fouling properties that are “new” coals to the user and designer‚ the choice of permissible characteristics and dimensions of the furnace and gas ducts the conventional reference material is not sufficient‚ especially for high slagging and fouling coals. In these cases additional and more profound research of composition and properties of the mineral part are necessary. There are successful results of defining of coals slagging properties based on insight into the mechanism of deposit formation and on using the experimental data‚ concerning this knowledge. The present report adds to and expands the information earlier published by Engineering Foundation Conference (Alekhnovich‚ Bogomolov‚ 1994). The developed by UralVTI method is successfully used in Russia and‚ in a new version‚ the main prerequisites and conditions have been preserved. The specification and development of certain issued thereof is connected with new tasks and new information. In particular‚ boiler-building factories formulated the task of evaluating the slagging situation in implementing of reburning. They also supplied the information concerning the design and experience of operating the boilers that had not been analyzed according to the first version of this method‚ including combusting of coals not characteristic for Russia.

1. TERMS‚ MECHANISM OF DEPOSITS FORMATION AND USED PREREQUISITES Various deposit types are formed on the surfaces of boilers during PF combustion‚

depending on temperature zones and coal properties. At high temperatures of gas flow and the tube surface quick growing deposits are formed due to the sticking properties of fly ash particles. There are sufficiently clear marginal conditions (temperature of gases and surfaces)‚ below which such deposits are not formed. The temperature of gases‚ below that the deposits on the non-cooled surface are not formed is conventionally called the temperature of the start of slagging‚ As the marginal temperatures depend on coal combustion and experimental conditions it is more correct to use a term “the temperature of the start of slagging ” as coal index and the term “the temperature of active slagging ” in more general cases. We use the following terms to define deposition in boilers: slag deposits‚ slagging—means the deposits‚ the process of their forming and the accompanying problems at temperatures that are higher than the marginal ones. Slag deposits‚ when a noticeable excess of the marginal conditions takes place‚ have the composition that is much closer to the average composition of fly ash and acquire more

Prognosis of Slagging and Fouling Properties of Coals Based on Widely Available Data

457

strength due to sintering. Fouling deposits and fouling means the deposits and the process of their forming at temperatures below the crucial ones. Depending on the structure or the element responsible for their forming‚ fouling deposits are further subdivided into iron enriched‚ calcium based (or sodium based) and friable (fine dispersible) deposits. At the temperatures of the flue gases that are higher than the slagging deposits are formed directly on the non-cooled surfaces. On the cooled surfaces the slagging‚ as a rule‚ is of the phase character. Initially a layer of fouling (or initial‚ primary) deposits is formed and then the proper slagging deposits are produced. The performance of

the furnace and the super-heater concerning the conditions of slagging‚ for the most part‚ depends on the properties of the initial layer. There are practically no problems due to the process of self-cleaning‚ if the primary layer consists of friable fouling and fouling on sodium base. During the forming of the primary layer consisting of iron enriched deposits and sulphur-bounded ones the slagging deposits stick firmly on the tube surface. The deposits of similar content and properties may be‚ simultaneously‚ formed according to various patterns‚ from particles of various content and aggregate states. Depending on the type of coal and temperature zones the part of particles‚ stuck as a result of this or that mechanism changes. Thus‚ in our research‚ the iron enriched deposits were formed from the melt of low-fusible eutectics (the phases of pyrrhotite—wustite—phyllite)‚ the products of insufficient oxidation of pyrite‚ melts of phyllite‚ alumo-silicate particles with iron enriched film and ferric particles with the film of phyllite. The forming of sulphur-bounded deposits of various strength and intensity is due to fastening of CaO (include the mechanism of a chemosorption) and its sulphation‚ sticking of CaO-CaS eutectics and Ca enriched alumo-silicates. Fouling on the sodium base‚ as seen in the literature‚ takes place due to sublimation and further condensing of Na‚ for the most part on the surface of small particles. Deposit formation of various types could be a competing process or an accompanying one. For instance‚ for two investigated coals with a low potential of forming of iron enriched deposits‚ the experiments showed the forming of such deposits when the rate of friable deposition reduced or regular destruction of loose layer took place. We also assume; that the formation of calcium based deposits and fouling on sodium base is the competing processes. In particular‚ we combine it with the fact‚ that in the USA‚ for example‚ the problem of forming of calcium based deposits appeared only in the mid-80’s (Hurley et al.‚ 1996) after the subbituminous coals with a relatively low sodium content were used. By the way‚ the successful use of additions in China (Cui et al.‚ 1994)‚ which contain decomposing compounds of alkaline metals‚ could also be connected with the intensifying of the formation of loose deposits on their basis instead of strong slag deposits. In addition to the above stated ideas and prerequisites we could say‚ that the worked out methods are characterized by the following: — to draw conclusions we used the results of researches of slagging properties of 40 coals. Permissible heat-release-rates and gas temperatures in the boilers have been picked by the results of 53 boiler design-coal analyze; — for approximate prognosticating we used the averaged statistic data depending on the types of alkali‚ sulphur and iron in coals. The dependencies were

obtained based on the results of processing large experimental data and from publications. For instance‚ we have analyzed 262 samples (30 coals from

458

A. N. Alekhnovich et al.

various coal basins) to find out the amount of pyrite sulphur within its general content; — the result‚ characterizing deposits of this or that type‚ if possible‚ is assessed on

the basis of various initial data; — for various temperature zones in the boiler the choice of properties of the slag deposits could be different. It is assumed‚ that in the zone before the downdraft duct of the boiler these are the marginal temperatures for super-heater and platens—the strength of slag deposits and for furnace walls—the combination of characteristics of the initial layer and the excess of the combustion temperature over (or viscosity).

2. SLAGGING AND FOULING INDICES In our methods we used a separate assessment of potential of coals to form various deposits as well as the properties of slagging deposits. To range the coals and to pick out the characteristics of the furnace we also used a combination of these indices.

2.1. Tendency to Form Iron Enriched Deposits Iron enriched deposits are most typical of coals containing pyrites‚ although cases of them forming with predominance of iron in the siderite have also been established. The experimental research included the ranging of coals based on the results of the inspection of boilers and experiments with the help of probes‚ as well as defining of the properties of coals in the laboratory (gravity fractionating‚ content of iron and sulphur in pyrite form‚ etc.). It has been noted‚ that the best result is the iron content in pyrite form in the heavy fraction where is the fraction of iron in a pyrite form. Forming of the iron enriched deposits on the base of the pyrite is the result of its noncomplete combustion and is more characteristics for low-rank coals that are combusted

after rough grinding at a low temperature of the flame. The content of pyrite sulphur on the ash base and the adiabatic combustion temperature have been used in the first version of the method to define the amount of pyrite‚ the coal type and so the iron enriched index‚ In the new version we had objective to take into consideration the maximum quantity of independent factors‚ even those that are statistically irrelevant for the combination of investigated coals. The method of regressive analysis has been used to define the dependency of the tendency of a coal to form iron enriched deposits on such factors as coal rank‚ iron content in its pyrite form (pyrite sulphur calculated against the ash base)‚ iron content in non-pyrite form and the adiabatic combustion temperature The equations used to calculate indexes are shown in Table 1. The value of reduced pyrite sulphur is calculated according to following equations: (lignite‚ brown coals)‚

(light black coals)‚

(lean black coals)‚ where

is reduced contribution of non-pyrite iron‚

Prognosis of Slagging and Fouling Properties of Coals Based on Widely Available Data

459

The meaning of according the experimental results and calculations against are shown in Fig. 1 The coal has a very high potential for forming iron enriched deposits at and a low one at

In the situation of the absence of experimental data concerning the content of pyrite iron or sulphur the value can be approximately evaluated against the general sulphur content The use and certain successes in using such an evaluation are defined

460

A. N. Alekhnovich et al.

by the fact‚ that with the increase of the content of pyrite sulphur increases and for the coals of the majority of coal fields and similar type coals there is a statistical correlation of these values. The analysis of the large amount of experimental data shows‚ that the ratio of varies over a wide range and‚ among other factors‚ depends on the coal rank. This ratio increases in series of lignite—brown coal—high-volatile coals— low-volatile coals. For coals of various types the averaged regression equations have been defined. They look like: % (lignite‚ brown coals type Bl)‚ % (brown coals B2‚ B3 types)‚ % (light black coals)‚ ‚ % (lean black coals).

For the investigated coals a more reliable evaluation is made on the basis of individual prognosticating equations for each coal basin. We should emphasize here‚ that even when the equations are statistically reliable‚ the evaluation of slagging properties of a certain coals‚ especially when it is not known properly‚ against the general sulphur

content‚ could result in a significant error.

2.2. Tendency to Form Sulphur-Bounded Deposits As it has been noted earlier‚ the formation of calcium based deposits is considered as a competing process for the fouling on the basis of active alkalis. In general index has the following form: where characterizes the content of Ca‚ able to form such deposits and characterizes the tendency to form fouling of the basis of active sodium. In previous developments index was defined according to the part of particles

in ash with CaO content higher‚ than 12%. To define the part of particles with CaO content higher‚ than 12% the results of SEM analysis of fly ash or the results of the definition of the composition of various dense-separated fractions of coal could be used‚ or the dependency Now it has been made the attempt to use the content of soluble Ca during its chemical fractionating. The chemical fractionating of coals with a tendency to form calcium based deposits as well as the coals without such tendencies is made under the methods‚ formulated by Benson and Holm (Benson et al.‚ 1996). The method of fractionating has been added a fourth stage—i.e. dissolving of the residue in the nitric acid in order to develop an additional index for the definition of the tendency of coals in forming iron enriched deposits. According to the different mechanisms the sulphur-bounded deposits can be formed as on the base of organically coordinated calcium so on the base of calcium carbonates. So the experimental data have been processed in two modifications. The amount of water and ammonium acetate soluble calcium is used in one of them‚ and summary removed calcium during chemical fractionation is used in another. The dependencies of approximation are

Requirements for the Approval of Veterinary Therapeutics or Growth Enhancers Used in Fish Production

461

When there are not the chemical fractionation data‚ the approximate evaluation can be used according to amount of total calcium CaO and ash on dry mass

Expert ranging and approximate evaluation of sulphur-bounded deposits index on the base of are shown in Fig. 2. When expert ranging the value was decided for all coals if calcium based deposits were not found in boilers. The index was decided if the deposits were found only at the gas temperature higher than

800°C and

if the deposits are formed in the zone with the gas temperature

2.3. Tendency to Form Deposits on the Basis of Active Sodium To prognosticate the low-temperature fouling on the basis of active sodium water and ammonium acetate soluble is used. Due to changes on the norm scale‚ the equation used to define differ not on principle from published one (Alekhnovich‚ Bogomolov‚ 1994): In the absence of the data concerning the soluble Na‚ an approximate evaluation is calculated against the equation: The equations of this type have been derived following the conditions of restricted changes within in alumosilicates of the mineral matter and‚ specifically inefficient amount of soluble for the majority of coals. The coefficient of the equation has been defined by the existing experimental data for Russian coals and as well as the published data (Pollok et al.‚ 1983‚ McCollor et al.‚ 1996) for American coals (Fig. 3)

2.4. Characteristics of Slag Deposits To characterize the slag deposits we use the temperature of the start of slagging the dependency of deposit strength on the temperature (sintering strength for laboratory

462

A. N. Alekhnovich et al.

Calculated

%

tests) and other indices‚ defining the difference between the combustion temperature of fuel and the temperature of the start of slagging (slagging index) or viscosity of its mineral matter. The temperature of the start of slagging

is experimentally defined by means of

probes with non-cooled working zone. In the approximation it could be evaluated using

the ash chemical composition or the relation “viscosity—temperature”. In a more general way than was previously published the meaning of according to the ash chemical composition is evaluated as: where:

The part of double- and triple valence iron in the amorphous phase and is defined by means of Mossbauer’s spectroscopy in the sample of fly ash. The ratio depends‚ in a complex way‚ on the composition of the melt‚ its temperature and oxidation potential of flue gas (Mysen‚ 1987). Considering the above and the unstable character of the temperature and the composition of flue gas in the boiler duct‚ we failed to define an accurate correlation on the composition of the mineral matter. The coefficient of the linear correlation in various types of indices was 0.7-0.75. But nevertheless the resulting correlation shows the tendency for changing of from the composition and the correlation from NBO/T when is taken as an approximate evaluation: where

For coals with content less‚ than 10–12% the meaning of may be evaluated according to the experimental or calculated dependency of “viscosity—temperature”. Viscosity of supercooled melt corresponds to the meaning of temperature according to the experimental procedures accepted in Russia and Pa-s during the calculation according the viscosity model (Senior‚ Srinivasachar‚

Prognosis of Slagging and Fouling Properties of Coals Based on Widely Available Data

463

1995). The equation depending on bulk chemical composition of mineral matter and the evaluation of concerning the viscosity are true for coals with of the ratio When we have a similar mineral matter of the coal and a similar design of a furnace the slagging problems arise when we burn a high rank coal due to a higher combustion temperature. This effect is taken into consideration when we choose‚ as the indices of

slagging‚ the viscosity values η at the temperature say Then the index of slagging deposits looks like where “a” and “b” are empiric coefficients. At present‚ due to the absence of experience in using modern viscosity calculation models

this physically reliable index is used in the method as an auxiliary or a reference one. Index is used as main one. This index characterizes the difference between the combustion temperature and the temperature of the start of slagging When it is expressed linearly‚ when we substitute the combustion temperature by and when the temperature of the start of slagging evaluated depending on ash chemical composition the index looks like:

the index corresponds to the earlier published one (Alekhnovich‚ Bogomolov‚ 1994): To characterize the slagging tendency of coals in the furnace it was used the combined index

or without taking into consideration the value of adiabatic temperature like

The ranging of investigated coals according to

in shown in Fig. 4.

it look

464

A. N. Alekhnovich et al.

3. USE OF DATA FOR ANALYSIS AND IN DESIGNING OF BOILERS The information about the necessary volume of research in the procedure of selecting gas temperatures permissible for slagging at the furnace exit and before the convection duct is published in the previous article (Alekhnovich‚ Bogomolov‚ 1994) and is not cited here. In principle‚ no changes have taken place in the procedure of selecting furnace heat-

release-rates permissible for slagging depending on the slagging properties. As the main alternative we use selecting of plan-area-heat-release-rate according to the index and burner-zone-heat-release-rates

following the combined index

At the same

time‚ the new version of the method uses more universal index of slagging properties and there is a possibility of taking into consideration a greater number of design characteristics of the furnace and its operation conditions. More accurate graphic picture of the relation between “slagging properties—furnace characteristics—slagging situation” are shown in Figs. 5 and 6. Figure 6 does not contain certain data‚ concerning the situations‚ when the are no primary iron enriched deposits on the furnace wall. The values of

and

are calculated according to the following formulae:

where Q‚ Gkal/h—boiler capacity; m—furnace plan area sizes‚ r—coefficient of flue gas recirculation; k1‚ k2‚ k3‚ k4—coefficients that depend on burners arrangement‚ number of burner layers‚ boiler capacity and fuel distribution; m2—area of burners zone. The location of marginal lines in Figs. 5 and 6 and their equations have been chosen in conformity with the following conditions:

Heat-release rate / Plan area‚

Prognosis of Slagging and Fouling Properties of Coals Based on Widely Available Data

465

Heat-release rate / Burner-zone area‚

—when the slagging properties of the coal are low‚ the furnace operates without cleaning means; — all the boilers exit the slagging zone when the loads are reduced down to‚ at maximum‚ 80 per cent of their nominal load. It is necessary to note that said slagging zone borders are “pretty crucial”. The per-

missible heat-release-rates according to Figs. 5 and6 in a number of cases are lower‚ than the recommended ones in Russia and they do not exclude non-optimum characteristics of the furnace and cleaning means. There are reported examples of successful operations of boilers at higher heat-release-rates‚ however‚ in this case such a choice should be justified specifically (efficient wall cleaning‚ their design etc.). Considerably higher meanings of could be acceptable in cases‚ when the forming of strong primary deposits is not characteristic (the location of slagging zone borders is not shown in the figures in this case). Higher meaning of is accepted also in case the meaning of is lower‚ than the marginal one.

In a number of boilers there are serious problems‚ connected with the slagging of steam super-heaters‚ though the calculated temperatures at the furnace exit was not higher‚ than the recommended and experimentally justified. The analyze of the opera-

tion of these boilers showed‚ that the measured meaning of

is higher‚ than projected

and recommended‚ though the calculations were done using the coefficients of thermal efficiency of walls and the calculation methods which are successfully used for those coals in other cases. It was defined that the slagging of the steam super-heater is the result of an intensive slagging of furnace walls due to higher meanings of burner-zone-heatrelease-rates chosen. The more is the excess of over the permissible value (Fig. 6) the more is the difference between the actual and calculated temperature (Fig. 7). The method of calculation and evaluation of slagging properties of coals‚ worked out by the UralVTI and boiler building factories‚ additionally to the data‚ stated here‚ contains a data bank concerning the known coals‚ recommendations for selecting other

466

A. N. Alekhnovich et al.

specifications of a boiler‚ alternative methods of evaluation and has the form of a computer program.

4. ANALYSIS OF REBURNING It is experimentally defined‚ that in installing of reburning of certain types of coals there is a noticeable increase of slagging properties of fly ash (Alekhnovich‚ 1996). The history of boilers equipped with reburning systems‚ at the same time‚ shows‚ that slagging not even increases but reduces. There are several factors influencing the slagging properties of ash. That means the

increase of the part of ash melt (glass) and its more uniform structure‚ the increase of the content of iron and‚ sometimes‚ slowing of pyrite burning. On the other hand‚ within the reburning scheme the distribution of fuel and air changes in the furnace. Both of these factors: the change of slagging properties and “fuel-to-air” distribution change are taken into account in the analysis of the new version of the method. The more general form of index has been used in this case when ratio KOFe is employed instead of K/O. So‚ the index is

Changing of slagging properties is accounted for by the changing of the temperature level ratio of depending on the initial meaning of this ratio and the amount of non-burned pyrite. As far as we can’t calculate all influencing factors (part of ash melt‚ its structure) it is assumed‚ that the changing of is the same‚ when we work with a low excess of air or‚ at least‚ the iron passed to category Changing in pyrite burning is calculated artificially by means of its content increase in the initial coal‚ however‚ due to higher meanings of in the combustion zone‚ it does not tell significantly on The results of the analysis according to burner-zone-heat-release-rates and index are shown in Fig. 8. The results show that‚ for the reported example‚ the changes of slagging properties of coal are lower‚ than their changes for various batches of used coal. Additionally‚ we show that the change of the fuel distribution influences the slagging situation more favorably‚ than a small worsening of slagging properties of fly ash.

Prognosis of Slagging and Fouling Properties of Coals Based on Widely Available Data

467

5. APPROXIMATE EVALUTION OF SLAGGING PROPERTIES OF “NEW” COALS Approximate evaluation of slagging properties of coals according to empiric indices based of the data from reference materials‚ strictly speaking‚ may be used only for a range of coals the information of which has been used to work out such indices. The use of

other coals does not exclude an error for some of them and this error will be the higher the more the relationship of the ash composition of the analyzed coal differs from the conventional average values. At the same time the possibility to use the worked out method for “unknown” coals is of a certain interest. This was the goal of evaluating the slagging properties of the coals‚ reported at The Engineering Foundation Conferences held in Solihull (1993) and in Watervilley Valley (1995)‚ when there are the data concerning their ranging and ash composition. The evaluation of three Great Britain coals showed‚ that their slagging properties are close and the fact was proved by the published results of test burning. Bentink coal

has the minimum slagging properties and Daw Mill coal has the maximum slagging properties which corresponds with the evaluation made by British researchers (Gibb‚ 1996). The evaluation of slagging and fouling properties of 4 American coals‚ which were burnt in one boiler and were assessed under PCQUEST program (Zygarlicke et al.‚ 1996) showed the following. As for the slagging properties of three coals out of four their ranging coincides with the experimental evaluation (Fig. 9); for Black Thunder coal the evaluation made by UralVTI as well as by EERC suggests higher slagging properties. As for the high temperature fouling the ranging done according to our methods practically coincides with the American. As for the low temperature fouling the evaluation done by UralVTI coincides with that by EERC. The evaluation by UralVTI of the potential of high temperature fouling for low rank coals described in publication (McCollor et al.‚ 1996) coincides completely with the experimental one (the content of in Colstrip coal was not defines since it was not analyzed). The properties of coals investigated in the work (Richards et al.‚ 1996) shows the possibility of forming calcium sulphate deposits‚ notwithstanding a higher tendency to form Na based deposits and yields the result which coincides with the existing practical results concerning the ranging of fouling properties of coals.

468

A. N. Alekhnovich et al.

Thus‚ the analysis showed a high enough universal character of the worked out method of evaluating of slagging and fouling properties of coals.

CONCLUSION Ranging of coals according to the slagging properties of similar type and investigated coals could be made on the basis of the available reference data. However‚ to define the slagging and fouling properties of a random coal it is necessary to carry out additional laboratory research.

Along the boiler dact various types of deposits are formed depending on temperature zones and properties of coals. We distinguish the following types of deposits: slag‚ iron enriched‚ sulphur-bounded‚ fouling based on active sodium and fine dispersed (loose). The method has been suggested to define the tendency (potential) of a coal to form deposits of various types. It is taken into consideration‚ that similar deposits could be formed according to different patterns as well as the fact that the forming of deposits of various types may be a competitive process. The latter means‚ that in specific conditions there are deposits the rate of forming of that is higher. On the basis of a large experimental material it has been worked out methods of the approximated evaluation of slagging properties using the widely available data and recommendations concerning the use of information on slagging properties of coals for choosing the characteristics of the boiler.

REFERENCES Alekhnovich‚ A.M. (1996). “The Effect of Operation Conditions on Slagging”. In L. Baxter and R. DeSollar (Eds.). Application of Advanced Technology to Ash-Related Problems in Boilers. New York and London: Plenum Press‚ 485–493. Alekhnovich‚ A.N.‚ Bogomolov‚ V.V. (1994). “Slagging Property Indices of Coals and Their Use When Designing Boilers”. In I. Williamson and F. Wigley (Eds.). The Impact of Ash Deposition on Coal Fired Plants. Washington: Taylor & Francis‚ 725–732.

Alekhnovich‚ A.N.‚ Gladkov‚ V.E.‚ Bogomolov‚ V.V. (1996). “Slagging Prediction when Using the Chemical Composition of Fly Ash Individual Particles and the Slagging Probability Model”. In L. Baxter and

Prognosis of Slagging and Fouling Properties of Coals Based on Widely Available Data

469

R. DeSollar (Eds.). Application of Advanced Technology to Ash-Related Problems in Boilers. New York and London: Plenum Press‚ 557–565. Allan‚ S.E.‚ Erickson‚ T.A.‚ McCollor‚ D.P. (1996). “Modeling of Ash Deposition in the Convective Pass of a Coal Fired Boiler”. In L. Baxter and R. DeSollar (Eds.). Application of Advanced Technology to AshRelated Problems in Boilers. New York and London: Plenum Press‚ 451–470. Benson‚ S.A.‚ Steadman‚ E.N.‚ Zygarlicke‚ C.J.‚ Erickson‚ T.A. (1996). “Ash Formation‚ Deposition‚ Corrosion and Erosion in Conventional Boilers”. In L. Baxter and R. DeSollar (Eds.). Application of Advanced Technology to Ash-Related Problems in Boilers. New York and London: Plenum Press‚ 1–15.

Borio‚ R.W.‚ Patel‚ R.L.‚ Morgan‚ M.E.‚ Kang‚ S.G.‚ Erickson‚ T.A.‚ Allan‚ S.E. (1993). “The Coal Quality Expert: A Focus on Slagging and Fouling”. Pre-print of second Annual Clean Coal Technology Conference. Atlanta‚ USA.—Sept. 7–9. Cui‚ Y‚ Gong‚ D.‚ Zhou‚ Y‚ Li‚ X. (1994). “A Study of Testing Additives and Application of Chinese Coals”. In I. Williamson and F. Wigley (Eds.). The Impact of Ash Deposition on Coal Fired Plants. Washington: Taylor & Francis‚ 703–714. Gibb‚ W.H. (1996). “The U K Collaborative Research Programme on Slagging Pulverised Coal-Fired Boilers:

Summary of Findings”. In L. Baxter and R. DeSollar (Eds.). Application of Advanced Technology to Ash-Related Problems in Boilers. NewYork and London: Plenum Press‚ 41–65. Halt‚ R. (1996). Correlating the Slagging of a U t i l i t y Boiler with Coal Characteristics. In L. Baxter and

R. DeSollar (Eds.). Application of Advanced Technology to Ash-Related Problems in Boilers. New York and London: Plenum Press‚ 451–470. Hurley‚ J.P‚ Nowok‚ J.W.‚ Strobel‚ T.M.‚ O’Keefe‚ C.A.‚ Bieber‚ J.A.‚ Dockter‚ B.A. (1996). “Rales and Mechanisms of Strength Development in Low-Temperature Ash Deposits”. In L. Baxter and R. DeSollar (Eds.). Application of Advanced Technology to Ash-Related Problems in Boilers. NewYork and London: Plenum Press‚ 83–96. McCollor‚ D.P.‚ Zygarlicke‚ C.J. (1996). “Mechanisms of Ash Fouling during Low Rank Coal Combustion”. In L. Baxter and R. DeSollar (Eds.). Application of Advanced Technology to Ash-Related Problems in Boilers. NewYork and London: Plenum Press‚ 223–235. Mysen‚ B.O. (1987). “Redox equilibria and coordination of Fe2+ and Fe3+ in silicate glasses from 57 Fe

Mossbauer spectroscopy”. Journal of Non-Crystalline Solids‚ v. 95–96‚ N I ‚ 247–254.

Pollock‚ W.H.‚ Goetz‚ G.I.‚ Park‚ E.D. (1983). “Advancing the art of boiler design by combining operating experience an advanced coal evaluation techniques”. Proc. Amer. Power Conf.‚ v.45‚ 102–117.

Richards‚ G.H.‚ Harb‚ J.N.‚ Baxter‚ L.L. (1996). “Investigation of Mechanisms for the Formation of Ash Deposits for Two Powder River Basin Coals”. In L. Baxter and R. DeSollar (Eds.). Application of Advanced Technology to Ash-Related Problems in Boilers. NewYork and London: Plenum Press‚ 293–308. Senior‚ CM..‚ Srinivasachar‚ S. (1995). “Viscosity of Ash Particles in Combustion Systems for Prediction of Particle Sticking”. Energy & Fuels. Vol.9‚ 2‚ 277–283. Zygarlicke‚ C.J.‚ Galbreath‚ K.C.‚ McCollor‚ D.P‚ Tomen‚ D‚L. (1996). “Development of Fireside Performance Indices for Coal Fired Utility Boilers”. In L. Baxter and R. DeSollar (Eds.). Application of Advanced Technology to Ash-Related Problems in Boilers. NewYork and London: Plenum Press‚ 617–636.

This page intentionally left blank.

THE SLAGGING BEHAVIOR OF COAL BLENDS IN THE PILOT-SCALE COMBUSTION TEST FACILITY

Alexander N. Alekhnovich‚ Vladimir V. Bogomolov‚ Natalja V. Artemjeva‚ and Vladimir E. Gladkov Ural Heat Engineering Institute Chelyabinsk‚ Russia

1. INTRODUCTION Due to the reduction of economic links between the republics of the former Soviet

Union and new economic conditions‚ in particular the increases in supplies of coal on short-term contracts‚ the problems of burning alternative coals‚ coal switching and blending have become a high priority task for power stations in Russia. Especially interesting from the practical as well as the scientific point of view is the national program of expanding the use of coals from Kansk-Achinsk Basin. This basin‚ located in Siberia‚ contains vast deposits of coal‚ which are extracted by means of powerful machinery in the opencast collieries. The coal is characterized by a low ash content and production costs. Ash content on dry basis is less 12% in the coal of the largest opencast colliery in Berezovo. At the same time the coal there contains ash of lignite type with the CaO content up to 46–60% in the Berezovo opencast colliery and has high slagging properties. This coal is burned successfully in specially designed boilers of P-67 type (2‚650 t/h‚ 800MW) by Podolsk Machine-Building Factory (Vasilyev et al.‚ 1994) and E500 (500 t/h) by Barnaul Boiler Plant. However‚ outside the region of conventional usage of these coals the power stations are equipped with boilers designed for burning lowslagging coals with acid ash (acid / base ratio It is clear that this situation‚ to a certain extent‚ is similar to the situation when sub-bituminous coals of Powder River Basin‚ USA were used (Hurley et al.‚ 1994). On the basis of the previous experience it is known‚ that coal switching and blending may result in greater slagging as compared to burning each of these coals separately. For instance‚ the increased slagging during the burning of Kuznetsk coals with low slagging properties in boilers that usually burn the coals of Kizel and Donetsk Basins. The Kuznetsk coal is characterized by the fact‚ that certain batches of it contain low melting ash (temperature of the start of slagging )‚ however‚ there are no hard‚ Impact of Mineral Impurities in Solid Fuel Combustion‚ edited by Gupta et al. Kluwer Academic / Plenum Publishers‚ New York‚ 1999.

471

472

A. N. Alekhnovich et al.

selectively condensed primary deposits during its combustion. Unlike the Kizel and

Donetsk coals which contain more high-melting ash ( and 1‚030°C)‚ but when they are burnt hard iron enriched deposits are formed on the surfaces of furnace walls and a super-heater which are difficult to remove. Based on the results of research using the combustion test facility (Alekhnovich et al.‚ 1996) it was defined that certain indices of slagging properties of coals do not fall under the additive laws. Thus‚ when even a small fraction of window glass powder was added to the high-melting ash of the Donetsk coal‚ the temperature of the start of slagging reduced considerably and this reduction was greater than could be expected by the average chemical composition of the blend ash. The objective of the research‚ carried out by UralVTI was to obtain experimental data concerning the slagging properties of coals in the volume which could be obtained during short-term experiments (less‚ than 24 hours) on the combustion test facility. To obtain a larger material for generalization‚ we tested the coal blends‚ the production and burning of which in reality is not planned.

2. SCOPE AND METHODS OF RESEARCH The research was carried out on the combustion test facility by UralVTI‚ which has 50 t/h coal imput. The circuit diagram of the test facility is shown in Fig. 1 and the diagram of the combustion chamber and the ducts was published early (Alekhnovich et al.‚ 1996 a). The facility is equipped with a pulverizer system and has a non-isothermal vertical combustion chamber with an up-draft duct. The gases from the combustion chamber enter the horizontal “experimental” gas duct‚ wherein they are continuosly cooled from 1‚200 down to 700°C. The “experimental” duct is fitted with a set of probes for investigating the process of slagging and fouling‚ including the device for measuring the tensile strength of deposits in situ.

The blend was prepared by mixing portions of coals which had been preliminary crushed and dried before they were loaded in to the coal bunker.

During these tests we defined the rate of slagging‚ the temperature of the start of slagging and the strength of deposits. The slagging of heat surfaces usually takes place in stages. Initially a primary layer is formed. The proper slag deposits are formed when its surface temperature runs into the critical value. The formation of the primary layer is prolonged process and it can’t be reproduced in the pilot-scale test facility. Non-cooled probes are used in our investigations to study slag deposits. They simulate the real boiler tubes having the deposit layer. The slagging rate was tested with the help of non-cooled probes made from heatresisting steel tubes Use of ceramic or steel probes does not change the results‚ as it was determinated previously. The location of the probe was changed step by step along the horizontal “experimental” duct. Its temperature was measured continuously with the help of the thermocouple embedded into the probe. The flue gas temperature in the zone of the probe location was determined as The

temperature difference was determined sometimes in all cross sections where the probes were located with the help of the shielded exhausted thermocouple. The time of probe residence in a gas duct was recorded and it ranges between 10 and 15 minutes. The slagging rate “g” was calculated as

The Slagging Behavior of Coal Blends in the Pilot-Scale Combustion Test Facility

473

where m— deposit mass‚ gram; 1—probe length where deposits were collected in meters.

The dimensionless slagging coefficient was calculated as

where F is the duct cross section‚

was also used during analytical treatment. It

is ash content in pulverized coal and

is

coal input into the test facility. There are marginal temperatures of a surface and gas low that the slag deposits are not formed. It was studied early‚ that the marginal temperatures are depended as on coal mineral matter properties so on outer conditions‚ including the value of the difference It was suggested (Alekhnovich‚ Bogomolov‚ 1988) to use the term “the temperature of the start of slagging ” as the slagging index of coals and to determine it under the fixed outer conditions. In other cases the term “actual slagging temperature ” is used. The term “the temperature of the start of slagging” also is used in this paper as most commonly used. Really the value was determined in the tests as the difference was not constant and specified. The strength of deposits has been determined in hot conditions with the help of

A. N. Alekhnovich et al.

474

special device for measuring the tensile strength of deposits in situ (Fig. 2). Sintering tests were carried out using the fly ash‚ collected on the facility exit.

The chemical composition of mineral matter of coals‚ blends and fly ash was defined by the x-ray-fluorescent method. The initial coals were fractionated in heavy liquids and underwent x-ray-phased analysis. In a number of tests the chemical composition of fly ash individual particles was defined by SEM method. For a series of tests during the burning of coal blends with a maximally different chemical composition of mineral matter (Ekibastuz coal—Berezovo coal) it was prepared polished sections of the deposits and was made elemental composition maps of deposits using Comebax unit. The coals of different basins and coal-fields were used during these tests. Their specifications are shown in Table 1. The coal of Berezovo was obtained from different sources and the characteristics of its mineral matter in various series of tests varied greatly. The acid/base ratio (K/O) varied within the range of 1.1–4.4‚ where

and

For other coals one averaged sample was used. In Table 1 we can see‚ that the characteristics of tested coals changed within a rather wide range from very slagging coal of Berezovo to non-slagging coal of Ekibastuz

The complete range of tests was made during the burning of individual coals and blends. The composition of tested blends is shown in Table 2.

3. RESULTS OF AND DISCUSSION For all tested compositions the character of changes of the temperature of the start as a function of the fraction of coal with higher slagging properties is similar. When a part of such coal is insignificant decreases and afterwards‚ as its fraction increases the decrease of the temperature of the start of slagging becomes smooth. The effect of a sharp decrease of when we add small quantities of coal containing low of slagging

The Slagging Behavior of Coal Blends in the Pilot-Scale Combustion Test Facility

475

476

A. N. Alekhnovich et al.

melting ash is more noticeable for coals with higher value of (coals of Ekibastuz and Podmoskovnij). The results of the main series of tests are shown in Fig. 3. On the ordinate axis the value of the relative temperature of the start of slagging is shown‚ which means the ratio of of blends and of a less slagging coal The dependencies of the rate of forming of slag deposits on the temperature during the combustion of blends are an intermediate position between the dependencies of individual coals forming the blend. The characteristic examples of the measured rate of slagging for blends made of coals with different and close values of are shown in Fig. 4. In burning of blends of Podmoskovnij coals with Berezovo coal it was measured a higher rate of slagging than the one for individual coals. Within the temperature range between of one coal and of other coal the rate of deposit forming is higher‚ than the rate which could be obtained only due to the use of more melting coal considering its fraction (Fig. 4) That means‚ that the mineral matter of the high-melting coal takes part in slagging under the circumstances when the temperature is lower then its the temperature of the start of slagging. The results concerning t sl and the rate of slagging are not unexpected and could be explained from various points of view. Let us assumed that there is extensive interaction between mineral matter of one coal in the blend and the other coal during burning. Fly ash particles of new average composition are formed in this case. We had earlier defined the linear dependency of and ratio K/O (Alekhnovich‚ Bogomolov‚ 1988). The change of the acid/base ratio K/O of the ash of the average composition and calculated depending on the part of ash of the added coal is of the same character as experimental value of (Fig. 5). During the forming of new compositions the possibility the increase the rate of slagging at temperatures lower‚ than of coals with higher value becomes evident. In reality‚ as was admitted by the majority of researchers‚ the interaction of the

The Slagging Behavior of Coal Blends in the Pilot-Scale Combustion Test Facility

477

mineral matter between different coal particles even of one coal is not great before bonding in deposits. When we speak of blends it means‚ that the fly ash particles of two coals in a blend behave independently before they get into the deposit. It corresponds with the results of testing the chemical composition of individual particles of fly ash of a number of blends. According to these tests there is no significant amount of particles of a new composition and the experimental distribution of the particle content for blends is similar to the values which is obtained when we sum up the particles of fly ash of individual coals in a corresponding proportion. The example of the obtained result for the blend of Chelyabinsk coal with Berezovo coal in shown in Fig. 6. The analysis of the elemental composition maps of deposits also shows‚ that they are formed from fly ash particles characteristic for individual coals. In the absence if interaction of the mineral matter composing the blend of coals the results‚ concerning and may be explained by a matrix hypothesis. Its essence lies

478

A. N. Alekhnovich et al.

in the fact‚ that a part of particles has sticking properties high enough to keep the particles in the deposits that do not posses such properties. In burning of blends the particles of fly ash formed from coals with lower value of act as the matrix particles. The calculation of the slagging rate and using the probability model (Alekhnovich et a!.‚ 1996 b) showed the complete correspondence of kind with the test. In a number of cases

The Slagging Behavior of Coal Blends in the Pilot-Scale Combustion Test Facility

479

there was a quantitative correspondence as well. An example of the results of calculations is shown in Fig. 7. The tensile strength of deposits in their hot state has been measured during the research. This strength may be characterized as the strength of forming deposits‚ since the time from the installation of the probe till the loads were used did not exceed 20 30 minutes. In reality the deposits become hard due to sintering. The results of the process were imitated during the tests on sintering of fly ash. It is defined‚ that the correlation “strength —temperature” for blends is located between the correlation for corresponding individual coals. An example of the results obtained in shown in Fig. 8. For blends of all coals the results were approximately the same with certain experimental deviations which were not crucial. Unlike this‚ the results of sintering tests for blends of various coals are different in principle. For some of them the correlation’s “strength during sintering—temperature” are located between the correlation’s for individual coals‚ similar to the results concern-

ing strength. In other cases (blends of Chelyabinsk and Kuznetsk coal type G with Berezovo coal) the strength during sintering of blends is considerably greater‚ than for coals

forming the blend. The results of sintering tests of fly ash of two blends are shown in Fig. 9. The principal differences in sintering fly ash‚ obtained after burning of blends of various coals with the coals of Berezovo‚ to our mind‚ are connected with the differences in mineral matter of these coals. Following the results of X-Ray crystal analysis we defined‚ that the coals‚ where there were anomalies of sintering‚ there are‚ for the most parts the inclusions of sodium feldspar‚ and for other coals potassium feldspar. The

first have an unlimited mixability with anortyte the latter—limited latter—limited one. one. The The forming of anortyte when burning the Berezovo coal with high content of CaO is experimentally grounded.

480

A. N. Alekhnovich el al.

CONCLUSION The tests of the combustion test facility showed the following characteristics of slagging properties of blends. The temperature of the start of slagging decreases considerably even when a small part of coal with the lower temperature is present. Further increase of the part of such a coal changes only slightly. Within the range of from one coal to the of other coal the slagging rate in

The Slagging Behavior of Coal Blends in the Pilot-Scale Combustion Test Facility

481

burning of the blend is higher‚ than the amount of deposits‚ which may be formed only from coal particles with low The strength of the formed deposits during the burning of the blend is no higher‚ than the one for the coal in the blend with higher slagging properties. The sintering of the fly ash may be much higher‚ than the sintering of fly ash of individual coals. Preliminary recommendations for forecasting the slagging properties of blends have been formulated basing on the results of the research.

482

A. N. Alekhnovich el al.

ACKNOWLEDGEMENTS This work was financed by the United Power Grid of Russia and is published with the permission of UPG. The authors would like to thank Mr. V.V. Demkin (Department

of Science and Technology of UPG) for his kind attention and help in investigation setting up.

The Slagging Behavior of Coal Blends in the Pilot-Scale Combustion Test Facility

483

REFERENCES Alekhnovich‚ A.N.‚ Bogomolov‚ V.V. (1988). “Temperature Conditions of the Start of Slagging when Burning

Coals with Ash Acidic Composition”. Thermal Engineering‚ v.35‚ N 1‚ 20–24. Alekhnovich‚ A.N.‚ Bogomolov‚ V.V.‚ Artemjeva‚ N.V. (1996 a). “Investigation of Some Slagging Problems at the Rigs”. In L. Baxter and R. DeSollar (Eds.). Application of Advanced Technology to Ash-Related Problems in Boilers. New York and London: Plenum Press‚ 529–539. Alekhnovich‚ A.N.‚ Gladkov‚ V. E.‚ Bogomolov‚ V.V. (1996b). “Slagging Prediction when Using the Chemical

Composition of Fly Ash Individual Particles and the Slagging Probability Model”. In L. Baxter and R. DeSollar (Eds.). Application of Advanced Technology to Ash-Related Problems in Boilers. New York and London: Plenum Press‚ 557–565. Hurley‚ J.P.‚ Benson‚ S.A.‚ Mehta‚ A.K. (1994). Ash Deposition at Low Temperatures in Boilers Burning HighCalcium Coals“. In I. Williamson and F. Wigley (Eds.). The Impact of Ash Deposition on Coal Fired Plants. Washington: Taylor & Francis‚ 19–30. Vasilyev‚ V.‚ Belov‚ S.‚ Maidanik‚ M. (1994). “Slagging‚ Fouling‚ and Cleaning of Boiler Burning KanskoAchinsky Brown Coal”. In I. Williamson and F. Wigley (Eds.). The Impact of Ash Deposition on Coal Fired Plants. Washington: Taylor & Francis‚ 63–73.

This page intentionally left blank

IN SITU MEASUREMENTS OF THE THERMAL CONDUCTIVITY OF ASH DEPOSITS FORMED IN A PILOT-SCALE COMBUSTOR Alien L. Robinson‚ Steven G. Buckley‚ Gian Sclippa‚ and Larry L. Baxter Combustion Research Facility Sandia National Laboratories

Livermore‚ CA 94551-0969

1. INTRODUCTION The formation of ash deposits reduces heat transfer rates to furnace wall tubes‚ superheater tubes‚ and other heat transfer surfaces in coal-fired power plants. The magnitude of this reduction largely depends on the thickness‚ thermal conductivity‚ and absorbtivity of the deposits. Each of these parameters is determined by the physical and chemical characteristics of the deposit‚ which‚ in turn‚ depend on the combustion conditions‚ the inorganic composition of the fuel‚ the processes by which the deposits are formed‚ reactions within the deposits‚ and reactions between the deposits and the furnace gases. This paper examines the thermal conductivity of ash deposits. Raask [1985] and Wall et al. [1993] provide reviews of the subject. It is important to recognize that heat transfer through a deposit occurs by conduction through the solid‚ conduction and convection through the gas in the void spaces‚ and by radiation. Since the contribution of each of these mechanisms to the total heat transfer rate is difficult to quantify‚ a lumped parameter‚ the “effective” thermal conductivity‚ which accounts for heat transfer by all the different modes‚ is used to characterize the heat transfer rate through a deposit. The effective thermal conductivity is defined as‚

where Q is the total heat transfer rate through the deposit‚ and dT/dy is the temperature gradient within the deposit. In this paper‚ the phrase “thermal conductivity” is used to refer to the effective thermal conductivity of the deposit. Several studies have measured the thermal conductivity of ash deposits [Abryutin

and Karasina‚ 1970; Anderson et al.‚ 1987; Boow and Goard‚ 1969; Golovin‚ 1964; Mills Impact of Mineral Impurities in Solid Fuel Combustion‚ edited by Gupta et al. Kluwer Academic / Plenum Publishers‚ New York‚ 1999.

485

486

A. L. Robinson et al.

and Rhine, 1989; Mulcahy, et al., 1966]. The reported values span several orders of magnitude from 0.012 [Golovin, 1964] to [Wall et al., 1993]. Two studies report thermal conductivity values less than that of air, suggesting a Knudsen mode of heat conduction [Boow and Goard, 1969; Golovin, 1964], in which the mean free path of a gas molecule is larger than the characteristic pore dimension of the deposit, resulting in noncontiuum heat transfer. The effective thermal conductivity of ash deposits is thought to largely depend on the deposit physical structure [Wall et al., 1993]. Highly porous deposits of loose, unsintered, particulate matter generally have low values of thermal conductivity. Solid, sintered deposits have high values of thermal conductivity. Little experimental data exists to quantitatively define the relationship between deposit physical structure and thermal conductivity. Examinations of this relationship are generally based on theoretical analysis and a qualitative examination of experimental data. Since deposit structure is difficult to quantitatively measure, measurements of thermal conductivity are often presented as a function of deposit temperature, particle size, and deposit composition. The reported results are not entirely consistent regarding the influence of these parameters on deposit thermal conductivity. Changes in thermal conductivity appear well-correlated with deposit temperature because of the relationship between temperature, sintering, and deposit structure [Anderson et al., 1987; Mulcahy et al., 1966]. At temperatures below that at which the deposit begins to sinter (900 to 1,200 °C) the thermal conductivity is relatively constant; at temperatures above the onset of sintering the thermal conductivity rapidly increases with temperature. Sintering the deposit irreversibly increases its thermal conductivity. The reported measurements are less consistent with regards to the effect of particle size on thermal conductivity. The study by Boow et al. [1969] suggests that the thermal conductivity of an ash deposit varies linearly with the log of the mean particle size. The smaller the median diameter of the ash particle the lower the thermal conductivity. However, the measurements reported by Anderson et al. [1987] do not show any variation in thermal conductivity with particle size. Finally, results reported by Boow et al. [1969] suggest a dependence between the thermal conductivity and the presence of atoms that form mixed silicates (Fe, Al, C), so called coloring agents. The consistent shortcoming of the reported thermal conductivity measurements is that they are based on post mortem analysis techniques which destroy or significantly alter the physical structure of a deposit. Typically, powdered ash samples are examined; these samples are generated using a laboratory ashing furnace, captured fly ash from a power plant, or a pulverized deposit from a boiler. Although these samples may have the same chemical composition as boiler deposits, the morphology of the deposit has been destroyed. A few studies have examined intact, sintered deposits, however the thermal shock associated with removing them from a boiler likely altered the physical structure of the deposits. Consequently, because of the relationship between deposit physical structure and thermal conductivity, the reported measurements of thermal conductivity may not accurately characterize the effect of ash deposit on heat transfer rates in real furnaces. In this paper, we report in situ, time-resolved measurements of thermal conductivity of ash deposits formed on a temperature-controlled deposition probe in the Multifuel Combustor at Sandia National Laboratories. We describe the experimental technique. We discuss results from experiments conducted with an Illinois #6 coal and a blend of Illinois #6 coal and wheat straw. We compare the measured values with theoretically derived minimum and maximum thermal conductivity values.

In Situ Measurements of the Thermal Conductivity of Ash Deposits Formed in a Pilot-Scale Combustor

487

2. METHODS

2.1. Experimental Facility Experiments to measure the thermal conductivity of fly ash deposits were conducted using the Multifuel Combustor (MFC) at Sandia National Laboratories. A schematic of the MFC is shown in Fig. 1. The MFC is a pilot-scale (~ 30kW)‚ 4.2m high‚ down-fired‚ turbulent flow reactor that simulates gas temperature and composition his-

tories experienced by particles in pulverized-fuel combustion systems. The combustor has a circular cross section with an ID of 15cm. A more detailed discussion of the MFC is available in the literature [Baxter and Mitchell‚ 1992]. Pulverized fuel is injected into the MFC pneumatically via a water-cooled lance inserted through the side of the furnace (see Fig. 1). For these experiments‚ the fuel feed rate was set to maintain an oxygen concentration of 4% by volume (dry basis) in the combustion products at the exit of the reactor. Fuel was injected at the top of the combustor just below the natural gas burner‚ ~ 4.2m above the test section. Based on this distance and the gas flow rate‚ the calculated residence time of a fuel particle in the combustor was approximately 1 s. This is comparable to the residence time available for particle combustion in commercial systems. The natural gas burner was not fired during these experiments. Electric heaters are used to maintain the temperature of the walls of the combustor. For these experiments‚ the wall temperature was controlled at a setpoint 1‚150°C. The gas temperature in the flame zone (approximately the top 1 m of the combustor) is higher than the wall temperature. Below the flame zone‚ the gas temperature inside the reactor is at equilibrium with the wall temperature. The combustion products rapidly cool in the final‚ unheated section of the combustor (the last 0.6m of the combustor). The temperature of the gases exiting the combustor was ~950°C.

2.2. Fuel Selection two different fuels: Illinois #6 coal‚ a common high-volatile bituminous coal‚ and straw‚ a herbaceous biofuel. The coal is fired in pulverized form‚ commercial grind‚ 70% though

488

A. L. Robinson et al.

a 200 mesh Samples of biofuel were ground to pass through a 0.5-mm mesh. Results from two experiments are reported here‚ one firing 100% coal‚ and a second firing a blend of 50% coal and 50% straw by mass. The straw was blended with the coal to promote deposit formation.

2.3. Deposit Formation Ash deposits are collected on an air-cooled‚ stainless steel probe located in the test section of the MFC. The test section is the open region between the exit of the reactor and the exhaust duct (see Fig. 1) which provides optical access to deposition probe. The cooling air flow rate through the probe was chosen so that the average surface temperature of the probe was 500°C at the beginning of an experiment. The cooling air flow rate remained constant throughout each experiment. The deposition probe simulates a superheater tube in a commercial power plant. Since the gas and particle velocities through the test section of the MFC are roughly a factor of 4 smaller than typical convective pass velocities in the MFC versus in a power plant)‚ we cannot match both the Reynolds and Stokes Numbers found in typical power plants. For these experiments‚ the outside diameter of the deposition probe (2.2cm) was chosen to match the Stokes number of fly ash particles striking a heat exchanger tube in a commercial power plant because deposit formation in coal-fired power plants is dominated by inertial impaction. Consequently‚ the Reynolds number of the deposition probe used in this study is roughly a factor of 10 smaller than that of a typical superheater tube in a commercial power plant. Matching Stokes numbers ensures that size distribution of ash particles striking the deposition probe in the MFC will be the same as the size distribution of particles hitting a superheater tube. Therefore‚ the ash deposits formed in the MFC should have approximately the same physical structure and chemical composition as deposits formed in commercial coal-fired power plants. The difference in Reynolds number indicates lower heat transfer rates to the deposition probe in the MFC in comparison to a typical superheater tube. However‚ the surface temperature of the deposition probe is actively controlled to match those found in utility boilers. Particle impaction creates a multidimensional deposit on the upstream side of a superheater tube or deposition probe. The complex geometry of such a deposit makes it difficult to accurately determine its thermal conductivity. Therefore‚ we rotate the deposition probe at a speed of 0.25 rpm to generate a uniform‚ one-dimensional ash deposit. We assume that the rotation of the test probe does not affect deposit formation because the rotational velocity of the probe is four orders of magnitude smaller than the velocity of the particles striking the probe surface. Although an ash deposit is collected over an approximately 15-cm-long section of the deposition probe‚ for thermal conductivity analysis we only examine the center 6.4cm of the deposit‚ where the deposit thickness is most uniform. We refer to this section of the deposit and the deposition probe as the test section. The center of the test section is aligned with the centerline of the combustor to minimize axial temperature gradients within the deposit.

2.4. Instrumentation To determine the thermal conductivity of the deposit‚ we measure the thickness of the deposit‚ the surface temperature of the deposit‚ the surface temperature of the deposition probe‚ and the heat flux through the deposit. These measurements are made in situ

In Situ Measurements of the Thermal Conductivity of Ash Deposits Formed in a Pilot-Scale Combustor

489

while the deposit is formed on the deposition probe. The surface temperatures and heat flux measurements are recorded at 5-s intervals; the thickness of the deposit is measured approximately every 24 min. The surface temperature of the probe is measured using four type-K thermocouples embedded into the outside of the probe wall. To monitor the azimuthal variation in the probe surface temperature, the thermocouples are embedded 90° apart. To monitor the lengthwise variation in surface temperature, the thermocouples are distributed axially along the test section. The thickness of the deposit is measured using a range-finding laser (Selcom, Model 2207-32/180-B) mounted on a precision bearing. The bearing is set up so that the laser scans horizontally along the probe axis. The system measures the location of a surface located in the exhaust stream of the combustor with a resolution of under typical coal combustion conditions. By comparing results of successive scans along the same line or face of the probe, we determine the thickness and growth rate of the deposit. Scans are taken along 4 different lines each located 90° apart to verify that the deposit thickness increases uniformly. Three optical pyrometers (Accufiber, Model 100C) monitor the surface temperature of the deposit. The pyrometers focus at the center of the test section, 45°, 110°, and 170° below the flow stagnation point. Since the pyrometers detect emitted, scattered, and reflected radiation along their line of sight, radiation scattered and emitted by particles and radiation reflected by the deposit surface interfere with the surface temperature measurements. Shields prevent particulate matter and nearly all of the scattered and reflected radiation from passing through the pyrometer line of sight. The shields extend from the lens of pyrometer to within mm of the surface of the deposit. We fit a sinusoid to the temperature measurements to determine the surface temperature profile of the entire deposit. Periodically we change to the axial position of the pyrometer to verify that there is not an axial temperature gradient along the deposit surface. Using a laser pyrometer, we measure the spectral emissivity of the bare probe at the beginning of an experiment and of the deposit at the end of each experiment. The emissivity of the surface is determined based on the reflected light from the surface. The deposit emissivity measurement is made post mortem immediately after removing the deposition probe from the combustor test section. The measurement is generally made before the temperature of the deposition probe has decreased by more than 100°C. The heat transfer rate through the deposit is determined using two thermocouples mounted along the centerline of the deposition probe. The thermocouples are mounted 6.4cm apart and define the outside edges of the deposit test section. To ensure that the air flow through the probe is well-mixed the inside surface of the probe has been rifled and screens are mounted immediately upstream of each thermocouple. The heat transfer rate is calculated based on the measured mass flow rate and the temperature increase of the cooling air.

2.5. Data Interpretation For analysis, the deposit is considered to be a cylindrical shell of uniform thickness. The deposit geometry is defined by the outside diameter of the deposit probe and the measured deposit thickness. Assuming steady-state, two-dimensional heat transfer through the deposit and uniform deposit thermal conductivity, the deposit temperature distribution is described by,

490

A. L. Robinson et al.

The significance of the second-term of Eqn. (2) depends primarily on the thickness of the deposit. For most deposits, the radial terms dominate. Using Eqn. (2), we numerically solve for the temperature distribution within the deposit. The boundary conditions are the measured temperature distribution on the inside and outside surface of the deposit. Based on the deposit temperature distribution, we calculate the average temperature gradient at the inside edge of the deposit,

Combining the result from eqn. (3) and the measured heat transfer rate we obtain the effective thermal conductivity of the deposit,

where Q is the measured heat transfer rate through the deposit, and is the area of the inside surface of the deposit ( is the radius of the inside surface of the deposit, 1.1 cm, and L is the length of the test section, 6.4cm).

3. RESULTS AND DISCUSSION In this section, we first examine the complete set of measurements recorded while cofiring Illinois #6 coal with straw to illustrate how we determine the various parameters required to evaluate the thermal conductivity of an ash deposit. We then examine

and compare the time series measurements of thermal conductivity of the ash deposit formed while tiring Illinois #6 coal and the deposit formed while cofiring Illinois #6 coal with straw.

3.1. Surface Temperature, Deposit Thickness, and Heat Flux Measurements Results from thickness scans of the deposit formed while cofiring Illinois #6 coal and straw are shown in Fig. 2. These scans were taken along the same face of the probe; recall that thickness scans are taken along four different faces to verify that the deposit grows uniformly. Although a total of 8 scans along each face were recorded during the experiment, results from only 3 scans are shown in Fig. 2 for visual clarity. The scan taken at the beginning of the experiment is indicated by the line labeled t = 0. The variation in this scan illustrates the magnitude of the noise, in the measurement due to fly ash and burning char particles passing through the beam path of the scanning laser and beam steering. Subsequent scans clearly indicate the deposit growing with time. The edges of the test section are indicated by heavy vertical lines. Although deposit thickness increases at the edges of the deposition zone, within the test section the height of the deposit is relatively uniform.

Determination of the deposit thickness requires accurate measurement of both the location of the probe and the deposit surface. We refer to the location of the probe surface as the baseline. Unfortunately, we cannot use the results from the t = 0 scan as a baseline for subsequent scans because the probe thermally deforms over the course of an

In Situ Measurements of the Thermal Conductivity of Ash Deposits Formed in a Pilot-Scale Combustor

491

experiment. We determine a baseline for each scan by removing deposit from the outside edges of the deposition zone. This does not disturb the deposit in the test section because the deposition zone is approximately 15-cm wide whereas the test section is only the center 6.4cm of the deposition zone. Consequently, each scan records the location of the probe surface on both sides of the test section. The baseline is determined by fitting a straight line to the measured probe surface location on each side of the test section. By subtracting this baseline from the thickness scan, we calculate the thickness of the deposit. The average thickness of the deposit is calculated by averaging the measurements within the test section. Time series of the parameters required to estimate effective thermal conductivity are shown in Fig. 3. Deposit thickness is shown in Figure 3a. Average probe and deposit surface temperature are shown in Figure 3b. The heat transfer rate through the deposit is shown in Fig. 3c. These measurements were made while cofiring Illinois #6 coal and straw. The deposit thickness results shown in Figure 3a are based on the previously described analysis of the thickness scans. The symbols indicate the average thickness measurement for each probe face. For a given set of scans, the deposit thickness varies by less than between the different faces. The line represents the average thickness for the entire test section as a function of time. The deposit grew throughout the 3-hr experiment, reaching a maximum thickness of almost 1.7mm. Initially the thickness grows rapidly with the growth rate tapering off at increased time. This measured change in the deposit growth rate could be due to change in deposition rate or to sintering of the deposit. The deposit and probe surface temperatures are shown in Fig. 3b. The line labeled “probe surface” is the average of the measurements of the 4 embedded thermocouples. The measurements of the individual thermocouples sinusoidally fluctuate about this average. The line labeled “deposit surface” is the average deposit surface temperature calculated from the optical pyrometer measurements and the measured emissivity of the oxidized deposition probe, 0.85, and of the deposit, 0.75, respectively. The measured surface emissivity of the deposit agrees with previously reported values of emissivity [Wall et al., 1993]. We assume that 12 min into the experiment enough deposit has formed to reduce the surface emissivity from 0.85, the measured bare probe emissivity, to 0.75,

492

A. L. Robinson et al.

the measured deposit emissivity at the end of the experiment. For the conditions of these experiments, this change in emissivity has little affect on the calculated temperature— changing the emissivity from 0.85 to 0.75 increases the surface temperature by less than 5°C. At the beginning of the experiment, before deposit formation, the “deposit surface” temperature measured by the pyrometer agrees with the probe surface temperature measured with the thermocouples. As the deposit grows, the deposit surface temperature increases while the probe surface temperature falls. The constant cooling air flowrate in combination with the insulating effect of the deposit causes the large decline in the probe surface temperature. To maintain a constant probe surface temperature, the flow rate of the cooling air would have to decrease, which would cause a large increase in the deposit surface temperature. However, during these experiments, this flow rate was held constant to best simulate a utility boiler. The measured heat transfer rate through the deposit is shown in Fig. 3c. As the deposit grows the heat transfer rate decreases. Over the course of the 3-hr experiment, the heat transfer rate decreased by 20%. We analyze a surface temperature data to estimate the temperature difference across the deposit. To illustrate this analysis, ten minutes of time-resolved measurements of the probe and deposit surface temperature are shown in Fig. 4. The probe surface temperature measurements oscillate sinusoidally with the temperature traces from the 4 different thermocouples each being 90° out of phase. The deposit surface temperature measure-

In Situ Measurements of the Thermal Conductivity of Ash Deposits Formed in a Pilot-Scale Combustor

493

ment is relatively constant because each pyrometer is focused on one point in space, not one point on the probe surface. The variation in the deposit surface temperature is caused by particles (fly ash and burning char) passing through the line of sight of the pyrometer beam steering. As described earlier, shields are used to minimize the magnitude of this noise, but the shields do not extend all the way to the deposit surface. To estimate the temperature difference across the deposit, we first determine the azimuthal distribution of the probe and deposit surface temperatures. Figure 5 presents these distributions estimated from the 10-min period shown in Fig. 4. This analysis requires aligning the probe surface temperature measurements (each thermocouple signal is 90° or 1 min out of phase), and fitting a sinusoid to the measured deposit surface temperature data. The peak probe and deposit surface temperatures are assigned a probe orientation of 0°, the leading edge of the deposition probe. To determine the deposit surface temperature profile, we fit a sinusoid based on the average temperatures measured by each

pyrometer and the angular orientation of the pyrometers. In a separate experiment, we made surface temperature measurements at many angles to verify that a sinusoid accurately represents the azimuthal variation of the deposit surface temperature.

3.2. Thermal Conductivity of Ash Deposits Before examining the experimental results, we establish a realistic range of ash deposit thermal conductivity. The thermal conductivity of air is the lower limit of this range. An upper limit for this range is determined by treating the thermal conductivity of the gas and solid phases of the deposit in parallel,

494

A. L. Robinson et al.

where is the porosity of the deposit, is the thermal conductivity of the solid phase of the deposit, and is the thermal conductivity of the air within the deposit. To calculate these limits, we use the average deposit temperature, the measured deposit porosity, and assume the solid phase of the deposit has a thermal conductivity of (this value corresponds to amorphous at 800 K). We determine the porosity of

the deposit at the end of each experiment. The value is calculated based on the volume of the deposit (estimated from the last thickness scan), the deposit mass (determined by scraping the deposit off the probe), and assuming a density of the solid material within the deposit of These limits are intended to provide a rational basis for comparison; they are not theoretical minimum and maximum values for thermal conductivity of ash deposits. Knudsen conduction could reduce the thermal conductivity of deposit to values less than that of air. Other heat transfer modes through the deposit (e.g. radiation) could result in effective thermal conductivity values greater than those predicted by our analysis. Based on the measured deposit thickness, heat flux, and deposit temperature difference, we calculate the thermal conductivity of the ash deposit using Equations (2)-(4). Although the time series measurements of surface temperature and heat flux are continuous, we only calculate the thermal conductivity at 24min intervals immediately before the thickness scan.

Results for the deposit formed while cofiring Illinois #6 coal and straw are shown in Fig. 6. Figure 6a shows the average temperature difference and the average temperature gradient across the deposit. Figure 6b shows the effective thermal conductivity for the deposit formed while cofiring at an average deposit temperature of The

thermal conductivity of the ash deposit formed while cofiring increases throughout the experiment, starting at a value of and reaching a value of after three hours. The increase thermal conductivity may arise from the deposit sintering. The deposit porosity at the end of the experiment was 0.85. Results from the experiment conducted while firing Illinois #6 coal are shown in

In Situ Measurements of the Thermal Conductivity of Ash Deposits Formed in a Pilot-Scale Combustor

495

Fig. 7. Figure 7a shows the average temperature difference and the average temperature gradient across the deposit. Figure7b shows the effective thermal conductivity at an average deposit temperature of

The thermal conductivity of the deposit formed

while firing Illinois #6 coal varies between 0.12 and The deposit porosity at the end of the experiment was 0.88. Comparing the thermal conductivity results shown in Figs. 6 and 7 indicates that there is not a significant difference between thermal conductivity of the ash deposit formed while firing 100% Illinois #6 coal and formed while cofiring Illinois #6 coal and straw. The measured values of thermal conductivity fall within the estimated range of likely values. The thermal conductivity of both deposits is a factor of 2 to 3 greater than the thermal conductivity of air ( at the average temperature of the deposit). The thermal conductivity of these deposits is more than an order of magnitude smaller than the upper limit, (calculated using equation and

). We hypothesize that the measured thermal conductivity values are much closer to the lower limit because these deposits formed of highly porous, unsintered particles,

4. CONCLUSION The results described in this paper demonstrate the capability to measure in situ the thermal conductivity of ash deposits formed in the MFC. The measurement technique does not disturb the natural physical structure of the deposit. This capability is significant because the structure is thought to largely determine the thermal conductivity of the deposit. For the fuels examined to date, the measured thermal conductivity values of the ash deposits are a factor of 2 to 3 greater than air. Except for the growth rate of the deposit, there is little difference between the measured thermal conductivity of

496

A. L. Robinson et al.

the deposit formed while firing Illinois #6 coal and the deposit formed while cofiring Illinois #6 coal and straw.

To date, we have only measured the thermal conductivity of relatively low temperature, highly porous, unsintered ash deposit. These deposits should have the lowest values of thermal conductivity. In the future, we plan to conduct higher-temperature experiments to determine the thermal conductivity of deposits during sintering.

5. ACKNOWLEDGEMENTS The authors thank Don Hardesty for management of the project. This work was sponsored by the U.S. DOE FETC, Advanced Research and Technology Development Coal Utilization Program.

6. REFERENCES Abryutin A. A., and Karasina, E. S. (1970). “Thermal conductivity and thermal resistance of ash deposits in boiler furnaces.” Teploenergetika (Thermal Engineering), 17(12), 46–50. Anderson, D. W., Viskanta, R., and Incropera, F. P. (1987). “Effective thermal conductivity of coal ash deposits at moderate to high temperatures.” Transactions of the ASME Journal of Engineering for Gas Turbines and Power, 109, 215–221.

Baxter, L. L., and Mitchell, R. E. (1992). “The release of iron during the combustion of Illinois #6 coal.” Combustion and Flame, 88(1), 1–14.

Boow, J., and Goard, P. R. C. (1969). “Fireside deposits and their effect on heat transfer in a pulverized-fuellired boiler: Part III. The influence of the physical characteristics of the deposit on its radiant emittance and effective thermal conductivity.” Journal of the Institute of Fuel, 42(412–419). Golovin, V. N. (1964).“Investigation of tube fouling in a TP-90 boiler.” Teploenergetika, 11(3), 23–28. Mills, K. C., and Rhine, J. M. (1989). “The measurement and estimation of the physical properties of slags formed during coal gasification 2. Properties relevant to heat transfer.” Fuel, 68, 904–910. Mulcahy, M. F. R., Boow, J., and Goard, P. R. C. (1966). “Fireside deposits and their effect on heat transfer

in a pulverized-fuel-fired boiler Part I: The radiant emittance and effective thermal conductance of the deposits.” Journal of the Institute of Fuel, 39, 385–394. Raask, E. (1985). Mineral impurities in coal combustion: Behavior, problems, and remedial measures. Washington: Hemisphere Publishing Corporation.

Wall, T. F., Bhattacharya, S. P., Zhang, D. K., Gupta, R. P., and He, X. (1993). “The properties and thermal effects of ash deposits in coal-fired furnaces.” Progress in Energy and Combustion Science, 19, 487–504.

LOW CORROSIVITY OF COAL CHLORINE Elliott P. Doane1 and Murray F. Abbott2 1

Kerr-McGee Coal Corporation Manager, Energy Technology 123 Robert S. Kerr Street Oklahoma City, OK 73102 2 CONSOL Inc. Research and Development 4000 Brownsville Road Library, PA 15129

1. INTRODUCTION Coal chlorine (Cl) has been associated for many years with fireside fouling and corrosion problems in pulverized coal (p.c.)-fired utility boilers in the United Kingdom (UK). More recently Cl in municipal waste has been a source of corrosion in municipal waste-to-energy plants worldwide. In the United States (US), maximum guidelines or even limits ranging from 0.25-0.30 wt.% Cl have been utilized by most boiler manufacturers [Abbott et al., 1994], reportedly based on the extensive UK experience as well as other data. This caused many US utilities to avoid buying such coal, even though tens of millions of tons of it from the Illinois Basin have been burned for decades by other utilities without any adverse effects [Abbott et al., 1994; Wright and Krause, 1994]. Moreover, critical analyses of the laboratory and plant data has not either defined the mechanism through which coal chlorine may participate in corrosion of utility boilers or assigned it a definite role in such corrosion [Stringer and Banerjee, 1991]. The preponderance of current evidence shows that the high-temperature corrosivity of the hydrogen chloride (HC1) from Cl in Illinois Basin coals is negated by the coal sulfur (S) combustion products [e.g. Krause, 1991]. This mechanism is operative under both waterwall and superheater/reheater (SH/RH) conditions. Long-term experience burning Illinois Basin coals without corrosion attributable to Cl supports the contention that the Cl content of these coals is not a significant factor in high-temperature corrosion. Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

497

498

E. P. Doane and M. F. Abbott

2. UTILITY EXPERIENCE BURNING ILLINOIS BASIN STEAM COALS The most compelling evidence that Cl in Illinois Basin coals is not a factor in boiler equipment corrosion is its long history of utilization for steam generation with no apparent chloride-related fireside corrosion problems. More than 1.3 billion tons of Illinois Basin steam coal has been burned by utilities in the past two decades. Therefore, an analysis of coal purchased by utilities burning Illinois Basin steam coals was completed to identify boilers that burned a high proportion of deep-mined (more than 400 feet depth) Illinois coal, which is more likely to be high-Cl, over many years [Abbott et al., 1994]. Those power stations identified as having burned 30% or more of deep-mined Illinois steam coals on average over an eight-year period from 1983-1990 are listed in Table 1. Estimated coal Cl content is provided for most plants. These power plants utilize boilers manufactured by Babcock & Wilcox Company, ABB-Combustion Engineering, Inc., and Foster Wheeler Energy Corporation with rated capacities from 94 to 952 MW. Threefourths of them started up prior to 1975. Thus, this long-term experience covers most p.c.-fired boiler designs. Although Table 1 covers an eight-year period, many of these stations have actually burned these coals far longer. If combustion of Illinois Basin coals with Cl levels

at or above 0.3% causes significant high-temperature corrosion or operating problems, it would have been observed at these power stations over this time period. Accordingly, a knowledgeable boiler consulting company, Stringfellow Energy Consultants, was engaged in 1991 to interview appropriate personnel at each of these power stations. Each was asked what problems were experienced over the years burning Illinois Basin coals, and whether any of these problems could be attributed to coal Cl content. None of the power station personnel contacted identified any operating problems attributable to coal Cl. A more recent EPRI survey reached essentially the same conclusions [Wright et al., 1994].

Low Corrosivity of Coal Chlorine

499

3. CHLORINE IN UK AND US COALS Most of the Cl in both US Illinois Basin and UK coals is present as chloride ions weakly bound to positively charged sites on the coal surface [Chou et al., 1994; Pearce and Hill, 1986]. There is not enough Na present in most coals to balance the Cl content stoichiometrically, as shown in Fig. 1. The data in Fig. 1 are for 44 commercial UK coals [Gibb and Angus, 1983] and for 34 currently marketed Illinois coals [Demir et al., 1994]. Coal Cl is rapidly evolved as HC1 during combustion. Each 0.1 wt.% Cl in bituminous coal corresponds to about 80ppmv in the flue gas.

There is a strong correlation between Cl and Na in UK coals, even though their relationship is not stoichiometric. The data on UK coals in Figure 1 show a linear correlation coefficient r = 0.75. The r value given in the original article was 0.82, but recent reanalysis have obtained a value of r = 0.75, [Doane and Abbott, 1995; James et al., 1995]. The latter reference also reports that an analysis of a detailed study of chlorine in UK coals in the late 1950’s gives a linear correlation coefficient r = 0.88. This strong correlation between Cl and Na in UK coals was essential to the past use in the UK of coal chlorine, which could be readily determined, as an index of alkali metals content, for which

analyses were not commonly available. As Cutler and Raask [1981] put it: “The chlorine content of coals has thus been used to indicate the equivalent amount of non-silicate alkali metal in the coal as a criterion for the corrosion propensity of the coal.” There is no meaningful correlation between Cl and Na for the 34 Illinois coals in Figure 1, for which r = 0.18. Therefore, Cl content would not be a valid index of alkali metal content, as is the case with UK coals. Consequently, acid-soluble alkali metal content, not Cl

content, has been the criterion for predicting corrosion propensities of coals in US and Japanese studies [Abbott et al., 1994]. In addition to these differences in the relationship of Cl with Na, the UK coals forming the basis of UK experience have higher ash contents than do Illinois coals. The average ash content (dry) for the 44 UK coals [Gibb and Angus, 1983] is 19.2%, while the average ash content (dry) for the 34 Illinois coals [Demir et al., 1994] is 10.2%. This difference is due mainly to the fact that virtually all the Illinois coals are washed, while most of the UK coals are not. The effect of conventional gravity separation and froth

500

E. P. Doane and M. F. Abbott

flotation processes on Illinois No. 6 coal is to lower significantly its slagging and fouling tendencies, and to increase the furnace exit gas temperature at which fluid slag is first observed by about 200°F [Harrison and Parkinson, 1983]. The higher ash content of the U K coals on which their furnace corrosion experience is based may have adversely impacted corrosion as well. It will be seen in the discussion of corrosion mechanisms that normal boiler conditions drive deposit composition toward sulfates in preference to chlorides. But a recent thermodynamic study [Seifert and Born, 1995] shows that a molar excess of calcium (Ca) to sulfur (S) results in increased S and Cl in the mineral phases, up to at least 800°C. The higher ash content of the UK coals makes their average molar Ca/S ratio 0.15, which is about twice that for the Illinois coals. Obviously, this ratio in coals from both sources is far below stoichiometric. But it is another difference between UK and US Illinois coals that may provide a clue to the apparently benign corrosion performance of washed Illinois coals.

4. FURNACE WALL CORROSION Small-scale studies simulating waterwall conditions have not demonstrated a correlation between HC1 in p.c. combustion gases and high corrosion rates of mild steel.

Experience with commercial boilers, including those with lowburners, generally has not shown a significant effect of coal Cl. Current thinking is that severe furnace wall corrosion rates result from poor combustion conditions, with high H 2 S levels (correlated with high CO levels), high heat flux, slag deposits, and especially flame impingement.

4.1. UK Experience Furnace wall corrosion was initially correlated linearly with coal Cl content, using data from Rugely Station in the UK [Lees and Whitehead, 1983]. Almost since commissioning, fireside corrosion rates exceeding l00nm/hr had been recorded in all five boilers at this Station. The Cl content of the coal burned there increased from 0.35% to 0.65% from 1967 to 1977, and increased corrosion rates were observed. However, the correlation with coal Cl failed as more data became available from a wider range of p.c.-fired boilers. A recent examination of all available UK furnace waterwall data [James et al, 1995] concludes that increases in coal Cl content increase the corrosion rate only where corrosion rates already exceed l00nm/h with low-Cl coal, already a high corrosion rate by most standards. There is no apparent influence of coal Cl on furnace wall corrosion where corrosion rates are below this threshold value. In UK and US utility experience, waterwall corrosion rates of l 0 0 n m / h or more require poor combustion conditions with flame impingement, which combines high heat flux, high CO levels, slag deposits, unburned coal particles, and an aggressive sulfidizing environment. Such conditions should not be present in a properly designed and well-operated boiler.

4.2. Low-

Burners

Lowburners are designed to carry out a substantial portion of devolatilization and combustion in a fuel-rich flame under reducing conditions. In flue gas from coal combustion, reducing conditions result in high concentrations of CO and and a sulfidizing environment. The corrosion studies discussed later show that corrosion rates in simulated coal flue gases are an order of magnitude higher under reducing than under

Low Corrosivity of Coal Chlorine

501

oxidizing conditions. Yet in most studies high levels of HC1 do not increase significantly the already high corrosion rates under reducing conditions, even in the strongly reducing conditions in coal gasifiers. It was feared in the UK that installation of low-NOx burners would increase furnace wall corrosion due to increased flame length, higher amounts of unburned coal, and locally reducing conditions. Such concerns have proved largely groundless [James et al., 1995]. Experience in the UK with corner-fired units fitted with lowburners in the 1970’s has shown dramatic reductions in corrosion rate of

the front wall. Similar results have been reported on a corner-fired supercritical boiler, although this coincided with a reduction in load factor. Although initial experience with lowburners in front-fired units was not as encouraging, secondary air adjustments have resulted in lower corrosion rates than before in these units as well. This is attributed to the fact that the lowburners contain the flame vortex within a layer of excess oxygen and promote a more diffuse flame that reduces peak incident heat fluxes. No adverse effects of coal Cl have been reported with properly adjusted lowburners. Actually, this is consistent with the earlier UK experience that showed no effect of Cl as long as corrosion rates were less than l00nm/h. Coal Cl content has not been shown to be a factor in corrosion of furnace walls in the US, with or without lowburners. This may be because the design and operation of boilers in the US has been adequate to keep waterwall corrosion rates below l00nm/h. However, rapid corrosion has been observed on furnace walls near lowburners in several US boilers. When lowburners were installed in four 676 MW wallfired supercritical steam generators at Gibson Station, furnace wall corrosion increased to the point of becoming a significant maintenance problem [Urich and Kramer, 1996].

The units were burning Illinois Basin coal from the nearby Wabash Mine, containing

0.15–0.2% Cl. A computational fluid dynamics study showed that increased corrosion was occurring throughout those areas where reducing conditions existed. Gas analyses taken at the furnace side walls at the same plant showed a good linear correlation between CO and concentrations, with 5% CO corresponding to 300 to 360ppmv of [Gabrielson and Kramer, 1996]. Chlorine was ruled out as a major contributor. Modifications that changed flow patterns within the furnace to eliminate reducing zones at the walls successfully minimized corrosion.

4.3. Laboratory and Combustion Rig Tests The greater control over variables in laboratory tests led the Central Electricity Generating Board in the UK to carry out extensive studies in the late 1970s and early 1980s of the effect of HC1 on corrosion of mild steel and on scale morphologies summarized in Doane and Abbott [1995] and James et al. [1995]. Unfortunately, these studies never reproduced the highest rates of waterwall corrosion or the scale morphologies observed in some UK boilers. Under simulated waterwall oxidizing conditions, average corrosion rates over l,000h at 400°C and 500°C were only 40–85nm/h. Varying HCl concentration from zero to 2,000ppmv increased corrosion rates moderately at 400°C, but had no effect on corrosion rates at 500°C. Under simulated waterwall reducing conditions, corrosion rates remained below 200nm/h at 400°C, but were about 400–600nm/h at 500°C. Varying HCl concentration from zero to l,000ppmv had no effect on corrosion rates at either temperature. Later work at the University of Leeds using a small methane fueled combustion rig has been reported [James et al., 1995]. Higher heat flux increased corrosion rates sharply in the presence of HCl, or both together. Although the addition of HCl caused

502

E. P. Doane and M. F. Abbott

enhanced rates of attack under reducing atmospheres, the presence of and the formed from it tempered the effect of HC1 on corrosion. This is entirely consistent with

other results discussed later in this paper.

Results of a recent combustion rig study [Davis et al., 1997] confirm only that expo-

sure to strongly reducing conditions, consistent with flame impingement, accelerate metal loss. The innovative corrosion methodologies used in this study constitute a valuable tool for the investigation of fireside corrosion. However, the experimental design and data analysis are open to criticism. With regard to experimental design, the three coals tested do not provide a wellbalanced set of values for sulfur and chlorine. It would be better to use six different coals to provide a more complete representation of coal compositions than to include what are essentially replicate runs. Mapping the sulfur and chlorine compositions more uniformly, and over a greater range, would strengthen the validity of the results. With regard to data analysis, the functional form of the correlation does not consider that corrosion is a consequence of local conditions. The contribution of ash deposits to the corrosion mechanism should be included, especially when the coal ashes exhibit significant differences in iron content. Although the (CO) and do not occur together, the time weighted combination implicitly assumes that the corrosion reactions are of the same order, which first needs to be established. The analysis of the data develops curves that are extrapolated below the range for coal chlorine studied, and con-

sequently suggests a pronounced minimum corrosion rate near the extreme limit of the compositions tested. The maximum corrosion rate for sulfur occurs within the variability of the analysis of a single coal. It may well be that all the detail observed in the data analysis is an artifact of curve fitting with limited data. It would be better to focus on

the magnitude of errors related to the estimated values to bring out the validity of the data. The recent work of Kung et al. [1996] does not confirm even the moderate effect of coal Cl under substoichiometric conditions in the most recent UK studies. The corrosion of several commercial alloys was investigated in a laboratory retort at 371°C and 482°C for 1,000 hours. The reducing gas environment contained 5.6% CO, 0.6% 500ppm and either no HC1 or 350ppm HC1, corresponding to about 0.44% Cl in coal. Little or no effect of the added HC1 was observed at 371°C, and a definite beneficial effect was found at 482°C. Corrosion experience in equipment following coal gasifiers also shows little effect of coal Cl under extreme reducing conditions at furnace wall temperatures. In coal gasification plants, virtually all the sulfur is present as and the 0.1 wt.% Cl in coal which would correspond to about 80ppmv HC1 in boiler flue gas would result in 250 to 500 ppmv HC1 in an oxygen-blown gasifier. Pressurized gasifier operation greatly increases the partial pressure of this HC1 compared to that in boiler flue gas. Yet there appears to be little or no effect of HC1 on corrosion. Bakker and Perkins [1991] believe chlorine effects on high-temperature corrosion in gasifiers are probably of secondary importance. They found that addition of 600ppm HC1 to a simulated syngas containing 44.4% CO,

29.6% and 0.6% at 350°C reduced corrosion rates of several alloys. General material experience with syngas coolers has been very encouraging, whether or not the coal is high in Cl [Bakker, 1993]. Corrosion studies on stainless steel have failed to demon-

strate a clear effect of HC1 contents up to 5,000ppmv on the corrosion rate. Recently [Bakker, 1997] showed significantly increased corrosion of alloys with 400ppmv HC1 at 540°C in strongly reducing syngas compositions, although the effect disappeared whenever calculated oxygen partial pressures exceeded the extremely low level of atm.

Low Corrosivity of Coal Chlorine

503

4.4. Role of S and Cl in Furnace Wall Corrosion In studies of incineration of municipal waste high in Cl but low in S, Krause et al.

[1979] found high corrosion rates and high Cl levels in ash deposits on metal surfaces at furnace wall temperatures. When either sufficient elemental S or high-S coal was fired with the waste, the Cl content of the ash deposits declined to below the detection limit. Corrosion rates correspondingly declined to essentially the rates observed when firing coal alone. This is consistent with the similar effect of or observed under simulated furnace wall conditions in the University of Leeds work discussed under Laboratory and Combustion Rig Tests. The effect on initial (8 h) corrosion rates on A106 carbon steel is illustrated in Fig. 2 [Krause et al., 1979]. Blends of up to 74 wt.% municipal refuse with coal containing 3% S in a stoker-fired boiler caused essentially the same corrosion

rate as that for firing 100% coal. At 74wt.% refuse, the sulfur-to-chlorine (S/C1) ratio is only 2.2, which would result in a SO2/HC1 ratio of 2.4. This indicates that the minimum S/C1 ratio in coal required to negate high temperature coal Cl corrosion of A106 steel is below 2.2. A laboratory study of the effects of HC1 and chlorides on the high temperature oxidation of steel and alloys [Grabke et al., 1995] confirms that a molar ratio of

HC1 = 2.0 is sufficient to reduce the corrosion rate to the baseline value measured

with no HC1. Figure 3 below from that study shows weight gains at 500°C of steel preoxidized and then coated with flyash from a waste incinerator. The Figure also indicates that a ratio of /HC1 = 1.0 is insufficient.

5. SUPERHEATER/REHEATER (SH/RH) CORROSION Alkali iron trisulfates, and possibly pyrosulfates, are considered to be the major corrosive species in SH/RH corrosion in coal-fired boilers [Latham et al., 1991]. A UK correlation with Cl content may be confounded with sodium (Na) content, because of their strong correlation in UK coals. Numerous studies show that HC1 from coal Cl has little influence in the presence of excess at SH/RH conditions, because the sulfate salts are the thermodynamically stable species [Daniel, 1991; Seifert and Born, 1995]. Recent

studies do not show that coal Cl releases more alkali metals from the ash [Barnes et al.,

504

E. P. Doane and M. F. Abbott

1994; Muchmore et al., 1995]. So Cl in Illinois Basin coals would not be expected to play a role in SH/RH corrosion.

5.1. Role of S and Cl in SH/RH Corrosion The corrosive alkali iron sulfate SH/RH deposits do not ordinarily contain detectable Cl, because chloride salts are not thermodynamically stable under these conditions. Equilibrium calculations for simulated combustion conditions for several different coals show that no chloride salts would be expected in coal ash deposits on SH/RH tubes [Daniel, 1991; Seifert and Born, 1995]. The instability of chloride salts under con-

ditions similar to those in the SH/RH region was also directly demonstrated in studies of corrosive deposits formed in a high-velocity Mach 0.3 burner rig burning jet fuel

[Santoro et al., 1985]. Aqueous solutions of sodium chloride and sodium nitrate were injected individually into the combustor as fine spray droplets. The deposits on an inter nally cooled rotating collector at temperatures from 500°C to 1,000°C contained only sodium sulfate, despite the extremely short residence times and relatively low sulfur oxides concentrations involved. As Hancock [1987] puts it: “There is not the slightest doubt that sulfation [of chloride salts] does occur at high temperatures in atmospheres containing the oxides of sulfur . . . ” Less appears to be required to suppress HC1 corrosion of higher alloys used for SH/RH tubes than for carbon steel [Krause, 1991]. The molar ratios of /HCl needed to minimize corrosion rates are about 1.5 with 316 SS and about 0.7 with alloy 825. The temperature range in both cases is 371 to 593°C. Most coals have S/C1 weight ratios exceeding 2.2, or a molar /HCl ratio over 2.4. The question of how much chloride can be tolerated in deposits at SH/RH temperatures has also received some attention. Laboratory electrochemical studies [Cutler

Low Corrosivity of Coal Chlorine

505

et al., 1983] in an alkali sulfate melt at 600°C showed no effect on corrosion rate of adding one anion percent Cl to the melt. With a synthetic ash in a flue gas atmosphere, Rehn [1985] observed little effect of 3% added Cl at 675°C on corrosion. The 3% Cl increased corrosion at 790°C, where it caused the synthetic ash to fuse. These results appear to indicate that higher concentrations of Cl on the order of 1% by weight may not be corrosive in deposits under SH/RH conditions. This is in contrast to the previously cited work of Krause et al. on refuse incineration, which tends to put the safe upper limit of Cl content in deposits at furnace wall temperatures at around 0.1%.

5.2. Coal Cl and Release of Alkali Metals One claim is that the role of coal Cl in SH/RH corrosion is to release a portion of the alkalis during the combustion process by the reaction of HC1 with aluminosilicate clay minerals in coal particles [Gibb and Angus, 1983]. The vapor phase alkalis may then react with sulfur oxides in the flue gas to form potentially corrosive alkali sulfates, which are the thermodynamically preferred alkali species. Gibb and Angus carried out combustion studies of 44 UK coals in a bomb calorimeter. Both the raw coals and their ashes were finely ground and extracted with boiling water to obtain soluble Na and K. The water-soluble Na and K in the ash were considered to be indicative of the amount of each alkali metal that would be released upon combustion in a PC burner. Table 2 presents a summary of the Gibb and Angus data, and does not in fact show that Cl content has much to do with the fraction of alkali metals in the ash that is water-soluble. On average, combustion dramatically reduces the fraction of Na that is water-soluble, while the percent water-soluble K shows little change. Essentially that same statement can be made about the 5 low Cl coals and the 12 high Cl coals. The correlation in Fig. 7 in Gibb and Angus [1983] is simply a natural consequence of the strong correlation of water-soluble in the coal with Cl and the invariance of the portions remaining water-soluble in the ash. Direct analyses of alkali metals in flue gases show no correlation with coal Cl content. Barnes et al. [1994] selected 14 coals from several mining regions in the UK. These coals were burned in a laminar flat-flame burner at 1,400–1,800° K and the concentrations of Na and K in the gas phase were measured optically. Typically the fraction of K released was less than 5% of the fraction of Na released. The correlations between Na release and coal Na content, Na/ash ratio, or coal Cl content were not statistically significant at any of the conditions investigated. This was also true of K release and coal Cl content. In the US, actual measurements of Na and K in gases from a small gasifier at 850 and 950°C were made using two Illinois coals, one containing 0.06% Cl and the other containing 0.4%) Cl [Muchmore et al., 1995]. The Na and K contents of the two coals were similar. About 21% of the Na and 12% of the K were found in the gas phase.

506

E. P. Doane and M. F. Abbott

No significant difference in the release of K was observed between the two coals, and the release of Na was slightly lower with the high-Cl coal. Accordingly, both UK and US measurements of alkali contents in coal combustion and gasification gases confirm that coal Cl does not affect the amount of alkali metal released to the gas phase.

5.3. Superheater/Reheater Corrosion Indexes The three most widely-used published correlations for SH/RH corrosion were dis cussed previously [Abbott et al., 1994; Doane and Abbott, 1995]. The index used in the UK applies to the 300 series stainless steels and Esshete 1250. Attempts to develop a similar correlation for low alloy ferritic steels were unsuccessful [James et al., 1995]. The UK index includes gas and metal temperature terms applied to Cl contents above 0.06 wt.%. In contrast, both the other indexes include Na, K, Ca and magnesium (Mg) terms, and one has a S term, but they do not have a Cl term. In developing the index of Borio et al. [1969], no correlation of corrosion rate with Cl was observed over the range of 0.01–0.25wt.% Cl in the coals utilized. The index of Shigeta et al. [1986] has an additional term involving the content of the flue gas. Although this index was developed to predict corrosion rates for 347SS at 700°C, the rates for various temperatures and for other alloys can be predicted by applying appropriate correlation factors from graphs in

the original article [Shigeta et al., 1986]. To summarize, the UK index of SH/RH corrosion contains a Cl term but no Na or K terms, while the other two indexes contain Na and K terms but no Cl term. The strong correlation of Cl with Na in UK coals, coupled with the accepted role of alkali iron trisulfates in SH/RH corrosion, may mean that Cl content is simply an indicator of alkali metals content in UK coals. James et al. [1995] admit that, while the mechanistic role of alkali metals in high temperature corrosion is well understood, the determination of empirical relationships has been frustrated by the lack of routine analyses of alkali metals in UK coals. The temperature dependence of corrosion can often be expressed concisely in terms of an activation energy Table 3 gives values estimated from data from several

Low Corrosivity of Coal Chlorine

507

studies of corrosion related to firing coal or municipal refuse. The data for 300 series stainless steels are averaged. First, the Shigeta data show the metal temperature dependence expected for the classical alkali iron trisulfate corrosion bell curve—a rapid increase in corrosion rate with 55kcal/g-mole up to about 700°C, with equally rapidly declining corrosion rates above that temperature. Although the metal temperature squared term in the UK equation precludes a constant activation energy, the range of values is comparable to that of the Shigeta correlation. This suggests that both are dealing with a consistent phenomenon i.e., sulfate salt corrosion. The purpose of one study at temperatures of over 600°C was to evaluate corrosion in waste incineration, so a synthetic flue gas with = 1.8 molar was the highest ratio used [Mayer et al., 1983]. As discussed previously, ratios below 1.8 are not

sufficient to mitigate HC1 corrosion, and observed corrosion rates were indeed high. Up to 700°C, the for the corrosion observed in this study is only about 7.5kcal/g-mole, in sharp contrast to roughly 40 to 60kcal/g-mole for sulfate corrosion over a similar temperature range. Above 700°C, this HC1 corrosion does not display the bell-curve behavior of the sulfate corrosion. Instead, corrosion increases sharply with temperature, with an estimated

of 36 kcal/g-mole. It appears that the corrosion observed by Mayer et al.

[1983] involves oxidative HC1 corrosion, and the process described by the UK index does not. It also appears that the temperature dependence of corrosion rates is low whenever metal temperatures are 500°C or less, regardless of whether there are sufficient sulfur compounds to mitigate HC1 corrosion. For either simulated [Mayer et al., 1983] or actual [Krause, 1991] high-Cl waste incineration, this limited dependence on temperature seems to extend to at least 600°C. Although the highest of 12kcal/g-mole at these lower temperatures occurs under strongly reducing conditions [Latham et al., 1991], much of the UK p.c. coal rig data with = 3 kcal/g-mole was also taken under strongly reducing conditions [Davis et al., 1997]. Accordingly, the temperature dependence of corrosion rates at metal temperature of 500°C or less are unsatisfactory for differentiating among these corrosion mechanisms.

One of the features of the UK SH/RH index is the inclusion of a gas temperature term reflecting the observed increase in corrosion with higher heat flux, even at the same metal temperature. Interestingly, waste incineration data for 310 SS at a metal temperature of 593°C reveal a similar dependence on gas temperature [Krause, 1991]. The values of 8 and 14kcal/g-mole are fairly comparable, especially considering the different conditions. To summarize, formation of alkali iron trisulfates is generally accepted as the main cause of corrosion of SH/RH tubes in coal-fired boilers. The corrosion indexes of Borio and Shigeta emphasize the importance of alkali metals and contain no Cl term. Activation energies suggest that the UK corrosion index also represents alkali iron trisulfate corrosion, even though it contains no alkali metals terms. The theory that higher Cl contents release more Na and K to the gas phase is not supported by experimental results. Thus, it seems increasingly likely that the Cl term in the UK index serves as a substitute for the alkali metals, as originally thought in the UK.

6. ABB-CE SUPERHEATER CORROSION TEST EXPERIENCE A recent paper highlighted ABB research experience on the effects of fuel chloride on boiler corrosion [Plumley et al., 1994], The four examples that were presented were

508

E. P. Doane and M. F. Abbott

discussed in Doane and Abbott [1995]. Additional observations on two of the examples are presented here.

6.1. Perchloroethylene Addition The original paper [Borio et al., 1969] reported corrosion data on 321SS probes from 300h combustion tests on seven coals. For two of these tests perchloroethylene was added to increase the Cl content to 1.32% and 1.65% based on coal. The other five coals had Cl contents varying from 0.05% to 0.38%. The two coals with perchloroethylene were quite corrosive. Naturally, added perchloroethylene may not behave the same during combustion as an equivalent amount of coal Cl. The other five coals exhibited relatively low corrosion rates that showed no correlation with

coal Cl content. In fact, the coal with 0.38% Cl had the lowest corrosion rate of all the coals. If we consider the S/C1 ratio in these tests, the five coals with low corrosion rates

have S/C1 ratios of 4.2 or higher, which is more than enough to negate any effect of the Cl. The S/C1 ratio for 1.32% Cl coal is 1.4, which probably is insufficient. The S/C1 ratio for the 1.65% Cl coal is 2.0, which is borderline or perhaps sufficient. But another factor comes into play at such high levels of Cl. The postulated mechanism by which sufficient minimizes HC1 corrosivity is by forming alkali sulfates instead of alkali chlorides. A simple chemical equation representing these reactions is:

If this reaction approaches equilibrium in the ash deposits at furnace temperatures, then the concentration of (K, Na) Cl will increase as the partial pressure of HC1 increases, even though the mole ratio of remains constant. In fact, if the partial pressure of HC1 doubles, the mole ratio of must also be doubled in order to

keep the concentration of (K, Na) Cl constant. If this is the correct mechanism, then

substantially higher ratios of S/C1 in coal would be required at high values of coal Cl to keep alkali metal chloride concentrations sufficiently low to minimize their corrosive effect.

6.2. High-Cl Illinois Coal Eight high-temperature alloys were evaluated in 300 h combustion tests with two low-Cl bituminous coals, generally at 732°C (1,350°F) and 899 °C (1,650 °F). Corrosion rates on an Illinois No. 6 coal containing 0.45% Cl were determined similarly, generally at 593°C (1,100°F) and 871 °C (1,600°F). The S/C1 ratio for this coal was about 4. The results were presented as a histogram [P lumley et al., 1994], but it is difficult to interpret the effect of coal Cl because the temperatures are different. Table 4 contains all data from the original report [Plumley et al., 1979–80] at 732°C and 899°C comparing the two lowCl coals and the high-Cl coal. At 732 °C, corrosion with the 0.45% Cl coal is equal or lower for three of the alloys, and worse for two other alloys, compared to the low-Cl coals. At 899 °C, corrosion with the 0.45% Cl coal is equal or lower for two of the alloys, and worse for two other alloys, compared to the low-Cl coals. This does not justify concluding that the high-Cl Illinois coal is more corrosive.

Low Corrosivity of Coal Chlorine

509

7. CONCLUSIONS Current evidence shows that the primary cause of rapid furnace wall corrosion is poor combustion of the coal, not Cl content. Poor combustion produces local areas with high heat flux, high CO levels, excessive slagging, and an aggressive sulfidizing environment. Laboratory and combustion rig studies show high rates of corrosion under reducing waterwall conditions whether or not HC1 is present. This is confirmed by the long-term favorable UK experience burning high-Cl coals with low-NOx burners. Similarly, US experience with low-NO x burners also has not identified coal Cl as a factor in corrosion. Under oxidizing furnace wall conditions, corrosion rates are low to moderate, i.e., less than 20% of rates under reducing conditions in small-scale studies. Utility boiler

experience both in the UK and the US shows no effect of coal Cl on the low corrosion rates under oxidizing furnace wall conditions. The cause of accelerated SH/RH corrosion is molten alkali-iron trisulfate attack at the base of fouling deposits. Chlorides are not ordinarily found in SH/RH deposits from coal combustion, and indeed they should not be present according to thermodynamic equilibria. Sufficient (if oxidizing) and H2S (if reducing) in combustion gases minimizes high-temperature corrosion by the HO formed from coal Cl. If the (S/C1) weight ratio in coal is over about 2.2 at Cl contents common in coal, sulfates are formed in preference to chlorides in deposits at elevated temperatures. Then the situation becomes one of corrosion by sulfates or sulfides, with coal Cl playing little or no role. This is not true for most municipal refuse combustion, which does not contain enough S to form sulfates in preference to chlorides. Accordingly, the role of coal Cl has long been questionable. Experimental evidence does not support the theory that coal Cl releases Na and K to the flue gas phase. So it appears that coal Cl is not a significant factor in SH/RH corrosion.

Even if currently mined US and UK coals are basically identical, the evidence makes it unlikely that Illinois coals containing 0.4–0.5 wt.% Cl are corrosive in boilers. But Illinois coals contain on average about one-half the potentially corrosive Na and K

and about one-half the ash of UK commercial coals. At least two independent surveys of utilities burning high proportions of high-Cl Illinois Basin coals have not found corrosion problems ascribed to coal Cl. And despite utility combustion of well over a billion

E. P. Doane and M. F. Abbott

510

tons of Illinois Basin coals over the past two decades, there is no convincing evidence of Cl corrosion. Consequently, utilities should not hesitate to use the higher-Cl Illinois Basin coals.

8. REFERENCES Abbott, M. F., Campbell, J. A. L., and Doane, E. P. (1994). “Utility Experience Burning High-Chlorine Illinois Coals.” Power-Gen '94, Book I I I , Vol. 1, pp.1–18. Bakker, W. T. (1997).“The Effect of Chlorine on Corrosion of Stainless Steels in Gasifiers.” NACE Corrosion 97 Conference, Paper No. 134, 21 pp.

Bakker, W. T. and Perkins, R. A. (1991). “The Effect of Coalbound Chlorine on Coal Gasification Plants.” in Chlorine In Coal, J. Stringer and D. D. Banerjee (eds.), Elsevier Science Publishers B.V., Amsterdam, pp. 63–80.

Bakker, W. T., Kip, J. B. M., and Schmitz, H. P. (1993). “Materials Testing in Syngas Coolers of Coal Water Slurry Fed Gasifier.” Materials at High Temperatures, 11(1/4), pp. 81–89. Barnes, D. I., Wardle, D. G., and Kean, D. (1994). “Sodium and Potassium Release from UK Coals During the Early Stages of Pulverised Coal Combustion.” in The Impact of Ash Deposition on Coal Fired Power

Plains, J. Williamson and F. Wigley (eds.), Taylor and Francis, Washington, DC, pp. 325–336. Borio, R. W, Hensel, R. P, Ulmer, R. C, Grabowski, H. A., Wilson, E. B., and Leonard, J. W. (1969). “The Control of High-Temperature Fire-Side Corrosion in Utility Coal-Fired Boilers.” Office of Coal Research R&D Report No. 41.

Chamberlain, John (1983). Private Communication with M. F. Abbott. C hou, M.-l. M., Pan, W. P., Huggins, F. E., Lytle, J. M., Liu, L., Huffman, G. P., and Ho, K. K. (1994). “Chlorine in Coal and Boiler Corrosion.” Eleventh Annual International Pittsburgh Coal Conference Proceedings, pp. 359–364. Chou, M.-l. M., Lytle, J., Pan, W. P., Liu, L., Huggins, F. E., Huffman, G. P., and Ho, K. K. (1994). “Chlorine and Ash Composition in Two Illinois Coals and Two British Coals.” Fourth International Conference on the Effects of Coal Quality on Power Plants, Charleston, S.C. Cutler, A. J. B. and Raask, E. (1981). “External Corrosion in Coal-Fired Boilers: Assessment From Laboratory Data.” Corrosion Science, Vol. 21, No. 12, pp. 789–800. Cutler, A. J. B., Grant, C. J., Laxton, J. W, Price, D. D., and Stevens, C. G. (1983). “Laboratory Measurements of the Corrosion of Superheater Materials in Atmospheres Simulating the Combustion of High Chlorine Coals.” in Corrosion Resistant Materials for Coal Conversion Systems, D. B. Meadowcroft and

M. I. Manning (eds.), Elsevier Science Publishers, New York, pp. 159–177. Daniel, P. L. (1991). “The Effect of Coal Chlorides on Furnace Wall Corrosion.” in Chlorine in Coal, J. Stringer and D. D. Banerjee (eds.), Elsevier Science Publishers B.V., Amsterdam, pp. 207–217. Davis, D. J., James, P. J., and Pinder, L. W. (1997). “Combustion Rig Studies of Fireside Corrosion in Coal Fired Boilers.” Fifth International Conference on the Effects of Coal Quality on Power Plants, Kansas City, KS. Demir, I., Harvey, R. D., Ruch, R. R., Damberger, H. H., Chaven, C., Steele, J. D., and Frankie, W. T. (1994). “Characterization of Available (Marketed) Coal from Illinois Mines.” Final Report to the Coal Research Board, IDENR, pp. 22 & 23. Doane, E. P. and Abbott, M. F. (1995). “High-Chlorine Illinois Basin Coals: Challenge or Opportunity?” Engineering Foundation Conference, The Economic and Environmental Aspects of Coal Utilization VI,

Santa Barbara, CA. Gabrielson, J. E. and Kramer, E. D. (1996). “Measurement of Reducing Gases Formed in Low NO X Combustion.” 1996 Joint Power Generation Conference, EC-Vol. 4/FACT-Vol. 21, Vol. 1, pp. 19–23. Gibb, W. H. (1983). “The Nature of Chlorine in Coal and its Behavior During Combustion,” in Corrosion Resistant Materials for Coal Conversion Systems, D. B. Meadowcroft and M. I. Manning (eds.), Applied Science, London, pp. 25–45. Gibb, W. H. and Angus, J. G. (198.3). “The Release of Potassium from Coal During Bomb Combustion.” Journal of The Institute of Fuel, Vol. 56, pp. 149–157. Grabke, H . J., Reese, E., and Spiegel, M. (1995). “The Effects of Chlorides, Hydrogen Chloride, and Sulfur Dioxide In the Oxidation of Steels Below Deposits.” Corrosion Science, Vol. 37, No. 7, pp. 1023–1043. Hancock, P. (1987). “Vanadic and Chloride Attack of Superalloys.” Materials Science and Technology, Vol. 3,

pp. 536–544.

Low Corrosivity of Coal Chlorine

511

Harrison, C. D. and Parkinson, J. W. (1983). “Physical Coal Cleaning to Reduce Illinois No. 6 Seam Coal’s Slagging Potential.” Mining Engineering, Vol. .38, No. 8, pp. 830–833. James, P. J., Robinson, M. T., Gibb, W. H., and Pinder, L. W. (1997). “Effect of Fuel Chlorine on the Fireside Corrosion of Ferritic Furnace Wall and Superheater/Reheater Tubing—A U.K. Perspective.” Volume 2 in Possible Effects of the Chlorine Content of Coal on Fireside Corrosion in Pulverized Coal-Fired Boilers, E P R I TR-108107.

Krause. H. H. (1991). “Effects of Flue-Gas Temperature and Composition on Corrosion from Refuse Firing.” NACE Corrosion 91 Conference, Paper No. 242, 6 pp.

Krause. H. H., Vaughan, D. A., Cover, P. W., Boyd, W. K., and Oxsley, R. A. (1979). “Corrosion and Deposits from Combustion of Solid Waste. Part V I : Processed Refuse as a Supplementary Fuel in a Stoker-Fired Boiler.” Trans. ASME, J. Eng. for Power, Vol. 101, pp. 592–597. Kung, S. C., Daniel, P. L., and Seeley, R. R. (1996). “Effect of Chlorine on Furnace Wall Corrosion in U t i l i t y Boilers.” NACE Corrosion 96 Conference, Paper No. 165, 7 pp. Latham, E., Meadowcroft, D. B., and Pinder, L. (1991). “The Effects of Coal Chlorine on Fireside Corrosion.” in Chlorine in Coal, J. Stringer and D. D. Banerjee (eds.), Elsevier Science Publishers B.V., Amsterdam, pp. 225–246. Lees, D. J. and Whitehead, M. E. (1983). “The influence of Gas and Deposit Chemistry on Fireside Corrosion of Furnace Wall Tubes in Coal-Fired Boilers.” in Fouling of Heat Exchanger Surfaces, R. W. Bryers (ed.). Engineering Foundation, New York, NY, pp. 69–104. Mayer, P., Manolescu, A. V, and Thorpe, S. J. (1983). “Influence of Hydrogen Chloride on Corrosion and Corrosion-Enhanced Cracking Susceptibility of Boiler Construction Steels in Synthetic Flue Gas at Elevated Temperatures.” in Corrosion Resistant Materials for Coal Conversion Systems, D. B. Meadowcroft and M. I. Manning (eds.), Applied Science Publishers, London, pp. 87–103. McNallan, M. J. (1991). “Corrosion Behavior of Metallic Materials in Chlorine-Containing Gaseous Environments.” in Chlorine in Coal, J. Stringer and D. D. Banerjee (eds.), Elsevier Science Publishers B.V., Amsterdam, pp. 175–192. Meadowcroft, D. B. (1987). “High Temperature Corrosion of Alloys and Coatings in Oil- and Coal-Fired Boilers.” Materials Science and Engineering, Vol. 88, pp. 313–320.

Muchmore, C. B., Hippo, E. J., Chen, H. L., Joslin, J. L. Jr., Wang, L., Hughes, A., Sivanandan, S., Daman, E., and Banerjee, D. D. (1995). “Distribution of Sodium, Potassium and Chlorine Between Solid and Vapor Phases Under Coal Gasification Conditions.” in Coal Science, J. A. Pajares and J. M. D. Tascon (eds.), Elsevier Science Publishers B.V., Amsterdam, pp. 819–822. Pearce, W. C. and Hill, J, W. F. (1986). “The Mode of Occurrence and Combustion Characteristics of Chlorine in British Coal.” Progress in Energy and Combustion Science, Vol. 12, pp. 117–162. Plumley, A. L., Accortt, J. I., and Roczniak, W. R. (1979–80). “Evaluation of Boiler Tube Materials for Advanced Power Cycles,” Vol. I, II, and III and Supplement I, NTIS PB82-116823, -31, -49 and -56. Plumley, A. L., Borio, R. W, Levasseur, A. A., and Roczniak, W. R. (1994). “ABB Research Highlights on the Effects of Fuel Chloride on Boiler Corrosion.” Eleventh Annual International Pittsburgh Coal Conference Proceedings, pp. 385–391. Raask, E. (1983). “Reactions of Coal Impurities During Combustion and Deposition of Ash Constituents on Cooled Surfaces,” in The Mechanism of Corrosion by Fuel Impurities, H. R. Johnson and D. J. Littler (eds.), Butterworths, London, pp. 145–154. Rehn, I. M. (1985). “Improved Materials for Coal-Fired Boiler Superheater and Reheater Tubes Subjected to Fireside Corrosion.” in High Temperature Corrosion in Energy Systems, M. F. Rothman (ed.), The Metallurgical Society of AIME, Warrendale, PA, pp. 377–388. Santoro, G. J., Gokoglu, S. A., Kohl, F. J., Stearns, C. A., and Rosner, D. A. (1985). “Deposition of Na2SO4 From Salt-Seeded Combustion Gases of a High Velocity Burner Rig.” in High Temperature Corrosion in Energy Systems, M. F. Rothman (ed.). The Metallurgical Society of AIME, Warrendale, PA. pp. 417–434. Seifert, P. and Born, M. (1995). “Effect of Fuel Composition on Chloride Corrosion in Furnaces.” VGB Tech. Ver. Grosskraftsverksbetr., [Tech. Bar.] VGB-TP 1995, Paper 9, 21 pp. Shigeta, J.-I., Hamao, Y., Aoki, H., and Kajigaya, I. (1986). “Development of a Coal Ash Corrosivity Index for High Temperature Corrosion.” ASME Paper 86-IJPGC-FACT-3. Stringer, J. and Banerjee, D. D. (eds.) (1991) Chlorine in Coal, Elsevier Science Publishers B.V., Amsterdam. Urich, J. A. and Kramer, E. (1996). “Designing Solutions for Low NO, Related Waterwall Corrosion.” 1996 Joint Power Generation Conference. EC-Vol. 4/FACT-Vol. 21, Vol. I , pp. 25–29. Wright, I . G. and Krause, H. H. (1997). “Survey of the Effects of Coal Chlorine Levels on Fireside Corrosion in Pulverized Coal-Fired Boilers.” Volume 1 in Possible Effects of the Chlorine Content of Coal on Fireside Corrosion in Pulverized Coal-Fired Boilers, EPRI TR-108107.

512

E. P. Doane and M. F. Abbott

Wright, I. G., Dooley, R. B., and Mehta, A. K. (1997). “U.S. Perspective on the Effects of Fuel Chlorine on the Fireside Corrosion of Furnace Wall and Superheater/Reheater Tubing.” Volume 3 in Possible Effects of the Chlorine Content of Coal on Fireside Corrosion in Pulverized Coal-Fired Boilers, EPRI TR-108107. Wright, I. G., Ho, K. K., and Mehta, A. K. (1994). “Survey of the Effects of Coal Chlorine Levels on Fireside Corrosion in Coal-Fired Boilers.” Fourth International Conference on the Effects of Coal Quality on Power Plants, Charleston, S. C.

LABORATORY STUDIES ON THE INFLUENCE OF GASEOUS HCl ON SUPERHEATER CORROSION Keijo Salmenoja1, Mikko Hupa, and Rainer Backman2 1

Kvaerner Pulping Oy P.O. Box 109, FIN-33101 TAMPERE Finland 2 Åbo Akademi University Department of Chemical Engineering Lemminkäisenkatu 14-18B, FIN-20520 TURKU Finland

1. INTRODUCTION High temperature corrosion in the superheater (SH) area is one of the most important factors limiting the increase of thermal efficiency of modern power and recovery boilers. Due to high temperature corrosion, steam outlet temperatures have been limited

generally below 550 °C. In incinerators and in boilers burning high alkaline and high chlorine fuels, used steam outlet temperatures are in the range 450–500°C [Neumann and Kautz, 1997]. Molten alkali salt corrosion limits the steam temperatures of recovery boilers below 500°C. Chlorine (Cl) can cause SH corrosion either by gaseous hydrogen chloride (HC1) or by alkali chlorides in the deposits [Bryers, 1996]. Corrosion from gaseous HC1, however, is limited to high flue gas temperatures, high concentrations and reducing atmospheres. This type of corrosion is typically found in the lower furnace [Husemann, 1992]. Alkali chlorides are known to lower the melting temperatures of deposits significantly [Krause, 1987; Hupa et al., 1990] and molten phases are therefore generally blamed for the rapid corrosion of the superheaters [Kofstad, 1988]. A liquid deposit in direct contact with the tube metal surface dissolves effectively the alloy resulting in rapid corrosion. However, it was recently reported that rapid corrosion of Fe-Cr alloys by chlorine is also possible without any molten phases present on the tube metal surface, probably by a reaction sequence involving volatile chlorine species [Salmenoja et al., 1996]. Expanded use of different biomass based fuels (biofuels), such as straw and forest Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

513

514

K. Salmenoja et al.

residues have increased corrosion problems in many industrial boilers. This is mainly associated with the high potassium (K) and Cl content in the biofuels. Rapid corrosion was encountered in the SH area of a bubbling fluidized bed (BFB) boiler burning a mixture of bark and chlorine containing sludge from biological effluent treatment plant (biosludge). Maximum measured corrosion rate was around 0.8 mm/month [Salmenoja et al., 1996]. The original tube material was CrMoV 12 1 with 12wt.-% chromium (Cr) and 1.0wt.-% molybdenum (Mo). The steam outlet temperature was 530 °C corresponding to maximum SH material temperatures in the range 560–580°C. Alloy was replaced with an austenitic stainless steel NiCrCeNb 32 27 (AC 66), which contains 27wt.-% Cr and 32.0wt.-% nickel (Ni). Since then the boiler has been running without corrosion problems [Salmenoja et al., 1995]. In corroded tubes high chlorine concentrations were found in the interface between the alloy and the oxide scale. A new iron oxide layer had been formed on the original tube surface, between the oxide scale and the deposit layer. The formation of the outer iron oxide layer can be either due to solid phase diffusion of iron ions or diffusion of gaseous iron chlorides through the oxide scale. Diffusion of iron ions through the oxide scale is a well-known phenomenon in high temperature oxidation [Birks and Meier, 1983]. However, in this case temperatures may have been too low to achieve the observed high corrosion rates by the diffusion of Fe ions. The role of volatile iron chlorides has been discussed in several papers [Fruehan, 1972, Jacobson, 1986; Stott et al., 1988; Strafford et al., 1989a; Strafford et al., 1989b, Bramhoff et al., 1990; Karlsson et al., 1990; Chu et al., 1993; Grabke et al., 1995],

although no generally accepted mechanism has been presented. Speculated reaction scheme contains; 1) volatile iron chloride formation in the alloy, 2) diffusion of the

formed chlorides through the oxide scale and 3) oxidation of iron chlorides near the scalegas interface. Oxidation of the iron chloride releases chlorine back to the system and may then be again available for further reactions with Fe. This is the so-called chlorine circulation theory and was first introduced by Vaughan et al. (1977). The objective of the present work is to study the influence of gaseous HC1 on the oxidation and corrosion of powdered metal and binary Fe-Cr alloy samples in the laboratory. Simultaneous TGA-DTA analyzer was used to find evidence of the formation of volatile iron chlorides during chlorine induced corrosion.

2. EXPERIMENTAL Simultaneous DTA-TGA analyzer SDT 2960 (TA Instruments) was used in the tests. The experimental set up is shown in Fig. 1. In the analyzer the TG mass change is measured by a taut-band meter movement located at the rear end of each ceramic arm. The balance arm is maintained in the horizontal reference position by regulation of the current through the transducer coil. An infrared LED light source and a pair of photodiodes detect movements of the arm. The DTA measurement is made by a pair of Pt/PtRh thermocouples which go through the ceramic arms to a platinum sensor located in the bottom of the sample and the reference holder. The furnace with ceramic lining provides uniform heating up to 1,500°C. Purge gas is introduced through the balance

chamber. A special pipe line for reactive gases introduces the gases directly into the furnace 30mm before the sample holders. The maximum sample weight is 200mg with a balance sensitivity of 0.1 µg. Heating rates from 0.1 to 100°C/min can be used up to 1,000°C, while 0.1 to 25°C/min can be

Laboratory Studies on the Influence of Gaseous HCl on Superheater Corrosion

515

used up to 1,500°C. The minimum log-interval is one point per 0.5 seconds. The TG and

DT signals are recorded with a computer and plotted with a chart recorder. Approximately a 15mg sample was placed in a 100µ l platinum (Pt) cup and was allowed to react either in oxidizing or reducing conditions in the presence of 0.5-1.0% HC1. Reducing conditions were attained with 1–2% carbon monoxide (CO). Purge gas in oxidation tests was air and in reducing conditions nitrogen (99.5%). Diluted CO (4.78% in ) and HC1 (4.88% in ) were used as reactive gases. Total gas flow through the ther-

mobalance was l00ml/min, unless otherwise mentioned. Furnace temperature was increased from room temperature up to 900°C at a rate of 10°C/min. After the tests,

reacted samples were collected and analyzed with energy dispersive X-ray analyzer (EDXA).

In the interrupted tests the temperature for isothermal conditions (650–670 °C) was

chosen according to the oxidation tests. At this temperature the rate of the oxidation of the alloy was high enough. Temperature was in the first step raised from room temperature up to the isothermal temperature at a rate of 25°C/min. A higher temperature increment was used to avoid excessive pre-oxidation of the powders in the heating state. After

the system had reached a steady state, HO was introduced into the thermobalance. The length of the HO exposure was decided according to the weight gain of the sample. When the weight increase was 0.5%, 1.5%, and 4.5%, the HC1 flow was disconnected. After a certain period, the HC1 flow was turned on again. In some tests the sample was exposed several times to HC1. Exposed samples were studied with scanning electron microscope (SEM) and analyzed with EDXA. Several powdered binary Fe-Cr alloys with various Cr contents, including T 22

(with 2.25wt.-% Cr),

(12% Cr) and austenitic stainless steel AC 66 (28% Cr) were

used in the tests. Due to the wide particle size range of the powder, atomized AISI 410 powder (12.5% Cr) with a more even particle size distribution was used to study the behavior of the binary Fe-12Cr alloys. Carbon steel, 100% Fe and 100% Cr powders were

516

K. Salmenoja et al.

used as references. Iron powders with two different particle sizes were used to compare the effect of particle size. Table 1 gives a list of the alloys used with their average compositions and particle sizes.

3. RESULTS 3.1. HC1 Exposure Tests under Constant Conditions Basic oxidation tests were first made with Fe. Iron powders with different particle sizes behaved similarly, except small differences in the reactivity. Fe powder with

a smaller average particle size was more reactive, as was expected. Figure 2 shows a typical oxidation curve of Fe powder in air. Weight gain of the Fe sample starts at around 250°C. This is connected to a low and broad peak in the DT signal. A somewhat larger peak is located at around 450°C with a simultaneous change in the slope of the weight gain curve. Rapid oxidation of Fe powder starts at around 575 °C. This can also be seen in the DT signal as a strong exothermic peak. Fe samples were usually almost

Laboratory Studies on the Influence of Gaseous HCl on Superheater Corrosion

517

totally oxidized after the tests. Oxidation of the sample to means a weight gain close to 43%. Chromium is known to increase the oxidation resistance of the alloys. This was tested both in air and in air in the presence of 0.5% HCI. Figure 3 shows the behavior of all the tested alloys in air, including Fe and carbon steel. Only Fe powder and carbon steel showed significant weight gain below 600 °C. Alloys and AC 66 were only slightly oxidized during the tests. Figure 4 presents corresponding results in the presence of 0.5% HCl in air. Oxidation of the samples starts at around the same temperature as in air. All the other materials, except and AC 66, show significant weight gain below 600 °C. Rapid oxidation of high chromium alloys starts only above 600 °C.

518

K. Salmenoja et al.

In reducing conditions in the presence of HC1, only Fe samples showed weight loss during the tests. Chromium containing alloys and AISI 410 showed a slight weight gain during the whole temperature range. No weight loss was observed. According to these tests, it seems that even with 2% CO the oxygen partial pressure is still high enough to oxidize chromium in the alloys. Chromium samples, however, showed a relatively high weight gain during the tests. Part of the weight gain comes from chromium chloride

formation, since chromium chloride was detected with EDXA from the exposed

Cr samples.

3.2. Tests with Interrupted HCl Exposure Figure 5 shows the behavior of Fe powder in the presence of 1.0% HC1 and 1.0% CO. In this test temperature was first increased 10°C/min up to 700°C. Then the temperature was kept constant at 700°C. Below 550°C the weight of the sample was increasing. Above 550°C the weight started to decrease. When the sample was kept at isothermal

conditions, weight decreased at a constant rate as long as the HC1 flow was on. When

the HCl flow was turned off, the weight loss ceased immediately. When the HC1 flow was turned on again, weight of the sample starts to decrease without any delay and at the same rate as before.

Figure 6 shows the weight gain curve of a 17.6489mg powder sample in an interrupted test in the presence of 0.5% HC1 in oxidative conditions. When the HC1 flow

was turned on at 40min, a slow weight gain of the sample was seen. No peaks was detected in the DT signal. At 70min the HC1 flow was turned off and only air was allowed to flow through the thermobalance. The rapid oxidation continued and only a slight

change in the slope of the weight gain curve could be seen. At 90min, when HC1 was turned on again, the rate of the weight gain returned close to the original level. At 115 min, when HC1 was turned off again, a slight change is seen in the slope, but weight still increases to grow rather quickly.

Laboratory Studies on the Influence of Gaseous HCl on Superheater Corrosion

519

Alloy AISI 410 behaved differently in interrupted tests, as shown in Fig. 7. When HC1 exposure was started at 40min, it took some 20 minutes before the rapid weight gain started. This was accompanied with a strong exothermic peak in the DT signal. When the HC1 flow was turned off, a slight change in the rate of the weight gain could be seen and the DT signal went down very rapidly. However, the weight gain rate and the DT signal did not readily return to their original levels. After about a 15 minutes delay the DT signal and the rate returned to the original level. This delay was not seen in reducing conditions.

520

K. Salmenoja et al.

When the HC1 flow was turned on again at 120min, no changes in the weight nor in the DT signal occurred. At 130min the HC1 concentration in the thermobalance was increased from 0.5% to 1.0%. This initiated a rapid increase in weight and caused a strong exothermic peak in the DT signal. The HC1 flow was turned off at 135 min, but the weight still continued to grow at a relatively high rate.

4. DISCUSSION According to the present study, HC1 enhances the oxidation of Fe, Cr, and binary Fe-Cr alloys. However, the rate of oxidation in high Cr alloyed Fe-Cr steels is very low below 600 °C. Therefore, significant corrosion of high chromium alloys in the presence

of HC1 in oxidative conditions occurs only when the metal temperature is above 600 °C. Increasing chromium content in the alloys appears to be beneficial in resisting the oxidation in the presence of HC1. Powdered Fe samples generated volatile iron chlorides in the presence of HC1 in reducing conditions. This was seen as a loss in the weight of the Fe samples and as a decrease in the DT signal during the test. This is consistent with earlier studies of Hupa et al. (1990). Iron chlorides were also detected with EDXA in the exposed samples. Small

clusters of iron chloride grains could be seen in SEM micrographs. Strong neck formation between the particles in the exposed samples implies that melting had occurred during the tests in reducing conditions. Maximum temperature in the tests was 700 °C and the melting temperature of is 677 °C. The melting temperature of is 304 °C. The temperature is defined as the temperature where the vapor pressure of the component is equal to atm [Fontana and Staehle, 1976]. At this temperature, the volatilization of the component will become significant. According to Fontana and Staehle (1976), the temperature for is 536 °C. Corresponding temperature for is 167 °C. In this study, the temperature onset where the weight of the Fe samples started to decrease in reducing conditions varied between 540 and 550 °C, depending on the sample and HC1 concentration. This is very close to the temperature for Vapor pressures of different components in reducing conditions in the presence of

HC1 were checked with equilibrium calculations. The computer program ChemSage [Eriksson and Hack, 1990] was used in the calculations. According to equilibrium calculations, oxygen partial pressure in reducing conditions (1.0% CO and 0.5% HC1) is between and atm in the temperature range from 400 to 700 °C. Corresponding partial pressure is between and atm. At these conditions, is the stable phase. According to the equilibrium calculations, for in the test conditions is 530 °C. Partial pressure of in the whole temperature range is between and atm. The formation of appears to be very limited. The temperature f o r ' is 741 °C and 611 °C for [Fontana and Staehle, 1976]. Melting point of and is 814 and 1152°C, respectively. Maximum temperature during the runs was 850°C and the samples were sintered after the runs. Sintering have most probably been caused by molten since some melting was observed in the SEM micrographs. According to the equilibrium calculations, is the stable phase at the test conditions. When is formed some volatility should have been seen in the tests. However, no pronounced changes in the weight nor in the DT signal was observed during the runs with chromium. In interrupted tests the length of the incubation period varied depending on the

Laboratory Studies on the Influence of Gaseous HCl on Superheater Corrosion

521

HCl concentration and the sample. With lower HCl concentrations the incubation time was longer. Differences in the incubation time between x20 and AISI 410 samples may partly originate from different fabrication processes of the powders. The x20 powder was made by grinding, but the AISI 410 powder was made by atomizing. In atomized powders the oxide layer on the particles can be relatively thick. According to the tests with alloy x20, a short HCI exposure is enough to start enhanced oxidation. When enhanced oxidation once starts, it continues until the sample is totally oxidized. With AISI 410 samples this memory effect was considerably shorter. After the HCl exposure, oxidation continued only around 10 minutes before the enhanced oxidation stopped. With higher HCl concentrations, the rate of oxidation is higher and oxidation ceases more rapidly after the HCl exposure. The particle size distribution of the x20 powder was very broad from small particles up to large 700 µm particles with an average particle size of around . . The particle size distribution of AISI 410 powder was more even with an average particle size of This means that a 15 mg sample of x20 contains 30–40 particles and a 15 mg AISI 410 sample around 100,000 particles, on the average. Assuming that all the particles are spherical and the oxide layer growing on the particles is shrinking core model [Levenspiel, 1993] can be used to calculate the growth of the oxide layer. When the weight increase of the sample is 10%, the oxide scale on AISI 410 particles is around thick. Corresponding oxide scale thickness on x20 particles is around If we assume that the diffusion in both oxide layers is equal, this means that diffusion time through the oxide layer on x20 particles is 15 times longer than with AISI 410 particles. This results in a longer memory effect in the x20 samples. If the memory effect is 10 minutes long with AISI 410 samples, then according to the above calculations, it would be around 150 minutes with the x20 samples. Since the oxidation time in the present tests was not that long, therefore alloys x20 and AISI 410 behaved differently.

5. CONCLUSIONS Thermogravimetric tests made with powdered metals and binary Fe-Cr alloys showed that increasing Cr content increases the resistance to oxidation in oxidative conditions produced by oxygen or by oxygen and gaseous HCl. In reducing conditions both and was formed with pure metal samples. A short HCl exposure to x20 and AISI 410 alloys render possible a rapid subsequent enhanced oxidation of the samples without any HCl present. This oxidation rate was close to the original rate with HCl. After a certain period, enhanced oxidation ceased and the rate returned to the level obtained in air. This memory effect was not observed in reducing conditions nor with pure Fe and Cr samples. Observed memory effect may result from two different processes. Firstly, once the chlorine has penetrated the oxide scale on a particle, it may form a mixture of iron and chloride which is molten at low temperatures. This molten iron chloride layer moves forward and eats the iron from the alloy. It also continuously disturbs the normal transport of oxygen and prevents the formation of a dense oxide layer. A porous oxide scale potentiates the rapid subsequent oxidation of the alloys. After all the iron in the alloy is oxidized, the enhanced oxidation stops. A second possibility is the formation of volatile iron chlorides inside the particle. Formed iron chlorides diffuse outwards and becomes oxidized when they reach areas

K. Salmenoja et al.

522

where the oxygen partial pressure is high enough. The reason why the oxidation stops after a certain period may two folded. First, diffusion of iron chloride from the particle decreases the amount of chlorine available, and finally all the chlorine will be consumed. Secondly, oxidation of metal chlorides releases chlorine back to the system. Part of this chlorine may escape from the particles and therefore result in a gradual decrease in the oxidation rate.

6. ACKNOWLEDGEMENTS Imatran Voima Foundation is greatly appreciated for the financial support to this work.

7. REFERENCES Birks, N., Meier, G. (1983). Introduction to High Temperature Oxidation of Metals. London: Edward Arnold Publishers Ltd. Bramhoff, D., Grabke, H. J., Reese, E., Schmidt, H. P. (1990). “ E i n f l u ß von HCl and auf die Hochtemperaturkorrosion des Crl Mo-Stahls in Atmosphären mit hohen Sauerstoffdrücken”, Werkstoffe and Korrosion, 41, 303–307. Bryers, R. W. (1996). “Fireside Slagging, Fouling, and High-Temperature Corrosion of Heat-Transfer Surface due to Impurities in Steam-Raising Fuels”, Prog. Energy Combust. Sci., 22, 29–120. Chu, H., Datta, K., Gray, J. S., Strafford, K. N. (1993). “Corrosion Mechanism of Fe(Ni)25CrAlX Type Alloys in a Bioxidant Environment at Elevated Temperatures”, Corrosion Science, 35 (5–6), 1091–1098. Eriksson, G., Hack, K. (1990). “ChemSage—A Computer Program for the Calculation of Complex Chemical Equilibria”, Metallurgical Transactions, B 21, 1013–1023. Fontana, M. G.,Staehle, R. W. (1976). Advances in Corrosion Science and Technology, Vol. 5. New York: Plenum Press. Fruehan, R. J. (1972). “The Rate of Chlorination of Metals and Oxides: Part I. Fe, Ni and Sn in Chlorine”, Metallurgical Transactions, 3 (10), 2585–2592. Grabke, H. J., Reese, E., Spiegel, M. (1995). “The Effects of Chlorides, Hydrogen Chloride, and Sulfur Dioxide in the Oxidation of Steels below Deposits”, Corrosion Science, 37 (7), 1023–1043. Hupa, M., Backman, R, Skrifvars, B-J., Hyöty, P. (1990). “The Influence of Chlorides on the Fireside Behavior in the Recovery Boiler”, Tappi Journal, 73 (6), 153–158. Husemann, R. U. (1992). “Korrosionsercheinungen und deren Reduzierung an Verdampfern und Überhitzerbauteilen in kommunalen Müllverbrennungsanlage”, VGB Kraftwerkstechnik, 72 (10), 918–927. Jacobson, N. S. (1986). “Reaction of Iron with Hydrogen Chloride-Oxygen Mixtures at 550°C”, Oxidation of Metals, 26 (3/4), 157–169. Karlsson, A., Møller, R, Johansen, V. (1990). “Iron and Steel Corrosion in a System of and Alkali Chloride. The Formation of Low Melting Point Salt Mixtures”, Corrosion Science, 30 (2–3), 153–158. Kofstad, P. (1988). High Temperature Corrosion. New York: Elsevier Applied Science. Levenspiel, O. (1993). The Chemical Reactor Omnihook. Corvallis, Oregon: OSU Book Stores Inc. Neumann, J., Kautz, H. R. (1997). “Auswertung der internationalen Literatur zur Hochtemperaturkorrosion in Kohle- und Müllkraftwerken”, VGB Kraftwerkstechnik, 77 (4), 329–334. Salmenoja, K., Mäkelä, K., Backman, R. (1995). “Corrosion in Bubbling Fluidized Bed Boilers Burning Chlorine Containing Fuels”. In the Proceedings of the 8th International Symposium on Corrosion in the Pulp

& Paper Industry, Stockholm, Sweden, May 16–19, 198–206. Salmenoja, K., Mäkelä, K., Hupa, M., Backman, R. (1996). “Superheater Corrosion in Environments Containing Potassium and Chlorine”, Journal of the Institute of Energy, 69, 155–162. Stott, F. H., Prescott, R., Elliott, P., Al'Atia, M. H. J. H. (1988). “Assessment of the Degradation of Metals and Alloys in Air-2% Chlorine at high Temperature”, High Temperature Technology, 6 (3), 115–129. Strafford, K. N., Datta, P. K., Forster, G. (1989a). “High-Temperature Chloridation of Binary Fe-Cr Alloys at 1000°C”, Materials Science and Engineering, A120, 61–68.

Laboratory Studies on the Influence of Gaseous HCl on Superheater Corrosion

523

Strafford, K. N., Datta, P. K., Forster, G. (1989b). “The High Temperature Chloridation of Iron, Nickel, and

some Iron-Nickel Model Alloys”, Corrosion Science, 29 (6), 703–716. Vaughan, D. A., Krause, H. H., Boyd, W. D. (1977). “Chloride Corrosion and Its Inhibition in Refuse Firing”. In the Proceedings of the international Conference on Ash Deposits and Corrosion from Impurities in Combustion gases, Henniker, New Hampshire, June 26–July 1, 473–493.

This page intentionally left blank

THE ROLE OF ALKALI SULFATES AND CHLORIDES IN POST CYCLONE DEPOSITS FROM CIRCULATING FLUIDIZED BED BOILERS FIRING BIOMASS AND COAL Bengt-Johan Skrifvars, Tor Laurén, Rainer Backman, and Mikko Hupa Åbo Akademi University, Finland Peter Binderup-Hansen, I/S Midtkraft, Denmark

1. INTRODUCTION High amounts chlorine and sulphur in a fuel is generally connected with ash related

operational problems in the boiler in which the fuel is fired. The problems occur as fireside deposits in different locations of the fluegas channel or as corrosion problems. Sulphur and chlorine together with alkali and earth alkali metals are known to strongly affect the thermal behaviour of the ash. First melting temperatures as low as 515 °C may be found if unsuitable amounts of alkali, sulphur and chlorine is present in the ash. A vast experience on the matter exists from coal firing [Sarofim and Helble 1994; Bryers 1992; Harb and Smith, 1990] as well as from firing different types of waste sludges [Backman et al., 1987] Backman et al., 1996; Salmenoja et al., 1996]. Forrest derived fuels such as wood, bark or forest residue (branches and tops) contain usually low amounts of sulphur and chlorine. The low potential for sulphur dioxide emissions from combustion of these kind of fuels as well as the indication of a

fairly well behaving ash in most kind of combustion systems, are generally considered as two important advantages for the fuels. Sometimes these general indications are applied also on any type of biomass. This may, however, lead to serious errors since other biomasses such as straw or annually grown energy crops may contain significant amounts of both chlorine and sulphur [Nordin 1993]. In Fig. 1 the amount of chlorine and sulphur is shown for a number of different fuels, including coal, peat, wheat straw and forest derived fuel. Fluidized bed combustion is regarded as a very flexible combustion system with a capacity to burn a wide range of fuels. From the ash behaviour point of view the low

combustion temperature of some 800–900 °C is favourable compared to conventional pulverised systems since operational problems due to a low melting ash are assumed to be avoided. The FBC technique has, however, limits. Common knowledge from conventional Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

525

B.-J. Skrifvars et al.

526

pulverised combustion systems may further not be directly applied on FBC combustion

due to the key differences between the technologies, the combustion temperature difference and the presence of the bed in the FB case. In this paper we report results from an extensive measurement campaign where ash behaviour was studied in three different circulating fluidized bed boilers firing both forest derived fuels, straw, peat and/or coal. We focus here on the deposit formation in the studied boilers, especially on the chlorine and sulphur in the deposits. We compare the cases where forest derived fuel with low sulphur and chlorine is fired in a CFB with the cases where coal or peat is fired in the same CFB. We also compare the deposits formed in these cases with a case where a high chlorine straw is fired in a CFB boiler together with coal.

2. EXPERIMENTAL The boilers in which the measurements were conducted were i) an CFB boiler firing coal or wood chips, ii) an CFB boiler co-firing straw and coal and iii) a 125 CFB boiler co-firing forest residue and coal. A schematic view of the sampling locations in the boiler is shown in Fig. 2.

2.1. The

CFB Boiler

The CFB boiler is a hot water boiler producing heat to a local district heating net. The boiler has been used at several occasions before for emission studies [Lyngfelt et al., 1995; Åmand 1994] as well as for ash chemistry studies [Skrifvars et al. 1997]. In these tests coal peat and wood was fired separately at three different occasions [Skrifvars, Backman et al., 1997]. Deposit samples were collected from the post cyclone fluegas channel of the boiler using the air-cooled probe sampling technique [Jackson and Raask, 1961]. Front and back side deposits were analysed separately both quantitatively and with SEM/EDAX. Also the fuel entering the boiler was analysed. CO and as well as temper-

The Role of Alkali Sulfates and Chlorides in Post Cyclone Deposits

527

atures in the bed and in the fluegas at the deposit sampling points were further monitored simultaneously together with boiler operating parameters such as fuel and air feed. The fuel analyses as well as a summary of the fluegas emissions and some key operating parameters during the sampling periods are presented in Table 1.

2.2. The

CFB Boiler

The CFB boiler is a cogeneration plant producing high pressure steam (92 bars/505 °C) for power production [Hansen, 1997]. The boiler co-fires straw and coal. A similar kind of deposit sampling technique was used in this boiler for collecting deposit samples from the post cyclone fluegas channel right in front of the first heat exchanger package (SH 3). Also here the fuel entering the boiler as well as fluegases exiting the

528

B.-J. Skrifvars et al.

boiler were monitored during the test campaign. Details of the measurements is found in the literature [Hansen, 1997]. Table 2 summarises the fuels used in these tests as well as some operating parameters.

2.3. The

CFB Boiler

The CFB boiler produces both high pressure steam (110 bars/540 °C) for power production and heat to a local district heating net [Laurén et al., 1997]. During the test campaign the boiler was firing forest derived residue fuel (Scandinavian wood branches and tops) both alone and together with coal. The share of coal during the cocombustion tests was 20% on a thermal base. Two combustion conditions were tested,

high bed temperature, 920–930 °C and low bed temperature, 840–870 °C. Deposits were

collected in an identical way as was described above. Also fuel analyses were performed

as well as the continuous collection of operational data and fluegas emissions. These data are summarised in Table 4.

3. RESULTS

3.1. The

CFB Boiler

The quantitative wet chemical analyses of the short term deposit samples collected in the post cyclone fluegas channel at a fluegas temperature around 600 °C of the CFB boiler during the wood chips, peat and coal combustion periods are

The Role of Alkali Sulfates and Chlorides in Post Cyclone Deposits

529

shown in Fig. 3. The probe surface temperature was set at 520 °C to simulate a superheater tube. The analyses results are shown as bars and the analysed elements, expressed as weight-% in the sample of their corresponding oxides (except for chlorine), are identified in the legend. The major elements in the wood chips combustion case were silicon, calcium and

potassium on the front side deposits and calcium, potassium and chlorine on the back side deposits. On both sides also sulphur was found. Further more all deposits collected during the wood chip combustion period contained chlorine. In the two back side

B.-J. Skrifvars et al.

530

deposits the chlorine content was more than 10% by weight. In one of the back side deposits the element iron dominated which probably is an indication of a sampling error, i.e. part of the sampling probe material followed with the deposit to the analysis. A special feature for these deposits is further that none of the analyses results added up to 100% oxides. This indicates probably that calcium was partly present as calcium carbonate in the deposit and that the missing part consequently was carbon as carbonate. In the two other combustion cases, i.e. coal and peat, the elements silicon, aluminium, iron and calcium dominated the deposit composition almost independently of fuel and deposition side. All deposits contained sulphur but no chlorine was detected. Figure 4 shows the analyses of short term deposits collected further downwards in the fluegas channel at fluegas temperature of approximately 280°C. Here the probe surface temperature was set at 250°C. The analyses results showed a similar feature as those found in the hotter sampling location, i.e. significant amount of chlorine especially on the back side of the probe during wood chip combustion but no chlorine in the coal and peat combustion cases.

3.2. The

CFB Boiler

The analyses results of the short term deposit samples collected from the superheater region in the CFB boiler are summarised in Fig. 5. The probe surface temperature was set at 520°C. The fluegas temperature at the sampling location varied

between 833–909°C depending on the running condition (Table 4). The deposit sample analyses were made using semi-quantification of SEM/EDAX since the deposit amounts collected during the sampling was so low that no quantitative wet chemical analyses could be made. The major elements in these deposits were calcium, potassium and sulphur in all the tested combustion conditions. A small amount of chlorine appeared in the deposits

The Role of Alkali Sulfates and Chlorides in Post Cyclone Deposits

531

during 100% forest residue combustion when the bed temperature was decreased some 50°C. The corresponding decrease in the fluegas temperature at the sampling location was 20°C. Very small differences could be found in the deposit composition when shifting from 100% forest residue combustion to co-combusting with coal. A small increase in the silicon content may be seen but otherwise the composition is almost the same. No chlorine seemed to be present in any of the samples during the co-firing tests. The analyses results of the short term deposit samples collected from the economiser region in the

CFB boiler are summarised in Fig. 6. Here the probe

surface temperature was set at 340°C. The fluegas temperature at the sampling location varied between 571–587°C depending on the running condition (Table 4). The deposit sample analyses were again made using semi-quantification of SEM/EDAX since also here the deposit amounts collected during the sampling was so low that no quantitative wet chemical analyses could be made. The major elements in these deposit samples were the same as in the hotter sampling location, except for chlorine, which appeared in all the deposit samples collected at this colder sampling location. The highest amount, approximately 10wt-%, was found in the deposit collected during the 100% forest residue run at high bed temperature. The lowest amount, approximately 2wt-%, was found in the deposit collected during the cofiring run at low bed temperature.

3.3. The

CFB Boiler

The quantitative wet chemical analyses of selected short term deposit samples collected in the post cyclone fluegas channel in front of the screen tubes of the CFB boiler during the straw + coal A and straw + coal B combustion tests are shown in Fig. 7. The probe surface temperature was set at 525°C. The major elements in these

532

B.-J. Skrifvars et al.

deposits were potassium, sulphur and chlorine together with silicon, aluminium and calcium. A detailed description of the deposits is found in the literature [Hansen, 1997].

4. DISCUSSION A comparison of the chlorine levels in all the deposits collected from the CFB boilers showed that chlorine was found in almost every deposit collected at the hotter

The Role of Alkali Sulfates and Chlorides in Post Cyclone Deposits

533

sampling location and always found in the deposits at the colder location when biomass was combusted. During peat or coal combustion again no chlorine could be found in any of the deposits. This seems to be the case even at very low chlorine levels entering the boiler with the fuel. Sulphur seemed to be found in the deposits independently of which fuel was fired. The chlorine and sulphur levels found in the different deposit samples are shown in Figs. 8 (hotter sampling location) and 9 (colder sampling locations). A closer examination of the collected deposits in the CFB using SEM/EDAX analyses of cross sectioned deposits revealed that in those deposits collected from the wood combustion test period chlorine was found together with potassium in a scale just next to the metal surface. In one case further a clear corrosion attack on the tube could be seen (Fig. 10).

534

B.-J. Skrifvars et al.

Also the calculated molar ratios of chlorine and sulphur to alkali in the presuperheater deposits (Fig. 1 1 ) and economiser deposits (Figure 12) suggest that sulphur and chlorine would be found almost exclusively as alkali sulphates and chlorides during firing of wood and forest residue or combined firing of forest residue with coal or straw

with coal. The molar ratio of the deposits collected from the hotter sampling location in boiler makes an exception to this. Here the molar ratio combined with the ele-

The Role of Alkali Sulfates and Chlorides in Post Cyclone Deposits

535

mental analyses suggest an excess of sulphur in the deposits. Also in the coal and peat firing cases in the boiler a clear excess in sulphur can be seen. Looking more closely on this and comparing the CFB boiler with the 125 CFB boiler one finds that the fluegas temperatures at the hotter sampling location was somewhat higher in the boiler (875–892°C) than in the boiler (605 °C). One can speculate why it seems as if chlorine would escaped the deposits in the hotter sampling location of the boiler while it doesn’t in the boiler. One possibility is that the chlorine was present in the boiler as gaseous KC1 in the hotter sampling location and in an initial stage condensed directly on the probe metal surface, creating a thin sticky layer of deposit on the probe. With time, however, the chlorine would be diluted in the deposit by other ash particles sticking on the surface. The chlorine may also have been released again from the deposit by reacting with sulphur or by penetrating further down into the metal. A further condensation from the gas phase would have stopped due to an increased temperature gradient in the deposit. In the 18 boiler KC1 may again have condensed already before reaching the probe and consequently continued to be transported to the surface by diffusion as small micron-sized particles. Thermodynamic multi-component, multi-phase phase equilibrium calculations seem to support this kind of a speculation. The calculations indicate that gaseous KC1 would start condensing around 640°C. Such a calculation is shown in Fig. 13. Here a case is shown where all the alkali and sulphur from a typical forest residue fuel is firstly released into the gas phase and then cooled down. Comparing the fluegas temperatures at the sampling locations for the two boilers it shows up that it in the boiler would be high enough for KC1 to be found as gas while it in the boiler would be so low that KC1 would be in a condensed form. A similar indication, i.e. that chlorine would mainly be found as KC1, is also given for the deposit analyses of the coal and straw fired CFB boiler [Hansen, 1997]. The share of chlorine in the fuel entering the boiler is some 10 times higher and the share of potassium is approximately doubled in this boiler compared to the other cases. Indications of the possibility to drive out chlorine from the deposits or block the

536

B.-J. Skrifvars et al.

chlorine to enter the deposits by adding sulphur to the system were seen in these tests. Similar indications have been reported earlier [Henriksen et al., 1995; Michelsen et al. 1996]. The indications were here, however not very consequent. Figure 14 shows measurements from the and boilers where the chlorine level in the deposits is plotted as a function of the levels found at the same time in the stack. In the 18 boiler the trend is clear. As long as sulphur was found in the fluegas no chlorine was detected in the deposits. One should however note that this trend also goes hand in hand with the used fuels, i.e. the trend may therefor also be caused by the fuel rather than by the sulphur entering the boiler with the fuel. One could speculate that it would be the

form in which sulphur, chlorine and alkali is found in the wood, i.e. as easy volatilized

elements, that causes the effect shown in Fig. 14. In the coal and peat firing cases again

alkali would predominantly be found in a mineral form which would make the alkali less available for the chlorine and sulphur. Also in the 125 MW th boiler co-firing with coal at a 20% level decreased the chlo-

The Role of Alkali Sulfates and Chlorides in Post Cyclone Deposits

537

rine somewhat in the economiser deposits but the effect was small and chlorine remained in the deposits throughout the tests. In the CFB boiler no comparison could be made since the combustion tests were all performed at co-firing conditions. The analyses show, however, clearly that both chlorine and sulphur remained in the deposit. An indication of a possible sulphating reaction of the chlorine in the deposits is further given by the fact that mature deposits, collected at a boiler shut down period from the same location as where the short term deposit samples were taken, showed clearly lower chlorine levels than the short term deposits [Hansen, 1997]. This is also in accordance with earlier experiences [Henriksen et al., 1995].

From a deposit stickiness point of view a co-existens of alkali sulphate and chlorine in a deposit is worse than either of the pure components. The melting behaviour changes dramatically when both sulphur and chlorine is present together with alkali compared with a case where only either of the anions are present with alkali. This is illustrated in Fig. 15. A drop in the first melting temperature from 920°C for a chlorine free alkali sulphate to 515°C for a alkali sulphate chlorine ash takes place already with a very

538

B.-J. Skrifvars et al.

low amount of chlorine. A similar behaviour is seen when a small amount of sulphur is added to an alkali chlorine material. Also the operating experiences from the boiler seem to indicate this. Earlier experiences in this boiler indicated that a high sulphur coal used together with straw gave worse deposit problems than a low sulphur coal [Henriksen et al., 1995; Hansen 1997]. The conclusion of this would be that chlorine do react with sulphur, but that it seems to be very hard to get all chlorine out of the deposit. The consequence of this may be a worse deposit build up in the boiler rather than a better, slower one.

CONCLUSIONS The results from these deposit analyses showed that chlorine sulphur and alkali would have a major role in the deposit formation chemistry of a circulating fluidized bed boiler firing biomass also at very low levels of chlorine and sulphur entering the boiler. Alkali sulphates and chlorides have a low first melting temperature, 515°C, which may enhance deposit formation. Chlorine in a deposit may further cause corrosion of the heat exchanger tube on which the deposit is situated. A sulfation of deposits, driving out the chlorine as HC1, was shown in the

CFB. A comparison of the chlorine levels in the deposits with the measured in the fluegas showed that if was found in the fluegas no or low levels of chlorine was found in the deposits. Indications of this were also seen in the CFB. Excess in the fluegas could, however not, inhibit the chlorine to reach the deposits in the post cyclone fluegas channel of the CFB:s if the input amount was high such as in the case with straw firing. Neither could Chlorine be completely inhibited to reach the deposits by co-firing with coal in any of the tested conditions. Only by shifting to 100% coal or peat could this be achieved. More work is needed to quantitatively tie the deposit chemistry with the fired fuel.

Especially the fuel analysis should be developed. It is clear that standard fuel analyses

do not provide sufficient data for making reliable predictions of ash behaviour.

ACKNOWLEDGEMENTS This work has received financial support from the Vattenfall AB, the Nordic Energy Research Program, Norrköping Energi och Miljö AB, the ELSAM Research and Development Programme and the Finnish National Research Program LIEKKI-2. This is gratefully acknowledged. The support from the personnel at the Chalmers University of Technology boiler, the Norrköping Energi och Miljö AB boiler and the I/S Midtkraft boiler is also acknowledged.

REFERENCES Backman, R., Hupa, M., Uppstu, E: Tappi J. 70(6), (1987). Backman, R., Skrifvars, B-J., Hupa, M., Siiskonen, P., Mäntyniemi, J: Journal of Pulp and Paper Science (JPPS) 22 (1996) 4: J119–J126.

The Role of Alkali Sulfates and Chlorides in Post Cyclone Deposits

539

Bryers, R. W: Fireside behaviour of mineral impurities in fuels, from Marchwood 1963 to Palm Coast 1992, in Inorganic transformations and ash deposition during combustion (Ed: S. Benson), United Engineering Trustees Inc. 1992. Hansen, P. F. B: Deposit formation in the convective path of a Danish 80MWth CFB-boiler co-firing straw

and coal for power generation, presented at the Engineering Foundation Conference on the Impact of Mineral Impurities in Solid Fuel Combustion, Kona, Hawaii, USA, November 1997. Harb, J. N., Smith, E. E: Prog. Energy Combust. Sci 16(3), 169 (1990). Henriksen, N., Larsen, O. H., Blum, R., Inselman, S: “High-Temperatrure corrosion when co-firing coal and straw in pulverized coal boilers and circulating fluidised bed boilers”, presented at the VGB conference “Corrosion and corrosion protection in power plants”, November 29–30, 1995, Essen, Germany. Jackson, P. J., Raask, E: J. Inst. Fuel 34, 275 (1961). Laurén, T., Skrifvars, B-J., Backman. R., Hupa, M: Deposit formation in the convective path and the catalysator in the Norrköping Energioch Miljö AB CFB boiler, Result report, Åbo Akademi University, 1997. Lyngfelt, A., Åmand, L-E., Leckner, B: Low NO and emissions from circulating (fluidised bed boilers,

Proc. of the 13th ASME International FBC Conference, Orlando FE, USA, 1995, pp. 1049–1057. Michelsen, H. P., Larsen, O. H., Frandsen, F, Dam-Johansen, K: “Deposition and high temperature corrosion in a 10 MW straw fired boiler”, presented at the Engineering Foundation Conference “Biomass Usage

for Utility and Industrial Power”, April 29–May 3, 1996, Snowbird, Utah, USA. Nordin, A: On the chemistry of combustion and gasification of biomass fuels, peat and waste; environmental

aspects, PhD Thesis, University of Ume, 1993. Salmenoja, K., Mäkelä, K., Hupa, M., Backman, R: J. Inst. Energy, 69(9), 155 (1996). Sarofim, A., Helble, J.J: Mechanisms of ash and deposit formation, in The impact of ash and deposition on

coal fired plants (Ed:s J. Williamson, F. Wigley), Taylor & Francis, 1994, pp. 567–582. Skrifvars, B-J., Sfiris, G., Backman, R., Widegren-Dafgrd, K, Hupa, M: Energy and Fuel 1 1 (1997) 4: 843–848.

Skrifvars, B-J., Backman, R., Hupa, M., Sfiris, G., Åbyhammar, T., Lyngfelt, A: Fuel 77 (1/2), 65 (1997). Åmand, L-E: Nitrous oxide emission from circulating fluidized bed combustion, Dr Thesis, Chalmers University of Technology, 1994.

This page intentionally left blank

FLY ASH DEPOSITION ONTO THE CONVECTIVE HEAT EXCHANGERS DURING COMBUSTION OF WILLOW IN A CIRCULATING FLUIDIZED BED BOILER Terttaliisa Lind 1 , Esko I. Kauppinen 1 , George Sfiris2, Kristina Nilsson 2 , and Willy Maenhaut 3 1VTT 2

Chemical Technology, Aerosol Technology Group, Espoo, FINLAND

Vattenfall Utveckling AB, Stockholm, SWEDEN

3 University

of Gent, Gent, BELGIUM

INTRODUCTION In the search for renewable energy sources, the combustion of biomass has gained wide attention lately. The biomass combustion offers a possibility to produce electricity and heat with no net input of carbon dioxide to the environment. This together with the recirculation of the ash back to the soil creates a sustainable energy cycle which is a desir-

able alternative for the future energy markets. In some cases, though, the ash forming

species contained in the biomass fuels cause deposition problems in the biomass fired plants which lead to lowered energy production and in the worst cases to unscheduled

shut-downs.

Extensive work has been done on the ash deposition problems in coal-fired power plants and it has been thoroughly reviewed by Bryers, 1996. The deposition problems in biomass combustion, even though many times as severe as in coal combustion, have gained less attention so far. Only lately have several studies also addressed the deposition in biomass combustion and co-combustion of biomass with coal (Hansen et al., 1997; Miles et al., 1996; Hein et al., 1995; Nordin and Skrifvars, 1996; Skrifvars et al.,

1997). As it now seems, the ash deposition may be one of the critical issues to be solved in the biomass combustion in order to increase its use in the energy production in the future.

Biomass fuels contain incombustible constituents mainly in the form of inorganic and organic salts as well as organically bound in the plant structure. Compared to coal, the mineral content of biomass fuels is typically very low. The composition of the ash varies greatly between different biomass fuels, but calcium and potassium compounds Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

541

542

T. Lind et al.

are usually the dominating constituents. The ash composition also varies depending on the part of the plant that is used in combustion, e.g. silicon is found in large quantities in the bark and potassium in the fast-growing parts of the plants. During the combustion, the inorganic constituents are transformed into ash particles. The combustion technology and the conditions as well as the fuel determine the characteristics of the formed ash. The ash volatilisation is usually more extensive in biomass combustion than in coal combustion. Because of the wide variety of the biomass fuels and combustion technologies that are utilised, it is difficult to draw general conclusions from the deposition studies that have been carried out so far. In this study, we determined the deposition of the fly ash particles in the convective back pass in a circulating fluidized bed boiler during combustion of Swedish willow Salix. Fluidized beds are a suitable method to burn many biomass fuels because of their flexibility in the acceptable fuel characteristics, e.g. moisture content and the particle size. The ash particles are created as a result of the fuel particle devolatilization and the subsequent char combustion in a suspension in which the fuel particles are mixed with the bed particles and air. Bed particles may be sand, dolomite or calcite, depending on the fuel. The interaction of the ash particles with the bed particles makes the fluidized bed ash particles quite unique when compared with the ash particles from other combustion processes. This combined with the relatively low combustion temperatures (typically

800–900°C in biomass combustion) determines the ash particle characteristics from fluidized beds (Valmari et al., 1998a; Valmari et al., 1998b).

METHODS

Plant Description The deposition measurements were carried out at a full-scale circulating fluidized bed boiler in Nässjö Kraftvarmeverk, Sweden, Fig. 1. The unit is a cogeneration plant fuelled with solid biomass. The plant has been operated since 1990. The capacity of the

boiler is 26MW heat and 9MW electricity. The tests were carried out during three days in March 1997. The fuel during the experimental period was a Swedish willow Salix, Table 1. All the process parameters were collected once a minute and stored by a computer. The boiler conditions were kept as stable as possible during the measurements. The circulating fluidized bed combustion (CFBC) average process values are given in Table 2. The bed temperature was about 750°C and the temperature at the top of the furnace about 880°C. CO varied from 100 to more than 500ppm. Fuel, fly ash, bottom ash and bed sand samples were collected

Fly Ash Deposition onto the Convective Heat Exchangers

543

during the measurements and analysed for the content of Na, Al, Si, P, S, Cl, K and Ca, Table 3. The fuel sample was ashed at 550°C prior to the analysis (except for the elements Cl and S).

Mass Size Distributions The fly ash deposition in the convective back pass was determined by measuring the fly ash particle size distributions at two locations in the flue gas channel: measurement location #1 downstream of the cyclone and upstream of the convective back pass at the flue gas temperature of 830°C, and location #2 downstream of the convective back pass and upstream of the electrostatic precipitator at the flue gas temperature of 160°C,

544

T. Lind et al.

Fig. 1. The fly ash particle size distributions were determined by collecting size-classified fly ash samples with an 11-stage, multijet compressible flow Berner-type low-pressure

impactor (BLPI) (Kauppinen, 1992; Kauppinen and Pakkanen, 1990). The deposition was calculated as the difference in the two size distributions. No soot-blowing was carried out in the boiler during the measurements. The collections with the impactor upstream of the convective pass were carried out

after the dilution with a dilution probe. In the dilution probe, the sampled gas was simultaneously diluted and cooled. The dilution probe consisted of three coaxial stainless steel tubes. Cool, dry dilution air was introduced into the sample through a porous inner tube (pore diameter ) which was covered by another tube. The dilution air flow was regulated with a critical orifice. Cooling air cooled both the probe surface and the dilution air. With the dilution air and the cooling air, the temperature of the sample coming out of the dilution probe could be regulated. The impactor was heated up in an oven to 120°C in order to avoid water condensation during the collection. A precutter cyclone was used before the dilution probe in-stack to collect particles larger than about 3µm aerodynamic diameter. The sample flow through the pre-cutter cyclone was 4.5 slpm and the dilution air flow 16 slpm, resulting in a dilution ratio of 1:4.6. Downstream of the convective pass the impactor was inserted inside the flue gas channel. Thin aluminium (Al) and polycarbonate (poreless Nuclepore, NP) films were used as impaction substrates. Prior to the collection, the substrates were greased with Apiezon L vacuum grease to prevent particle bounce and heated in an oven for 6 hours at 150°C to avoid substrate mass loss during collection. In total, 3 BLPI samplings were carried out both upstream and downstream of the convective pass. Two collections at both locations were carried out using Nuclepore collection substrates to be used for elemental analysis.

Elemental Analysis The impactor and cyclone-collected samples were analysed for Na, Al, Si, P, S, Cl, K and Ca. The elements Na and Al were measured by instrumental neutron activation analysis (INAA) (Maenhaut and Zoller, 1977), Si, P and S by particle-induced X-ray emission (PIXE) (Maenhaut et al., 1981a; Maenhaut et al., 1981b), and Cl, K and Ca by both techniques.

Fly Ash Sampling for Electron Microscope Fly ash samples were collected from measurement point # 1 for particle characterisation with scanning electron microscopy (SEM) and elemental analysis with energy dis-

Fly Ash Deposition onto the Convective Heat Exchangers

545

persive X-ray analysis (EDX). The samples were collected using a sampling probe developed at VTT Aerosol Technology Group to collect particle samples from high tempera-

tures. In the probe, the collection of the particles is based on thermophoresis. The sampling probe enables sampling from temperatures of up to 1,200°C, while avoiding particle agglomeration on the sample. The particle density on the sample is optimised for SEM. With the probe, the particles are collected on a small copper grid which is coated with a thin carbon layer normally used for transmission electron microscope (TEM) samples. The sample is cooled with pressurised air while inserting the probe into the duct or furnace and the air stream is turned off for the sampling time, typically a few seconds. After sampling, the air stream is again turned on and the probe taken out from the duct. The sample grid may then be mounted on a SEM or TEM sample holder and the analysis can be made without coating the sample. The morphology of the particles collected during the measurements in Nässjö Kraftvärmeverk was studied using a field-emission SEM (LEO 982 GEMINI) with magnifications of 10,000–200,000 and accelerating voltages down to 0.7 kV. The EDX analyses were carried out with a NORAN Voyager 3 analyser.

Particle Size Analysis with Laser Diffraction The size distribution of the cyclone-collected particles from measurement point #1 was determined using laser-diffraction method Coulter LS 130 at the Mineral Processing department of VTT Chemical Technology. For the analysis, the sample was homogenised, weighed and dispersed in sodium pyrophosphate solution to create a stable dispersion to be used in the analysis. Ion-exchanged water was added and a suspension was created by sonicating in ultrasonic bath for 1 min. Water was not expected to solve particles in suspension since only coarse fly ash particles were analysed with laser-diffraction. Only small fraction of the coarse fly ash particles was water soluble according to ion-chromatography analysis. The laser-diffraction analysis was carried out using polarisation intensity diffraction scattering (PIDS) mode. The particle size distribution obtained with laser diffraction was combined with the size distributions obtained with BLPI to cover the size range 0.01–200 µm at measurement location #1. The particle-size distributions obtained by laser diffraction were found to correspond to the aerodynamic particle size, i.e. particle density This was determined by comparing the size distributions of the fly ash particles in the same sample obtained with laser diffraction to those obtained with the impactor method. In combining the size distributions, the aerodynamic size range 0.01–2 µm was covered by the impactor results, the aerodynamic size range by the cyclone-collected sample and the size distribution in the aerodynamic size range was calculated as a sum of the mass concentrations from the impactor and the cyclone. This was done by adding the mass concentrations from the impactor to the mass concentrations from the cyclone sample in each impactor size class. This was carried out for impactor stages and

RESULTS

Fly Ash Particle Size Distributions The fly ash particle concentrations in the flue gas upstream and downstream of the convective back pass were _ and _ respectively, Table 4.

546

T. Lind et al.

The fly ash particle mass size distributions had clearly two modes at both locations, Fig. 2. The fine particle mode was at 0.2–0.3 µm (aerodynamic diameter) at both locations whereas the coarse mode had a peak at about 6 µm at location #1 and at about at location #2. The concentrations of the fine mode were at both locations very close to each other, and in locations #1 and #2, respectively. However, due to a higher total concentration in location #1, the fine mode represented 8% of the total particle concentration at location #1 and 19–25% at location #2. It has to be pointed out that at location #1 the mass size distributions were measured after diluting the flue gas and hence the size distributions contain the gas phase species that were condensed during the dilution. These species include at least the chlorides of sodium and potassium which condense at a temperature much lower than the temperature at the measurement location #1, i.e. 830°C. In the flue gas, these species condense in the convective back pass on the surfaces of other particles. These measurements show that the same amount of the gas phase species were condensed in the dilution and in the flue gas channel during a slower cooling because the fine mode particle concentrations in the both locations are very close to each other.

Morphology of the Fly Ash Particles The fly ash particle morphology was studied at location #1 from the samples collected in-duct. The coarse fly ash particles were found to be agglomerates,

Fly Ash Deposition onto the Convective Heat Exchangers

547

Fig. 3a, that consisted of up to thousands of primary particles. Fig. 3b shows the surface of the coarse ash particle shown in Fig. 3a. The primary particles are about 100–200 nm in size. The irregular particle shape of the coarse fly ash particles may play an important

role in the deposition. So far, most of the deposition studies including the studies on

sticking probability have only concentrated on spherical particles. The irregular particles may have a sticking probability very different from the one of a spherical particle with the same aerodynamic diameter. The fine fly ash particles , Fig. 3c, were found to be single spherical particles. They contained mainly sulphates and chlorides of potassium with some calcium and sodium, Table 4.

Deposition The deposition of the fly ash particles in the convective back pass between two sootblowing periods as a function of particle size in the size range 1–100 µm was calculated by dividing the mass concentration in each size class at location #2 by that at location #1, Fig. 4:

deposition = 1– (concentration(#2)/concentration(# 1))

The fly ash particle deposition increased with particle size from about 5% at up to about 90% at The slight decrease in the deposition in the last data point at may be due to re-entrainment of the large fly ash particles from the heat exchanger tube surfaces. However, it may also be due to measurement uncertainties as this is the largest size class collected at location #2. The total deposition calculated from the total particle concentrations was 56–64%. The deposition in the size range . was not determined as the dilution at location #1 resulted in condensation of gas phase species

onto particles during the sampling. Hence, the fine particle size distribution measured

548

T. Lind et al .

Fly Ash Deposition onto the Convective Heat Exchangers

549

with the impactor did not represent the fine particle size distribution in the flue gas channel. Figure 4 also shows the deposition curve which is expected if the deposition would only occur by inertial impaction (Wessel and Righi, 1988). The calculations for inertial impaction were done for a tube with a diameter of 33.7mm using particle aerodynamic diameters and the flue gas velocity 10m/s. By comparing the deposition curve measured in this study to the one expected from inertial impaction it can be concluded that the particles in the size range were deposited much more efficiently than expected from the inertial impaction only. Hence, the particles in the size range must have deposited by other mechanism than inertial impaction, presumably by turbulent deposition. The fly ash deposition in the convective back pass was studied in the same boiler with another biomass fuel (Valmari et al., 1998b). The studied fuel was a Swedish forest residue, and the deposition in the convective pass was found to be 70–80%. The amount of particles collected in the electrostatic precipitator during sootblowing was found to agree with the amounts of the particles deposited between sootblowing periods.

Elemental Concentrations and Deposition The concentrations of the elements Na, Mg, Al, Si, P, S, Cl, K and Ca in the flue gas in the fine particle and the coarse particle modes are given in Table 4. Clearly two classes of elements can be distinguished: the elements which were enriched and the elements which were depleted in the fine particles. K, Cl and S as well as Na to some extent showed enrichment in the fine particles. Mg, Al, Si, P and Ca were clearly depleted in the fine particles. The fine particle mode was formed from the species that were released to the gas phase during the combustion and condensed while the gas was cooling down. At least part of the fine mode particles measured with the impactors at the location #1 at the flue gas temperature 830°C were in the gas phase at the measurement location and were condensed during the dilution. The condensation resulted in the enrichment of the released species in the fine particles. The penetration of the different elements in the convective back pass was calculated from the values in Table 4 by dividing the concentration of each element at location #2 by the value at location #1. The deposition was calculated separately for the fine particles and the total amount of particles, Table 5. Here, it has to be noted that the fine particle concentration at location #1 also includes species that were in the gas phase at the measurement location and were condensed during the dilution. The penetration for the small particles varies from 68% for Ca and P to more than 200% for Si and Al. The high values for Si and Al are probably due to the very low total concentrations in the fine particles leading to large uncertainties in the measured values. In general, the penetration values for fine particles are 70–140%. The large variation is probably due to the fact that not all species were condensed at location # 1 as well as to the heterogeneity in fuel and some variation in the combustion conditions. The total elemental penetrations exhibit large differences between different elements. The penetration is clearly smallest for Si with only 4–5% of Si passing the convective pass. Smaller than average penetrations (average 36–44%) are also observed for Al, Ca and P. These are all elements that were enriched in the large particles. The penetration is higher than average for Na, S, Cl and K, the species that were mainly found in

550

T. Lind et al.

the fine particles. The total amount of Cl in the particles was larger at location #2 than

at location # 1. Hence, the total penetration for Cl is more than 100%, presumably because more Cl was condensed during the slow cooling of the flue gases in the flue gas channel than in the rapid cooling during the dilution at location #1. The deposition results seem to be in contradiction to some earlier studies on biomass combustion fly ash deposition in stoker-fired boilers (Bryers, 1996). These studies showed that the volatile species K, S and Cl were enriched in the deposit whereas in our study they were actually deposited less than the other, non-volatile species. This leads to the conclusion that also other factors than the deposition rate of the particles may affect the deposit characteristics. Such factors may be e.g. the uneven removal of the deposit in the soot-blowing or chemical reactions of the deposit with gas phase species. In this

study, we did not measure the concentration of HC1 in the flue gas. It is quite possible that HC1 as well as would react with the deposit resulting in its enrichment in the deposit. Also, soot-blowing may remove only a part of the deposit while the innermost layer remains on the tube surface. The elemental concentrations as a function of particle size, Fig. 5, were determined at both measurement locations. K, Cl and S concentrations show a clear increase with decreasing particle size, quite in agreement with their enrichment in the fine particles due to the condensation of gaseous species. Ca and P concentrations are very low in the particle size range and after a rapid increase in the size range the concentrations remain constant. Si concentration increases continuously with increasing particle size. The concentrations of all the elements in each size class at location #2 were very close to the concentrations at location #1. This means that no preferential deposition of any of the elements occurred in the convective pass because this would have resulted in the decrease in the concentration of such an element from location #1 to location #2. Hence, the deposition seems to be governed mainly by the particle size so that the coarse particles are deposited more effectively than the fine particles, and the particle composition is not a determining factor. In Fig. 5, the concentrations of Cl and K are actually slightly larger at location #2 than at location #1 due to the heterogeneous condensation of their gaseous species on the particles in the convective pass.

Fly Ash Deposition onto the Convective Heat Exchangers

551

T. Lind et al.

552

CONCLUSIONS The deposition of fly ash onto the convective back pass in the CFBC of Swedish willow (Salix) between two soot-blowing periods was found to be 56–64%. The deposition was mainly determined by the particle size. The particle composition had no effect

on the deposition. The deposition increased rapidly with increasing particle size. For particles and about 90% for the particles. The elements which were mainly in the coarse fly ash fraction, Si, Al, Ca and P, were deposited most efficiently, whereas the deposition was smaller than average for K, S and Cl. Large amounts of K, S and Cl were found to be released during the combustion and were therefore found in the fine fly ash particle fraction.

example, it was found to be about 5% for the

ACKNOWLEDGEMENTS This work was funded by the CEC through a contract JOR3-CT95-0001 which is gratefully acknowledged. We want to thank Mr. Raoul Jarvinen and all the personnel at Nässjö Kraftvärmeverk for their contribution to the measurements. W. Maenhaut is

indebted to the “Fonds voor Wetenschappelijk Onderzoek—Vlaanderen” for research support. He is also grateful to J. Cafmeyer and M.-T. Fernández-Jiménez for assistance in the chemical analyses.

REFERENCES Bryers, R.W. (1996). “Fireside slagging, fouling, and high-temperature corrosion of heat-transfer surface due to impurities in steam-raising fuels.” Prog. Energy Combust. Sci., 22, 29–120.

Hansen, P.P., Lin, W. and Dam-Johansen, K. (1997). “Chemical reaction conditions in a Danish 80 Mwth CFBboiler co-firing straw and coal.” Proceedings of ASME 14th FBC Conference, 287–294. Hein, K.R.G., Heinzel, T., Kicherer, A. and Spliehoff, H. (1996). “Deposit formation during the cocombustion of coal-biomass blends.” Applications of Advanced Technology to Ash-Related Problems in Boilers (L. Baxter and R. DeSollar, eds.), Plenum Press, New York, 1996, 97–116. Kauppinen, E.I. (1992). “On the determination of continuous submicron liquid aerosol size distributions with

low pressure impactors.” Aerosol Sci. Technol., 16, 171–197. Kauppinen, E.I. and Pakkanen, T.A. (1990). “Coal combustion aerosols: a field study.” Environ. Sci. Technol., 24, 1811–1818. Maenhaut, W. and Zoller, W.H. (1977). “Determination of the chemical composition of the South Pole aerosol by instrumental neutron activation.” J. Radioanal. Chem., 37, 637–350. Maenhaut, W., Selen, A., Van Espen, P., Van Grieken, R. and Winchester, J.W. (1981a). “PIXE analysis of aerosol samples collected over the Atlantic Ocean from a sailboat.” Nucl. Instr. and Meth., 181, 399–405.

Maenhaut, W., Cornelis, R., Cafmeyer, J. and Mees, L. (1981b). “Analysis of skeleton remains, ascribed to Mary of Burgundy, and of soil samples, recovered from the central tomb of the Church of Our Lady, Bruges.” Bull. Soc. Chim. Belg., 90, 1115–1125. Miles, T.R., Miles, T.R.Jr., Baxter, L.L., Bryers, R.W., Jenkins, B.M. and Oden, L.L. (1996). “Boiler deposits from firing biomass fuels.” Biomass and Bioenergy, 10 (2–3), 125–138. Nordin, A. and Skrifvars, B.-J. (1996). “Quantification of deposit formation rates as functions of operating conditions and fraction of biomass fuel used in a converted PC boiler (100MW).” Applications of Advanced Technology to Ash-Related Problems in Boilers (L. Baxter and R. DeSollar, eds.), Plenum Press,

New York, 117–127. Skrifvars, B.-J., Sfiris, G., Backman, R., Widegren–Dafgard, K. and Hupa, M. (1997). “Ash behaviour in a CFBC boiler during combustion of Salix.” Energy & Fuels, 11, 843–848.

Valmari, T., Kauppinen, E.I., Kurkela, J., Jokiniemi, J.K., Sfiris, G. and Revitzer, H. (1998a). “Fly ash formation and deposition during fluidized bed combustion of willow.” J. Aerosol Sci., 29, 445–459.

Fly Ash Deposition onto the Convective Heat Exchangers

553

Valmari, T., Lind, T.M., Kauppinen, E.I., Sfiris, G., Nilsson, K. and Maenhaut, W. (1998b). “A field study on ash behaviour during circulating fluidized bed combustion of biomass: Parts 1 and 2”. Submitted to Energy & Fuels.

Wessel, R.A. and Righi, J. (1988). “Generalized correlations for inertial impaction of particles on a circular cylinder.” Aerosol Sci. Technol., 9, 29–60.

This page intentionally left blank

ASH BEHAVIOUR IN BIOMASS FLUIDISED-BED GASIFICATION Antero Moilanen, Esa Kurkela, and Jaana Laatikainen-Luntama VTT Energy P.O. Box 1601, FIN-02044 VTT Espoo, Finland

1. INTRODUCTION In biomass combustion, ash deposit formation is a common problem and has been studied by a number of researchers [Miles et al., 1995; Baxter and DeSollar, 1995; Nordin et al., 1995; Bryers, 1996; Miles et al., 1996; Moilanen et al., 1996]. There are also some

observations about operational problems in fluidised-bed gasification processes, caused by ash. In pressurised steam-oxygen gasification of peat, ash deposits have been formed in the upper part of the gasifier and in the cyclones [Moilanen, 1993]. Furthermore, straw ash has been found to cause both bed sintering and deposit formation in air-blown gasification [Kurkela et al., 1996], These problems were difficult to overcome in strawalone gasification. In fact, the gasification temperature had to be reduced to below 800–850°C, which resulted in poor carbon conversion and high tar concentrations. On the other hand, co-gasification of coal and straw (up to 50wt% straw) could be carried out without any signs of ash problems in spite of high operation temperatures of the order of 950–980°C [Kurkela et al., 1996]. One variable of significance that was observed to prevent the detrimental behaviour of ash in the gasification process was carbon conversion and measures to achieve this [Kurkela et al., 1996; Skrifvars et al., 1995]. The completeness of fuel carbon conversion is dependent on the reactivity of residual char and the operating conditions. If the reactivity is high, ash is formed rapidly and, consequently, deposits are also formed rapidly. The gasification reactivity of the biomasses has been observed to vary within wide limits [Moilanen and Kurkela, 1995; Moilanen and Saviharju, 1997]. Ash and its composition can be regarded essential in this respect and they vary significantly in different biomasses [Wilèn et al., 1996]. The ash components, mainly alkaline metal, contribute catalytically to the rate of gasification, which may increase or decrease as a function of conversion depending of the behaviour of catalytically active substances. However, it is rather unknown, in detail, how these substances react during the gasification of biomass chars. Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

555

556

A. Moilanen et al.

The aim of this study was to obtain data on the detrimental ash behaviour of different biomass types in fluidised-bed gasification, and on the basis of these data to determine the process conditions and measures of preventing this kind of behaviour. Different types of biomass fuel relevant to energy production, such as straw and woody biomasses were used.

2. MATERIALS AND METHODS Two different experimental methods were used in this study: laboratory thermobalance studies and bench-scale fluidised-bed gasification.

2.1. Laboratory Studies Different woody biomasses and straws were used as sample materials in this study. The samples are listed in Table 2. Analytical data on these biomasses have been published previously by Wilén et al. [1996] and Kurkela et al. [1996]. The most detailed studies were carried out with willow, wheat straw and alfalfa. The ash compositions of these feedstocks are presented in Table 1. A pressurised thermobalance (PTG, Fig. 1) was used to study the characteristic reaction behaviour of the fuel and its ash. Elevated pressure was applied, as the elevation of pressure has been proven to contribute to ash sintering at lower temperatures than in ambient pressure [Moilanen and Kurkela, 1997]. In the present tests, the effect of temperature and pressure as well as the gasification agent on gasification reactivity and ash sintering were determined using the following reaction conditions: temperature 850°C (isothermal measurement), pressure 1 or 30 bar, steam as gasification agent. If needed tests were also carried out at lower temperature of 750°C. With some feedstocks also was used as gasification agent.

For a test, the fuel sample (particle size below 0.2mm, sample amount 100–200mg)

was placed in the cylindrical sample holder having the wall made of wire mesh. In an isothermal test, after the adjustments were completed (i.e. temperature, pressure, gas composition), the sample was lowered into the reactor with the winch system (Fig. 1). During

Ash Behaviour in Biomass Fluidised-Bed Gasification

557

the measurement, the gasification temperature and weight of the sample were monitored. The weight change recorded during the first 60 second period was due to several effects including buoyancy and gasification, and this part of the weight-time curve was removed in the evaluation process of the results. The sintering degree of ash residue after the reaction was inspected under microscope.

2.2. Fluidised-bed Reactor Studies The behaviour of ash in fluidised-bed gasification was studied in the bench-scale atmospheric fluidised-bed reactor (AFB, fuel feed rate 0.5kg/h, bed diameter 5cm, freeboard diameter 10cm, electrically heated jackets, Fig. 2). In this reactor, ash agglomeration and deposit formation was monitored both in the bed and freeboard by collecting samples from the reactor after the tests. The operation parameters of the reactor were selected to obtain differences in the ash agglomeration behaviour both in the bed zone as well as in the freeboard. The main operational parameters comprised temperature, bed material and the use of steam. The gasification agent was introduced only as the primary fluidization gas and no secondary or tertiary air was used. The selected operation conditions were in many respects similar to those of real large-scale gasifiers. Temperatures varied from 750 to 840°C, and aluminium oxide, limestone, dolomite and coal coke were used as bed materials. Most of the tests were carried out with Danish wheat straw–95.

3. RESULTS AND DISCUSSION 3.1. Thermobalance Tests The ash residues from the thermobalance experiments were studied by microscopy using the following classification criteria (Fig. 3, Table 2): 1. Non-sintered ash residue: ash structure resembling the original fuel particles, easily crumbling when touched (no asterisk: o)

558

A. Moilanen et al.

2. Partly sintered ash (different degrees in this group): particles contained clearly fused ash (1 or 2 asterisks: *, **); 3. Totally sintered ash: the residue was totally fused to larger blocks (3 asterisks: ***).

The thermobalance tests carried out in high pressure steam showed that under pressure ash sintering with some of the feedstocks was much stronger than in atmospheric pressure. Willows (Finnish and Swedish), spruce bark and alfalfa had this type of behav-

iour. The ash residues were totally sintered when willow and spruce bark samples were gasified under 30 bar steam and at 850°C, and the same result was obtained even at 750°C. With alfalfa, the ash residue was completely sintered even at 700°C. On the other hand in atmospheric pressure, no sintering was observed at 850°C with willow and spruce and clearly weaker sintering with alfalfa.

Ash Behaviour in Biomass Fluidised-Bed Gasification

When gasifying in

559

at 850°C, a different sintering behaviour was observed: the

ash residues of the willow were strongly sintered, but the ash of spruce bark sintered less in these conditions. The sintering of alfalfa was also weaker under than in steam gasification. The reason for this behaviour is assumed to be connected with the lack of silicon in these ashes (see Table 1). The ash of pine bark differed from that of spruce bark in having no sintering tendency in these tests. This observation is in consistence with the tests in the fluid-bed gasifier in which pine bark has shown no ash problems. The other biomasses, like pine saw dust indicated strong sintering after gasification in the thermobalance. This was somewhat surprising, because the saw dust ash had not shown any deposition or agglom-

eration problems in the fluidised bed gasification. The reason for this can be that the ash content of saw dust is very small, and hence, ash problems are avoided due to the dilution of ash in the process.

In earlier fluidised-bed gasification tests of forest residues at VTT (mainly spruce based), ash deposits were formed in the cyclones of the gasifier. To find out the reason

for this behaviour, this feedstock was characterised by separating different parts: needles,

bark and stem. According to the results, the bark seemed to be the most critical component of forest residue feedstock. The bark ash of forest residue B in Table 2 showed strong sintering under pressure, similarly as measured with the separately studied spruce bark sample. The ash from needles seemed to have only a weak sintering tendency in pressurised gasification, and stem ash indicated no signs of sintering.

The wheat straw ash generally showed strong agglomeration in gasification tests performed earlier at VTT, but this behaviour seemed to be dependent on the quality of the straw used as feedstock [Kurkela et al., 1996]. In the thermobalance tests of this study, the wheat straw ashes sintered strongly when gasified at 850°C both under 1 bar and also

560

A. Moilanen et al.

under 30 bar steam. Even at 750°C strong sintering was observed in some samples. The use of as gasification agent resulted in less sintering. To find out the reason for the differences observed in the ash behaviour of wheat straws, the relationship between the chemical composition, sintering and gasification reactivity of ash was studied in further detail for this biomass. Table 1 presents the chemical composition of the ash of Danish wheat straw (= 95), willow and alfalfa. Accordingly, the potassium content of this straw ash was very high; 30.1% of in ash. The wheat straw samples previously used in the gasification tests (indicated as Danish wheat straws A and B) had a content of 8.1 and 18.1% in ash, respectively [Kurkela et al. 1996]. Their gasification reactivity behaviour, i.e., gasification rate vs. conversion, is shown in Figs. 4 and 5. The reactivity is expressed in form of instantaneous gasification reaction rate, %/min (i.e. mass change rate divided by residual ash-free mass), and the conversion (given as fuel conversion, % ash-free) is the reacted part of the whole biomass including pyrolysis. The reactivity behaviour of wheat straws showed that the gasification rate can be increasing or decreasing as a function of conversion, depending on the quality of wheat straws B and = 95 showed an increasing trend and straw A a decreasing trend. This difference seems to be associated with the potassium and silicon ratio in the ash. Potassium is known to be a strong catalyst in gasification reactions, and the potassium silicate formation could decrease its activity, as indicated e.g. by Kannan

and Richards [1990]. The gasification reactivity behaviour has been shown to affect the sintering of ash in such a way that the retarding gasification rate prolongs the achievement of the total conversion and, as a consequence, ash is also formed slower and the carbon material prevents ash particles from agglomerating. This phenomenon indicates that the control of carbon conversion may contribute to ash sintering in gasifiers [Moilanen and Kurkela 1995; Kurkela et al., 1996]. Figure 4 shows that for wheat straw A, which had a retarding trend of gasification rate, the time required for achieving the ash conversion of 100% was more than four times longer than for straw B (which had the trend upwards). But, when straw A was gasified totally the resulting ash residue was also totally sintered (Table 2). The effect of potassium content on the reactivity behaviour was studied more closely by washing the straw samples with water. The theory behind this was that rain

Ash Behaviour in Biomass Fluidised-Bed Gasification

561

had probably influenced straw A by leaching out potassium. Henriksen et al. [1997] have also reported decreasing gasification reactivity after washing straw. Thus, straw B was washed with distilled water, and a result similar to straw A was obtained (Fig. 5). The effect of the intensity of water leaching was also tested: the washing procedure was carried out with two methods (indicated as I and II in Fig. 5). Sample I was washed with water at ambient temperature and sample II with boiling water. The potassium content after washing was not determined, but the ash content was reduced from 4% to 3% in sample I and to 2.2% in sample II (the amount of ash was taken from the thermobalance tests). The reactivity reduced clearly to the same level as that of straw A after washing independent of washing intensity.

3.2. Results of the Fluidised-bed Reactor Tests In the bench-scale fluidised-bed reactor (AFB), nine gasification tests were carried out. Danish wheat straw =95 was used as the fuel in tests AFB 1-8, while test AFB9 was carried out with spruce bark. A mixture of air and steam was used as the gasification agent in all tests except test AFB2, which was carried out without air. The air to fuel ratio was selected so that it corresponded to 31–33% of the air to fuel ratio of stoichiometric combustion. Approximately 45 vol% of the fluidising gas was steam in all experiments. The use of relatively high air ratio and high steam feed was necessary to achieve high enough carbon conversion (typical to real-scale gasifiers) without recycling the cyclone fines. The test conditions and the main results are presented in Table 3. The total test time used for the steady state gasification set point is given in Table 3. The time when the first changes in process parameters (i.e. sudden variations in the bed temperature and/or a pressure rise in the reactor) during the test indicated that agglomeration took place is also given. Attention was paid to deposit formation in the freeboard and in the bed. The size and strength of agglomerates were different in every test depending on the operating conditions. Some test runs had to be interrupted, because so large agglomerates were deposited in the bed that fluidising was no more possible. Examples of such agglomerates are shown in Fig. 6. The agglomerates were mainly large and irregular, except for one test, in which they were quite regular in size (AFB6).

562

A. Moilanen et al.

Ash Behaviour in Biomass Fluidised-Bed Gasification

563

In every test, some agglomerates or deposits were found either in the bed or on the freeboard walls (or both). The material deposited on the freeboard wall was not strongly sticked in any of the tests, and it was easy to brush off from the wall (Fig. 7). The amount of freeboard deposit was determined after each test. The amounts are given as milligrams against gram of fed feedstock in Table 3. In test AFB2, the amount was highest and in tests AFB4, 5, 8 and 9 the amounts were very small, in fact almost negligible. The microscopic study of the deposits indicated that in all cases the deposits consisted of particles containing molten ash particles, fine bed material particles sticked to this ash, and also unreacted char particles were present in different amounts. An example of this type of molten ash particle (cross section) is shown in Fig. 7.

At test runs AFB1 and 2, the effect of gasification agent with alumina bed material was tested. The use of steam prevents peak temperatures which can lead to ash

melting. In the test runs AFB3-5, limestone bed material was tested at various temperatures, and in tests AFB6-8, different bed materials were compared. In AFB9 spruce was used as feedstock. In both tests carried out with alumina bed, AFB1-2, large agglomerates were collected from the bed. The agglomerates were less porous in test AFB1, where air was used. Freeboard deposits also occurred in large amounts in these tests. When limestone bed material was tested (AFB3-5) at various temperatures, bed agglomerates were formed at temperatures of 800°C and above. No agglomeration was detected at

564

A. Moilanen et al.

750°C, although particles containing molten ash were seen in the fairly thin freeboard deposit. In the limestone test series, most severe bed agglomeration occurred at about 800°C (AFB4). The freeboard deposits in these tests were clearly thinner than in the previous runs. Tests AFB6 and 7 were carried out at the same temperature level as the tests AFB13. In test AFB6, performed with precalcinated limestone at 830°C, round agglomerates of regular size were obtained, the size of which was growing to such an extent that fluidising was no more possible (Fig. 6). In test AFB7 dolomite was used as bed material and only a few small agglomerates were formed. The temperature level was the same as in the previous test. When using coal coke (AFB8) large bed agglomerates were obtained. These phenomena indicate that the interaction of ash and bed material in the fluidised bed contributes significantly to the agglomeration. No bed agglomerates were detected in the gasification test with spruce bark (AFB9). However, some particles containing fused ash were also seen in the fairly thin freeboard deposit.

3.3. Discussion of AFB Results The tests with alumina bed material resulted in strong agglomeration in the bed. In the first test (AFB1), applying a mixture of air and steam, it was difficult to maintain

the temperature steady at 830°C and large variations were observed in the bed temperature. This was a sign of starting fluidising malfunctions due to agglomeration (in about 1 hour after start). However, fluidisation did not collapse totally until after four hours, when there were already large agglomerates in the bed. The second test (AFB2) assured that no temperature peaks occurred, and therefore the air feed was excluded (heating only

with electric resistances) and only steam was used as gasification agent. Gasification continued for a longer time, but large agglomerates were detected also in this test. However, these agglomerates were porous and very permeable by gas, which probably was the reason for that no dramatic changes in the process parameters were detected during the test run. The contact between straw ash and alumina bed particles forming agglomerates was studied more closely with microscopy. An example of bed agglomerate formed when gasifying straw in the aluminium oxide bed is shown in Fig. 8. In the agglomerate, the bed material particles were glued relatively strongly to each other by molten ash. SEM analyses taken from the bed material moving from the bed material grain towards the sticking ash material indicate the composition at different locations. Point 3 (Fig. 8) presents an analysis of the gluing zone and indicates that it consisted mainly of potassium and silicon, and some calcium. According to the chemical composition of ash (Table 1) potassium and silicon are the major CO components in this wheat straw ash. According to the phase diagrams the binary system has the lowest melting temperature of 740°C (with the composition of 33wt% and 68wt% ) [Levin et al., 1985]. In the next test, alumina was changed to limestone bed material. In AFB3, the temperature level was the same as in the previous two runs. Also in this test, large agglomerates were formed in the bed already within 2 hours after the start. Despite the fact that agglomeration took place also in the limestone bed, the contact between ash and bed material differed from that in alumina bed. The agglomerates in the limestone bed were more brittle due to the friable structure of limestone particles (but firm enough to resist bed attrition). In test AFB4 (limestone bed material), bed agglomeration was even more intense than in AFB3 although the test temperature was lower (795°C). An explanation for this could be the formation of fused carbonate between uncalcined limestone and straw ash, which can be connected with the slower rate of calcination in this run. The

Ash Behaviour in Biomass Fluidised-Bed Gasification

565

566

A. Moilanen et al.

eutectic temperature of the mixture of potassium carbonate and calcium carbonate is about 750°C [Levin et al., 1985]. However, this theory of carbonates need to be investigated in further detail with bed materials and ashes. When pre-calcined limestone was used (AFB6), a large amount of round ash agglomerates of fairly regular size was formed in the bed (Fig. 6). The agglomerates were formed by molten ash, which encapsulated

bed material particles inside. In this run, there was an increase in the amount of freeboard deposit due to easy attrition on calcined limestone. The deposit contained, in addition to molten ash particles, relatively high amounts of limestone particles sticked to them. When dolomite was used (AFB7) only slight agglomeration was observed in the bed. This was probably due to the calcination of forming MgO, and therefore affecting the structure of bed particles. The coke bed material was selected as the co-gasification of coal and straw had earlier been successful [Kurkela et al., 1996]. The preventive mechanism in this test was supposed to be due to that ash melt could be absorbed by the porous coke. However, this was not the case in this test. This can be understood on the basis of the measurements by Raask [1985], who observed that ash melt does not wet carbon surfaces. Possible reasons for the difference in the agglomeration behaviour may lay in differences in the operational conditions of these gasification tests. The co-gasification test was carried out under pressure and at higher temperature (around 1,000°C). At this temperature the gasification reactivity of coal char is significantly higher than at the temperature used in this study (abt 800°C). This low reactivity was also seen in the bed material (removed after the test), which consisted mainly of unreacted coke. In the co-gasification test, however, the coal reacted totally, so that coal ash was able to become mixed with straw ash. Furthermore, the fluidisation and bed agitation conditions were more intense in the pressurised PDU gasifier, which may have inhibited agglomerate formation.

4. CONCLUSIONS Characteristic data on ash behaviour can be obtained by laboratory tests and can be used for planning, for example, run conditions for tests with pilot or PDU equipment. Ash sintering can be much stronger in pressurised conditions than under atmospheric pressure. This phenomenon seems to be related to the silicate content in ashes: when the silicon content was low the ash sintering was stronger in pressurised conditions than in atmospheric conditions. However, the chemistry need to be studied more in detail to understand the behaviour observed. The water-soluble constituents of straw (potassium, calcium) affect the relationship of ash sintering and reactivity. Both bed agglomeration and deposit formation in the freeboard occurred in the tests carried out using straw with problematic ash. The intensity of agglomeration was dependent on operational parameters. It was possible to affect the formation rate of agglomerates and deposits by the choice of gasification conditions and bed material. The agglomeration of bed particles is a combination of the chemical characteristics and the particle sizes of bed material and ash. The strength of the agglomerates seems to be dependent on the friability of the bed material particles. The alumina bed material particles were firm enough to form strong agglomerates resisting fluidisation, while the limestone particles (depending on the quality and the calcination degree) were friable leading to brittle agglomerates or large regular size particles in the fluidising bed.

Ash Behaviour in Biomass Fluidised-Bed Gasification

567

ACKNOWLEDGEMENTS The financial support from the Finnish National Research Programme LIEKKI2 and VTT Research Programme PROGAS through the Technology Development Centre Finland (Tekes), companies Carbona Inc., Elkraft, Foster Wheeler Energy Oy and Imatran Voima Oy as well as VTT is gratefully acknowledged.

REFERENCES Baxter, L. & DeSollar, R. (eds.) (1995). Application of advanced technology to ash-related problems in boilers. New York: Plenum Press.

Bryers, R. W. (1996). “Fireside slagging, fouling, and high-temperature corrosion of heat-transfer surface due to impurities in steam-raising fuels.” Proc, Energy Combust. Sci., 22, 29–120.

Henriksen, U., Jacobsen, M. J., Lyngbech, M. J. & Hansen, M. W. (1997). “Relationship between gasification reactivity of straw char and water soluble compounds present in this material”. In: A. V. Bridgwater and D. G. B. Boocock (eds.). Developments in thermochemical biomass conversion. Glasgow: Blackie Academic & Professional. Pp. 881–891.

Kannan, M. P. & Richards, G. N. (1990). “Gasification of biomass chars in carbon dioxide: dependence of gasification rate on the indigenous metal content.” Fuel, 69, 747-753. Kingery, W. D., Bowen, H. K. & Uhlmann, D. R. (1976). Introduction to ceramics. 2nd. Ed. New York: John

Wiley & Sons. 1032 p. Kurkela, E., Laatikainen-Luntama, J., Ståhlberg, P. & Moilanen, A. (1996). Pressurised fluidised-bed gasification experiments with biomass, peat and coal at VTT in 1991–1994. Part 3. Gasification of Danish wheat straw and coal. Espoo: VTT Energy. 41 p. + app. 5 p. VTT Publ. 291.

Levin, E. M., Robbins, C. R. & McMurdie, F. (1985). Phase diagrams for ceramists. Vol. I. 5th print. Washington: The American Chemical Society. 601 p. Miles, P. E. T. R., Miles, T. R. Jr., Baxter, L. L., Bryers, R. W., Jenkins, B. M. & Oden, L. L. (1995). Alkali deposits found in biomass power plants. Summary Report. Golden, CO: National Renewable Energy Laboratory. 82 p. + 35 app. Miles, T. R., Miles, T. R. Jr., Baxter, L. L., Bryers, R. W., Jenkins, B. M. & Oden, L. L. (1996). “Boiler deposits from firing biomass fuels.” Biomass and Bioenergy, 10 (2–3), 125–138. Moilanen, A. (1993). Studies of peat properties for fluidised-bed gasification. Espoo: Technical Research Centre

of Finland. 34 p. + app. 35 p. VTT Publ. 149. Moilanen, A. & Kurkela, E. (1995). “Gasification reactivities of solid biomass fuels.” Preprints of papers presented at the 210th ACS National Meeting, Chicago, III., 20–24 August 1995. Am. Chem. Soc. Div. of Fuel Chem., 40 (3), 688–693. Moilanen, A. & Kurkela, E. (1997). Characterisation ash deposit formation in willow fluidised bed gasification.

Espoo: VTT Energy. To be published. Moilanen, A. & Saviharju, K. (1997). “Gasification reactivities of biomass fuels in pressurised conditions and product gas mixtures.” In: A. V. Bridgwater and D. G. B. Boocock (eds.). Developments in thermochemical biomass conversion. Glasgow: Blackie Academic & Professional. Pp. 828–837. Moilanen, A., Sipilä, K., Nieminen, M. & Kurkela, E. (1996). “Ash behaviour in thermal fluidised bed conversion processes of woody and herbaceous biomass.” In: P. Chartier, G. L. Ferrero, U. M. Henius, S. Hultberg, J. Sachau and M. Wiinblad (eds.). Biomass for energy and the environment. Kidlington: Elsevier. Pp. 706–711. Nordin, A., Öhman, M., Skrifvars, B.-J. & Hupa, M. (1995). “Agglomeration and de-fluidization in FBC of biomass fuels—mechanisms and measures for prevention.” In: L. Baxter and R. De Sollar (eds.) Application of advanced technology to ash-related problems in Boilers. New York: Plenum Press. Raask, E. (1985). Mineral impurities in coal combustion; Behaviour, problems, and remedial measures.

Washington: Hemisphere Publishing Corporation. 484 p. Skrifvars, B.-J., Hupa, M., Moilanen, A. & Lundqvist, R. (1995). “Characterization of biomass ashes.” In: L. Baxter and R. De Sollar (eds.) Application of advanced technology to ash-related problems in Boilers. New

York: Plenum Press. Wilén, C., Moilanen, A. & Kurkela, E. (1996). Biomass feedstock analyses. Espoo: VTT Energy. 25 p. + app. 8 p. VTT Publ. 282.

This page intentionally left blank

BENCH-SCALE BIOMASS/COAL COFIRING STUDIES Deirdre Belle-Oudry and David C. Dayton National Renewable Energy Laboratory 1617 Cole Boulevard Golden, CO 80401-3393

1. INTRODUCTION The threat of increased global warming has subjected the use of fossil fuels to increasing scrutiny in terms of greenhouse gas and pollutant emissions. As a result, the use of renewable and sustainable energy resources, such as biomass, for electricity production has become increasingly attractive. The use of dedicated biomass feedstocks for electricity generation could help reduce the accumulation of greenhouse gases because carbon dioxide is consumed during plant growth. The agricultural and wood products industries generate large quantities of biomass residues that could also provide fuel for electricity production. Increasing the use of these waste biomass fuels could alleviate the burdens of waste disposal in the agricultural and wood products industries. One solution to increasing the use of biomass to produce electricity is to build dedicated biomass power plants that use 100% biomass fuel. As of July 1995, there were 360 biomass plants and approximately 1% of the grid connected electricity in the United States was generated from renewables [Cogen, 1995]. The initial capital investment to build new biomass power plants to increase this percentage is high. Another scenario for increasing the use of biomass to produce electricity is to cofire biomass and coal in extant coal-fired power plants. Coal-fired power plants are used to produce most of the electricity in the United States and if biomass were cofired at low percentages in a small number of coal-fired power plants, the use of biomass for power production could dramatically increase. Cofiring biomass and coal can increase the use of sustainable fuels without large capital investments and take advantage of the high efficiencies obtainable in extant coal-fired power plants. Fuel diversity is another advantage of biomass/coal cofiring. Cofiring reduces the need for a constant supply of biomass that would be required in a biomass power plant. Cofiring biomass and coal is therefore a viable way to manage the increasing emissions of greenhouse gases and may also reduce other pollutants from power-generating facilities. Biomass and coal have fundamentally different fuel properties. For instance, biomass is a more volatile fuel than coal and has a higher oxygen content. Coal, on the other hand, Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

569

570

D. Belle-Oudry and D. C. Dayton

has more fixed carbon than biomass. In general, biomass contains less sulfur than coal, which translates into lower sulfur emissions as higher blending ratios of biomass are used. Wood fuels tend to contain very little ash (on the order of 1 % ash or less) and consequently increasing the ratio of wood in biomass/coal blends can reduce the amount of ash that needs to be disposed. A negative aspect of biomass is that it can contain more potassium and chlorine than coal. This is particularly true for some grasses and straws. Several utilities have tested biomass/coal cofiring in utility boilers [Boylan, 1993; Gold and Tillman, 1993]. Several issues, however, remain regarding how blending biomass and coal will affect combustion performance, emissions, fouling and slagging propensities, corrosion, and ash saleability [Tillman and Prinzing, 1995]. Many of these issues have been highlighted in a comprehensive, multi-laboratory study known as the European Commission’s APAS Clean Coal Technology Programme [Bemtgen, Hein, and Minchener, 1995] that was formed to study the feasibility and technical challenges associated with replacing coal partially by biomass and converting wastes (such as sewage sludge) to energy in coal-fired power plants. In an effort to further address issues that face biomass/coal cofiring, representatives from the National Renewable Energy Laboratory, Sandia National Laboratories Combustion Research Facility, and the Federal Energy Technology Center have embarked on a collaborative effort to study many of the fireside issues that face biomass/coal co-combustion such as ash behavior, particle capture efficiency, carbon burnout, and emissions, and reactivity. This paper describes bench-scale biomass/coal cofiring experiments in support of this collaborative effort.

2. EXPERIMENTAL APPROACH The combustion behavior, gaseous emissions, and alkali metals released during the combustion of several biomass/coal blends were investigated with a direct sampling, molecular beam mass spectrometer (MBMS) system [Evans and Milne, 1987] in conjunction with a high-temperature quartz-tube reactor that has been described in detail in the literature [Dayton, French, and Milne, 1995; Dayton and Milne, 1996]. The biomass and coal samples, including the blends, were received from L. Baxter of the Sandia Combustion Research Facility. For this study, results are presented for blends of Eastern Kentucky coal with Red Oak wood chips, Danish Wheat Straw, and Imperial Wheat Straw (from California). The proximate, ultimate and ash analyses for the pure fuels are presented in Table 1. Blends are reported as a percentage on an energy input basis, based on the higher heating value of the feedstock. The blends investigated during this study consisted of 5%, 15%, and 25% biomass, on an energy input basis, with the Eastern Kentucky coal. Twenty to fifty milligrams of the pure fuels and blended samples were loaded into hemi-capsular quartz boats that were placed in a platinum mesh basket attached to the end of a ¼-in. diameter quartz rod. This quartz rod can be translated into a heated quartzor alumina-tube reactor enclosed in a two-zone, variable-temperature furnace. Furnace temperatures were maintained at 1,100°C and a mixture of 20% in He was flowed through the reactor at a total flow rate of 3.0 standard liters per minute. The residence time of the combustion products in the reactor before sampling was 0.7–0.9 second. Gas temperatures near the quartz boat were measured with a type-K thermocouple inserted through the quartz rod. The actual boat temperature and the flame temperature were not measured.

Bench-Scale Biomass/Coal Cofiring Studies

571

3. RESULTS AND DISCUSSION The MBMS results for the pure fuels and the blends were similar to results obtained in the past for biomass and coal combustion [Dayton and Milne, 1996]. All the samples exhibited multiple phases of combustion, including the devolatilization and char combustion phases. The char combustion phase for coal was generally longer than the char combustion phase during biomass combustion. Combustion of the blends showed a similarly longer char combustion phase compared to combustion of the pure biomass samples. Triplicate samples of the pure fuels and blends were studied to establish experimental reproducibility. After identifying the combustion products of interest, ion intensity-versus-time profiles were integrated to determine the total relative amount of a given product released during the combustion event. In this way, relative amounts of several products measured during combustion of the pure fuels and blends could be compared. These results were normalized to the

signal intensity measured before the

sample was inserted into the hot zone of the reactor and the sample weight. Relative species concentrations represent the averages of the triplicate samples, and the reported error bars are one standard deviation.

572

D. Belle-Oudry and D. C. Dayton

3.1. Biomass/Coal Blends with Eastern Kentucky Coal Summaries of the MBMS results for combustion of Eastern Kentucky coal, the pure biomass fuels, and the blends of the biomass fuels with Eastern Kentucky coal in the above mentioned ratios in 20% in He at 1,100°C are shown in Figs. 1 and 2. These figures display the relative amounts of 7 combustion products as determined by integrating the time-versus-intensity profiles for ions measured during the combustion event. The main isotope of carbon dioxide at was not measured during these cofiring studies because the signal saturated the detector at the sensitivity required to detect the minor combustion products. Carbon dioxide production could be monitored by measuring the signal intensity at which corresponds to the isotope. This isotopic species accounts for 1.17% of the total carbon dioxide based on the natural abundance of the carbon and oxygen isotopes. Figure 1 displays the relative amounts of the species detected at and during the combustion of the pure fuels and the coal/biomass blends in 20% in He at 1,100 °C. The most was detected during the combustion of the wheat straws and the most and was released during combustion of the coal and the coal blends. In this study, blends of Red Oak and Imperial Wheat Straw with Eastern Kentucky coal were studied in two different ratios to

investigate the possible effect of blending ratio on the combustion products. The data in Fig. 1 suggest that varying the blending ratio has little effect on the amount of CO(g), and released during combustion. The sulfur release is consistent with the fact that the Eastern Kentucky coal has the highest sulfur content of the pure fuels investigated. Except for the Red Oak, the amount of NO(g) released during combustion of

Bench-Scale Biomass/Coal Cofiring Studies

573

the fuels and blends was similar. The most NO(g) was released during the combustion of the 85% Eastern Kentucky coal/15% Imperial Wheat Straw blend.

Figure 2 shows the relative amounts of various chlorides detected at and during combustion of the pure fuels and the Eastern Kentucky coal/biomass blends. The most HCl(g) was detected during combustion of the wheat straws. The most I and were detected during combustion of the Imperial Wheat Straw. This is consistent with the fact that the Imperial Wheat Straw has the highest levels of potassium (2.75 wt% dry basis), sodium (1.66 wt% dry basis), and chlorine (2.17 wt% dry basis) of the pure fuels investigated in this study. Figures 3 to 6 display the relative amounts of NO(g), HCl(g), and KCl(g), respectively, detected during the combustion of the Eastern Kentucky coal/biomass blends compared to the expected amounts of these products based on the combustion results for the pure fuels. Figure 3 shows that any decrease in the amount of observed because of blending biomass with coal was merely caused by diluting the amount of sulfur in the fuel blend compared to the pure coal. The Eastern Kentucky coal contains 0.89 wt% (dry basis) sulfur, the biomass samples contain, on a dry basis, 0.33 wt% (Imperial Wheat Straw), 0.17 wt% (Danish Wheat Straw), and 0.02 wt (Red Oak) sulfur, respectively. Therefore, blending any one of the biomass fuels with the coal reduces the amount of sulfur in the fuel blend. The amount of detected during combustion of the blends was consistent with expectations based on the values calculated from the amount of detected during combustion of the pure fuels. The overall amount of detected was greater for those blends with higher ratios of coal, but these levels of were consistent with the expected amount.

574

D. Belle-Oudry and D. C. Dayton

These results appear to be inconsistent with previous studies published in the literature [Nordin, 1995; Nordin and Ohman, 1996] where significant (up to 95%) sulfur retention by alkali metals in the biomass in coal/biomass blends was observed. A closer examination of these previous studies suggests that the results for emissions during combustion of biomass/coal blends are consistent. The previous studies were conducted in various fluidized bed combustors where the gas/solid interaction times are cosiderably longer than those in the quartz-tube flow reactor used in this study. The bed temperatures established in the previous studies (550–850°C) were also lower than the 1,100°C furnace temperature established in the experiments presented in this study. The earlier studies [Nordin, 1995] clearly show that sulfur retention by the biomass ash decreases sharply with temperature above 850°C. In addition, the earlier studies [Nordin and Ohman, 1996] claim that in large scale biomass/coal cofiring in a pulverized fuel boiler, sulfur emissions were reduced according to the percentage of biomass used in the fuel blend. The gas/solid interaction times are considerably shorter in pulverized fuel boilers compared to fluidized bed combustors. Therefore, these observations from earlier work are consistent with emissions measured for the residence times and temperatures established in the present study.

Figure 4 displays similar data for the amount of NO(g) released during combustion of the Eastern Kentucky coal/biomass blends. All of the NO(g) measured in these studies came from the conversion of fuel-bound nitrogen to NO(g) since He was used as a diluent in the reactor atmosphere instead of Therefore, the amount of NO(g) measured in these laboratory experiments does not contain a contribution from thermal which can be substantial in commercial-scale pulverized fuel systems. Within the errors

of the measurements, the amount of fuel-bound nitrogen converted to NO(g) during combustion of the blends was consistent with expectations based on the amount of NO(g) released during combustion of the pure fuels. Once again, the blending ratio may

affect the relative amount of NO(g) released but there were no synergistic effects during

Bench-Scale Biomass/Coal Cofiring Studies

575

combustion of the blends to suggest that any more or less NO(g) was released than expected based on the combustion results for the pure fuels. The wheat straws have less fuel-bound nitrogen, on a dry basis, (1.08 wt%—Imperial Wheat Straw and 1.04 wt%— Danish Wheat Straw) than the Eastern Kentucky coal (1.49 wt%). The Red Oak contains very little fuel-bound nitrogen on a dry basis (0.20 wt%). In reality, higher blending ratios of biomass with coal may have a more substantial effect on production because more biomass in the fuel blend may decrease flame temperatures, catalyze reactions, or have other effects on the combustion process that can affect thermal production in utility boilers. Figure 5 shows the measured and calculated amount of HCl(g) released during combustion of the Eastern Kentucky coal/biomass blends. The amount of HCl(g) detected during the combustion of the Eastern Kentucky coal/Red Oak blends should be minimal because neither the Red Oak nor the coal contain very much chlorine. Therefore, the amount of HCl(g) detected during the combustion of the Eastern Kentucky coal/Red Oak blends was consistent with the expected values based on the combustion results for the pure fuels. For the Eastern Kentucky coal/wheat straw blends, however, more HCl(g) was released during combustion of the blends compared to the expected amounts of HCl(g) determined from the combustion results for the pure fuels. The amount of Imperial Wheat Straw in the blend does not appear to affect this result. Figure 6 summarizes the amount of KCl(g) detected during combustion of the Eastern Kentucky coal/biomass blends compared to the amount of KCl(g) expected to be released based on the combustion results for the pure fuels. Converse to the HCl(g) results, the amount of KC1 vapor detected during combustion of the 85% Eastern Kentucky coal/15% Imperial Wheat Straw blend was less than expected. For the 95% Eastern Kentucky coal/Imperial Wheat Straw blend this conclusion is not as evident given the large error bar on the measured value. The results for the 85% Eastern Kentucky

576

D. Belle-Oudry and D. C. Dayton

coal/Danish Wheat Straw blend were also not as conclusive because of the large error bar on the measured value. The Red Oak has very low potassium and chlorine contents, which suggests that the measured values for KC1 vapor detected during combustion of the Eastern Kentucky coal/Red Oak blends may be near or below the detection limits for KCl(g). Similar data for the relative amounts of NaCl(g) detected during the combustion of the Eastern Kentucky coal/biomass blends were also obtained. These data suggest that less NaCl(g) was detected during the combustion of the Eastern Kentucky coal/wheat straw blends than expected based on the combustion results for the pure fuels. The amount of NaCl(g) released during combustion of the Eastern Kentucky coal/Red Oak blends was negligible and insensitive to the blending ratio.

3.2. Thermochemical Equilibrium Calculations An equilibrium analysis of the biomass/coal blend combustion was undertaken in an attempt to explain some of the observations made during the batch combustion experiments. These calculations were not an attempt to model the batch experiments preformed in this study. The calculations were performed using a modified version of STANJAN [Reynolds, 1986], a thermodynamic equilibrium computer code that minimizes the Gibbs free energy of the system by the method of element potentials with atom population constraints. The theory relates the mole fraction of each species to element potentials for each independent atom in the system. This method is particularly suited to large systems because the element potentials and the total number of moles in each phase are the only variables. Gas-phase species were treated as ideal gases in the calculations, and the condensed phase was assumed to be an ideal solution. This simplified treatment of the condensed phase may not accurately represent reality, and caution should be exercised in overinterpreting the calculated condensed-phase species mole fractions [Blander, et al., 1997].

Bench-Scale Biomass/Coal Cofiring Studies

577

Information about the mechanics and mathematics of the program is available in the lit-

erature [Van Zeggemon and Storey, 1970]. The main program has been modified to accept as many as 600 species and 50 phases [Hildenbrand and Lau, 1993]. A comprehensive database of species and related thermodynamic data was used to predict the equilibrium gas- and condensed-phase compositions given an initial temperature and pressure as well as the mole fractions of the following atoms in the fuel or blend: Al, Ba, C, Ca, Cl, Fe, H, He, K, Mg, Mn, N, Na, O, P, S, and Si. Equilibrium product compositions were calculated for the Eastern Kentucky coal, the three biomass fuels, and the three 85% coal/15% biomass blends. The equilibrium calculations suggest that the mole fractions of NO(g) and for the blends are consistent with the linear combination of NO(g) and calculated for the pure fuels. This signifies that any difference in the amounts of and NO(g) measured during combustion of 85% Eastern Kentucky coal/15% biomass blends was caused by dilution. This is consistent with the experimental observations. The equilibrium results for HC1 and KC1 vapors showed similar trends as the experimental results. Less gas-phase KC1 was predicted from the compositions of the blends versus the amount of KC1 calculated from the ratios of the equilibrium results for the pure fuels. Conversely, more gas phase HC1 was predicted from the equilibrium calculations on the compositions of the blends compared to the amount of HC1 calculated from the combinations of HC1 predicted from the pure fuels. Within the limitations of how realistically the equilibrium calculations treat the condensed phase, the effect of blending coal and biomass on the composition of the ash can be interpolated from the equilibrium calculations. For instance, the amount of condensed-phase KC1 calculated from the composition of the blends was lower than the amount of condensed-phase KC1 predicted from the linear combinations of the equilibrium results for the pure fuels. The results of the calculations suggest that the concentrations of the alkali aluminosilicates were enhanced when coal and wheat straw

578

D. Belle-Oudry and D. C. Dayton

were blended compared to a simple ratio based on the equilibrium results for the pure

fuels. Therefore, it can be inferred from the equilibrium results that the reduction in KCl(g) and subsequent increase in HCl(g) that was observed during the combustion of wheat straw/coal blends results because potassium is being sequestered by the coal mineral matter. Less potassium is available for volatilization, therefore, the volatile Cl is released as HCl(g).

4. CONCLUSIONS The MBMS results for the relative amount of detected during combustion of the coal/biomass blends suggested that any decrease in the amount of observed because of blending coal and biomass was the result of diluting the sulfur in the fuel blend. The conversion of fuel-bound nitrogen to NO(g) during combustion of coal/biomass blends met expectations based on the combustion results for the pure fuels.

This may not be the case, however, in full-scale biomass/coal cofiring applications. The MBMS experiments were isothermal and the NO(g) detected was the result of fuel bound nitrogen conversion to NO(g) because He was used as the diluent in the reactor atmosphere instead of N 2 eliminating the formation of thermal NOx. Of course, in a full-scale cofiring situation there are many other engineering factors that contribute to formation. Utility boilers are far from isothermal and addition of biomass in a pulverized coal-fired boiler can significantly change the flame structure and characteristics. The addition of biomass can potentially have positive effects for reduction in commercial facilities because of the high oxygen and volatiles content of biomass. Adding biomass can also reduce flame temperatures leading to lower levels of thermal and the high moisture content of some biomass may also be effective for reduction at full-scale. Therefore, the bench-scale experiments described herein suggest that there are no synergistic interactions between the pure fuels in a given biomass/coal blend all else being equal. In full-scale cofiring applications, where all else is not equal, this conclusion may be quite different

The chlorine released during the combustion of the coal/biomass blends did appear to be affected by blending the two fuels beyond just a dilution effect. The amount of KCl(g) and NaCl(g) detected during the combustion of the coal/wheat straw blends was lower than expected. The results of the equilibrium calculations suggest that potassium from the biomass was sequestered in the ash by the coal mineral matter as alkali aluminosilicates. The high concentrations of alumina and silica in the coal tend to interact with the large amount of potassium in the wheat straws. With less potassium available for volitilization, the amount of HCl(g) released during the combustion of the coal/wheat straw blends was higher than expected based on the combustion results for the pure fuels. Blending coal and high chlorine- and alkalicontaining fuels apparently affects the chlorine equilibrium in such a way that cannot be explained based on just mixing the pure fuels. Some other chemical interactions between alkali metals in biomass and mineral matter in coal affects the partitioning of chlorine in the gas phase between alkali and hydrogen chlorides. A reduction in gas-phase alkali chlorides could potentially lead to lower than expected fouling and slagging in commercial-scale systems when cofiring a high alkali containing biomass fuel with

coal. In fact, these effects with alkali metals may be enhanced in continuous feeding systems where the gas/solid contact time would be longer compared to the batch experiments discussed above.

Bench-Scale Biomass/Coal Cofiring Studies

579

5. ACKNOWLEDGMENTS The authors acknowledge support from the Solar Thermal and Biomass Power Division of the U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy. Special thanks go to Richard L. Bain and Thomas A. Milne for both programmatic and technical support and guidance. Dr. Larry Baxter and Dr. Alien Robinson, Sandia National Laboratories, supplied the biomass and coal samples.

6. REFERENCES Bemtgen, J.M., Hein, K.R.G., and Minchener, A.J. (1995). APAS Clean Coal Technology. Volume 1: Co-Utilization of Coal, Biomass, and Waste, Volume 2: Combined Combustion of Biomass/ Sewage Sludge and Coals, and Volume 3: Co-Gasification of Coall Biomass and Coall Waste Mixtures. Published at the Institute for Process Engineering and Power Plant Technology, University of Stuttgart. Blander, M, Milne, T., Dayton, D., Backman, R., Blake, D., Kuhnel, V., Linak, W., Mann, M., and Nordin, A. (1997). “Chemistry of the Combustion of Biomass: A Round-Robin Set of Calculations Using Available Computer Programs.” Proceedings from the Engineering Foundation Conference on the “Impact of Mineral Impurities on Solid Fuel Combustion,” November 2–7, 1997, Kona, HI. Boylan, D.M. (1993). “Southern Company Tests of Wood/Coal Cofiring in Pulverized Coal Units.” Proceedings from the conference on Strategic Benefits of Biomass and Waste Fuels, March 30 April 1, 1993 in Washington, DC EPRI Technical Report (TR-103146). pp. 4–33. Cogen Database, Utility Data Institute, July 1995. Dayton, D.C., French, R.J., and Milne, T.A. (1995). “The Direct Observation of Alkali Vapor Release During Biomass Combustion and Gasification 1. The Application of Molecular Beam/Mass Spectrometry to Switchgrass Combustion.” Energy and Fuels, 9(5). 855–865.

Dayton, D.C. and Milne, T.A. (1996). “Laboratory Measurements of Alkali Metal Containing Vapors Released During Biomass Combustion.” In L. Baxter and R. DeSollar (Eds.), Application of Advanced Technologies to Ash-Related Problems in Boilers, New York: Plenum Press, pp. 161–185. Evans, R.J. and Milne, T.A. (1987). “Molecular Characterization of the Pyrolysis of Biomass: I. Fundamentals,” Energy and Fuels 1, pp. 123–137. Gold, B.A. and Tillman, D.A. (1993). “Wood Cofiring Evaluation at TVA Power Plants EPRI Project RP 3704–1.” Proceedings from the conference on Strategic Benefits of Biomass and Waste Fuels, March 30–April 1, 1993 in Washington, DC. EPRI Technical Report (TR-103146). pp. 4–47. Hildenbrand, D.L. and Lau, K.H. (1993). Thermodynamic Predictions of Speciation of Alkalis in Biomass Gasification and Combustion, SRI International, Inc. Final Report for NREL Subcontract XD-2-112231, Menlo Park, CA, February. Nordin, A. (1995). “Optimization of Sulfur Retention in Ash when Cocombusting High Sulfur Fuels and Biomass in a Small Pilot Scale Fluidized Bed,” Fuel 74(4), pp. 615–622. Nordin, A. and Ohman, M. (1996). “Sulfur Capture by Cocombustion with Biomass Fuels—Gathered Experiences of Process Optimization and Emission Minimization,” Proceedings of the Third Nordic Conference on SOX and NOX From Heat and Power Generation, in Lyngby, Denmark, March 13–14, 1996. Reynolds, W.C. (1986). “The Element Potential Method For Chemical Equilibrium Analysis: Implementation in the Interactive Program STANJAN,” Department of Mechanical Engineering, Stanford University. January. Tillman, D.A. and Prinzing, D.E. (1995). “Fundamental Biofuel Characteristics Impacting Coal-Biomass Cofiring.” Proceedings from the Fourth International Conference on the Effects of Coal Quality on Power Plants, August 17–19, 1994 in Charleston, SC. EPRI Technical Report (TR-104982) pp. 2–19. Van Zeggemon, F. and Storey, S.H. (1970). The Computation of Chemical Equilibria, Cambridge, England, Chapter 2.

This page intentionally left blank

IRON IN COAL AND SLAGGING The Significance of the High Temperature Behaviour of Siderite Grains During Combustion

Gary Bryant, Christopher Bailey, Hongwei Wu, Angus McLennan, Brian Stanmore†, and Terry Wall

Cooperative Research Centre For Black Coal Utilisation Department Of Chemical Engineering University Of Newcastle Callaghan, 2308, NSW, Australia † Department of Chemical Engineering University Of Queensland St Lucia, 4097, Qld, Australia

1. INTRODUCTION Accurate prediction of the severity of ash deposition on heat transfer surfaces in pulverised fuel fired furnaces has been the driving force of many studies. These studies

initially attempted to develop engineering indices, which could be used to predict the severity of deposition from a particular coal, based on bulk analysis of the coal prior to utilisation [Raask, 1985]. The indices developed were only applicable to certain coal ash

chemistries and found little application for other coals. A number of studies were then undertaken to determine the behaviour of specific minerals [Srinivasachar, 1990a, 1990b],

which allowed interpretation of experimental and industrial scale data in terms of

mineral specific reactions. Iron in coal has long been linked with slagging. Consequently, several engineering indices make some allowance for iron content when predicting deposition severity [Juniper, 1995]. The Northern Hemisphere coals upon which these indices are based generally contained pyrite as the principal form of iron. In Australian coals iron is predominantly in the form of carbonates, hence the engineering indices developed for the Northern Hemisphere coals would only loosely apply to Australian coals. Siderite limestone dolomite where MgO 4055wt.%) and ankerite are the principal carbonate minerals found in Australian coals. The variable compositions found for siderite in Australian coals led Patterson et al. 1994 to classify four main types. Siderite (where is the Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

581

582

G. Bryant et al.

purest form, although minor amounts of CaO, MgO or MnO may be present. Magnesium

(magnesian) siderite occurs where MgO substitutes for 8–35wt.% of in the mineral lattice. Calcium (calcian) siderite occurs where CaO substitutes for 8–22wt.% of in the mineral lattice. Magnesium calcium (Magnesian Calcian) siderite occurs when CaO substitutes for 8–18 wt.% of in the mineral lattice, and MgO substitutes for 8–15 wt.% of in the mineral lattice. For ankerite the composition is extremely variable with up to 60wt.% of MgO being substituted by The mineral matter present in coal can be classified as either excluded mineral matter or included mineral matter. Excluded mineral matter is that portion of the mineral matter which is separated from the carbonaceous portion of the coal during pulverising, while included mineral matter is intimately associated with the combustible material in the coal. Excluded mineral matter is exposed to more oxidising conditions and lower particle temperatures compared to included mineral matter in the same furnace/environment. Further the low probability of interaction with other particles in the post combustion flue gases allows excluded mineral matter to be considered independently of other particles in the hot combustion gases. Reactions involving included mineral matter must take into account mineral associations within a given coal particle, as coalescence of minerals in close proximity in a coal particle is likely to occur.

1.1. Decomposition Reactions of Carbonates The behaviour of siderites has been studied previously, with thermogravimetric (TG) and differential thermal analysis (DTA) being the usual techniques. The general reaction for the decomposition of carbonates is [Warne, 1979]:

where M = Fe, Ca, Mg, Mn, the subsequent oxidation of wustite (FeOn) in an oxidising atmosphere being [Vasyutinski, 1969]:

Each carbonate species has a characteristic threshold temperature (depending upon the decomposition atmosphere) where thermal decomposition will commence. These are summarised in Table 1 [Kulp, 1951;Patterson, 1991; Hurst, 1993]. From Table 1 it is evident that there is a wide range of temperatures for the decomposition of siderite. However, all decomposition temperatures encountered are significantly less than the temperatures used in this experimental program or those found in a pulverised fuel fired boiler. Coal contains almost entirely ferrous iron [Fe(II)], during pyrolysis the ferrous iron content does not change. At the completion of combustion and deposition the ferrous iron accounts for 20% of the total iron present in coal ash [Ram et al., 1995]. Wustite, produced as a result of the initial decomposition of siderite, is unstable below 572°C [Raask, 1985], and depending upon the temperature and local environment oxidises to hematite The overall rate of decomposition depends on many factors including the partial pressure of the temperature, particle size and the extent of diffusion of through the particle [Bryers, 1987]. In pulverised fuel fired boilers siderite and ankerite have been reported to fragment when heated rapidly [Raask, 1985; Bryers, 1996] yielding fume particles 0.1 to in diameter. In contradiction to these reports, a recent paper concluded that siderite did not fragment in the flame, but melted to produce

Iron in Coal and Slagging

583

non-sticky particles when fired at 1,400°C under both reducing and oxidising conditions, with typical particle residence times of 30ms [Ten Brink, et al. 1994]. The siderite type was not specified, nor whether the particles were included or excluded.

1.2. Decomposition of Pyrite There have been several studies on the decomposition reactions of excluded pyrite. These include oxidising conditions [Srinivasachar and Boni, 1989; Srinivasachar et al. 1990; Ten Brink et al., 1996], and reducing conditions [Ten Brink et al., 1996]. Included pyrite transformations under combustion conditions have been reported [Bool et al., 1995; Ten Brink et al., 1996]. The decomposition and oxidative pathways for excluded and included pyrite are shown in Fig. 1. The decomposition of pyrite is very rapid taking 400ms for a particle with a high degree of fragmentation occurring during

this process [Srinivasachar et al., 1990]. Of particular interest is the Fe-O-S melt phase which is formed. This species has a very low melting point (~1,080°C), which resulted in

100% of particles adhering to a deposition probe at temperatures above this value [Srinivasachar et al., 1990].

1.3. Ternary Phase Diagrams As discussed earlier, excluded and included mineral matter are expected to undergo different reactions due to different time-temperature histories and local environments. Excluded siderite can be represented by the mixed siderite system, where the major inorganic oxides are calcium, magnesium and iron, and the resulting ternary phase diagram for the system, with isotherms, [Sheel, 1975; Patterson et al., 1994] is presented in Fig. 2. The region where potentially sticky particles would form (say a liq-

uidus temperature occurs where and Most siderites and mixed siderites would, upon decomposition, have a composition that would fall outside this region. Also presented on Fig. 1 are the regions corresponding to the compositions found for siderite and mixed siderites contained in Australian bituminous coals discussed earlier. By inspection of Fig. 2, most siderite types contain less than 30% by weight CaO, so the determining factor for potentially slagging particles is the MgO content. Included siderite residues, alternatively, are expected to react with other minerals within the burning char particle, in particular quartz and clays. The ternary phase diagram for the system, with isotherms [Nürnberg, 1981], is presented

584

G. Bryant et al.

in Fig. 3. The region in which mineral matter from Australian bituminous coals would fall is indicated. Increasing iron levels in a given coal particle results in an area where the composition of mineral matter could be expected to fall, and this has been observed in other studies [Kennedy and Wall, 1996]. The region where potentially slagging particles

Iron in Coal and Slagging

585

would form (say a liquidus temperature < 1,400°C) occurs were by weight for a particle with a ratio between 1.5 and 4.7. This represents a significantly larger proportion of the ternary phase diagram than encountered for excluded mineral matter.

2. EXPERIMENTAL

2.1. Samples of Excluded Siderite Coal siderite mineral fractions were obtained from three Australian bituminous

coals by density separation at a specific gravity of 2.0, resulting in siderite enriched mineral fractions. Sample A was predominately siderite, sample B was magnesium siderite with some siderite inclusions and a significant amount of pyrite, and Sample C contained siderite with elevated levels of manganese and significant levels of calcite. Chemical

compositions of coal siderite mineral fractions used in the experimental program

expressed as weight percent elemental oxides are given in Table 2. A point count conducted as per Australian Standards [AS 2061 1989, AS 2856 1986] is presented in Table 2 for each sample. Table 2 indicates that the siderite enriched mineral samples contain significant quantities of siderite. Illustrative average compositions determined by SEM EDS analyses of siderite grains, for each of the samples used in the investigation are also presented in Table 2.

2.2. Samples of Included Siderite Samples of coals containing included siderite were prepared by dry sieving (63–

size bin) the float fraction obtained from the density separation conducted to collect excluded siderite. A coal analysis of the included material is presented in Table 2.

A point count conducted on a polished cross section of the included minerals, also presented in Table 2, indicated that the samples contain predominantly quartz and clays with very low levels of siderite.

G. Bryant et al.

586

2.3. Samples of Pure Siderite Two pure mineral samples (samples D and E) were also used, these samples were prepared by crushing well characterised rock samples. Sample D was siderite and sample E magnesium siderite. Illustrative average compositions determined by SEM EDS analyses of siderite grains, for each of the samples used in the investigation, are also presented in Table 3. Comparison of siderite compositions from Table 2 and Table 3 show that coal siderite sample A and pure mineral sample D have similar compositions as do coal siderite sample C and pure mineral E.

2.4. Combustion Tests A laminar flow drop tube furnace with air as carrier gas was used in the investiga-

tion. Furnace set temperatures for the experiments were 1,100°C and 1,600°C. The samples were fed into the furnace at a rate of Ash products were cooled rapidly

Iron in Coal and Slagging

587

in a quench probe, before passing through a cyclone and aerosol filter arrangement designed to separate coarse and fine ash particles

2.5. Analytical Techniques Chemical analyses were conducted using XRF spectroscopy and SEM EDS spectroscopy. Mineral compositions were determined by SEM EDS and point count techniques. Particle morphology was determined by use of SEM imaging, whilst particle size distributions were determined using laser diffraction techniques.

3. RESULTS AND DISCUSSION 3.1. Excluded Siderite Reactions Table 4 shows the chemical composition of ash products, as determined by XRF, for each of the samples at the two combustion temperatures used in the study. From Tab. 4, the chemical composition of combustion residues collected from the furnace under oxidising conditions have similar compositions to the feed sample, with the exception of sample B, which has lost sulphur as a result of the decomposition of pyrite. There appears to be no partitioning of species as a result of reactions at elevated temperatures.

588

G. Bryant et al.

3.1.1. Melting of Excluded Siderite. Figure 4 shows the types of particles resulting from combustion experiments on excluded siderite samples at 1,100°C and 1,600°C, with Table 5 providing particle compositions. Excluded coal siderite residues and pure mineral residues examined by SEM imaging are dry and porous at a combustion temperature of 1,100°C, as shown in Fig. 4. This observation is to be expected due to the low probability that particles will interact with other particles in the hot combustion gases. Melting

of siderite residues of particular compositions has occurred at a combustion temperature of 1,600°C. From Fig. 4 dry porous particles contain MgO levels in excess of 5 wt.%,

whilst particles which have undergone significant softening and melting contain MgO levels less than 5 wt%. 3.1.2. Fragmentation of and Fume Formation from Excluded Siderite. Scanning electron micrographs of products from the combustion of coal residues at 1,100°C and 1,600°C indicate that for all three coal siderite fractions studied there is extensive fracturing of the mineral grain. This is consistent with the hypothesis of rapid evolution weakening the lattice. Although there is weakening of the siderite lattice, there is only minimal evidence for the formation of iron oxide fragments. Particle size distributions for the coal derived excluded siderite residues, which are presented in Fig. 5, indicate that there is minimal fragmentation of excluded mineral matter during combustion of sample A and sample B. Sample C contains a significant amount of calcite which has been well documented to fragment during rapid heating [Bryers, 1996]. Comparison of the proportion of material collected on the filter paper for each sample at 1,100°C and 1,600°C, as shown in Table 6, indicates that increasing combustion temperature, adjusting for losses in the collection system has little effect on the observed extent of fragmentation. Comparisons between the coal residues and pure mineral residues indicate that the pure

Iron in Coal and Slagging

589

mineral fractions fragment more than the coal mineral fractions at the same combustion temperature (1,600°C), although the total degree of fragmentation was observed to be less than 2wt.% for all samples studied. SEM imaging of filter paper residues indicated zero fume formation for the coal mineral samples at 1,100°C, as all the particles on the filter papers were angular in shape. At 1,600°C the only particles collected by the aerosol filter were larger than so these particles were most likely melted iron oxide fragments rather than condensed iron oxide fume. The absence of particles less than indicates there is no evidence for the formation of fume from excluded siderite grains under oxidising conditions.

590

G. Bryant et al.

3.2. Included Siderite Reactions Table 7 shows the chemical composition of ash products as determined by XRF for included siderite samples at the two combustion temperatures used in the study.

3.2.1. Fragmentation of Included Siderite. Primary fragmentation of the coal (thermal shock of heated coal particle and rapid devolatilisation), and secondary fragmentation of the char (combustion and disintegration of char bridges) are the most likely sources for fragments of char containing siderite. Shedding of ash particles from the surface of a burning char is another mechanism by which small included siderite residues may enter the combustion gases. The contribution of the shedding of ash to the final ash particle size distribution under normal pulverised fuel firing has been found to be insignificant [Helbe and Sarofim 1989]. Thus the formation of fine ash from fragmentation of included siderite is a coal dependant variable. 3.2.2. Coalescence of Included Siderite. Figure 6 shows the types of particles resulting from combustion experiments on included siderite samples at 1,100°C and 1,600°C, with Table 8 providing the particle compositions. A variety of particle types resulting from combustion at 1,100°C was observed, some particles being dry and porous at a combustion temperature of 1,100°C, as shown in Fig. 6, whilst others have completely melted. The observation of melted particles is expected as the ternary diagram indicates a liquidus temperature of 1,150°C for eutectic compositions [Nürnberg, 1981] and the burning char particle being in excess of 200°C above the prevailing gas temperatures [Timothy et al., 1982].

At 1,100°C melted particles are characterised by elevated levels of as determined by SEM EDS analysis, this is shown in Tab. 8. Similar observations are made at 1,600°C, and although there is a significantly higher degree of melting, dry porous phases are still present and are characterised by low levels of as shown in Fig. 6 and Table 8.

3.3. Thermodynamic Equilibrium Calculations Thermodynamic equilibrium calculations were conducted using FACT [Bale et al., 1996]. A pure siderite mineral was exposed to gases and solids for two scenarios. First,

Iron in Coal and Slagging

591

excluded siderite exposed to an oxidising environment, and second included siderite exposed to carbon and low partial pressures of oxygen. The resulting iron species distribution as a function of temperature is shown in Fig. 7. Excluded siderite is expected to decompose and oxidise to hematite (Fe2O3) at equilibrium and as temperature increases, wustite levels increase, as does the Fe(II)/Fe(III) ratio. Increasing the Fe(II)/Fe(III) ratio has been shown to decrease slag viscosity and melting temperatures, these parameters being important in relation to ash behaviour in conventional pulverised fuel and integrated gasification combined cycle (IGCC) technologies. For included siderite, however, at low temperatures wustite is the most stable species and as temperature increases wustite melts to form part of the slag followed by the formation of iron fume. Temperatures associated with the formation of fume for both the oxidising and reducing scenarios are higher than would be expected for excluded siderite grains, therefore, from thermodynamic equilibrium calculations, fume is only expected to form from included siderite.

3.4. Comparison of The Decomposition and Reactions of Siderite and Pyrite Excluded pyrite decomposes and reacts to form an iron oxy-sulphide melt phase, which has a liquidus temperature between 900–1,080 °C. The time spent in the molten state is found to constitute approximately 80% of the total time required for oxidation of pyrite to solid magnetite [Srivinasachar and Boni 1989]. Excluded siderite residues used in this study however are not predicted to form melt phases until a temperature of 1,380°C, unless calcio-wustite phase (melting point 1,150°C) forms as a result of the decomposition of siderite. Excluded pyrite has been reported to fragment and form fume during decomposition [Bryers 1996], whilst fragmentation or fume formation from

592

G. Bryant et al.

excluded siderite has not been observed in this study or reported elsewhere [Ten Brink et al. 1994]. Included pyrite decomposes and reacts forming the iron oxy-sulphide melt which may stick to existing surfaces. Alternatively, this oxy-sulphide melt may contact and react further with silicates and alumino-silicates to form sticky particles. For the case where the iron oxy-sulphide melt phase does not contact silicates or alumino-silicates, oxidation to a dry hematite phase is the result. After combustion and oxidation of the iron containing minerals is complete, excluded siderite residues behave similarly to excluded pyrite residues. Included siderite residues have higher melting temperatures, and, if lost from the burning char before further reaction would not be sticky in the combustion gases. However, if included siderite were to contact and react with included silicates and alumino-silicates, sticky particles would result for particular Fe-Si-Al composi-tions. A schematic representation of the behaviour of siderite is presented in Figure 8.

Iron in Coal and Slagging

593

3.5. Future Work Mössbauer spectroscopy will be conducted on excluded and included mineral matter residues in order to establish the Fe(II): Fe(III) ratio. Residence times in the drop tube furnace are short, therefore most iron is expected to be in the form of Fe(II) or wustite.

4. CONCLUSIONS Ternary equilibrium diagrams have been used to interpret the behaviour of included and excluded siderite grains during combustion. For excluded siderite grains these diagrams indicate that particles are sticky at combustion temperatures when MgO levels are low (<2wt.% at 1,400°C or < 5wt.% at 1,600°C), and this prediction has been supported by experiments. For included siderite grains, reaction with clays present in the burning char particle results in iron aluminosilicate glass phases which are molten when FeOn levels are high (>35wt.% at 1,400°C and > approximately 15wt.% at 1,600°C), and this prediction has also been supported by experiments. There was little evidence for fragmentation of excluded siderite grains with fragments forming less than 2wt.% of all excluded mineral residues. Formation of fume has not occurred for excluded siderite residues. However thermodynamic equilibrium calculations predict significant fume formation from included siderite. The production of fume was predicted to increase with increasing temperature and decreasing oxygen partial pressure. Included and excluded siderite, upon heating decompose to wustite. For excluded mineral residues the wustite rapidly oxidises to hematite. For included siderite residues, this wustite can react with char to form fume at elevated temperatures, react with clays present in the char to form alumino-silicates which may form sticky particles, or, if unreacted, oxidise upon burnout to hematite. There are similarities between the behaviour of included pyrite and siderite in terms of formation of melt phases at lower temperatures.

594

G. Bryant et al.

Included siderite is expected to be less troublesome, as the proportion of melt phase for included siderite is much less than that for included pyrite at a given temperature. For excluded minerals, siderite is expected to be much less troublesome due to the much higher liquidus temperatures when compared to excluded pyrite residues.

5. ACKNOWLEDGEMENTS The authors wish to acknowledge the financial support provided by the Cooperative Research Centre for Black Coal Utilisation, which is funded in part by the Cooperative Research Centres Program of the Commonwealth Government of Australia, and The Australian Coal Association Research Program (ACARP).

6. REFERENCES AS 2061-1989, 1989, Preparation of coal samples for incident light microscopy. AS 2856-1986, 1986, Coal—Maceral analysis. Bryers, R.W., (1987). Symposium On Slagging And Fouling In Steam Generators, 63. Bryers, R.W., (1996). Prog. Energy Combust. Sci., 22, pp 29. Juniper, L., (1995). ACTC Combustion News, February Edition. Helbe, J.J. and Sarofim, A.F., (1989). Combustion and Flame, 76, pp 183–196. Hurst, H.J., Levy, J.H., Patterson, J.H., (1993). Fuel, 72, pp 885. Kennedy, E.M and Wall, T.F., (1996). CCSEM Analysis of Coal Minerals, Ash and Deposits, Proceedings, Workshop On Coal Characterisation For Existing And Emerging Technologies, CRC For Black Coal Utillisation, Feb 14–15, ppB38–B42. Kulp, J.L., Kent, P., Kerr, P.P., (1951). The American Mineralogist, 36 (9–10), pp 643. Patterson, J.H., Hurst, H.J., Levy, J.H., (1991). Fuel, 70, pp 1252. Patterson, J.H., Corcoran, J.F., Kinealy, K.M., (1995). Fuel, 73, pp 1753. Raask, E., (1985). Mineral Impurities in Coal Combustion, Springer-Verlag, New York. Scheel, R., (1975). Sprechsaal Keram., Glass. Baustoffe., 108 (23–24), pp 685. Nürnberg, K., (1981). Slag Atlas, Verlag Stahleisen, Germany. Ram, L.C., Tripathy, P.S.M., Mishra, S.P., (1995). Fuel Processing Technology, 42, 47–60. Srinivasachar, S., Boni, A.A., (1989). Fuel, 68, 829–836. Srinivasachar, S., Helble, J.J., Boni, A.A., (1990a). Prog. Energy Combust. Sci., 16, pp 281–292. Srinivasachar, S., Helble, J.J., Boni, A.A., Shah, N., Huffman, G.P., Huggins, F.E., (1990b). Prog. Energy Combust. Sci., 16, pp 293–302. Ten Brink, H.M., Eenkhorn, S., Weeda, M., (1994). Flame Transformations of Coal Siderite in The Impact of Ash Deposition in Coal Fired Plants, Proceedings Engineering Foundation Conference, J. Williamson and F. Wigley Ed, Taylor and Francis, pp 373. Timothy, L.D., Sarofim, A.F., Beer, J.M., (1982). Characteristics of single particle coal combustion, Proceedings, 19th Symp. (Int.) Combustion, The Combustion Institute, Pittsburgh, 1123–1130. Urbain, G., Cambier, F, Deletter, M., Anseau, M.R., (1981). Trans. J. Brit. Ceram. Soc., 80, pp 139. Vasyutinski, N.A., (1969). Mineral Sb., 22(4), pp 407. Warne, S.St.J., (1979). Differential Thermal Analysis of Coal Minerals, In Analytical Methods for Coal and Coal Products Vol III, Ch52, Edited by C. Karr Jr., Academic Press, New York. Wibberley, L.J., (1980). PhD Thesis, University of Newcastle.

FRACTIONATED HEAVY METAL SEPARATION IN BIOMASS COMBUSTION PLANTS— POSSIBILITIES, TECHNOLOGICAL APPROACH, EXPERIENCES Ingwald Obernberger and Friedrich Biedermann Institute of Chemical Engineering, Technical University of Graz A—8010 GRAZ, Inffeldgasse 25, AUSTRIA

1. INTRODUCTION Coincident to forcing the thermal energy utilization of biomass (wood chips, sawdust, bark, straw, cereals), the amounts of combustion residues (ash) increase. Therefore, it is necessary to find ways of utilizing the ashes produced in a sustainable manner. Previous research has shown that the elementary cycle of nature within the process of thermal energy utilization from biomass is disturbed by dry and wet deposition of heavy metals on the forest ecosystem caused by environmental pollution. By separating a side stream rich in heavy metals (the so-called filter fly-ash—precipitated in electrostatic filters, fibrous filters or flue gas condensation units) it should be possible to recycle the major part of the ash produced, the so-called “usable ash”. The usable ash represents a mixture of bottom ash and fly-ashes collected before filter fly-ash precipitation takes place (see Fig. 1). The heavy metal levels measured in the usable ash fraction from state-of-the-art combustion plants show that the concentrations of Cd approach and in some cases even exceed the present Austrian limiting values for the utilization of biomass ashes on agricultural fields or forest soils [Bundesministerium für Land- und Forstwirtschaft, 1997; Salzburger Landesregierung, 1997]. The concentrations of Zn have also shown to be close to the limiting values. Consequently, the aims of technological development are to reduce heavy metal concentrations in the usable ash and to upgrade them in the filter fly-ash, by designing a combustion technology that allows for efficient heavy metal separation during the combustion process. Such a primary measure has the advantage that no further ash treatment is necessary, which reduces the operating costs of the plant and meets the requirements Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

595

596

I. Obernberger and F. Biedermann

for a decentralized closed-cycle economy within the system of energy production from biomass.

2. OBJECTIVES Based on previous biomass research [I. Obernberger, 1995] and the requirements for sustainable ash utilization, comprehensive investigations on the heavy metal fluxes and the influencing variables in biomass combustion plants were carried out. The research work covered the following points [F. Biedermann et al., 1997; I. Obernberger et al., 1997a]:

1. Test runs in a state-of-the-art grate-fired biomass combustion plant (Lofer, Austria) taking into account various operating parameters: Tests with different kinds of bio-fuels (bark, wood chips). Variation of the temperature in the combustion zones in order to check its influence on the heavy metal concentration in hot precipitated fly-ashes. Tests at different plant loads in order to check the influence of different fly-ash production rates on the heavy metal fluxes. 2. Determination of the concentrations of heavy metals and nutrients in the different ash fractions produced and in the bio-fuels used. 3. Evaluation and interpretation of the results in order to examine the possibilities and the potential of a new combustion technology with integrated fractionated heavy metal separation. 4. Implementation of a new technology with integrated fractionated heavy metal separation in a biomass district heating plant.

Possibilities, Technological Approach, Experiences

597

5. Test runs in the new biomass combustion plant with integrated fractionated heavy metal separation. 6. Evaluation of the results achieved in comparison with the results obtained from

the test runs performed in a state-of-the-art combustion plant in order to assess the potential of the new combustion technology as regards fractionated heavy metal separation. 7. Optimization of the new technology to enhance fractionation efficiency and to reduce investment costs. 8. Ecological evaluation of the ash quality produced using the new combustion technology.

3. METHODOLOGY Altogether, 11 large-scale tests were carried out in a state-of-the-art biomass combustion plant (Lofer, Austria) equipped with a moving grate furnace, a multi cyclone and a flue-gas condensation unit. During the test periods, samples of 7 different ash fractions precipitated at different temperatures between 1,000 and 40°C and of the bio-fuel used

were taken at regular intervals and analyzed for their contents of nutrients and heavy metals. Moreover, the amounts of biomass fired and ashes produced, the ash precipitation temperature, the amount of heat produced, the water content of the bio-fuel and the excess oxygen in the flue gas were measured. Figure 1 shows a scheme of the biomass combustion plant where the test runs were performed. Every test period lasted 3 days. The first day was necessary to adjust the combustion unit to the specified side constraints. After a pre-run of 24 hours the boiler, the combustion unit and the fly-ash precipitation units were carefully cleaned for the subsequent 48-hour test period.

In the autumn of 1996 a series of 8 test runs was performed at the new biomass combustion plant with integrated fractionated heavy metal separation (Stra walchen, Austria). The sampling methodology and the recording of operating data were similar to the ones already described. The different ash samples taken during the test runs in Stra walchen and the respective precipitation temperatures are shown in Fig. 2.

4. DESCRIPTION OF THE NEW COMBUSTION TECHNOLOGY WITH INTEGRATED FRACTIONATED HEAVY METAL SEPARATION Based on the research results achieved, the following points were considered for the design and development of a new combustion technology with integrated fractionated

heavy metal separation [F. Biedermann et al., 1997] (for details see section “Results”): • The amount of bottom ash should be increased as much as possible. • The major part of the fly-ash (particle size ) should be precipitated at temperatures above 850°C. • The precipitation of the fine fly-ash remaining in the flue gas should be carried out in an efficient way to avoid an increase of heavy metal emissions.

598

I. Obernberger and F. Biedermann

These objectives were to be achieved by a new design and process control of the furnace and a new fly-ash precipitation technology (see Fig. 2). A new geometry of the primary combustion chamber decreases the chamber volume over the last grate section and consequently raises the temperature in this section (influence of radiation). Furthermore, a relaxation zone in the primary combustion zone increases the residence time of the flue gas and enhances fly-ash precipitation on the grate. Moreover, a reducing atmosphere is achieved and the entrainment of ash particles decreased by means of section-separated and well-controlled distribution of the primary air over the grate. A second and bigger relaxation zone is situated in the secondary combustion

chamber. Larger fly-ash particles are precipitated there at temperatures between 900 and 1,100 °C. The furnace is equipped with “air staging” technology, which means that the gasification of wood and the combustion of the gases produced take place in different combustion chambers. The reducing atmosphere in the primary combustion zone should both increase heavy metal evaporation and reduce nitric oxide emissions. Furthermore, a high-temperature cyclone was placed behind the furnace. The precipitation temperature in this cyclone varies between 950 and 1,050 °C to guarantee low concentrations of volatile heavy metals (especially Cd) in the ash fraction produced. The usable ash consists of a mixture of bottom ash and hot-precipitated fly-ashes. The fine fly-ash particles leaving the boiler are effectively precipitated in a flue gas condensation unit with an aerosol electrostatic filter installed behind it for low-temperature precipitation (about 40 °C). This filter fly-ash fraction has to be disposed of or industrially utilized. The new technology was implemented in an Austrian biomass combustion plant (nominal boiler capacity ) which was put into operation in the spring of 1996.

Possibilities, Technological Approach, Experiences

599

5. RESULTS—IMPORTANT INFLUENCING VARIABLES FOR FRACTIONATED HEAVY METAL SEPARATION IN BIOMASS COMBUSTION PLANTS The following section discusses and compares the results of the test runs carried out in the state-of-the-art combustion plant (Lofer) and the plant with integrated fractionated heavy metal separation (Straßwalchen). The discussion focuses on the environmentally most relevant heavy metals in bio-fuels and ashes, which are primarily Cd and secondarily Zn.

5.1. Cd and Zn Concentrations in the Different Ash Fractions Sampled— Influence of Temperature on Heavy Metal Behavior Figure 3 and 4 show the average concentrations of Cd and Zn in the different ash fractions examined. The concentrations increase significantly with decreasing temperature of precipitation and decreasing ash particle size. The flue dust contains about 700 times as much Cd and 200 times as much Zn as the bottom ash. The new combustion technology shows a considerable reduction potential for Cd and Zn in the bottom ash. Compared to state-of-the-art combustion plants the Cd concentrations in the bottom ashes produced in the new plant equipped with integrated fractionated heavy metal separation are about 20 times lower. The Cd levels in the combustion zone fly-ash and in the fly-ash precipitated in the high-temperature cyclone are slightly lower than in fly-ashes precipitated at similar conditions in state-of-the-art combustion

plants (see Fig. 3). Furthermore, the investigations showed a significant dependence of the Cd concentrations in hot-precipitated fly-ashes on the temperature of precipitation (see Fig. 5). The Cd concentrations increased with decreasing temperature of ash precipitation. According to high-temperature equilibria calculations, Cd should be completely in the

600

I. Obernberger and F. Biedermann

gaseous phase at temperatures measured in secondary combustion chambers of biomass

combustion plants. At temperatures above 700°C no correlation between temperature and Cd concentrations should be observed under oxidizing conditions [J. Dahl et al., 1997]. This result of equilibrium calculations is in contradiction to the data achieved from the test runs and probably indicates the existence of a non-volatile Cd compound formed by surface reactions of the ash particles with the metal vapors, which was not taken into account in the equilibrium calculations [T. Lind et al., 1997]. This open problem is the subject of ongoing investigations [I. Obernberger et al., 1997b]. The average Zn concentrations in the bottom ash produced in the new plant with integrated fractionated heavy metal separation (Stra walchen) are 5 times lower than in the respective ash fraction from state-of-the-art combustion plants (see Fig. 4). The levels of Zn in the combustion zone fly-ash are lower than in hot-precipitated fly ashes produced in the state-of-the-art combustion plant (Lofer), whereas the Zn concentrations in the fly ash precipitated in the high-temperature cyclone are considerably higher. On an average they even exceed the respective concentrations found in the multi-cyclone fly-ash of the state-of-the-art combustion plant.

Possibilities, Technological Approach, Experiences

601

For hot-precipitated fly-ashes produced in the state-of-the-art combustion plant (Lofer) and for the combustion zone fly-ash (Straßwalchen) the investigations revealed no significant dependence of Zn concentrations on the temperature of precipitation (see Fig. 6). These results show that Zn compounds formed under oxidizing conditions should be in the solid or liquid phase. On the contrary, the Zn concentrations in fly ashes precipitated in the high-temperature cyclone increase considerably with increasing temperatures of precipitation (see Fig. 7). High-temperature equilibria calculations for Zn under oxidizing conditions show that at temperatures below 1,100°C ZnO should be formed, which is partly in the liquid phase (above 600°C) and mainly in the solid phase (see Fig. 8) [J. Dahl et al., 1997]. If the results of these equilibrium calculations and the low-pressure impactor measurements with subsequent particle analysis (see section “Influence of particle size on the heavy metal concentrations in fly-ash fractions”) are considered along with the behavior of Zn in a reducing atmosphere (see section “Influence of the gaseous phase around ash particles on the heavy metal fluxes”), the Zn vapors released in the primary combustion zone should partly condense (forming aerosols) and partly undergo surface reactions with fly-ash particles. These surface reactions, which are not known at present, are likely to increase with temperature. Increasing agglomeration of particles in the high-temperature cyclone (due to sintering and melting processes) could also be a reason for a rise in heavy metal concentrations. The exact behavior of Zn in the hot cyclone is not yet fully understood.

602

I. Obernberger and F. Biedermann

The concentrations of Cd and Zn in the condensation sludge (filter fly-ash) considerably exceed the respective Austrian limiting values for the utilization of biomass ashes on agricultural fields or forest soils. Consequently, this ash fraction has to be collected separately for subsequent disposal or industrial utilization. The highest heavy metal concentrations are found in the flue dust which shows the importance of an efficient dust precipitation technology.

5.2. influence of the Gaseous Phase Around Ash Particles on the Heavy Metal Behavior Figure 9 and 10 show the results of high-temperature equilibria calculations for Cd, Zn and their compounds in a reducing atmosphere (primary combustion zone) for moving grate furnaces burning wood chips. These calculations show that under reducing conditions high amounts of Cd are expected to be volatilized at relatively low temperatures due to the fact that elemental Cd

has a high vapor pressure [I. Obernberger, 1997; J. Dahl, 1995; A. Nordin, 1993; A.D. Tillmann, 1994] (see Fig. 9). Elemental Zn is expected to be volatilized as well, but the volatilization temperatures are higher than for Cd (see Fig. 10).

Possibilities, Technological Approach, Experiences

603

The equilibrium calculations tie in well with the results achieved from the test runs. The levels of Cd and Zn in the bottom ash are very low (see Fig. 11 and 12). This ash fraction is formed under reducing conditions on the grate. Furthermore, the Cd and Zn concentrations in the bottom ash have a significant influence on the combustion air ratio in the primary combustion zone (see Fig. 11 and 12). The concentrations decrease with decreasing primary air ratio (increasing reduction potential).

604

I. Obernberger and F. Biedermann

5.3. Influence of Particle Size on the Heavy Metal Concentrations in Fly-Ash Fractions The average concentrations of Cd in the sieve fractions of fly ashes increase with decreasing particle size (see Table 1). In the smallest sieve fraction the concentrations are three to five times higher than in the fraction. The results show that the particle size (particle surface) is the most important influencing variable for the precipitation of Cd on fly-ash particles at temperatures below 600 to 700°C. Moreover, the formation of aerosols (submicron particles) by condensation of gaseous heavy metal species is important. This was proven by taking aerosol samples from bark- and wood chip-fired combustion plants by means of low-pressure Berner impactors and analyzing the submicron particle fractions for their heavy metal concentrations (see Table 2).

5.4. Amount of Ash Fractions Produced Table 3 shows that state-of-the-art technology and the technology with integrated fractionated heavy metal separation produce similar amounts of bottom ash. Compared to state-of-the-art technology the amounts of fly-ashes produced in the new plant

Possibilities, Technological Approach, Experiences

605

are lower and the amounts of condensation sludge considerably higher (approx. 3 times, see Table 3). This is mainly due to the lower precipitation efficiency of the hightemperature cyclone for smaller particles as compared to multi-cyclones operating at temperatures of about 200 °C. Therefore, larger amounts of condensation sludge (filter fly-ash) have to be disposed of or industrially utilized, which increases the operating costs of the plant and results in a disadvantage of the new technology. In comparison to state-of-the-art technology, the amounts of flue dust produced by the new technology are lower due to the high dust precipitation efficiency of the aerosol electrostatic filter.

5.5. Distribution of Cd and Zn Among the Different Ash Fractions Produced In the bottom ash, which represents the largest ash fraction (see Table 3), only small amounts of Cd and Zn are bound (see Fig. 13). Less than 1% of Cd and only about 3% of Zn were found in the bottom ash of the plant equipped with the new technology. These results are considerably lower than those achieved by state-of-the-art technology, which demonstrates the high Cd and Zn fractionation efficiency of the new technology (especially the effect of a reducing atmosphere). Furthermore, the fly ashes produced in the new plant contained less Cd than the respective ash fraction from the plant equipped with state-of-the-art technology, whereas the amounts of Zn were slightly higher. The condensation sludge (filter fly-ash) precipitated in the new plant contains on an average 4 times as much Cd and about 2.5 times as much Zn as the condensation sludges from state-of-the-art combustion plants. Consequently, an upgrading effect of heavy metals in the filter fly-ash fraction was achieved by using the combustion technology with integrated fractionated heavy metal separation. The heavy metal emissions with the flue dust are lower for the new combustion technology due to efficient dust precipitation effected by an aerosol electrostatic filter installed subsequent to the condensation unit.

606

I. Obernberger and F. Biedermann

5.6. Concentrations of Cd and Zn in the Usable Ash Mixture Table 4 shows the calculation results for the concentrations of Cd and Zn in the usable ash (mixture of bottom ash and fly ashes except the filter fly-ash and the flue dust) for different combustion technologies. The results show that application of the new technology with integrated fractionated heavy metal separation considerably reduces the concentrations of Cd in the usable ash. The average concentrations approach the respective guiding values for soils valid in Austria. The Zn concentrations in the usable ash, however, are almost similar for both

technologies. This is mainly due to the high Zn concentrations in the fly-ash precipitated in the high-temperature cyclone. The Zn levels in the bottom ash are considerably reduced using the new technology (see Fig. 4).

6. CONCLUSIONS The test runs and investigations carried out revealed a high potential of fractionated heavy metal separation in biomass combustion plants by means of primary measures. Depending on where the ash precipitation takes place the following three influencing variables are considered to be of significance (see Table 5): the gaseous atmosphere around ash particles, the temperature of ash precipitation and the size of the fly-ash particles.

The concentrations of Cd and Zn in the ash at high temperatures depend on the reduction potential of the gaseous atmosphere around the ash particles (important for the bottom ash) and on the temperature of precipitation (important for fly ashes precipitated at high temperatures and in an oxidizing atmosphere) [I. Obernberger et al., 1997a]. Consequently, the temperature in the primary combustion zone should be as high as possible and the primary air ratio under-stoichiometric. In an oxidizing atmosphere (in the secondary combustion chamber) the temperature of ash precipitation is an important influencing variable at temperatures above 900°C (according to the experimental

results achieved). As soon as condensation or surface reactions take place (at temperatures below 700–900°C) fractionated heavy metal separation can only be achieved by particle size-selective ash precipitation.

Possibilities, Technological Approach, Experiences

607

At present the development of a combustion technology with integrated fractionated heavy metal separation is at a high level as regards the reduction of Cd and Zn in bottom ashes. Furthermore, the Cd concentrations in hot-precipitated fly-ashes are considerably lower than in fly-ashes generated by multi-cyclones, whereas the potential of hot fly-ash precipitation as regards the fractionation of Zn is rather low. Moreover, the amounts of filter fly-ash increase due to the lower precipitation efficiency of hot cyclones

as compared to the respective performance of multi-cyclones. Therefore, hot cyclones are considered not efficient enough for fractionated heavy metal separation. The further development of a combustion technology with integrated fractionated heavy metal separation will place special emphasis on modifications of the grate and the primary combustion zone. The amounts of bottom ash should be increased with the aim of precipitating the total amount of usable ash on the grate. This could be achieved by

recycling the fly-ashes collected in the multi-cyclone into the primary combustion zone. Appropriate laboratory-scale glowing experiments with fly ashes have already shown the high volatilization potential of heavy metals in contaminated fly-ashes under reducing

conditions and at high temperatures [J. Neuhold, 1995; J. Dahl et al., 1997b].

ACKNOWLEDGMENTS The research presented has been financially supported by the Austrian Fund for Innovation and Technology, by the cooperation of the Federal and the State Governments of Austria for Environmental Research and by the European Commission DG XII within the JOULE III project No. JOR3-CT950001 “Fractionated heavy metal separation in biomass combustion and gasification plants”.

REFERENCES Biedermann F., Obernberger I. (1997). “Possibilities and Efficiency of Fractionated Heavy Metal Separation in Biomass Combustion Plants”. In: Proceedings of the 4th European Conference on Industrial Furnaces and Boilers, April 1997, Porto, Portugal, I N F U B (ed.), Rio Tinto, Portugal. Bundesministerium für Land- und Forstwirtschaft (1997). Der Sachgerechte Einsatz von Pflanzenasche im Wald,

Richtlinie, Bundesministerium für Land- und Forstwirtschaft (ed.), Vienna, Austria. Dahl J., Ljung A., Obernberger I., Nordin A. (1997a). “Potential of heavy metal free ash recirculation, a comparison of experimental findings and chemical equilibrium results”, to be published in: Biomass and Bioenergy.

Dahl J., Obernberger I. (1997b). “Recovery of heavy metals by thermal ash treatment”, in: Obernberger I., Dahl J., Fractionated heavy metal separation in biomass combustion and gasification plants, 3rd six monthly

608

I. Obernberger and F. Biedermann

progress report, JOULE I I I project No. JOR3-CT950001, European Commission DG XII (ed.), Brussels, Belgium. Dahl J. (1995). Thermodynamik Investigation of the Faith of Heavy Metals in Biomass Combustion, diploma work within the ERASMUS- program, Institute of Chemical Engineering (ed.), Technical University of Graz, Austria. Lind T., Valmari T., Kauppinen E., Pyykönen J. (1997). “Fractionated heavy metal separation in biomass combustion and gasification plants”, work packages 5–8, in: Obernberger I., Dahl J., Fractionated heavy metal separation in biomass combustion and gasification plants, 3rd six monthly progress report, JOULE III project No. JOR3-CT950001, European Commission DG XII (ed.), Brussels, Belgium. Neuhold J. (1995). Laboruntersuchungen zur thermischen Abtrennung der Schwermetalle Cd, Zn, und Pb von Flugaschen aus Biomassefeuerungen, diploma work, Institute of Chemical Engineering (ed.), Technical University of Graz, Austria. Nordin A. (1993). On the Chemistry of Combustion and Gasification of Biomass Fuels, Peat and Waste, Environmental Aspects, Department of Inorganic Chemistry (ed.), University of Umeå, Sweden. Obernberger I., (1997). Nutzung fester Biomasse in Verbrennungsanlagen unter besonderer Berücksichtigung des

Verhaltens aschebildender Elemente, Schriftenreihe “Thermische Biomassenutzung”, Vol. 1, Institute of Chemical Engineering (ed.), Technical University Graz, Austria, dbv- Verlag, ISBN 3-7041-0241-5. Obernberger I. (1995). “Characterisation and Utilization of Ashes from Biomass Combustion Plants”, in: Pro-

ceedings of the Earth Conference on Biomass for Energy, Development and the Environment, Jan. 1995, Havanna, Cuba, EUROSOLAR (ed.), Bonn, Germany. Obernberger I., Biedermann F., Kohlbach W. (1997a). FRACTIO—Fraktionierte Schwermetallabscheidung in Biomasseheizwerken, final report, Institute of Chemical Engineering (ed.), Technical University Graz, Austria. Obernberger I., Dahl J. (1997b). Fractionated heavy metal separation in biomass combustion and gasification

plants, (first) twelve monthly progress report, JOULE III project No. JOR3-CT950001, European Commission DG XII (ed.), Brussels, Belgium. Salzburger Landesregierung (1997). Richtlinien für die Aufbringung von Asche aus Biomassefeuerungen auf landwirtschaftliche Böden (Asche RL), guideline, Salzburger Landesregierung (ed.), Salzburg, Austria. Tillmann D.A. (1994). Trace Metals in Combustion Systems, Academic Press, Enserch Environmental, Sacramento, California ISBN 0-12-691265-3.

DISTRIBUTIONS OF MAJOR AND TRACE ELEMENTS IN ENTRAINED SLAGGING COAL GASIFICATION PROCESS Kenichi Fujii1, Masamitsu Suda2, Tadayoshi Muramatsu3, and Masahiro Hara4 1

Kawasaki Heavy Industries 1-1, Kawasaki-cho, Akashi, 673, Japan 2 Chube Electric Power Company 3 Electric Power Development Company 4 Center for Coal Utilization, Japan

1. INTRODUCTION Out of concern for the global environment, technologies for coal-fired combined cycle generation are under development throughout the world today, aiming at highly efficient use of coal. Under these circumstances,the author and others [Hirao, 1993; Fujii, 1995] proposed a new gasification combined cycle generation system witch differs from the conventional IGCC system. We studied system configuration and conducted bench scale test. The R&D has been conducted between Center for Coal Utilization, Japan, Kawasaki Heavy Industries, Chubu Electric Power Company, and Electric Power Development Company, under the sponsorship of the Ministry of International Trade and Industry. Today, R&D is underway of pilot plant at Wakamatsu works of EPDC [Harada,1997]. This report deals with the basic concepts of the system and the results of bench scale test, in which distributions of major and trace elements were investigated experimentally and estimated using a computer simulation.

2. SYSTEM CONCEPT The mechanism of the system which the author and others are developing is shown in Fig. 1. Coal is partially burned (gasification) to separate and remove the ash Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

609

610

K. Fujii et al.

as molten slag. Dust char in the product gas is then removed by hot precipitator (approximately 700C). The gas is burned as-is in the gas turbine combustor. Finally at the gas turbine outlet, waste heat is recovered and the flue gas is denitrified and desulfurized. In comparison with the conventional IGCC system, the present system has the following features. 1) High plant efficiency

In making up combined cycles, the temperature (700C) of the gas introduced in the gas turbine is higher than in the conventional IGCC system (400C). This results in high plant efficiency. 2) Low plant cost and high reliability The product gas is not desulfurized before being introduced into the gas turbine. Desulfurization is carried out at the outlet of HRSG, using a technically proven, conventional large-capacity desulfurizer. This system’s environmental performance and reli-

ability are quite comparable to those of conventional state-of-the-art pulverized coal firing plant. Implementation of this system, however, requires the solution of technical problems not posed by conventional systems. These major technical problems include the alkali metal vapor contained in the product gas. This vapor is expected to increase owing

to the use of dedusting high-temperature gas. Also included is the problem of turbine blade hot corrosion resulting from the increased sulfur content in gases entering turbine which comes from locating the desulfurizer at the gas turbine outlet.

Distributions of Major and Trace Elements in Entrained Slagging Coal Gasification Process

611

To study these problems, we conducted the basic tests using 7T-coal/d CPC bench scale test plant. Combustion residues such as slag, char and bag-filter ash were analyzed carefully. This report describes the above-mentioned bench scale test results, and computer modeling of distributions of major and trace elements from CPC process.

3. STRUCTURE OF CPC Figure 2 shows the conceptual structure of the CPC. The basic concept is a twostage dry-feed entrained-flow gasifier. CPC is composed primarilyof the precombustor, the CPC main body, the slag-free duct and the reductor. The precombustor has a pulverized coal burner at the top, in which mixture of primary coal and recycled char are supplied downward together with air, so that the coal is burned in a fuel-rich gasification condition. Air is also fed from the side of the precombustor, to enhance combustion gas

temperature. High temperature partial combustion gas is then blown tangentially into the CPC main body, where coal is gasified mainly. Most of the coal ash is captured on the surface of molten slag, and discharged through the slag hole. The water-cooled selfcoating wall protects the CPC main body from high temperature gas. The slag free duct

is made of water-cooled flat panel to prevent adhesion of fly slag. Secondary coal is supplied in the middle of the slag free duct to lower product gas temperature quickly and raise the heating value of the product gas.

4. TEST FACILITY AND PROCEDURE Figure 3 shows the schematic flow diagram of the bench scale PCPC test facility. It comprises compressor, pulverized coal storage, lock hopper type pulverized coal feeder,

612

K. Fujii et al.

pressure vessel containing CPC and reductor, product gas cooler, cyclone, char recycle feeder, and many other auxiliary components. The test system provides a coal capacity of 7t/d, design pressure of 1.1 MPa, and normal operating pressure of 0.4 MPa. The normal operating pressure of the test facilities was selected much lower than the estimated actual operating pressure (2 to 3 MPa).

5. TEST RESULTS

5.1. Distribution of Major and Trace Elements Table 1 shows the total supply of coal and the total discharge of slag, char and bag-filter ash as a combustion residue after a long hours operation (101 hours). Mass balance analysis indicated that the ash recovery is 85% into slag, 4% into char, 10% into bag-filter ash and 1% into flue gas (chimney). 92% of char was recycled to gasifier for further gasification reaction and the reduction of char discharge (4% is the final discharge). Figure 4 shows the distribution of major and trace elements in residues and flue gas. Mass-balance study indicate that elements may be partitioned Into four broad group as is summarized in Table 2.

Distributions of Major and Trace Elements in Entrained Slagging Coal Gasification Process

613

5.2. Leaching Characteristics of the Residue It is necessary to understand the leaching characteristics of slag or ash from CPC system in order to ascertain the potential environmental consequences of dumping residues. Table 3 shows the result of leaching test of the CPC residues. Mineral content except for cadmium were detected. Nonetheless, almost of leachates were below detection limit. Leachates of bag filter ash in As and Se were only detected. Levels of leachates were just like the ground water limit value, less than the limit value of landfill disposal, demonstrating the safety of disposal.

614

K. Fujii et al.

5.3. Carbon Conversion Efficiency Figure 5 shows carbon conversion efficiency in case of two stage coal supply. The value of carbon conversion efficiency is a function of coal feed ratio (primary/secondary), total coal feed (stoichiometry), oxygen concentration, coal reactivity (coal brand), temperature, pressure and gasifier configuration.

6. MODELING 6.1. Vapor Pressure Figure 6 shows saturated vapor pressure of trace elements, regulated in Japan. These six elements except for Cr are vaporized easily under the moderate temperature. Table 4 shows mineral content of coal feed. As the content is small and the vapor pressure is high, fully vaporization temperature (saturation point) is extreamly lower than gasifier temperature above 1,500C, indicating easy gasification in a gasifier.

Distributions of Major and Trace Elements in Entrained Slagging Coal Gasification Process

615

6.2. Simulation Model In order to obtain the distribution characteristics of major and trace element in the CPC bench test facility, computer modeling was then examined. Outline scheme of simulation is shown in Fig. 7. Simulation model is composed of gasifier, reductor, cyclone, product gas combustor, bag filter and chimney. The model includes many operational parameter such as coal feed ratio (primary/secondary), temperature profiles, ash removal efficiency and its recycling ratio. Computer modeling

calculate the quantities of solid-state or gaseous elements passing each equipment. The indicated figure is an example in case of 50% vaporization element.

6.3. Experimental Prediction Figure 8 shows computer simulation results on the distribution of elements from the CPC bench scale test facility. The recovery of stable ash is in good agreement with the test results, shown in table 1. The horizontal axis refers to the extent of vaporization

in gasifier, which correspond to variations of elements, which is shown in Fig. 4. As the vaporization in gasifier is increased, from group 1 (weak vaporization) to group 2

616

K. Fujii et al.

Distributions of Major and Trace Elements in Entrained Slagging Coal Gasification Process

617

(medium vaporization), slag recovery shows steady reduction. Recovery of char and bagfilter ash show peak profile in case of medium vaporization group. In case of strong vaporization group, slag recovery is decreased and gas recovery becomes dominant.

Figure 9 shows the enrichment factor of slag, char and bag-filter ash. Each element’s enrichment factor is defined by (element recovery)/(stable ash recovery). As the vaporization in gasifier is increased, slag enrichment factor shows simple reduction, and enrichment factor of char as well as bag-filter ash shows peak profile. Peak profile is caused by the increase of recycle char, which is shown in Fig. 8.

6.4. Influence of Recycling Figure 10 shows ash recovery without char recycling. In this case, slag recovery decreases in proportion to vaporization ratio in gasifier. In most elements, recovery of

618

K. Fujii et al.

char and bag-filter ash exceed 50% like a fly-ash from pulverized coal firing boiler. Nonrecycling of char would increase the discharge of black char, which is difficult to be disposed without further treatment. Figure 11 shows ash recovery with fully recycling of char and bag-filter ash. In this extreme case, the quantities of recycling elements (especially, medium vaporizing elements) increase tremendously. It might cause another problem such as hot corrosion via Na, K, V, S, Cl. In consequence, slag recovery as well as gaseous recovery increases. Medium vaporize elements including Hg, Se, As, Cd, Pb might cause gaseous environmental impact.

CONCLUSIONS Distributions of major and trace element in entrained slagging coal gasification process were investigated experimentally with the CPC bench scale test plant and the computer modeling. The following conclusions were obtained. ( 1 ) Mass-balance analysis indicates that major and trace elements from CPC process may be partitioned into four broad group.

Group 1: stable mineral elements (less volatile) Ca, Mg, Si, Al, Ti, V, Fe Group 2: partially vaporize elements (medium volatile) Na, K, Cu, Nn, Ni, Cr Group 3: easy vaporize elements (more volatile) Pb, Zn, As, F, Cl, Hg Group 4: fully vaporize elements (combustible) S, C, N, H

Distributions of Major and Trace Elements in Entrained Slagging Coal Gasification Process

619

(2) Although coal contains some quantities of heavy metals, slag contains a negligible quantity except for Cr. Leachates from slag are extreamly small and less than national regulation limit in landfill disposal. (3) Computer modeling explained the transformational behavior of elements in gas and solid phases. (4) Recycling of residues gives large influences to the distribution of elements in gaseous and solid phases.

ACKNOWLEDGEMENTS Financial support of this research and development by the NEDO (New Energy and Industrial Technology Development Organization) and MITI (Ministry of International Trade and Industry) is greatly appreciated.

REFERENCES Clarke, L.B. (1991). Management of by-products from IGCC power generation. IEA Coal Research, London. Fujii, K. et al. (1995). “Development of Pressurized Coal Partial Combustor: Characteristics of 7T/d Gasifier with Various Coal-Feed Ratio.” Bulletin of JSME (in Japanese), 61(587), 350–3355. Harada, E. et al. (1997). “Research and Development of New Coal Based Combined Cycle Power Plant Concept.” International Conference on Power Engineering, 69–74.

Hirao, M. et al. (1993). “Development of Coal Partial Combustor for High Efficiency Power Generation.” International Conference on Power Engineering.

This page intentionally left blank.

BEHAVIOR OF INORGANIC MATERIALS DURING PULVERIZED COAL COMBUSTION Tsuyoshi Teramae, Toru Yamashita, and Takashi Ando Coal Research Laboratories Idemitsu Kosan Co., Ltd. 3-1 Nakasode, Sodegaura, Chiba, Japan Center for Coal Utilization, Japan (CCUJ), 6-2-31 Roppongi, Minato-ku, Tokyo, Japan

2. INTRODUCTION The behavior of ash at high temperature directly affects the performance and the operability of boiler in the coal utilization technologies such as pulverized coal combustion and coal gasification. Although the mechanism of ash formation, adhesion and deposition in a boiler is important subject, there are many points which must be solved. This study is being carried out by Center for Coal Utilization, Japan (CCUJ) as a part of the NEDO’s program related to the development of fundamental technologies for coal utilization. CCUJ started the study focused on the ash formation mechanism at the first step in order to construct the model of ash behavior. In this study, the change of composition, size and shape of individual ash particle under combustion were tracked by CCSEM (Computer Controlled Scanning Elec-tron Microscopy) to examine the thermal behavior of inorganics as they transformed into fly ash. Two coals, namely, Coal A (Australian coal) and Coal B (Japanese coal) were used for combustion tests which were carried out in a vertical type turbulent flow testing furnace. Chars and fly ash were sampled during combustion for analysis. The elements of Ca and Fe were focused in the CCSEM data analysis, because theses elements were particularly considered to affect the behavior of melting and coalescence. A computer simulation technique of coal combustion has been developed for more than ten years at Idemitsu Coal Research Laboratories [Satou et al., 1993]. The goal is to incorporate the thermal behavior model of inorganics into computer simulation code. Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al.

Kluwer Academic / Plenum Publishers, New York, 1999.

621

622

T. Teramae et al.

3. EXPERIMENTAL Coal A (Australian coal) and Coal B (Japanese coal) were used in the study. The two coals are different in ash composition and melting temperature. The coal properties for combustion tests are shown in Table 1. Combustion tests were conducted on the vertical type turbulent flow testing furnace (300mm in diameter and 2,500mm long ) in which self sustained combustion of 5kg/hr of pulverized coal was designed. A schematic of the furnace is shown in Fig. 1. Pulverized coal (–200 mesh, 80%) was fed from a single burner with of primary air. Secondary air was preheated to 350°C and introduced to the furnace. Two stage combustion air was also preheated to 350°C and introduced to the furnace at the distance, 1,660mm, from the burner. The furnace was equipped with 17 sampling ports every 15cm distance on the furnace wall. The ports are termed SP1, SP2,

SP3, . . . . . . , SP17 from top to bottom. Particles during combustion were collected by the sampling probes through SP3, SP4, SP5, SP6, SP8, SP11, SP14 and SP17. The distance of the ports from the burner were 560mm(SP3), 680mm(SP4), 820tnm(SP5), 940mm(SP6), l,180mm(SP8), l,560mm(SPl1), 2,000mm(SP14) and 2,600mm(SP17), respectively. Fly ash was collected in a cyclone.

Behavior of Inorganic Materials During Pulverized Coal Combustion

623

Unburned carbon in the chars and the fly ash sampled from each port was measured to grasp the combustion rate. Fig. 2 shows the change in the rate of unburned carbon in Coal A and Coal B. The rate of unburned carbon at “3,000mm” in the figure means the value of fly ash. The rate of unburned carbon of fly ash were 1.66% and 0.0% for Coal A and Coal B, respectively. The decreasing rate of unburned carbon in Coal B was larger than that in Coal A. It is considered that Coal B has more combustibility. The characteristics of raw coal (–200mesh, 80%), 3 chars (sampled from SP3, SP4 and SP6) and fly ash were analyzed by CCSEM. As unburned carbon of 3 chars from SP3, SP4 and SP6 decreased rapidly, it was considered that the samples were suitable for the investigation of ash formation during combustion. Data analysis related to the composition was focused on Si, Al, Ca and Fe. As shown in Table 2, each particle was classified into eight categories from compositional criteria. The weight proportion of each category was examined how it changed during combustion.

624

T. Teramae et al.

4. RESULTS AND DISCUSSION

4.1. Changes in Composition The changes of the ratio of each category during combustion are shown in Fig. 3 and Fig. 4. The weight proportion of calcium aluminosilicates (Si-Al-Ca) particles increased while Ca and Fe particles decreased in Coal A during combustion. The increase of Si-Al-Ca particles and decrease of Ca and Fe particles are also observed in Coal B, however, aluminosilicates (Si-Al) decreased and iron calcium aluminosilicates (Si-Al-CaFe) increased significantly. Si-Al-Ca and Si-Al-Fe ternary diagrams of Coal B are shown in Fig. 5 and Fig. 6. The dense region of the plots in the figures indicate the presence of many particles. Most of the plots existed near Ca and along the line between Si and Al in the figure. Minerals in the coal were considered to be mainly Quarts, Kaolinite and Calcite. As combustion progresses, the plots mostly appeared on the line between Ca and the point, 50% of Si and Al. It was clearly seen that Si-Al-Ca particles were produced under combustion. The

Behavior of Inorganic Materials During Pulverized Coal Combustion

625

plots mostly appear on the line from Fe to the line between Si and Al in the Si-Al-Fe ternary diagram of Coal B in the same manner as Si-Al-Ca diagram. This suggests that Quartz or aluminosilicates (Kaolinite etc.) in raw coal react Ca and Fe particles to produce Si-Al-Ca particles and Si-Al-Fe particles. It is considered that

626

T. Teramae et al.

Behavior of Inorganic Materials During Pulverized Coal Combustion

627

628

T. Teramae et al.

Si-Al-Ca-Fe particles were further produced under combustion. Fig. 7 shows a simplified model of the production pass to the Si-Al-Ca, Si-Al-Fe and Si-Al-Ca-Fe particles. Although it seems that there is more complicated route of the reaction of inorganic particles, it is decided here to assume a simple model. Figure 8 shows the relationship between the total rate of the formed particles (SiAl-Ca particles, Si-Al-Ca-Fe particles and Si-Al-Ca-Fe particles) and the combustion rate (combustion rate = 100 – unburned carbon). The values of the ordinate are the weight proportion of these particles against the total particles of Si, Si-Al, Si-Al-Ca, Fe, Si-AlFe and Si-Al-Ca-Fe. As it can be seen from Fig. 8, the reaction of particles with Si, Al,

Behavior of Inorganic Materials During Pulverized Coal Combustion

629

630

T. Teramae et al.

Ca and Fe proceeded during combustion, and the rate of the particles formed were larger in Coal B. The regression curves and equations are also shown in Fig. 8. Figure 9 shows the relationship between the weight proportion of the each category (Si-Al-Ca particles, Si-Al-Fe particles and Si-Al-Ca-Fe particles) and the combustion. The values of the ordinate are the weight proportion of the each category against the total weight of the three categories. It is seen in the figure that the rate of the Si-AlCa particles increased under combustion. The rate of Si-Al-Ca-Fe particles also increased while the rate of Si-Al-Fe particles decreased. The slopes of the curves of Coal A and Coal B in the figure are almost same.

4.2. Changes in Particle Size Particle size distributions of inorganic particles in raw coal, three chars and fly ash by CCSEM are shown in Fig. 10 and Fig. 11. Particles of about 3mm increased while coarse particles over 10mm decreased under combustion of Coal A. In Coal B, particles under 5mm and over 30mm decreased and particles of about 10mm increased. The size

range of increased particles was obviously different between Coal A and Coal B. In order to examine the changes of the particle size distribution in detail, particle size distribution of each category was shown in Fig. 12 and Fig. 13. The categories with

higher content were shown in the figures, because they affected the particle size distribution with all particles. As the Fig. 12 suggests that the behavior of Si-Al particles greatly affects the particle size distribution of Coal A. Particles of about 3mm increased while

coarse particles over 10mm decreased. It is considered that the fragmentation occurred during combustion. Ca particles are also shifted to the fine particle size. The contents of Si-Al, Si-Al-Ca and Si-Al-Ca-Fe particles in Coal B were large and affected to the overall particle size distribution. The Si-Al particles under 5mm decreased and particle content of 10 ~ 20mm was large in fly ash. The increase of Si-Al-Ca and Si-Al-Ca-Fe particles of 5 ~ 20mm was significant in Coal B. The cause of difference in the behavior of particle size distributions between Coal

Behavior of Inorganic Materials During Pulverized Coal Combustion

631

A and Coal B can be found in the particle size distribution change of Si-Al particles and the production rate of Si-Al-Ca and Si-Al-Ca-Fe particles. Figure 14 and 15 shows the weight proportion of included and excluded particles by particle size classifications. Raw Coal B has more included particles (52.0wt.%) than

632

T. Teramae et al.

Behavior of Inorganic Materials During Pulverized Coal Combustion

633

Coal A (36.3wt.%). Although the amount of included particles decreased under combustion of the two coals, the degree of decrease in Coal B is larger. Included particles of 23.2 wt.% in Coal A remained in Char 6. As it is described above, the increased particle size range of fly ash formed under combustion in Coal B is larger than Coal A. It is considered that the higher proportion of included particles and combustibility of Coal B promoted the reaction of Si-Al particles and Ca, Fe particles as compared to Coal A.

5. CONCLUSIONS From the analysis of CCSEM data of raw coal, chars and fly ash of Coal A and Coal B, the results regarding the behavior of inorganic particles during coal combustion was obtained as follows. 1) Aluminosilicates (Si-Al) particles such as Kaolinite react Ca and Fe particles to produce calcium aluminosilicates (Si-Al-Ca) particles and iron calcium aluminosilicates (Si-Al-Ca-Fe) particles. The production rate of these particles varies according to the coal type and the combustion rate. 2) The weight proportion of the fine particle region and of the coarse particle region decreased under combustion while the proportion of intermediate particle size between them increased. The increasing and decreasing behavior concerning weight proportion of size range also differed according to coal type. Fly ash obtained from the combustion of Coal B had coarser particles than that of Coal A. This was affected by the amount of productions of calcium aluminosilicates (Si-Al-Ca) and iron calcium aluminosilicates (Si-Al-Ca-Fe) particles.

634

T. Teramae et al.

3) The proportion of included particles in the raw Coal B was much more than

that in Coal A and Coal B was also more combustible. It is considered that the higher proportion of included particles and combustibility promoted the reaction of Si-Al particles and Ca, Fe particles and produced coarser fly ash

particles as compared to Coal A.

6. ACKNOWLEDGEMENTS This paper is published with permission of NEDO. The authors gratefully acknowledge the financial support of NEDO.

7. REFERENCE Satou, M., Tominaga, T. and Shinozaki, S. (1993). “Prediction of the flow, Combustion and Heat Transfer in a Industrial-Scale Boiler.” Proceedings of JSME-ASME International Conference on Power Engineering93, 337–341.

ENERGY PRODUCTION FROM CONTAMINATED BIOMASS Progress of On-Going Collaboration Projects*

Alexandra Grebenkov, Anatoli Iakoushev1, Larry Baxter2, Dave Allen3, Helle Junker4, and Jørn Roed5 1

Institute of Power Engineering Problems, Minsk, Belarus CRF/Sandia National Laboratories, Livermore, CA, USA 3 Wheelabrator Environmental Systems Inc., Anderson, CA, USA 4 ELSAMPROJEKT A/S, Fredericia, Denmark 5 RISØ National Laboratory, Roskilde, Denmark 2

1. INTRODUCTION The catastrophe at the Chernobyl NPP in the Ukraine resulted in radioactive fallout in many European countries and represents the single largest release of radioactive material in recorded history. The amount of radionuclide released from Chernobyl exceeds the sum of that released by the World War II atomic bombs used against Japan by more than a factor of 200. In Belarus, radionuclides have substantially contaminated 27 towns and 2,736 settlements with more than 2.1 million inhabitants, 17.5 thousand sq.km of arable lands, meadows and pastures, and 19 thousand sq.km of forests. According to the Belarus National Chernobyl Programme, about 15% of the State budget is allocated annually to elimination of consequences of the accident. In spite of this effort, social and industrial activity in contaminated regions are decreasing. Total control over the use of contaminated resources is impossible. Uncontrolled natural and anthropogenic transfer of radionuclides into other portions of the ecosystem occur by forest fires and habitual consumption of local wood by rural population for heating and cooking. Local populations also customarily collect game and food (mushrooms and berries) from the surrounding areas. While the government of Belarus strives to educate the population about the hazards of such activities, economic and social conditions are such that they *—Presented at the Engineering Foundation Conference on the Impact of Mineral Impurities in Solid Fuel Combustion, Nov. 2–7, 1997, Keauhou Beach Hotel, Kona, Hawaii. —Work is supported beelarator Environmental Systems Inc., the U.S. Department of Energy, the Governments of Belarus and Denmark Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

635

636

A. Grebenkov et al.

continue. In general, logs from contaminated forest used by wood processing industries sooner or later end up in a fire of municipal furnaces and domestic stoves without appropriate release control and ash handling. More than 200 thousand tons of radioactive ash

with specific 137Cs activity exceeding 10,000Bq/kg have been already collected and dispersed as a fertilizer by residents of contaminated regions. At present the Republic of Belarus suffers from an energy crisis connected with lack of indigenous fossil fuel resources. In the contaminated Gomel Province, where the most energy intensive industries are located, energy demand is higher than supply. In addition, about of the installed capacity that is operated with natural gas and fuel oil has reached the end of its designed life-time and should be rehabilitated. This power could be replaced with energy of bio-resources which would provide a pathway for remediation of affected territories. Belarus is ideally suited to bioenergy due to large area of productive industrial forest, flat landscape, well developed power distribution and district heating infrastructure, and a technically competent society. The energy potential represented by the contaminated Chernobyl area wood is approximately 25-PJ/year. Our proposed action is construction and operation of the biomass-fired power facilities, including the harvest of appropriate biomass resources. At the same time, radioactive contamination imposes certain limitations for the technologies applied that must be designed in view of financial, ecological and dose cost. In 1995 Sandia National Laboratory (SNLs) and Institute of Power Engineering Problems (IPEP) initiated a project to assess the economic, technological, and ecological aspects of conversion of biomass into heat and power in Belarus, with an emphasis on the contaminated regions. This project is funded by US Department of Energy through the Initiatives for Proliferation Prevention Program and by Wheelabrator Environmental Systems, Inc. The US government support for the project is intended to provide meaningful technical work to former scientists and engineers engaged in developing weapons of mass destruction as a means of preventing the proliferation of such technology to other countries. The US Government role is designed to be phased out as the project matures to a profitable enterprise through a three-phase process: a feasibility study (Phase 1), pilot-scale testing (Phase 2), and commercialization of the project on a basis of demonstration-scale facilities (Phase 3).

2. THE RESULTS OF PHASE 1 AND SCOPE OF PHASE 2 The first stage has been completed with the following principal outcomes: • The contaminated forests can be used as sustainable renewable fuel resources

which can be safely valorized to generate steam and electrical power along with significant remediation of contaminated lands; • The most feasible and cost effective option is related to power generation with heat co-generation that would allow to achieve an internal rate of return of more than 30—40%, and pay-back period less than 5 years. In case of direct electricity production this economic parameters are also very attractive for investment (Fig. 1,2); • Doses to personnel and local community are evaluated to be within acceptable range;

Energy Production from Contaminated Biomass

637

• Laboratory simulations, investigation of parameters in pilot-scale, commercialscale simulation and specified economic and risk analysis are required to obtain all necessary data to gain international consensus that the proposed action is feasible, safe, ecologically sound, and leads to remediation effect for forestry and forest ecosystems. The second stage deals with the following studies: • Study of combustion properties of forest litter, formation of radionuclides aerosols in combustion systems, secondary waste management (reduction of ash volume, disposal options), dose formation; • Safety analysis and risk assessment; • Specification of cost-benefit analysis of the applied actions, evaluation of the Belarus market and formulation of the commercial proposals; • Evaluation of site and size of demonstration or/and industrial scale plant.

To perform this stage IPEP committed in co-operation with SNLs and Wheelabrator Environmental Systems, Inc. to constructing and testing a fully functional, electric power-producing pilot facility consisted of the following main components:

638

A. Grebenkov et al.

• Gasifier of 100 kW designed and produced by IPEP; • Furnace of 3MW coupled with steam boiler and economizer designed by IPEP; • Power generating system based on suction gas diesel/ICE, steam-turbine generator of 400 kW, or turbine simulator; • Ash removal system designed by IPEP. The ash encapsulation pilot plant developed by Brookhaven National Laboratory is foreseen to be incorporated to the above facility; • Radionuclide release control system. Heat and power generated will be connected with local heating/electric grid.

3. BENCH-SCALE EXPERIMENTS AND MODELLING The bench-scale experiment was carried out at the Combustion Research Facilities/SNLs in order to testify the aerosol modelling and combustion properties of forest litter and duff. About 1.5 tons of actual litter, duff and wood were sampled in Belarus and sent to Sandia for this experiment. Two sample plots were selected to represent fuel samples from two main types of pine forest landscape and pine forest ecosystems in Belarus. Both plots are located in Berezinsky Leskhoz, about 65 miles north-east from Minsk city, on lowland and hillock surface of the first terrace of Berezina river. Both are characterized with wood stand of artificial origin, which consists of Pinus silvestris L of 40–60 years old, II–III bonitet, 16–18m height. There is intensive up-grow of pine trees on clearings without any other tree species. The ultimate and proximate analysis were carried out by IPEP to characterize these fuel materials. According to the results of modelling the contaminated biomass fired industrial facility must be provided with 99.95% of radionuclides collection efficiency for exhaust control system as a minimum. The particle control device for conceptual boiler design is proposed to be a baghouse with high-temperature fabric which has high durability, temperature tolerance, and collection performance. Similar product is available in Belarus. To determine the health effect caused by radionuclide release from a boiler the computer model is developed by IPEP which describes the relevant radionuclides and their

migration in atmosphere and in natural ecosystems. This allows to forecast of dynamics of ecological situation near any possible contamination aerosol sources of thermal origin. The logical structure of the computer code LOCMIGR intended for the description of dynamics of migration processes in actual situations is as follows: • formation of concentration fields inside of potential aerosols source with the following effluent its in atmosphere; • plume dispersion of an accidental effluent and near surface transfer of aerosols;

Computer code LOCMIGR is constructed on the basis of an evolution model of parameters description of transport flow and transported dispersible impurity phase. In

the basis of the given model the systems of conservation equations for separate phases are fixed which are numerically resolved together with the equations describing processes of interphase transfer and dynamics of interphase surfaces. The code will be verified and applied for evaluation of formation of concentration fields inside a potential source of

Energy Production from Contaminated Biomass

639

aerosols with the consequent affluent its in atmosphere, plume blurring of incident

affluent, near surface transfer of aerosols and its deposition on an enclosing relief.

4. SIMULATION AT THE INDUSTRIAL SCALE FACILITY In April 1997, a test burn was completed at the Shasta Power Plant operated by Wheelabrator Environmental Systems in northern California. Wheelabrator, SNL and IPEP personnel conducted the test burn with the primary objectives of demonstrating combustion properties of forest litter and duff at commercial scale evaluation of stack emissions. During the test the boiler rated was fired with normal wood fuel mixed with litter and duff collected in northern California. The composition of the combustion materials was similar to the composition of litter, duff and wood from the Chernobyl zone in Belarus, except for the radioactive contaminants. Two combinations, i.e. 20% and 50% of forest floor matter fed together with normal fuel, were tested. To simulate the 137 Cs and 90Sr radionuclides, solutions containing known concentrations of stable isotopes of Cs and Sr in the form of cesium sulfate and strontium nitrate were added to the fuel as it was fed to the boiler. Partitioning of the Cs and Sr, as well as the other inorganics and hydrocarbons, was determined in bottom and fly ash and in aerosols after the electrostatic precipitator in exhaust gas. During the test, the content of and were continuously monitored and recorded. The early results are as follows: • Fly ash mass fraction constitutes about 30% of total residues;

• Well over 99% of the cesium and strontium is retained in one of the ash streams; • Overall efficiency of the precipitator is at least 99.2% for Cs, the most volatile of the species. Cs concentrations in the flue gas were below detection limits using conventional techniques and are currently being investigated using advanced techniques. • concentrations increased slightly during combustion, but remained within permited limts using conventional measures (ammonia injection); • CO levels decreased notably during combustion of the test fuel; • Significant increases in ash loading on the grate and in the remaining streams were noted; These results allow us to assume that a boiler of conventional type can effectively burn the new biofuel material, and radionuclides emission does not exceed the permissible concentration.

5. DOSE AND RISK ANALYSIS The control of the radiological impact of the consequences of the Chernobyl accident relies largely on restrictions. Some laws have been set up which determined the status of the areas of radiation contamination, as well as the living conditions of the population. Most of the important activities such as the forest industry in Belarus suffer from these limitations. The question is the management of a situation which is less critical but more complex than it was shortly after the accident. The actually necessary measures aiming to decrease the essential part of radiological risk have been implemented. For the

640

A. Grebenkov et al.

time being the long term consequences of the accident affect more social and economic sectors which are not easy to precise. There is a need for efficient actions that may alleviate this burden. One possible option in this regard is the proposed project that involves the contaminated biomass into conversion process to obtain under the controlled conditions heat and electricity along with removal the source terms from forest environment. The objective of this task is to assess the potential risk to human health and the environment posed by combustion of radioactively contaminated biomass. Six potential scenarios were investigated according to the most probable exposure pathways: Scenario 1: Exposure due to routine forestry and recreation: • External exposure from trees and ground; • Inhalation of airborne dust; • Ingestion from forest food pathway.

Scenario 2: Exposure posed by domestic use of firewood: • Direct external exposure from a stove and hearth; • External exposure from ash residues handled; • Inhalation from a chimney and while handling hearth ash.

Scenario 2a: Exposure posed by use of ash as a fertilizer: • External exposure from a garden; • Ingestion from garden products pathway; • Groundwater/surface water pathway; Scenario 3: Exposure due to forest fire • External exposure from deposition; • Inhalation of airborne particles; • Ingestion from food chain.

Scenario 4: Exposure due to active forestry: • External exposure from trees and ground; • Inhalation of airborne dust; Scenario 5: Exposure from biomass fired boiler: • Direct external exposure from facility; • Inhalation from a chimney.

Scenario 6: Exposure from ash residues management and disposal: • External exposure from collection, packaging (immobilization) and transportation procedures; • External exposure from disposal facility.

Energy Production from Contaminated Biomass

641

The scenarios # 1–3 are the baseline option (“do nothing” option). The important issue is that in case of scenarios 2, 2a and 3 it is almost impossible to provide control of exposure and apply protection measures. The scenarios # 4–6 represent the active option when the proposed action is applied. To provide comprehensive analysis IPEP has collected all the relevant data needed. The following codes were used for dose and risk assessment: FORESTPATH or FORESTLIFE (developed by IPEP = others): • Prognosis of radionuclides migration in forest ecosystem; • Scenarios 1 and 4.

Universal Monte Carlo code MCNP4A (provided by LANL, modified by IPEP):

• External exposure dose rate; • Scenarios 1–6. LOCMIGR (developed by IPEP):

• Transport of radionuclides released from furnace and forest fire; • Scenarios 2, 3 and 5. RESRAD (provided by BNL):

• Inhalation dose from the 1 micron particles and ingestion dose; • Scenarios 1, 2a, 3 and 4. DOZA (developed by IREP)

• Multi-cameral model for inhalation dose; • Scenarios 1, 2, 3, 4, and 5 The quantitative risk analysis currently is being performed on a basis of the data obtained and the codes developed. The preliminary results of dose evaluation are as

follows: Exposure from routine forestry and recreation: • Forestry worker—3–6mSv/year (ext), 0.2–1.5mSv/year (int); • Resident—1.9–2.5 (ext), 0.2–0.5 (int). Exposure from domestic use of firewood: • Resident—<0.1 mSv/year (ext); <0.01 mSv/year (int)

Exposure posed by use of ash as a fertilizer: • Resident—<0.01 mSv/year (ext); <0.01 mSv/year (int).

642

A. Grebenkov et al.

Exposure from forest fire (excluding background exposure):

• Worker (fireman)—<0.1 mSv/year (ext); <0.1 mSv/year (int). Exposure from biomass fired boiler and ash management: • Worker—is to be less than 1 mSv/year (total); • Resident—is to be less than 0.1 mSv/year (total).

CONCLUSION The results obtained in this on-going research allow to assume the following:

• • • • •

A boiler of conventional type can effectively burn the new biofuel. Radionuclides emission may not exceed the permissible concentration. Ash residues can be effectively stabilized and handled. Proposed action will reduce risk and doses to population. Use of contaminated forest will lead to termination of its deterioration and improve social activity of local communities.

The following study is still to be carried out: • Aerosol/fly ash particles partitioning from any actual biomass fired boiler operation. This will help to address the specification of release control and radioactive

influence on surrounding area. • Advanced commercial biomass fired boiler design that can be used as demonstration scale facility. For a long term objective this is necessary in order to provide cost-benefit analysis of the commercial scale valorization of contaminated biomass. Boiler layout and boiler construction materials are also necessary to estimate exposure from installation in case of use of radioactive wood fuel for assessment of doses to personnel. • Feasibility of commercial use of a pilot plant. Financial analysis with minimum of uncertainties is necessary to ensure the real operation cost and to reduce cost of future decommissioning. • Radionuclides deposits on the boiler surfaces. It is very important to define in what form the radionuclides’ species are able to set down and to be fixed on the hot surfaces of boiler (grate, pipes, walls, etc.). This will help to address the adequate countermeasures while operation, and decontamination technologies while decommissioning. In order to utilize available expertize and to be able to share activities, but also to assure a reliable success of this very large and important project, in the beginning of 1997 ELSAMPROJEKT in co-operation with RISOE National Laboratory initiated supplementary project and submitted it to Danish Environment Protection Agency. In August, 1997 this project was approved by Danish Government and funding is open. The major task of the international project group now is effective co-ordination of joint efforts.

Energy Production from Contaminated Biomass

643

Both projects have been reviewed by the relevant Belarus Governmental Bodies, i.e. Ministry of Energy Savings & Supervision, Ministry of Emergencies, Ministry of Forestry, and Ministry of Environment Protection & Natural Resources. They have endorsed the strategy of contaminated wood valorization and approved the scope of the projects. Preliminary discussions have been held with local authorities and with prospective industrial clients for the purchase of power from biomass-fired plant. The competitive locations of a demonstration scale power plant (co-generation scheme) are basically

defined.

This page intentionally left blank.

TRIBOELECTROSTATIC COAL CLEANING Mineral Matter Rejection In-Line Between Pulverizers and Burners at a Utility

John M. Stencel, John L. Schaefer, Heng Ban, TianXiang Li, and James K. Neathery Center for Applied Energy Research, University of Kentucky 2540 Research Park Drive, Lexington, KY 40511 Telephone: (606) 257-0250 FAX: (606) 257-0302 email: [email protected]

INTRODUCTION With escalating demands on coal quality to increase boiler efficiency and decrease acid gas emissions, fewer coals can meet specifications without some form of beneficiation. Removal of the coal mineral matter prior to combustion can have multiple benefits. Reduction of the aluminosilicate clay minerals, which typically comprise 60–90% of the total coal mineral matter [O’Gorman, 1972], will produce a corresponding decrease in boiler erosion and fouling and will also decrease the amount of fly and bottom ash generated. Removing the pyritic minerals will decrease the emissions and lower or eliminate flue gas scrubbing requirements. The extent to which mineral matter is removed from coal by physical cleaning is dependent on the extent to which the coal is comminuted. In the US, coal is typically delivered to PC boiler sites having –3/4 inch size after which it is pulverized before combustion with about 70% passing a 200 mesh screen. Besides liberating mineral matter from the combustibles, such pulverization also liberates sulfur and nitrogen compounds from combustibles. If it were possible to take advantage of this liberation, power plant efficiencies may increase while the cost of operations and environmental impacts could decrease. A possible technique that could be applied to take advantage of the coal pulverization practised at all PC plants is triboelectrostatic beneficiation. It can beneficiate pulverized coal into carbon (combustibles) and mineral matter by using an electric field to divert the differentially charged particles [Gidaspow, 1987; Ciccu, 1989; Carta, 1970; Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

645

646

J. M. Stencel et al.

Alfano, 1985; Ban, 1994a; Ban, 1994b]. Differential charging is a consequence of particle-particle and particle-wall collisions and is dependent on differences in the surface properties of the constituent particles to be beneficiated [Lockhart, 1984]. The first known application of tribolectrostatic separation was described in the patent literature [Edison, 1892]. More recent research on triboelectrification and electrostatic separation has evaluated key scientific and engineering parameters that influence the triboelectrostatic processing of coal [Ban, 1994a; Ban, 1994b; Schaefer, 1996]. However, very little dry beneficiation research and development has been performed in comparison to the amount of effort in wet beneficiation research and development. Because most coal utility operations use dry coal and ash handling equipment, efficient and cost effective dry beneficiation systems may be expected to be more applicable to power production than are wet beneficiation systems. It is known that coal pulverization, in concert with the pneumatic transport of coal within burner pipes, imparts particle charge [Nieh, 1987]. However, the amount of this charge, its dependence on coal type and pulverizer conditions, and the polarity of the charge on combustible or mineral matter particles have not been measured within a utility operation. In laboratory settings, combustibles in bituminous coals attain a positive charge whereas mineral matter attains a negative charge [Ban, 1994a]. In-line (between pulverizers and burners), triboelectrostatic coal beneficiation has

yet to be demonstrated within a utility setting. Because no known published information was available on this subject, we initiated a study which measured the magnitude of particle charge and the extent to which mineral matter is liberated from combustibles within pulverized coal as it exits a pulverizer at two, coal-fired utility boilers. Complementary triboelectrostatic beneficiation testing was also performed using a laboratory-scale apparatus. Data from these utility and laboratory tests are presented and discussed herein.

EXPERIMENTAL The East Kentucky Power Cooperative (EKP—Spurlock station) and the Tennessee Valley Authority (TVA—Widows Creek station) offered the use of their facilities for this

study. At each station, sampling ports were installed in a burner pipe between a pulverizer and burner of the boilers. Each port was 3.8cm in diameter and included a ball valve and a compression fitting into which the tip of two sample probes (see Fig. 1 and Fig. 2) could be secured. The sampling probes enabled the extraction of pulverized coal samples from various positions across the diameter of the burner pipes. Although isokinetic sampling protocol was not strictly followed, an estimation of the flow velocities entering the probe tips relative to the flow velocities in the burner pipes suggested that isokinetic conditions were approximately attained. With one of the probes, the charge inherent to the coal/mineral matter particles within the burner pipe was measured. With the other probe, electrostatic separation experiments were performed to examine whether differential charge had been established on the particles; this probe also separated the combustibles and mineral matter within the pulverized coal stream. Figure 1 depicts the charge measurement probe. Its design was based on the Faraday cup principle and consisted of a sampling tube and a filter housing, electrically connected and screened from stray electrical and magnetic signals by a grounded shield. Between the grounded shield and the copper inner wall was a PVC dielectric. Pulverized coal from the burner pipe entered the tube through a hole at its tip with the assistance of suction created by an air eductor located at the air outlet of the filter housing. Because

Triboelectrostatic Coal Cleaning

647

the tube was electrically connected to the filter housing section, i.e., was also a part of the Faraday circuit, all charged particles entering the probe and captured within the bag filter contributed to the current measured by a picoammeter. The current measured by the picoammeter was recorded with the use of a computerized data acquisition system. Integrating the current over the time during which each sample was collected yielded a total charge for the captured pulverized coal particles. After a sample was collected, the filter bag was removed from the probe and sealed in a plastic container. The weight of these collected samples was measured upon returning to the laboratory where a small amount of each sample was submitted for proximate and ultimate analyses. By combining the total charge for the collected sample and its weight, a charge/mass (C/kg) ratio was calculated for each sample.

648

J. M. Stencel et al.

Figure 2 depicts the electrostatic separation probe used at the utility sites. It provided data on the presence of differential charge on the combustibles and mineral matter and gave an indicator of the beneficiation potential of the coal as it was pneumatically transported within the burner pipes. The separation chamber of the probe consisted of two removable, parallel copper plates. Pulverized coal flowed horizontally between and along the length of these plates and either became attached to the plates or was captured in a filter bag attached at the end of the plates. The outlet of the filter bag was connected to an air eductor which was used either to create a suction from the tube tip or a positive purge of air out of the tube tip. The analytical separator used during the laboratory experiments and its operation have been described elsewhere [Ban, 1994b]. Using the analytical separator, the beneficiation of combustibles and mineral matter in coal was quantified by collecting samples from the negative and positive plates of a parallel plate separator, and from a filter at the bottom of the separator. Because sample deposition on the plates was continuous, each plate was arbitrarily divided into four sections from which samples were collected. Weighing each of these samples and submitting them for proximate and ultimate analyses enabled data to be obtained which was plotted in a format of combustible recovery-versus-ash content. These data can be considered to represent a dry separation recovery curve, the results of which are compared to results from the in-line experiments at the utility sites.

RESULTS AND DISCUSSION Table 1 summarizes the charge results for both sites. While the Spurlock Station burner pipe was actually sampled at two locations, only results from the lower sample port are presented. Since this sample port was immediately above the pulverizer, it should allow for direct comparison to the Widows Creek results which were collected directly above the exhauster. At both sites, five different probe penetration distances were evaluated for particle charge. Preliminary tests were conducted at both sites to determine the optimum sample collection time. For the Spurlock Station this time turned out to be around 200 seconds while at Widows Creek samples were collected for around 60 seconds. Both pulverization systems showed little variation in the radial profile of average charge. The absolute value of average charge was slightly lower for Spurlock Station, compared to for Widows Creek. The charge polarity for the two sites was reversed.

The average charge represents a net value, indicating that at least one component of the pulverized coal is charging. In theory, the charge on the ash could be equal but opposite in polarity to the charge on the combustibles, yielding a net charge of zero for the pulverized coal samples. Even with zero net charge, a high degree of mineralcombustible beneficiation could be achieved. Many factors can affect the charge magnitude and polarity, including coal mineralogic and petrographic composition, charging surface composition and morphology, and charging conditions. Sampling procedures may also have an effect on the coal charge.

While the charge polarity was reversed at the two sites, the significant result from the charge measurement experiments was that both pulverization systems produced similar absolute values of charge. To determine whether that charge was sufficient to achieve combustible—mineral matter beneficiation, in-situ tests at the utility sites were performed. Table 2 summarizes results from these tests.

Triboelectrostatic Coal Cleaning

649

The total sample weights extracted from the burner pipes at the Spurlock Station tests were erratic, possibly due to roping inside the pipe. For the Widows Creek plant, a lower percentage of material was collected on the electrodes and a correspondingly higher

percentage of material was collected in the filter. Potential explanations for this finding include: insufficient particle charge; insufficient electric potential across the electrodes; or inadequate separator residence time. Considering the left electrode material to be reject, and the filter and right electrode to be cleaned product, ash removal ranged from 43–51% for Spurlock Station and 26–41% for Widows Creek. Combustible recovery ranged from 76–80% at the Spurlock Station and 81–86% for Widows Creek. In general, these results show that while the Widows Creek pulverized coal possessed a higher average charge, the separability was not as good as at Spurlock Station. During field test at both sites, pulverized coals were also extracted from the burner pipes for subsequent aboratory testing. Figure 3 shows the ash content as a function of combustible recovery for the laboratory tests. These tests were performed on a bench scale triboelectrostatic separator using an electric field strength of 2kV/cm. Comparing the laboratory results to the in-plant results at equivalent electric field strengths shows: for a 76% combustible recovery, the product mineral content was 8.0% and 7.7% for the Spurlock Station field and lab tests, respectively, and; for an 80% combustible recovery, the product mineral content was 11.4% and 10.2% for the Widows Creek field and lab tests, respectively. Therefore, the field tests, which were not optimized, produced separation results in very close agreement with those from the laboratory tests.

650

J. M. Stencel et al.

CONCLUSIONS This research has shown that both pressurized and suction configurations for utility pulverizers impart substantial and differential charge on mineral matter and combustible material in coal. The absolute value of the average charge per unit mass was the same order of magnitude for both utility systems and was in agreement with values obtained within a laboratory system. This differential charge can be utilized to beneficiate combustibles and mineral matter in-line between the pulverizers and burners under conditions typical to that of coal transport in the utility. While these results indicate the potential for on-site power plant pulverized coal beneficiation, further work is required to assess the process. Towards that end, and based on results from the power plant test series, a facility to simulate mineral mattercombustible beneficiation under conditions similar to those expected in burner pipes is under construction.

ACKNOWLEDGMENTS Partial funding from the USDOE, Grant DE-PS22-94PC94225, is gratefully acknowledged. Likewise, assistance from East Kentucky Power and the Tennessee Valley

Triboelectrostatic Coal Cleaning

651

Authority, in particular personnel at the Spurlock Station and Widows Creek Power Plants, is greatly appreciated.

REFERENCES Alfano, G.P., Carbini, M., Carta, M., Ciccu, R., Del Fal, C., Peretti, R., and Zucca, A. (1985). “Application of Static Electricity in Coal and Ore Beneficiation”, J. Electrostatics, 16, 315. Ban, H. (1994a). “Experimental Study of Particulate Charge Relating to Dry Coal Cleaning”, University of Kentucky, PhD Dissertation.

Ban, H., Schaefer, J.L., Saito, K., and Stencel, J.M. (1994b). “Particle Tribocharging Relating to Electrostatic Dry Coal Cleaning”, Fuel, 73, 1108. Carta, M., Ferrara, G., Del Fal, C., and Ciccu, R. (1970). “Process and Apparatus for Electrostatically Separating Ores”, US Patent 3, 493, 109. Ciccu, R., Peretti, R., Serci, A., Tamanini, M., and Zucca, A.(1989). “Experimental Study on Triboelectric Charging of Mineral Particles”, J. Electrostatics, 23, 157. Edison, T.A., (1892). “Method of and Apparatus for Separating Ores”, US Patent No. 476, 991. Gidaspow, D., Gupta, A., Mukherjee, A., and Wasan, D.T. (1987). Processing and Utilization of High Sulfur Coals II, Chugh, Y.P. and Caudle, R.D., editors, Elsevier, New York, pp. 271–281.

Lockhart, N.C. (1984). “Dry Beneficiation of Coal,” Powder Technology, 40, 17. Nieh, S., and Nguyen, T. (1987). “Measurement and Control of Electrostatic Charge on Pulverized Coal in a Pneumatic Pipeline”, Particle Science Technology, 5, 115. O’Gorman, J.V., Walker, P.E. (1972). “Mineral Matter and Trace Elements in U.S. Coals”, US Dept. Interior, Office of Coal Research and Development, Report No. 61.

Schaefer, J.L., Ban, H., Saito, K., Neathery, J.K., and Stencel, J.M. (1996). “Pulverized Induced Charge: Comparison of Utility Configurations”, 13th International Pittsburgh Coal Conference, 2, 873.

This page intentionally left blank.

DEVELOPMENT OF ADVANCED PFBC Makoto Takai and Masahiro Hara

Thermal Power Department, Electric Power Development Co., Ltd. 6-15-1 Ginza, Chuo-ku, Tokyo, 104 Japan Technical Development Department, Center for Coal Utilization, Japan 6-2-3 Roppongi, Minato-ku, Tokyo, 106 Japan

1. INTRODUCTION While the coal fired power plant releases relatively more greenhouse gases than other fossil fuels, coal is still the superior energy resource due to the world-wide supply potential and its cost advantages. In order to make the best use of a coal extensively hereafter, improvement of plant thermal efficiency is essential. This is the most realistic method for reduction of greenhouse gases. A-PFBC system has the advantages of higher plant efficiency, with potential of 46% HHV (net plant efficiency). This efficiency is about 10% (relative value) higher than PFBC, and also exceeds IGCC. Furthermore, A-PFBC system can utilize the R&D results on PFBC and IGCC projects.

Present status of A-PFBC development in Japan are as follows; — Basic research and conceptual design of large scale A-PFBC systems have been already done by CCUJ in 1992–1995.

— A-PFBC PDU (Process Development Unit) Test Project has been started by EPDC/CCUJ in 1996, and PDU test would be proceeded from end of 2000 to 2002. — PDU basic plan was completed in 1996, while it has been carrying out the

elementary test and site survey for detail design of PDU in 1997. This paper describes the development history, concept of series type A-PFBC system, and outline of PDU test project.

2. DEVELOPMENT HISTORY A-PFBC power generation technology has been considered and developed in various ways as the means to further improve heat efficiency. In other words, raise the inlet temperature of the gasturbine (G/T) while maintaining the features of PFBC power generation technology. Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

653

654

M. Takai and M. Hara

Various systems have been proposed as A-PFBC. In Japan, combining an oxidizer and a partial gasifier in series with both bubbling pressurized Fluidized bed was considered to be optimum. The concept and development process of the A-PFBC power generation technology will be described below.

2.1. Developing the Concept Regarding the rise of PFBC temperature to improve the heat efficiency, it is difficult to raise the temperature more than the current level (850°C) due to the generation of alkali vapor and the fusion of ash. As a possible solution, the concept of “topping” using such fuels as LNG has been proposed. However, if such a fuel is used, its supply must be based nearby and the combustion rate of other fuels are higher than coal, thus making this concept impractical. The A-PFBC (parallel type) power generation then tried coal partial gasification synthesis gas burning in topping combustor as its topping fuel. Because this type of power generation can secure high steam conditions and high efficiency, the Southern Company Services in the U.S. has adapted it as its Power System Development Facility (PSDF). Although British Coal in the U.K. was the first in the world to propose topping

combustion technology in the form of parallel type A-PFBC power generation, it is currently aiming for the so-called oxidization independent type. In this case, no flue gases from the oxidizer is introduced into the topping combustor. So, heat from the oxidizer is recovered only in the form of steam and thus, its efficiency is lower than the abovementioned parallel type and the series type described later. In Japan, as a further advanced system, the A-PFBC (series type) combustion technology has been proposed. One advantage of the A-PFBC (series type) is that the topping combustor necessary for the A-PFBC (parallel type) is not needed and basically the same

type G/T combustor for the current LNG c/c or IGCC can be used, and also the gas purification system can be formed in a simplified single line. The designation of this system (where the inlet temperature of the G/T is increased by combining the PFBC with a partial gasifier) varies depending on the country. Examples are shown below. UK: US: Germany: Japan:

Topping Cycle, Air Blown Gasification Cycle A-PFBC (advanced PFBC), 2nd Generation PFBC Hybrid Combined Cycle A-PFBC

In each country, research and development on A-PFBC has progressed to the point where it is widely known that A-PFBC enables higher efficiency than IGCC due to the raised temperature of the gas turbine and the improved steam conditions. Figure 1 shows the development history of A-PFBC.

2.2. Development History of Related Technologies A-PFBC is ideally positioned to make use of the R&D results and know-how being obtained in past and on going PFBC and IGCC projects, such as knowledge on equipment design, reaction behavior, trouble countermeasures and so on. Thus, it could be

Development of Advanced PFBC

655

expected to shorten the lead-up time of development compared with the development on quite new technologies. Figure 2 shows the transition of developed technologies related

to the A-PFBC system.

3. SYSTEM DESCRIPTION (A-PFBC SERIES TYPE) The advantage of this A-PFBC (series type) is that no topping combustor is needed

and gas purification system can be simplified (single system and relatively low temperature) and the higher plant efficiency. Figure 3 shows the basic configuration flow of the A-PFBC. The A-PFBC uses

PFBC as a combustor for unburnt carbon (char) in from a partial gasifier. The hightemperature gas of the PFBC charges the partial gasifier to make effective use of the thermal energy (sensible heat) inherent in the combustion gas as a heat source for the partial gasifier. Thus, the amount of coal charged to the partial gasifier can be reduced by the corresponding amount. Fuel gas discharged from the partial gasifier is desulfurized by a high-temperature desulfurizer using limestone and then cooled by a synthesis gas cooler. Thus, the gas cooler becomes a low corrosive gas atmosphere allowing high-temperature vapors to be recovered.

The cooled fuel gas then undergoes dust removal using a ceramic filter or other similar devices and is fed to a gas turbine to generate power.

656

M. Takai and M. Hara

Development of Advanced PFBC

657

3.1. A-PFBC Features This system comprises a pressurized fluidized bed (PFBC) and a fluidized bed type coal partial gasifier. Coal is gasified in a fluidized simplified gasifier and unburnt carbon (char) generated in the gasifier is completely burnt in the PFBC oxidizer. This system contributes to the gasification reaction by introducing this high-temperature combustion

gas into the partial gasifier. Table 1 shows the major features of A-PFBC System.

3.2. Results of Conceptual Designs Table 2 shows the major results of A-PFBC Conceptual Designs.

3.3. Reasons for High Efficiency The gross thermal efficiency can reach as high as 49.7% as shown above. The major reasons are described below.

658

M. Takai and M. Hara

— Raised inlet temperature of gas G/T: Improvement of G/T unit thermal efficiency. — Improvement of the steam condition: Because the limestone desulfurization method is applied as well as the PFBC, it is possible to recover high-temperature steam by a gas cooler. Limestone desulfurization allows desulfurization prior to the gas cooler so that its usage condition is under low corrosive atmosphere. Thus, the high-temperature steam can be recovered. — Utilization of limestone desulfurization method: Dry-type produces smaller loss compared to wet-type. This dry gasification desulfurization method produces no consumption loss of H2 caused (reduction of iron oxide when S is absorbed) in general iron oxide desulfurization (Nakoso, etc.) — Reduction of ash sensible heat loss: Dry ash is discharged to allow heat recovery.

3.4. Energy Balance Figure 4 shows the energy balance (at the time of 100% load) resulting from the

previously described conceptual design.

4. A-PFBC PDU PROJECT

4.1. Purpose The purpose is to verify the engineered practicability of a process in which partial gasifier/desulfurizer and oxidizer are combined.

Development of Advanced PFBC

659

4.2. Scale Coal charging amount: 15 tons/day class (5 MWt class)

4.3. PDU System Flow Figure 5 shows the PDU (Process Development Unit) System Flow.

4.4. Test Verification Items PDU test will be done for the purpose of verifying the various items as shown in

Table 3.

4.5. Schedule PDU test has been scheduled as shown in Table 4.

4.6. Suitable Coal for A-PFBC A-PFBC system is composed of the Fluidized Bed technology, thus it is preferable to use the coal which ash fusion temperature is relatively high, as not to generating

660

M. Takai and M. Hara

agglomerates. Suitable coals for IGCC (Entrained Flow) have relatively low ash fusion temperatures. The combination of both of the Entrained Flow and Fluidized Bed Technology allows use of various coal types. As for reference, Figure 6 and Figure 7 shows the correlation concerning clinker formation.

CONCLUSION — In order to make the best use of coal extensively hereafter, improvement of heat efficiency is essential and the most realistic method for reduction of greenhouse gases. — The A-PFBC is an effective system from the point of view of high efficiency. — Since no new problems have been found regarding the components of the system, it is highly realizable. — However, its performance has not been verified. Therefore, the practicability of the system should be evaluated and considered using the such tests as the PDU test. — PDU test will be done from end of 2000 to 2002. — In order to make use of various kind of coal widely, it needs both technologies (Fluidized Bed and Entrained Flow).

Development of Advanced PFBC

661

This page intentionally left blank

OPERATING EXPERIENCES OF 71MW PFBC DEMONSTRATION PLANT Hideki Goto and Syoichi Okutani Wakamatsu Coal Utilization Technology Research Center, Electric Power Development Co., Ltd.

1 Yanagasaki-machi, Wakamatsu-ku, Kitakyushu 808-01, Japan Technical Development Dept. Center for Coal Utilization, Japan 6-2-31 Roppongi, Minato-ku, Tokyo 106, Japan

1. INTRODUCTION PFBC system is an attractive technology for coal utilization due to the potential of

a higher electricity efficiency of combined cycle plants. Recently, research and development on PFBC system has been intensively proceeded in the world. Demonstration and/or commercial plants are now operated at Värtan in Sweden, Escatron in Spain, and

Tomato-Azuma and Wakamatsu in Japan. PFBC plant at Wakamatsu is a 71 MWe combined cycle plant consisting of 56.2 MWe steam turbine and 14.8 MWe gas turbine. The plant started generating in October

1993, as the fifth 70MW-class PFBC plant following Värtan P4, P5, Tidd and Escatron. After one year tuning operation and MITI inspection, we started demonstration program in January 1995. Different from other PFBC plants, ceramic tube filters (CTF)—hot gas filters—of 100% capacity are installed at Wakamatsu plant. The plant also has the secondary cyclones which are changeable to CTF system. We have operated the plant for 4,500 hours with CTF system and for 4,900 hours with 2 stage cyclone system. Now we are operating the plant with CTF for evaluating the characteristics of PFBC boiler and the performance of CTF system as well.

2. PLANT DESCRIPTION Wakamatsu PFBC plant is a combined cycle plant with steam turbine cycle and gas turbine cycle. Steam turbine is driven by the steam generated in the fluidized-bed combustion boiler which is installed in the pressure vessel and gas turbine is driven by the high-pressure exhaust gas from the boiler. Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

663

664

H. Goto and S. Okutani

Steam turbine cycle is similar to conventional power plant’s steam cycle. An ultrahigh temperature steam condition is used for efficient generation; 593 °C of main steam and reheat steam temperature. In the boiler, a mixture of crushed coal and limestone is fluidized and burned

efficiently with pressurized air. The boiler is composed with waterwall and evaporator tubes, primary- secondary- tertiary super heater tubes and reheater tubes are installed in the fluidizing zone. Crushed coal is mixed with limestone and water to the fuel paste, and the paste is injected into the boiler, just under the tubes, by the plunger pumps. Lime-

stone is fed into the boiler both for desulfurization and for bed material supply to keep the bed height. Approximately, the desulfurization rate in the combustor is more than 90%, so that scrubber or de-SOx facility is not necessary after the boiler. A boiler, seven primary cyclones, seven secondary stage cyclones and two bed material storage bins are installed in the pressure vessel. The tall and slender type pressure vessel is specially designed for Wakamatsu Plant. The boiler in the lower half of the vessel is connected with cyclones in the upper half by exhaust gas manifold pipe. Because of

the manifold pipe, gas and dusts are equalized to each cyclone. Figure 1 shows the installation in the pressure vessel and major specifications are as follows; — Boiler

Type:

Pressurized fluidized bed

Once-through Boiler (Bubbling type) Steam flow: Combustion Pressure: Bed Temperature: Rated Bed Height:

147 tons/hour (54001b./min) 1.1 MPa (142psig) 860°C (1580°F) 3.5m (11.5ft)

Fuels: Size:

Bituminous Coal, Paste feed W 7.1/D 4.3/H 7.9m (W 23.3/D 14.1/H 25.9ft)

— Pressure Vessel

Design Pressure: 1.5 MPa (213 psig) Design Temperature: 350°C (662°F) Size: ID 11/H 29.5m, t = 65mm (ID 36.1/H 96.8ft, t = 2.56 inch) — Gas Turbine Type:

Open-cycle, Ruggedized type

Rated Output:

Double-shaft 14.8MW

Inlet Gas Condition:

1.0 MPa (128 psig)

Rotor Speed:

HP 6,137 rpm LP 3,400–5,650 rpm

822°C(1512°F)

— Steam Turbine (existed) Type:

Tandem Compound

Rated Output: Steam Condition:

56.2 MW 10.2/2.7MPa (1465/365psig)

Rotor Speed:

3,600rpm

Double Flow 593/593°C (1100/1100°F)

Operating Experiences of 71MW PFBC Demonstration Plant

— Cyclones Number: 7 cyclones Inlet Gas Temperature: 860°C (1580°F) Inlet Gas speed: 30m/sec (98.4 ft/sec) Dust (in/out): (22/1.1 ppmw) — CTF Systems Number:

2 units

Inlet Gas Temperature: 850°C (1562°F) Number of Tubes: 243 tubes per unit Dust (in/out) (2.3/0.8 ppmw).

665

666

H. Goto and S. Okutani

A gas turbine drives generator and compressors which supply air to the boiler for combustion and fluidization. Exhaust gas from the boiler goes through the primary cyclones and CTF (or secondary cyclones) consequently to reduce the dust, and afterward into the gas turbine. The gas turbine at Wakamatsu has two shafts, HPT and LPT

Operating Experiences of 71MW PFBC Demonstration Plant

667

shafts; HPT drives a generator and HP compressor at constant speed whereas LPT drives LP compressor at variable speed to control air flow. Two CTF units are installed for dust collecting due to the extremely lower erosion damages to the gas turbine blades. Each CTF unit consists of three compartments and each compartment has 81 tubes inside. Exhaust gas flows down from top to bottom, from inside to outside of the tubes, in the meantime, dusts are collected inside of the tubes and clean gas goes out from the CTF unit. High pressure air is periodically blown to the

outside of the tubes in order to remove the collected dust down to the ash hopper. System diagram flow is shown in Fig. 2.

3. DEMONSTRATION PROGRAM Demonstration program started in January of 1995 and is divided into three phases. The purpose of each phase is following;

Phase 1: Stable Operation of the Plant Phase 2: Analysis of Operational Characteristics Phase 3: Establishment of Plant Reliability.

The schedule is delayed because of unexpected troubles. Planned schedule and actual is shown in Table 1. Operation mode and major tests from January of 1995 through December of 1997 are as follows; — January to April, 1995:

Operation with CTF system Boiler characteristics measurement Environmental characteristics measurement Coal changing test (Coal: # A, #B) — May to October, 1995: Operation with 2 stage cyclones Boiler characteristics measurement Environmental characteristics measurement

Coal changing test (Coal: #C, #D)

— October to December, 1995: Annual Inspection

668

H. Goto and S. Okutani

— December, 1995 to October, 1996: Operation with CTF system Plant reliability improvement Utility reduction test CTF operation improvement Start-up fuel injection improvement Coal changing test (Coal: # B, #D) — October to December, 1996: Annual Inspection — January to June, 1997: Operation with 2 stage cyclones (Coal: #B, #E) — July to December, 1997: Operation with CTF system (planned, Coal: # B, another)

4. OPERATION RESULTS Wakamatsu PFBC plant started generation on October 8th in 1993. After tuning operation and MITI inspection, we started demonstration and evaluation in January, 1995. We have operated for more than 10,900 hours since the first generation. The major operation results are as follows (as of August 31st, 1997: incl. operation with 2 stage cyclones);

— Operation Hours: — Total Generation: — Number of Start-up: — Coal Consumption:

9,448 hours 424,756 MWh 90 186,540 tons

— Coal Variation:

5

Because the plant is for demonstration, we have more inspections than commercial plant and also have more troubles and malfunctions; these caused lower availability

factor. Especially, at the beginning of tuning operation, many troubles caused fewer operation hours. In 1995, we improved the facilities and operation, but availability factor stayed at the level of 15% because of 1.5 month annual inspection, 1 month legal inspection of facility and 1 month work on gas turbine. In 1996, the availability factor was 19% for the same reason. We have burned five kinds of coal with variety of characteristics, e.g., high sulfur contents and low sulfur contents, high volatile and low volatile, and etc. Characteristics of coals are shown in Table 2.

4.1. Plant Efficiency Gross plant efficiency is about 37% and net efficiency is about 35%; relatively efficient, considering the plant scale and old existing steam turbine. Actual efficiency or

heat rate is shown in Table 3.

Operating Experiences of 71MW PFBC Demonstration Plant

669

4.2. Combustion Efficiency High combustion efficiency is expected in the PFBC boiler because of slow

combustion in the bed under high pressure. Actual combustion efficiency at Wakamatsu is as high as or higher than expected and stable between coals. Table 4 shows actual combustion efficiency by coals.

4.3. Environmental Characteristics One of the great advantage of the PFBC system is superior environmental characteristics with lower sulfur dioxide emissions and lower nitrogen oxide emissions. Sulfur dioxide emissions may be reduced in contact with limestone in the fluidizing bed. In the atmospheric fluidized bed combustor, calcines to form CaO and

and followed by sulfation of CaO:

670

H. Goto and S. Okutani

In the PFBC boiler, it is considered that dioxide at high pressure:

must be sulfated directly by sulfur

At Wakamatsu, both reactions (1) and (2) are estimated to occur in the PFBC boiler because a certain amount of CaO is found in the bed material and the fly ash according

to the analysis. Actual sulfur dioxide emissions are much less than 136ppm that is the value of the agreements, and desulfurization ratio is more than 90%. Reaction (1) and/or (2) indicates that more limestone makes less sulfur dioxide emissions. Table 5 shows the

difference of SOx emissions according to Ca/S molar ratio when burning #C coal. Nitrogen oxide emissions are suppressed because of the low combustion temperature; at 860°C of combustion temperature, i.e. bed temperature, thermal NOx cannot be generated but fuel NOx and prompt NOx is only generated. Actual NOx level is 100–200 ppm at the outlet of the boiler and 30–50 ppm at the outlet of SCR at 100% load. Dust or fly ash is captured by cyclones and CTF, so that dust density is controlled under at the stack. Actual dust density at the stack is and dust capture ratios are 93–97% at the primary cyclones, 98.5–99.5% at CTF and more than 99.9% in total.

4.4. Material Balance (Ash Balance) PFBC ash, composed with coal ash and limestone, is discharged from the boiler as bed ash and fly ash. Bed ash is removed from the bottom of the boiler when the amount of bed material is surplus to maintain the bed height. Most of the fly ash is captured by the primary cyclones (cyclone ash) and the rest is collected by CTF system (CTF ash). The percentage of bed ash to total ash varies from 10 to 30% in accordance with coal variation, limestone variation and Ca/S molar ratio.

Operating Experiences of 71MW PFBC Demonstration Plant

671

Example of material balance is shown in Table 6, Table 7 shows the average particle size and coal ash/limestone composition of ash. Conception drawing of gas and ash flow is shown in Figure 3.

4.5. CTF Performance The plant has been operated with ceramic tube filters for more than 4,500 hours to confirm the dust collection performance. Dust density is at the inlet of CTF

and

at the outlet; dust collecting efficiency is 98.5–99.5%. CTF units are

placed upstream of the gas turbine, so that dust density at the gas turbine is also 1–3 and allows to use the clean gas-type gas turbine. CTF collects the dust by filtration on the inner surface of the tubes when gas flows from inside to outside of the tubes. This filtration mechanism causes gas pressure drop; 30 kPa of design pressure drop. Actual pressure drop is about 24.5 kPa and stable at 100% load. Table 8 shows the CTF performance—dust collection efficiency and pressure drop.

5. MAJOR TROUBLES We had some troubles since we started the plant operation. Major troubles and its

countermeasures are listed below by components.

5.1. Boiler Spurge air pipes which supply air for fluidizing and firing from the bottom of the boiler were cracked by thermal stress. After the trouble, the design was changed to reduce the thermal stress.

672

H. Goto and S. Okutani

Stab tube at the outlet header of the primary super heater was cracked and broken by thermal stress. Connecting pipe works were changed after the trouble.

5.2. Fuel Supply System Fuel paste plug in the pipe and the nozzles occasionally occurred at the beginning of the tuning operation because of the bad paste quality. To improve the paste quality, we added the circulation line to the fuel preparation system to crush coal sufficiently. We also made the operational improvements to control the paste quality and to make the initial fuel injection during start-up successful, as stated in the latter chapter.

5.3. Cyclones Ash plug at the bottom of the primary cyclones occurred at low load because of insufficient gas flow for ash transportation. The bypass lines were installed to keep sufficient gas flow.

Ash plug in the ash removal pipe of the secondary cyclones occurred when burning #

C coal. The reason of this is considered that the deposit of sub-micron ash particle

burned and melted itself at the suction of the removal pipe. The design of the suction

nozzle of ash removal pipe was changed from L-type to straight type to prevent the ash deposit at the elbow of the nozzle.

5.4. CTF Filtration tubes were broken five times; the first three times at the beginning of tuning operation and the last two times in 1996. The countermeasures to the tube breakage of the first three times were as follows.

— Quality control of each tube by testing the strength, — Proper cooling water flow in the baffle plates to prevent the plate bending, — Reduced thermal stress of the tube near the baffle plates, — Improvement of the self-circulating blow-down system to keep the gas down

flow constant in the tube. The countermeasures are effective, but we suffered another kind of tube breakage in 1996. We executed the test of the operational improvement, as stated in the latter chapter.

5.5. Gas Turbine Strained stress increased at the LP turbine blades at resonance speed in summer 1995. The reason of this is considered below; — Increased gas stimulus to the blade because of the closing tendency of the inlet guide vanes,

— Dust density exceeded the standard code, — Unproper change of blade arrangement, — Decreased dumping factor at the blade wedge.

Operating Experiences of 71MW PFBC Demonstration Plant

673

LP turbine blades were replaced to increase the strength. HP turbine blades were eroded by the ash during 2 stage cyclone operation with #C coal. The secondary cyclones could not capture the sub-micron particles included in #C coal fly ash, so that too much dust flew to the gas turbine. It is concluded that the big quantity of quarts passing the gas turbine eroded the blades. We succeeded in increasing the particle size distribution and improving the secondary cyclone efficiency by adding more limestone into the fuel. By this countermeasure, the dust density was reduced to the 40–50% as ever. We are free from this problem with CTF system.

5.6. Hot Gas Pipe Gas leak and heat spot occurred at the flange, elbow piece and T-piece mainly because of thermal stress and/or pressure fluctuation. Improvement of hot gas pipe design and structure has been carried out.

6. OPERATIONAL IMPROVEMENTS We have improved the operation variously through the demonstration program. Some of improvements are as follows.

6.1. Paste Quality Control Paste quality is checked every eight hours during operation. Checking is made by visual observation, five finger test and pipe test. Test results such as observation, finger feeling and the shape of the paste by pipe test are recorded electronically in the network computer. Plant operators change the circulation ratio to the crusher and the amount of adding water according to the test results. Paste sample is taken to analyze fine contents and water contents once a day. The analysis data are also recorded in the network computer.

6.2. Initial Fuel Injection Fuel paste plugs in the fuel nozzles were experienced at initial fuel injection during every start-up in the early 1996. Initial fuel injection tests were executed many times, changing injection schedule and paste quality. Cold model tests were also executed to observe the fuel injection. Through these tests, it is concluded that paste plugs are related with initial paste flow; that is, paste plugs occur when paste flow is too small. The paste flow at initial fuel injection is increased to 4.0 tons/hour whereas the original is 1.2 tons/hour. After this improvement, paste plug problem has been solved.

6.3. CTF Ash Improvement CTF tubes were broken twice in the early 1996. At the inspection after the tube breakage, some ash deposits are found inside of the tubes. It is considered that ash deposit causes the temperature deferential on the tube and that subsequently the thermal stress damages the tube. CTF ash is very sticky because of its fine particles. Increasing the particle size is

674

H. Goto and S. Okutani

effective to reduce the ash stickiness. We investigated the methods for increasing the particle size and carried out “Cyclone plug operation” from August to October in 1996. Exhaust gas bypassed one of seven primary cyclones, so that larger particles were expected to fly into the CTF directly from the boiler. This operation was successful with stable pressure drop and increased particle size. After “Cyclone plug operation”, inside of the CTF was inspected and found clean with few deposits. This operation is effective to reduce the ash deposits, but it is necessary to consider the ash erosion in the hot gas pipe.

6.4. CTF Tube Breakage The plant is operating with the primary cyclone bypass pipe which leads a part of boiler outlet gas to the CTF directly in order to improve the CTF ash particle size distribution and to reduce the ash deposits. This is one of the most important test for the

stable CTF operation. On July 1997, all of tubes have been changed to the less expansion tubes which are more durable to thermal stress. Analysis of the gas flow and the dust flow in the CTF pressure vessel is being made by calculation and laboratory test to get information about the mechanism of ash deposit

inside the tubes. Furthermore, the effect of ash deposit to the tube breakage is being investigated at present.

CONCLUSION The plant has been operated for 6,000 hours with CTF system and for 4,900 hours with 2 stage cyclone system to date. We have obtained much information on operation and maintenance of the PFBC plant through three year demonstration. We have improved the operation as well as the facilities giving many solutions to the troubles. Although we are still straggling with some problems as stated in this paper, we confirm the superiority and the possibility of PFBC system from the viewpoint of generating efficiency and environmental adaptation.

We plan to operate the PFBC plant according to the demonstration program until December in 1997. Much more data and information on operation and maintenance are expected to be obtained. We wish our experience be informative and effective to the design of PFBC plant in the coming stage.

DEVELOPMENT OF AN INNOVATIVE FLUIDIZED BED CEMENT KILN SYSTEM Sadayuki Shinozaki 1 , Isao Hashimoto2, Katsuji Mukai 3 , and Kunio Yoshida4 1

Center for Coal Utilization, Japan 6-2-31 roppongi, minato-ku, Tokyo 106, Japan Phone 81-3-5412-2536, Fax 81-3-5412-2540 2 Japan Cement Association 3-1-1 Higashikawasaki-cho, Chuo-ku, Kobe 650-91, Japan Phone 81-78-682-5215, Fax 81-78-682-5539 3 Japan Cement Association 1 Kanda Mitoshiro-cho, Chiyoda-ku,

Tokyo 101, Japan Phone 81-3-3296-9797, Fax 81-3-3296-1395 4 Department of Chemical System Engineering, University of Tokyo 7-3-1 Hongo, Bunkyo-ku, Tokyo 113, Japan Phone 81-3-5800-6801, Fax 81-3-5684-8402

1. OUTLINE OF FLUIDIZED BED CEMENT KILN SYSTEM

1.1. System Configuration The configuration of the fluidized bed cement kiln system is shown in Fig.l, which consists of a suspension preheater(SP), a fluidized bed granulating kiln (SBK), a fluidized bed sintering kiln (FBK), a fluidized bed quenching cooler(FBQ) and a packed bed cooler(PBC). Each component of them as following:

(1) Suspension Preheater (SP) consisting of 4-stage cyclones for preheating and calcining raw meals, which applies the conventional technology. (2) Granulating Kiln (SBK) for granulating raw meals into granules of 1.5~2.5 mm in average diameter at Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

675

676

S. Shinozaki et al.

high temperature of 1,300°C level, without feeding seed-core clinkers. This is a key process of the system. (3) Sintering kiln (FBK) for completing the sintering of the granules produced in SBK at high temperature of 1,400°C level. (4) Fluidized Bed Quenching Cooler (FBQ) for quickly cooling down the cement clinkers sintered in FBK, from 1,400°C level to 1,000°C at the high temperature zone in the cooling process in order to get good quality cement clinkers. (5) Packed Bed Cooler (PBC)

Development of an Innovative Fluidized Bed Cement Kiln System

677

for cooling down the cement clinkers economically to the specified temperature of around 100°C. In the fluidized bed cement kiln system, the separate two kilns which have different functions as abovementioned, that is, SBK and FBK are applied instead of a conventional rotary kiln, in order to maintain high quality clinkers and to avoid agglomeration troubles in the high temperature level of 1,300°C~1,400°C. The combination of the separate two kilns gives good running condition as well as significant reduction of NOx emissions, in comparison with the rotary kiln system. For the cooling process of clinkers, FBQ and PBC are applied instead of a conventional grate cooler. The combination of FBQ and PBC assures the efficient heat recovery in the cooling process of clinkers, so that the quantity of cooling air reduces down to the quantity required for burning fuel in SBK and FBK, whereas much surplus cooling air is required in the grate cooler due to the inadequate efficiency of heat recovery. Therefore, the heat consumption as well as CO2 emissions reduce accordingly in the fluidized bed cement kiln system.

1.2. System Features (1) High flexibility for choice of various grades of coals (2) High efficiency of waste heat recovery (3) Significant reduction of NOx emission due to fluidized bed combustion (4) Low energy consumption resulting in less CO2 emission (5) High controllability of burning temperature assuring high quality of existing cements and possibility for producing new kinds of special cements (6) Less construction cost and area (7) Less running and maintenance cost, and (8) Stable operation

2. OUTLINE OF TESTING FACILITIES The principal specifications of main equipments in 20t/d pilot plant and 200t/d scale-up plant are listed in Table 1.

3. TEST RESULT The running tests with 20t/d pilot plant were carried out for about 4 years. Through the period of 4 years the important key technology of the system was developed and the characteristics of them were confirmed. Then, 200t/d scale-up plant was constructed for the purpose of confirming the scaling-up factors and operational characteristics of the system. The running tests with 200t/d has been proceeding successfully since February last year.

3.1. Development of Key Technology in SBK 1) Special Distributor At the first stage of the pilot plant, the spouted bed with single spout was applied

678

S. Shinozaki et al.

to SBK. However, the falling of particles through the throat and the excessive particle

entrainment hindered the stable operation of the pilot plant. In order to solve these problems, the special distributor was developed. It was designed to have many spouts to keep the required jet strength to maintain the good granulating condition, taking into consideration the function of the spouted bed. The spouts of the distributor were arranged not to form inactive areas on the distributor in order to avoid agglomeration/coating troubles. Thus, the fluidized bed granulating kiln was developed instead of the spouted bed, and it enabled two dimensional scale -up as well.

2) Classifying Mechanism It is very important to control the size of clinker granules discharged from SBK in order to prevent from agglomeration troubles in FBK. The size variation of clinker granules in SBK is caused by two factors: the granulating characteristics and the discharge method. As the solution to the latter factor, the discharge of clinker granules was modified from the overflow discharge to the bottom classifying discharge. The classifying mechanism was built in the discharge part located at the bottom side of SBK, which discharged the grown up granules up to the aimed size and turned back the smaller granules and particles into SBK for further granulation.

This classifying mechanism realized the superior size controllability and the sharp size distribution of the granules discharged from SBK as well as the stable size distribu-

tion of granules inside of it, which solved completely the agglomeration troubles of clinkers in FBK, even if at the high temperature of 1,400°C. Testing results are shown in Fig. 2 and Fig. 3.

Development of an Innovative Fluidized Bed Cement Kiln System

679

3) Raw Meal Blowing-in Device The gravity chute feeding system was applied for feeding raw meals into SBK originally. However, heavy coatings occurred on the inside wall, especially around the entrance of raw meals and the particle entrainment was also considerable. So that, the raw meal blowing-in device was applied for keeping quick dispersion of raw meals in the fluidized bed of SBK. It solved the coating troubles as well as the particle entrainment from SBK.

3.2. Result of Running Tests 1) Granulation in SBK Keeping the temperature of SBK at 1,300°C level and feeding raw meals in it, the granulation can be carried out continuously. In this case, it is not required to supply seed-cores of clinkers from outside, because the seed-cores for granulation are generated autogenuously in the system. The growing speed of granules is decided by the seed-core size produced and the coating speed, so that it is important to grasp those values for controlling the granule size.

680

S. Shinozaki et al.

The raw meals and clinker granules have a liquid-phase components at high temperatures around 1,300°C. The liquid-phase components are utilized as binder for granulation in SBK. When raw meals are fed into SBK, some parts agglomerate to form seed-cores and some parts adhere around the existing granules, causing them to increase in size. In this case the portion of seed-cores to be formed is very important to maintain the granulating condition. If it can be kept constant, the granulating condition can be also main-

tained constant, and SBK is operated stably. The portion of seed-cores to be generated depends on the temperature, the composition and the density of raw meals as well as the equipment conditions. The portion of seed-cores decreases usually almost proportionally to the granulation temperature on the condition of the constant density of raw meals, and it increases with the increase of raw meals density.

In the fluidized bed cement kiln system, the generation of seed-cores has been controlled to be kept constant, so that SBK has been operated stably without feeding seedcores from outside. So that, it is called “hot self-granulation”. Having studied the relation between the average diameter of granules and the weight percentage of each section of the size distribution, it was found that there was a strong correlation between the average diameter of granules (Dp50) and the weight percentage of section (R500). The Correlation obtained in both of the pilot plant and the scale-up plant is shown in Fig. 4. The two are almost the same, so that the size distribution of granules in SBK is expected to be maintained even with further scaleup of the system. 2) Clinkering Reaction in FBK

The average residence time in the scale-up plant has become half of that in the pilot plant. The residence time is one of the important factors which concerns with the clinker quality. The f-CaO content of clinker is the representative index to show the clinker quality, because it shows the proceeded degree of clinkering reaction. The f-CaO content of clinker is influenced by the sintering temperature and the residence time. The clinkering reaction and the crystallization accelerate with the sintering temperature increase. The target of the f-CaO is Less than 1.0%.

Development of an Innovative Fluidized Bed Cement Kiln System

681

According to the test results, the f-CaO of clinkers in the scale up plant was maintained to around 0.4% in half of the residence time required in the pilot plant at the same sintering temperature. And it was confirmed that the coal ash reacted completely with clinkers as well. The cross section of clinker is shown in Figure 7.

3) Heat Recovery Efficiency The waste heat recovery efficiency in the pilot plant and the scale-up plant was almost the same as shown in Fig. 5 and Fig. 6. The data shows that the heat consumption of the scale-up plant achieved the designed value. According to the figures, the heat recovery efficiency in FBQ increases with the increase of the specific cooling air quantity even over the designed point of 100%, so that the specific cooling air can be decided from an economical point of view, considering total gas balance of the system. The heat recovery efficiency in the PBC increases with the increase of specific cooling air quantity and saturates around the designed point of 100%. It shows that the designed point is almost the optimum point. 4) Pollutant Emissions (1) Emission The reduction of heat consumption as well as the use of low grade coal with low carbon content in the system enables to reduction of CO2 emissions considerably.

682

S. Shinozaki et al.

In the scale-up plant, the heat consumption decreased by which led to reduction of emissions by in comparison with conventional rotary kiln plants. According to the feasibility study on 3,000t/d cement plant which is an average commercial scale in Japan, emissions can be saved by by decreasing heat consumption, in comparison with the existing rotary kiln system one, although saving percentage of emissions decreases with the increase of capacity. (2) NOx emission The thermal NOx varies with fire flame temperature, combustion rate, excess air ratio, and so on. The combustion temperature in the fluidized bed cement kiln system is with no flame, whereas that of the rotary kiln system is with flame. Accordingly, the thermal NOx emissions reduce in the new system.

NOx emissions in the pilot plant are shown in Table 2. NOx emissions amounted to approximately 1/3 of that in the rotary kiln system when heavy oil was used, and around 1/2 when pulverized coal was used. NOx emissions in the scale-up plant have been still the same level. The level is high compared with the target of 200ppm. The figure will be attained by reducing the excess air.

4. CONCLUSION (1) The stable control for granule size in the granulating kiln has been accomplished by introducing the special distributor, the classifying mechanism of granule size and the raw meals blowing-in device. (2) The quality of clinkers in the scale-up plant was maintained enough at the half

of the residence time in the pilot plant at the same sintering temperature level. And coal ash reacted completely with clinkers.

Development of an Innovative Fluidized Bed Cement Kiln System

683

(3) The heat recovery efficiency in the scale-up plant was the same level in the pilot plant as expected. (4) emission reduced by in the scale-up plant in comparison with conventional rotary kiln plants, and (5) NOx emission in the scale-up plant is a little high comparing with the final target of 200ppm.

ACKNOWLEDGEMENTS Both of the pilot plant and the scale-up plant in this development project of the new technology have been subsidized by MITI. We wish to acknowledge the related personnel of MITI, CCUJ and JCA for their support and assistance.

This page intentionally left blank

THE INFLUENCE OF PRESSURE ON THE BEHAVIOUR OF FUEL CARBONATES Arvo Ots, Tõnu Pihu, and Aleksander Hlebnikov

Thermal Engineering Department Tallinn Technical University Tallinn, EE0026, Estonia

1. INTRODUCTION Combustion of solid fuels under the pressure is a prospective method to burn fuels with high content of carbonates. Utilization of solid fuels with high carbonate content in mineral matter has remarkable peculiarities burning in pressurized conditions1–3 . A fuel

that contains a considerable amount of carbonates is Estonian oil shale4. Burning the oil shale under the pressure causes some different problems from mineral matter. To observe the fuel mineral- and organic matter behaviour in atmospheric and pressurized combustion conditions an experimental system was designed and installed. This system enabled

to vary pressure, temperature and gas environment. Oil shale mineral matter behaviour during atmospheric combustion is well investigated and reported5–7. These results show that in the case of atmospheric pressure combustion, independent of combustion technology (PF, AFBC, etc.), the decomposition of carbonates occurs at very high rate. This pressure of in flue gas is in the range 0.014–0.016MPa and the temperature in AFBC is 800–900°C but in PF combustor the maximum temperature usually exceeds 1,400°C. Part of the CaO formed in the dissociation process of calcium carbonate stays in the free form and the rest of CaO will associate with sandy-clay minerals of fuel5. The dissociated from carbonates goes to the flue gas. Due to the carbonate decomposition the heating value of fuel decreases. The capturing of with ash takes place through CaO after the dissociation of calcium carbonate. In pressurized combustion conditions (total pressure 1.2–1.5 MPa) the partial pressure of in flue gas exceeds the equilibrium dissociation pressure of and calcite does not directly dissociate. In case of PFBC the combustion temperature is in the same level as in AFBC, but the pressure in the flue gas, depending on the total pressure, is in the range 0.13–0.15 MPa. Obviously the carbon dioxide partial pressure in the flue gas also depends on the carbonates decomposition rate. If the fuel consists minerals able to react with calcium oxide (for instance, sandyImpact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al.

Kluwer Academic / Plenum Publishers, New York, 1999.

685

686

A. Ots et al.

clay part minerals of fuel), then some part of calcite may associate with these minerals and to be freed, goes into the flue gas. Also it is possible that part of from calcite goes into the flue gas as a result of the direct reaction between and according to:

The

sources are organic and pyritic sulphur (combustible sulphur).

2. EXPERIMENTAL The aim of the experiments was to investigate the reactions between carbonate and sandy-clay minerals on the basis of Estonian oil shale in the pressurized and atmospheric combustion conditions. All experiments were performed in the pressurized combustion facility in Thermal Engineering Department of Tallinn Technical University. The experimental system (Fig. 1) is designed to burn up a portion of solid fuel under pressurized or atmospheric conditions. The experimental system consists of: gas

mixing-controlling system, reactor-heater, pressure-holding system and data acquisition and control system.

The gas mixing-controlling system made up of mass flow controllers (1, 2, 3, 4), calibrated accordingly to gas to be used, and of check valves (5). The mass flow con-

The Influence of Pressure on the Behaviour of Fuel Carbonates

687

trollers keep gas mixture composition in wide range of pressures (0.1–1.5 MPa) and flow rates. The reactor-heater is designed to meet working pressures up to 1.5 MPa and temperature up to 980°C. Two cylindrical electric heaters (13) were placed inside the reactor. Inside the heaters are placed quartz pipes that protect heating wire against sulphur compounds that may be present in gas mixture. To heat up and to keep given temperature for gaseous environment, the temperature controller (8), relay (9) and thermocouple (11)

were used. The thermocouple (10) measures the flue gas temperature during experiments. Two washer thermocouples (12) were used to measure the reactor body surface temperature (for safety reason). The reactor body and piping were made from stainless steel SS316. The system was supplied with a gas cooler (15) to cool down the flue gas and with filter (16) to remove tar from gas.

The pressure-holding system consists of needle control valve with electric activator (6) and pressure transducer (7). Opening of the needle valve is controlled by pressure controller (14). The main detail of data acquisition and control system (17) is an IBM High Speed 16-channel Plug-in Interface Card. This card is installed into the 486DX4/100 processor based PC. The card can handle analogue and digital signals simultaneously. The data during tests (temperatures, pressures and gas flow rates) were collected to the computer

HD.

2.1. Experimental Procedure Approximately 5 grams of material loaded into the sample holder was taken for each experiment. Stainless steel SS316 bound net with openings was used to make sample holder. In each run, the sample was loaded into the balance chamber (vertical chamber on the top of reactor with ball valves at both ends). From balance chamber the sample was dropped into reactor through which a gas mixture flows under predetermined

conditions (temperature, pressure and flow rate). Continuous gas analyses of flue gas were run during experiments. After a specified reaction time the reactor was cooled down together with the sample. The cooling process proceeds under pressure in environment in order

to prevent the decomposition of carbonates. The cooling rate is in the range 40–60 K/min. After falling the temperature below 250°C in reactor, the gas flow was shut down and reactor was depressurized. The sample was taken out from reactor. Mass of sample was determined before and after each experiment. Chemical analyses of the samples were made after experiments. Four sets of experiments were carried out: two with limestone (at atmospheric

pressure—0.1 MPa and at elevated pressure—1.2 MPa) and two with oil shale (0.1 and 1.2MPa). The temperature for all experiments was 850°C. The typical gas composition during tests was 86.1 vol. % 10.2 vol. % and 3.7 vol. % (gas mix No. 1). A gas composition of 94.6 vol. % and 5.4 vol. % (gas mix No. 2) was also used. Experiments were made with limestone from Vao open pit mine in Estonia and with samples prepared from oil shale received from Baltic Power Plant. The particle size of limestone and oil shale samples was

Chemical analyses of samples were done before (raw material) and after (residue) the experiments. The chemical compositions of tested materials are given in the Tables 1 and 2.

688

A. Ots et al.

3. RESULTS AND DISCUSSION

3.1. Experiments with Limestone Limestone was used as material for preliminary tests. The meaning of preliminary tests was to clarify the influence of pressure on the carbonate decomposition rate. The experiments with limestone samples were carried out at atmospheric and elevated pressure by temperature 850°C and in gas mix No. 1. After the experiment with limestone in pressurized conditions, the change of sample mass was not observed. This fact is explained with high carbon dioxide pressure exceeding calcium carbonate decomposition pressure according to the equilibrium curve for dissociation. Experiments in atmospheric conditions show a remarkable decreasing of limestone mass due to decomposition of calcium carbonate.

The Influence of Pressure on the Behaviour of Fuel Carbonates

689

The results of experiments were expressed in Fig. 2 as carbonate decomposition rate as a function of time. The carbonate decomposition rate of limestone in atmospheric conditions continuously increased with time.

3.2. Experiments under Atmospheric Pressure Experiments under atmospheric pressure were carried out at the temperature 850°C in the gas mix No. 1 and in gas mix No.2. In the first case the partial pressure is low (lower than dissociation pressure at 850 °C) and the intensive carbonate decomposition takes place. In the second case where the carbon dioxide pressure is high (higher than dissociation pressure at 850°C) the carbonate decomposition process is broken. The calculated value of carbonate decomposition rate on the basis of experimental data is 0.9 if the concentration in gas mix is 10.2 vol. % (Table 3).

Due to the high carbon dioxide concentration in gas mix No.2 (94.6 vol. % of ), the carbonates decomposition rate is much lower (0.29) comparing with decomposition

A. Ots et al.

690

rate of carbonates in medium containing 10.2 vol. % of This result is well correlated with calcium carbonate equilibrium curve depending on pressure. Under low carbon dioxide pressure, the calcium oxide from fuel mineral matter is produced mainly by thermal decomposition of limestone and dolomite. A part of the CaO reacts with sulphur dioxide according to the reaction: (2)

The following reaction between calcium oxide and pyrite is also possible:

(3)

Some part of calcium oxide stays in the free form. Part of produced CaO may react with components from sandy-clay minerals of fuel5,6 such as producing different minerals like and The ratio between CaO in free form and bounded with above mentioned minerals depends mainly on temperature. From the sulphation point of view, all minerals containing calcium oxide are able to react with sulphur dioxide. Due to the high carbon dioxide concentration in gas mix No.2, the carbonate

decomposition rate is much lower comparing with carbonate decomposition rate

with 10.2 vol. % concentration of The carbonate decomposition rate for the oil shale samples treated in gas mix with concentration 10.2 vol. % is about 0.9. This value and also the sulphur capturing factor with ash (about 0.6) show that the sulphur is not captured with ash totally. Part of the sulphur goes into gas, mainly as

3.3. Experimental Results under Elevated Pressure Experiments under elevated pressure (the total pressure was 1.2MPa, except one

experiment where the pressure was 1.7MPa) were carried out at the temperature of 850 °C in the oxygen-containing atmospheres. The experimental results are presented in Table 4. It is obvious that the pressure prevents the decomposition of carbonates and the sulphation based on reaction (1). The carbonate decomposition rate calculated on the basis of sulphur content in the oil shale is 0.09. The value of carbonate decomposition rate obtained in experiments is in the range 0.30–0.40 (Table 4). The decomposition rate due to calcium oxide sulphation was calculated on the basis of sulphate sulphur content in ash obtained from experiments (Table 4). The established in experiments sulphur capturing factor with ash was lower than 1.0. It means

that the total amount of the sulphur does not react with ash and part of sulphur goes to the flue gas mainly as

Actual carbonate decomposition rate is remarkable higher than the

decom-

position rate calculated on the basis of sulphation reaction (1). Consequently, the chemical reactions between carbonates and sandy-clay part minerals of fuel also must take place. Obviously the following reactions are possible:

The Influence of Pressure on the Behaviour of Fuel Carbonates

691

4. HEATING VALUE The starting point for the determination of the higher or lower heating values of solid fuels (HHV or LHV) is the experimentally established heating value (in laboratory). In the most cases it is the heating value measured in calorimetric bomb (calorific heating value). The heating value of the fuel, established in calorimetric bomb or by another

experimental method, does not correspond with heat release conditions in existing combustors. The fuel combustion is different in calorimetric bomb compared with fuel

burning process in combustors. In the fuel burning not only the oxidizing of organic components but also the conversion processes in the mineral matter take place. The conversion processes in mineral matter of fuel and its heat effects on the heating value depend, at first, on the qualitative and quantitative composition of inorganic part and, on the

second, on the combustion parameters such as temperature, composition of gas environment and pressure. Burning of fuels with high content of carbonates, the remarkable influence on the heating value of fuel produce the processes, connected with decomposition of carbonates. The most important heat effects that influence on the heating value of fuel are following:

i) endothermic heat effects from dissociation of carbonates; ii) exothermic heat effects from formation of calcium sulphate; iii) endothermic heat effects from formation of the double calcium oxide minerals

(e.g. reactions between the carbonate and sandy-clay minerals) Taking into account the additional heat effects in burning of fuels containing carbonates, the heating value of fuel is expressed by the formula:

where [QrL]—the LHV as received bases by the full dissociation of carbonates, taking into account full conversion of combustible sulphur (organic and pyritic) to sulphates and staying of rest calcium oxide in free form; —heat effect of non full dissociation of carbonates; minerals;

—heat effect from reactions between CaO and sandy-clay

A. Ots et al.

692

On the pressurized combustion conditions the combustible sulphur is fully converted to calcium sulphate and is taken into account in the term Heat effect of nonfull dissociation of carbonates depends on mineral carbon dioxide quantity in the fuel and on the decomposition rate of carbonates, and can be expressed by formula:

where -mineral carbon dioxide content in the fuels received bases, mass %; k CO2—decomposition rate of carbonates. The heat effect from reactions between CaO and sandy-clay minerals is expressed by the formula (for Estonian oil shale):

The heating value of the fuel strongly determined on the behaviour of carbonates in combustion and reflects different values of dissociation rate of carbonates depending on combustion technology (PF, AFBC, PFBC). The heat realise in fuel pressurized combustion is remarkably higher due to the lower decomposition rate of carbonates comparing with atmospheric burning, i.e. the

heating value of fuel in pressurized combustion is increased.

The dependence of the relative increasing factor of LHV

on carbonate decomposition rate and LHV Estonian oil shale as received bases is given in Fig. 3. Because the decomposition rate of carbonates lies in the region 0.3–0.4, then the oil shale LHV in pressurized combustion conditions is 5.5—8% higher (LHV as

The Influence of Pressure on the Behaviour of Fuel Carbonates

693

received bases is in range 7.5–8.5MJ/kg) compare with the conditions where full dissociation of carbonates takes place.

5. CARBONATE DIOXIDE CONCENTRATION IN FLUE GAS The carbonate decomposition rate in combustion process of fuel also influences to the concentration in the flue gas. The amount in the flue gas (in normal cubic meters per kg of fuel on dry basis) depends on organic carbon and mineral carbon dioxide

content in the fuel and can be calculated by formula:

where — vol. concentration in flue gas per kg of fuel, —organic carbon content in the fuel on dry basis, mass %; —mineral content in the fuel on dry basis, mass %; —carbonate decomposition rate. The decreasing factor of the carbonate dioxide concentration in flue gas may be expressed by the following formula:

where

—maximum vol. concentration in the flue gas vol. concentration in the flue gas by carbonate decomposition rate

The calculated values of the decreasing factor of the carbonate dioxide concentration in flue gas depending on the decomposition rate of carbonates and on the oil shale

LHV as received basis are given in the Fig. 4. According to Fig. 4 we can see that, depending on the LHV as received basis (7.5–8.5 MJ/kg), the concentration in flue gas should decrease 13–19% at pressur-

694

A. Ots et al.

ized burning of oil shale (decomposition rate of carbonates is in the range 0.3–0.4) comparing with conditions when the full decomposition of carbonates take place.

6. CONCLUSIONS During combustion of solid fuel with high content of carbonates under atmospheric pressure calcium carbonate is dissociated to calcium oxide and Part of formed CaO reacts with sulphur and sandy-clay minerals of fuel mineral matter and part stays in free form. During the combustion under elevated pressure due to high carbon dioxide pressure the carbonates do not decompose directly. The carbon dioxide will be freed in the chemical reactions between sulphur and sandy-clay minerals. Base on the combustible sulphur content of the fuel, the carbonate decomposition rate must not exceed the certain value (e.g. for Estonian oil shale 0.09). The established decomposition rate was in the range 0.3–0.4. This experimentally obtained carbonate decomposition rate is higher than the calcite decomposition rate calculated on the basis of sulphur capture possibilities with carbonates. This fact shows that the direct chemical reaction between the carbonates and sandy-clay minerals in fuel mineral matter occurs. The carbonate decomposition rate directly influences the heating value of fuel and

the emission. Basing on experimentally established carbonate decomposition rate one may conclude that the heating value, for instance, of oil shale increases approximately 5.5–8% and the carbon dioxide concentration in flue gas decreases about 13–19% compared with conditions when the carbonates decompose completely.

ACKNOWLEDGEMENTS This work is a part of the EU project Joule II Extention, PECO, Estonian Innovation Found and Estonian Research Found. Our special thanks belong to Indrek

Külaots and Teet Parve for advises and recommendations that helped us in designing laboratory equipment system. We are also very appreciate to Igor Tjurlik for help on first experiments at our laboratory system.

REFERENCES 1. Iisa, K. (1992). “Sulphur Capture under Pressurised Fluidised Bed Combustion Conditions”. Academic Dissertation. Abo Akademi University, Turku, Finland. 2. Yrjas, K. P., Külaots, I., Hupa, M. and Ots, A. (1995). “Sulphur Capture by Oil Shale Ashes under Atmospheric and Pressurized FBC Conditions”. In Kay J. Heinschel (Editor) 13th International FBC Conference New-York.

3. Yrjas, P. (1996). “Sulphur Capture under Pressurised Fluidised bed Combustion and Gasification Conditions”. Abo Akademi University, Turku, Finland. 4. Ots, A. (1996). “Utilization of Estonian Oil Shale at Power Plants.” In L. Baxter, R. DeSollar (Eds.), Applications of Advanced Technology to Ash-Related Problems in Boilers.” New York and London: Plenum Press.

5. Ots, A. (1977). “The Processes in the Steam Generators in Burning Oil Shale and Kansk-Achinsk Basin Coals”. Moscow:Energia, [in Russian].

The Influence of Pressure on the Behaviour of Fuel Carbonates

695

6. Ots, A., Arro, H., Jovanovic, L. (1980). The Behaviour of Inorganic Matter of Solid Fuels during Combustion. In: Fouling and Corrosion in Steam Generators. Beograd. 7. Arro, H., Prikk, A., Kasemetsa, J. (1997). “Circulating Fluidized Bed Technology—Test Combustion of Estonian Oil Shale”. Oil Shale, vol.14, No.3, Special, 215–217.

This page intentionally left blank.

MODELING OF ASH DEPOSIT GROWTH AND SINTERING IN PC-FIRED BOILERS Huafeng Wang and John N. Harb* Department of Chemical Engineering

Brigham Young University Provo, UT 84602 U.S.A.

1. INTRODUCTION The combustion of pulverized coal for electricity generation is a complex process, and one of the least understood and most costly aspects of this process is ash deposition on heat transfer surfaces. Ash deposition leads to reduced heat transfer in the boiler and corrosion of boiler tubes, which may result in reduced generation capacity and unscheduled outages. The extent of deposition typically depends on the amount, composition, and mode of occurrence of the inorganic matter in the coal, as well as the boiler design

and operating conditions. Since all coals have varying amounts and types of inorganic constituents, the deposition behavior of different coals will be different. Boiler design has historically involved the use of empirical data and proprietary correlations applicable to specific coals. The potential economic and/or environmental benefits of burning non-design coals has led to coal-switching, coal-blending, or coalcleaning. However, the impact of a new or modified fuel on ash deposition is difficult to predict. In fact, it is sometimes necessary to derate a boiler in order to manage deposition in day-to-day operation and to prevent catastrophic deposition from occurring. As our understanding of ash deposition has increased, development of computer models to predict deposition behavior has become more feasible. Models can be used to help diagnose and correct operational problems due to deposition, and to provide insight into the deposition behavior of a particular coal or coals. Ash formation and deposition models have recently been reviewed by Wang and Harb (1997) who discussed the advantages and limitations of several of these models. In spite of developments to date, accurate prediction of deposit properties as a function of time during deposit growth remains a difficult task. This paper is an extension of our previous efforts to model ash deposition. Previously, deposit properties were estimated by assuming that equilibrium was reached locally in the deposit (Richards et al., 1993; Wang and Harb, 1997). However, recent experimental data (West, 1996) clearly show that this is not the case. Consequently, a new Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

697

H. Wang and J. N. Harb

698

deposit property model based on viscous flow sintering has been developed and integrated into the comprehensive combustion code PCGC-3. The resulting model includes the effects of both ash chemistry and operating conditions on deposition in the radiant section of a boiler (slagging deposition). Similar to our previous work, this model has been applied to both a pilot-scale combustor and a utility boiler. A brief description of the model is provided first, followed by a discussion of the results from simulations of both the pilot-scale and full-scale facilities.

2. MATHEMATICAL MODEL This section reviews key aspects of the mathematical model presented in this paper. It begins with a brief description of the comprehensive combustion code which is used to simulate pulverized coal combustion. Methods for estimating particle impaction and sticking are then reviewed. Finally, a new sintering deposit model, the key contribution of this paper, is presented. This model is used to predict deposit properties as a function of time during deposition.

2.1. Combustion Code A modified version of PCGC-3, a comprehensive combustion code developed at Brigham Young University, was used in the present study (Hill and Smoot, 1993;

Wang and Harb, 1997). Simulations of pulverized-coal combustion were performed by first converging the gas flow field to a specified tolerance, a process which typically required 120–150 gas phase iterations. The flow field was then used for the particle (coal and ash) momentum, reaction and energy calculations. The mass, momentum and energy source terms resulting from the particle calculations were next used to update the gas phase calculations. This process was repeated until overall convergence was achieved, as measured by the amount of change in the gas phase following a particle calculation and the extent to which the overall mass and energy balances were satisfied. Between 10 and 20 particle iterations were required to converge a single combustion simulation with approximately 200,000 computational cells. The computation time required was equivalent to about 2–3 days of CPU time on a HP9000–755 computer workstation.

2.2. Particle Impaction and Sticking A statistical particle cloud model from the literature (Litchford and Jeng, 1991 and 1992) was incorporated into PCGC-3 as a postprocessor to estimate particle impaction rates at the walls. In the model, each computational parcel is characterized by a normal

(Gaussian) probability density function (PDF) in space. The instantaneous location of each computational parcel, determined stochastically, is taken to represent the mean of its corresponding PDF. The total number of computational parcels was chosen to control the size of the particle cloud so that the following two assumptions could be imposed without introducing significant error: (1) particles in one computational parcel have same

properties, and (2) once the mean location of a parcel hits the wall, the entire parcel is assumed to hit the wall. Calculations of particle impaction rates were performed for reacting particles, and carbon burnout was tracked along the particle trajectories. It was assumed that inorganic matter in each coal particle agglomerated to form a single ash

Modeling of Ash Deposit Growth and Sintering in PC-Fired Boilers

699

particle. An additional two to three days of CPU time on a HP9000–755 workstation were typically required to calculate the particle impaction rates. Once the impaction rates had been determined, it was necessary to estimate the fraction of the impacting particles which adhered to the walls. This was done on a particle-by-particle basis for a representative number of particles based on the physical characteristics (viscosity) of both the particle approaching the deposition surface and deposit surface itself (Walsh et al., 1990). A critical viscosity of Poise was chosen for determining the sticking probability of impacting particles after Richards et al. (1993). In this study, a viscosity model developed by Senior and Srinivasachar (1995) was used to calculate the viscosity of silicate glass particles (>15wt% Si) as a function of both temperature and composition. Oxide particles (e.g. or CaO) were treated as solids (infinite viscosity). Compositions of the fly ash samples used in the present study were measured with CCSEM. Alternately, the composition could have been provided by the output of any fly ash transformation model. However, both of these options provide compositions for thousands of fly ash particles, making particle-by-particle calculation of deposit build-up unfeasible. To circumvent this difficulty, a sorting program, developed by Slater et al. (1996), was used to classify the fly ash particles either from CCSEM or from an ash transformation model into a finite number of composition groups before calculation of particle sticking and deposit growth. Some of the fly ash particles have compositions which are readily associated with well-defined species such as iron oxide and silica. However, the majority of the fly ash particles are associated with compositions of amorphous phase. These amorphous phases with similar compositions were classified into same group. The average compositions and standard deviations of the classified groups were calculated for examination of these groups and for the definition of new groups. These groups, varied from one coal fly ash to another coal fly ash, were carefully defined in order to preserve the compositional features of the ash. Approximately 40 compositional groups were used for the simulations reported here. Comparison of the predicted bulk elemental composition with the measured composition revealed a deficiency of oxide particles in the inner layers of the deposit. It is believed that this deficiency was due to a sticking mechanism not accounted for in the present model. Specifically, it is possible for non-sticky particles to become trapped in an existing particulate deposit. Such a deposit could absorb the momentum of the incoming particles, preventing them from rebounding. An empirical correction was made to account for this effect.

2.3. Deposit Sintering One of the principal contributions of this paper is a model for the prediction of deposit growth and sintering. Viscous flow sintering is generally accepted as the key mechanism for densification and strength development in slagging deposits (Benson et al., 1992). Sintering models available in the literature generally assume particles of uniform size and composition (Frenkel, 1945; MacKenzie and Shuttleworth, 1949; Scherer, 1977). The influence of particle shape and size distribution on sintering rates has been considered (Coble, 1973; Chappell and Ring, 1986). The effect of the non-uniform composition on the sintering rate is also recognized as important, although it has not been discussed quantitatively in the literature (German, 1996). Unfortunately, ash deposits are formed from fly ash particles which come in a wide variety of sizes, shapes, and compositions. Thus, accurate prediction of the sintering rate in ash deposits is a very difficult task. Also, sintering predictions must be simple enough computationally to permit inte-

700

H. Wang and J. N. Harb

gration with the comprehensive combustion code for the simulation of ash behavior in large-scale boilers. This paper presents a simplified sintering model which attempts to represent the key features of sintering in ash deposits at a minimum computational cost. The model is based on the following assumptions: 1) fly ash particles which adhere to the surface to form the ash deposit are spherical, 2) particles which adhere to the surface can be divided into a finite number of particle groups (about 400 in this study) according to size and composition in order to represent the behavior of a large number of particles, 3) the extent of sintering of a particular particle group depends only on the size and composition of particles in that group, and 4) high viscosity particles (e.g. particles) influence the morphology and long range order of the deposit. The first assumption was made for simplification purposes and is not expected to have a large effect on the predicted results. Grouping of the particles (Assumption 2) was required in order to make the problem computationally tractable. It was important that the grouping be performed carefully in order to preserve the compositionally dependent features of the ash. The third assumption is a key simplification which states that, on the average, the extent of sintering of particles of a particular size and composition is the same as that which would be observed from a deposit made up entirely of those particles. This assumption makes it unnecessary to deal with individual pairs of particles of different size and/or composition. Instead, simple models such as those developed by Scherer (1977), Mackenzie and Shuttleworth (1949) can be used. The last assumption recognizes that there are particles which will not sinter until temperatures are sufficiently high and some mixing with basic “flux” components has taken place. These particles were defined as particles which had a viscosity greater than 1011 Poise at a temperature of 1,250K. The local extent of sintering as a function of time was calculated by dividing the deposit into discrete layers. The extent of sintering for each layer during the time step

was approximated by the mass average of the extent of sintering for each of the relevant particle groups. High viscosity particle groups were not allowed to mix with other particles and sinter until other groups of equal or smaller size had completely sintered. Once the sintering of a particular particle group was complete, the mass of this group was distributed among the high viscosity particle groups and assumed to completely mix. The high viscosity particle groups were subsequently allowed to sinter. This procedure permitted approximation of the local extent of sintering as a function of time, particle size, and composition. In order to model deposit growth and heat transfer, it was necessary to relate the thermal and physical properties to the overall sintering percentage of the deposit. The porosity and thermal conductivity of deposit were related to the overall sintering percentage and deposit temperature (Wang, 1998). The emissivity and absorptivity of the deposit surface was determined after Harb et al. (1993). These thermal and physical properties of the deposit were used to predict the deposit growth and heat transfer as discussed in the next section.

2.4. Deposit Growth And Heat Transfer Deposit growth is simulated for a specified period of time by dividing the total time into a number of finite steps. The amount of deposit which accumulates during the time step is referred as the deposit layer. Figure 1 illustrates the procedure used for

simulation of deposit growth. First, the extent of sintering which occurs during the time step is calculated for each layer of the existing deposit (layers 1 to i-1), and the deposit properties are updated for those layers. Next, the mass of the deposit which accumulates

Modeling of Ash Deposit Growth and Sintering in PC-Fired Boilers

701

during the time step is determined by the product of the particle impaction rate and capture efficiency. Sintering of new deposit layer (initial deposit porosity of 0.6) is then calculated and the properties of the new deposit layer are approximated. Finally, an iterative energy balance is performed to determine the heat flux and temperatures through the deposit (Richard et al., 1993). The entire process is repeated until the total time specified for deposit growth has been reached. 3. RESULTS AND DISCUSSION

3.1. Pilot-Scale Simulations The simulations performed for model evaluation were similar to those presented previously for an earlier version of the model where deposit properties were based on local equilibrium rather than viscous flow sintering (Wang and Harb, 1997). Validation was performed by comparison with experimental data from the Fireside Performance Test Facility (FPTF) of Combustion Engineering, Inc.’s ABB Power Plant Laboratories. This pilot-scale combustor is designed to operate at conditions that closely approximate the heat fluxes, particle residence times and temperatures found in full-scale facilities.

702

H. Wang and J. N. Harb

Heat flux measurements and/or deposit samples were supplied by Combustion Engineering, Inc. from work performed for the Coal Quality Expert (CQE) Program funded by the U.S. Department of Energy and EPRI (Thornock and Borio, 1992). Also, a limited number of deposit samples were analyzed at BYU using SEMPC (West, 1996). A 69 × 57 × 44 three-dimensional Cartesian grid was used to simulate the FPTF reactor with 36 cells on the surface of the cooled deposition panel. Since little information was available on the properties of the insulating walls, the temperature and emissivity of these walls were assumed to be constant in the simulation. Furthermore, these two parameters were chosen so that the simulated incident radiative heat flux at Panel 1 matched measured values. The combustion simulations were then performed, followed by calculation of the respective particle impaction rates and, finally, simulation of deposit growth. Four different simulations were performed (Table 1) to illustrate the effects of operating conditions and coal type on ash deposition. These simulations were performed for a 70/30 blend of Wyoming and Oklahoma coals, a 70/30 blend of the Wyoming coal and a cleaned Oklahoma coal, and a hvA Eastern Bituminous coal. Standard data on these coals are given in Table 2. Figure 2a provides a comparison of the predicted and measured net heat fluxes through Panel 1 during deposit growth for the 70/30 blend. As shown in the figure, the

net heat flux decreased significantly during the first 1–2 hours and remained relatively constant after 2–3 hours. This behavior demonstrates the important contribution of the initial layers of the deposit to the overall thermal resistance. Also shown in the figure are predictions made previously with a local equilibrium model (Wang and Harb, 1997). The new sintering model provides a better representation of both the shape and magnitude of the experimental results. The shape of the curve indicates that the model more accurately describes the relevant physical mechanisms. The precise agreement of the magnitude is somewhat fortuitous since there was no arbitrary adjustment of parameters to fit the data. Similar predictions for the same fuel at a lower firing rate and for a cleaned fuel are shown in Figs. 2b and 2c respectively. In all cases, the model appears to adequately predict

the experimental data. Note that the same set of physical parameters was used for all three simulations. Figure 3 shows the thermal resistance of the deposit as a function of time for both the new sintering model and the local equilibrium model. Also shown is a single experimental datum point estimated from the Coal Quality Expert (CQE) Program (Thornock and Borio, 1992). The thermal resistances predicted by the local equilibrium model are considerably higher than those calculated with the new sintering model. However, the new predictions are in much better agreement with the experimental data. The steep

Modeling of Ash Deposit Growth and Sintering in PC-Fired Boilers

703

change in the resistance during the first 100 minutes again illustrates the significant contribution of the initial layers of the deposit to the overall thermal resistance. In order to illustrate the potential impact of ash chemistry on the sintering behavior of the deposit, a simulation was performed under conditions identical to those of the clean coal fired at 3.6 MBtu/hr, except that the ash chemistry was arbitrarily changed to that of a hvA eastern bituminous coal (Table 2). The results shown in Fig. 4 indicate that the chemistry had a substantial impact on the thermal resistance of the deposit. A significantly thicker deposit (and hence a higher thermal resistance and temperature) was required before substantial sintering occurred with the bituminous ash chemistry. Finally, the ability of the sintering model to track changes in composition was evaluated by performing simulations of a hvA bituminous coal fired at 3.5 MBtu/hr. A deposit sample formed under similar conditions was also examined by SEMPC (West, 1996). A comparison of the predicted and measured data (Fig. 5) shows good agreement between the two. In particular, both the predictions and measurements show decreased amounts of Fe and Si rich phases in the middle region of the deposit due to increased sintering. The outer region (not shown) is somewhat more complex since temperatures are sufficiently high that additional phases crystallize out of a “melt.” Crystallization is not included in the current sintering model.

704

H. Wang and J. N. Harb

Modeling of Ash Deposit Growth and Sintering in PC-Fired Boilers

705

In summary, an ash deposit sintering model has been developed and integrated into the comprehensive code PCGC-3. Pilot-scale simulations were performed in order to validate the model, and it was found that the model adequately predicted the heat flux and the thermal resistance through a cooled panel designed to simulate the waterwall tubes

706

H. Wang and J. N. Harb

in a utility boiler. Predictions of deposit composition also look promising. This integrated model can now be used to predict the effects of operating conditions and ash characteristics on deposit growth and properties.

3.2. Utility Boiler Simulations (Goudey Station) The Goudey station power plant is operated by New York State Electric and Gas Co. (NYSEG). The boiler is a tangentially-fired, forced recirculation pulverized coal unit

with a nominal capacity of 85MWe. The firebox is about m and has 16 corner burners. three-dimensional Cartesian grid was used to simulate the boiler. As the initial deposit layer forms very quickly, a uniform initial deposit layer having a resistance of was assumed (Blokh, 1992). Heat transfer and ash deposition in the boiler are strongly coupled and a more sophisticated solution strategy than that used for the pilot-scale simulations was required for the large-scale simulations (Wang and Harb, 1997). In the present study, deposit growth was allowed to proceed for approximately two and half hours (simulated deposition time). The boiler was operated with excess air (4.5% in the effluent), a coal feed rate of 29,100kg/h, and an air flow of 487,400kg/h. Two simulations of the Goudey boiler were performed. The first (Simulation 1) was a simulation of actual boiler operation while burning an eastern bituminous coal

(different from the coal listed in Table 2 above). Simulation 2 was a hypothetical simulation performed under identical conditions as Simulation 1, but with a different ash chemistry. Consequently, differences between the results from the two simulations are due to chemistry only. In practice, differences in heating value of the coal, volatile content, reaction rate, ash loading etc. will have a significant impact on conditions in the boiler so that Simulation 2 does not represent a realistic situation. It is useful, however, for illustrating the sensitivity of the integrated ash deposition model to ash chemistry. Results of the boiler simulations are shown in Table 3 and Fig. 6. Differences in the ash chemistry were reflected in the magnitude of the average heat fluxes and in a 17 °C difference in the gas outlet temperature. Figure 6 shows the surface temperature of the deposit as a function of position for the right wall of the boiler. Differences in the surface temperature are due to local differences in the thickness/thermal resistance of the deposit. The above simulations illustrate the ability of the integrated deposition model to predict deposit growth in a utility boiler. In addition to identifying local regions where problems associated with slagging are expected to occur, the model predicts the effect of fly ash chemistry on the deposit formation, and the effect of deposit accumulation on the boiler performance.

Modeling of Ash Deposit Growth and Sintering in PC-Fired Boilers

707

4. CONCLUSIONS A deposit sintering model has been developed and integrated into the comprehensive combustion code PCGC-3. Comparisons of model predictions with experimental data from a pilot-scale facility indicate that the model adequately represents key physical processes. The model should therefore be useful for predicting the ash deposition

behavior of a variety of different coals. Utility boiler simulations were used to illustrate the sensitivity of the model to ash chemistry. Future work will focus on gathering more deposition data in order to further validate, improve, and apply the model.

708

H. Wang and J. N. Harb

5. ACKNOWLEDGMENTS Principal funding for this work was provided by the Advanced Combustion Engineering Research Center (ACERC). Funds for the center are received from the National Science Foundation, the state of Utah, 42 industrial participants, and the U.S. Department of Energy. Use of experimental data from Combustion Engineering, Inc. (ABB-CE) taken under the Coal Quality Expert program funded by the U.S. Department of Energy and EPRI is gratefully acknowledged. We also wish to express thanks to ABBCE for fellowship support (H. Wang) and for providing samples from the Fireside Performance Test Facility.

6. REFERENCES Blokh, A. G., Zhuravlev, Y. A., and Tiinkova, S. M. (1992). “An Analysis of Radiant Heat Transfer in Furnace Chambers Burning Slagging Coals.” J. Heat Mass Transfer, 35, 2849–2853. Chappell, J. S., and Ring, T. A. (1986). “Particle Size Distribution Effects on Sintering Rates.” Journal of Applied Physics, 60 (1), 383–391. Coble, R. L. (1973). “Effects of Particle-Size Distribution in Initial-Stage Sintering.” Journal of the American Ceramic Society, 56(9), 461–466.

Frenkel, J. (1945). “Viscous Flow of Crystalline Bodies Under the Action of Surface Tension.” J. Phys., 9(5), 385–390. German, R. M. (1996). Sintering Theory and Practice. New York, Wiley.

Harb, J. N., Slater, P. N., and Richards, G. H. (1993). “A Mathematical Model for the Build-Up of Furnace Wall Deposits.” In J. Williamson and F. Wigley (Eds.) The Impact of Ash Deposition on Coal Fired Plants. Washington, Taylor and Francis.

Hill, S. C., and Smoot, L. D. (1993). “A Comprehensive Three-Dimensional Model for Simulation of Combustion Systems: PCGC-3.” Energy and Fuels, 7, 874–883. Litchford, R. J., and Jeng, S. M. (1991). “Efficient Statistical Transport Model for Turbulent Particle Dispersion in Sprays.” AIAA Journal, 29(9), 1443–1451.

Litchford, R. J., and Jeng, S. M. (1992). “Statistical Modeling of Turbulent Dilute Combusting Sprays.” AIAA Journal, 30(10), 2549 2552. Mackenzie, J. K., and Shuttleworth, R. (1949). “A Phenomenological Theory of Sintering.” Proc. Phys. Soc. London, 62, 833 852.

Richards, G. H., Slater, P. N., and Harb, J. N. (1993). “Simulation of Ash Deposit Growth in a Pulverized CoalFired Pilot Scale Reactor.” Energy & Fuels, 7, 774–781. Scherer, G. W. (1977). “Sintering of Low-Density Glasses: I, Theory.” J. Am. Ceram. Soc., 60, 236–239.

Senior, C. L., and Srinivasachar, S. (1995). “Viscosity of Ash Particles in Combustion Systems for Prediction of Particle Sticking.” Energy & Fuels, 9, 277–283. Slater, P. N., Abbott, M. B., and Harb, J. N. (1996). “Algebraic Interpretation of Composition Phase Classification Criteria For CCSEM.” Symposium on Ash Chemistry, Spring 1996 National Meeting, New Orleans, LA, U.S.A., Match 24–28. Thornock, D. E., and Borio, R. W. (1992). Developing a Coal Quality Expert: Combustion and Fireside Performance Characterization Factors. Report prepared for CQ INC. and U.S. Department of Energy by Combustion Engineering, Inc. Wang, H. (1998). “Modeling of Ash Formation and Deposition in PC-Fired Utility Boilers,” Ph.D. Dissertation in progress, Brigham Young University, Provo, UT. Wang, H.,and Harb, J. N. (1997). “Modeling of Ash Deposition in Large-Scale Combustion Facilities Burning Pulverized Coal.” Prog. Energy Combust. Sci., 23, 267–282.

Walsh, P. M., Sayre, A. N., Loehden, D. O., Monroe, L. S., Beer, J. M., and Sarofim, A. F. (1990). “Deposition of Bituminous Coal Ash on an Isolated Heat Exchanger Tube: Effects of Coal Properties on Deposit Growth.” Prog. Energy Combust. Sci., 16, 327–346. West, J. (1996). Microanalylical Characterization of Ash Deposits From a Coal-Fired Pilot-Scale Combustor. M. S. Thesis, Chemical Engineering Department, Brigham Young University, Provo, UT 84602.

PREDICTING ASH BEHAVIOR IN CONVENTIONAL POWER SYSTEMS Putting Models to Work

Christopher J. Zygarlicke Energy & Environmental Research Center University of North Dakota PO Box 9018 Grand Forks, ND 58202-9018

1. INTRODUCTION Understanding of ash deposit formation mechanisms in conventional power systems is fairly extensive, and models have been devised that allow engineers and operators to minimize ash deposition problems. However, more work is required to bring ash behavior models to practical application in the coal-fired utility industry. With the advent of even more sophisticated coal-fired advanced power systems such as integrated gasification combined-cycle, entrained-flow gasification, and pressurized combustion, it is even more imperative that conventional power system ash behavior models have demonstrated real-world applications. One key factor in the formation of problematic ash deposits is the development of viscous flow properties in partially molten ash. The ability for material to flow and encapsulate or bond ash particles is the single most important factor in ash deposit formation in any system, and it is in this area that predictive modeling has shown the greatest potential for commercial application. This paper focuses on cases where ash deposition models and diagnostic tests have been put to practical use in conventional electric power systems to improve operation and save costs.

2. BACKGROUND A U.S. electric utility company has been experiencing severe ash slagging and fouling in its 425-MW Babcock & Willcox-type cyclone boiler while burning lignite coal from three seams in the Fort Union Group coal basin of North Dakota. The Energy & Environmental Research Center (EERC) took a dual approach, assessing both coal

quality and plant operation in order to define the ash deposition problem. A plan was implemented to formulate mine-planning, coal-blending, and boiler-operating strategies Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

709

710

C. J. Zygarlicke

to mitigate the slagging and fouling. Predictive modeling was used to assess coal quality

impacts. A historical review of the lignite mine and associated minemouth coal boiler operation revealed several areas in the mine that had reputations for causing severe slaggingand fouling-related deratings and outages. Samples of coal from these potential problem areas and other areas in the mine were analyzed in detail to assess coal quality. A boiler inspection and evaluation program was initiated at the power plant to provide up-to-date, objective information on ash deposition severity and the impacts of boiler operating conditions on ash deposition.

3. EXPERIMENTAL

3.1. Mine Coal Sampling and Analysis Samples of three lignite coal seams in each of three separate strip-mined pit areas were analyzed using standard American Society for Testing and Materials (ASTM) methods (proximate, ultimate, and ash elemental oxide composition), computer-controlled scanning electron microscopy (CCSEM), and chemical fractionation. Only two cores will be discussed in detail here: Core 1734, which is located on the west end of the

Orange Pit, and Core 1745, which is located on the east end of the Orange Pit (Fig. 1). Samples were also taken directly from the top of the mined pit exposures. Prior to analysis, the coals were pulverized in a Micron Powder Bantam Jet Mill. Sieve analysis was performed to verify that the particle-size distributions of the test coals were consistent with standard utility boiler practice (i.e., 70%–80% of the coal particles less than 200 mesh). The proximate, ultimate, and coal ash chemical analyses were conducted according to ASTM standards D3172, D3176, and D4326, respectively. A repre-

Predicting Ash Behavior in Conventional Power Systems

711

sentative portion of each pulverized coal was analyzed using the CCSEM method of

quantitative coal mineral analysis. The CCSEM method has been described by Zygarlicke and Steadman [1990] and Galbreath and others [1992], Interlaboratory studies of the method by Casuccio and others [1990] and Galbreath and others [1996] indicate that the repeatability relative standard deviation of CCSEM results (i.e., mineral proportions) is generally less than 20%.

The pulverized coals were also analyzed using chemical fractionation to deduce the distribution of elemental constituents between the mineral and maceral components of coal. Chemical fractionation is a selective leaching procedure based on the differences in solubilities of organically bound and mineral-bound inorganic constituents in stirred solutions of deionized water, 1M Ammonium Acetate and 1M HCl. Elements leached by water are primarily associated with water-soluble minerals (e.g., alkali

halides). Exchangeable ions, principally elements associated with salts of organic acids in lignites, are removed by HC1 removes elements associated with acid-soluble minerals (carbonates, oxides, and metastable sulfides) and organic coordination complexes, such as carboxylate groups on coal surfaces. Elements remaining in the final residue are presumably associated with insoluble silicate and sulfide minerals. The chemical fractionation analyses were conducted according to the procedure described by O’Keefe and Erickson [1994].

3.2. Boiler Sampling and Evaluation Boiler response to changes in coal quality was evaluated by a variety of means. Ini-

tially, a questionnaire was sent to the power plant’s operation staff requesting various design and operating data. This questionnaire was completed during a visit to the plant and provided a format for discussion. Because the plant personnel described a wide range of operating conditions, related to demand and boiler tube ash fouling, it was concluded that a boiler inspection program should be implemented. To accomplish this task, data

sheets were developed, with input from station personnel, to document the severity of boiler tube ash fouling on a shift-by-shift basis. A similar data sheet was developed to

document cyclone burner operation. These data sheets allowed the plant personnel and the EERC to track operating experience in a less subjective manner. In addition to these

data collection techniques, field sampling was performed to collect operating data, boiler tube fouling deposits, furnace exit gas temperature (FEGT) measurements, flue gas analyses, and particulate samples. The protocol for the field tests included using high-velocity thermocouple (HVT) measurements along with an infrared temperature monitor to establish the temperature

profile across the boiler at the entrance to the convective pass. Coal samples were obtained simultaneously from the feeders. Deposits were collected from the leading edge of the convective pass tube banks for analysis. The boiler had a permanently mounted Physical Sciences, Inc. (PSI) optical pyrom-

eter, which produced remarkably stable temperature readings, suggesting that the instrument readings are heavily damped or averaged. In contrast, the EERC pyrometer temperature fluctuations, although significantly low, show very rapid temperature readings up to 150°F above and below the average reading. Similar but less extreme temperature variations were also observed with the HVTs. It is concluded that the FEGT is probably fluctuating approximately 100°F above and below the “average” temperature measured by the installed boiler pyrometer. This is of particular significance, since 100°F excursions above average FEGT could produce quite low-viscosity ash material and lead to significant fouling deposition.

712

C. J. Zygarlicke

A portable gas analyzer was used in conjunction with the HVT to measure the concentration of flue gas components (O2, CO, SO2, and NOx). CO2 concentrations were calculated from the flue gas analysis information provided by plant operators. Samples of coal, fly ash, deposit, and slag collected during the field test were subjected to several analysis techniques. A representative coal sample taken during the time of field test measurements was analyzed in a way similar to that for the coal core samples using CCSEM, IMA (image analysis), and chemical fractionation, as well as proximate and ultimate analyses performed and the bulk ash chemistry determined by XRFA (xray fluorescence analysis). Selected deposit samples were examined by scanning electron microscopy point count (SEMPC), a technique analogous to CCSEM that is used on ash and deposits (Jones and others, 1989).

4. RESULTS 4.1. Coal Analysis Results The coal analysis results are summarized in Tables 1 and 2. Table 1 gives the composition of samples from Cores 1734 and 1745 from the Orange Pit. The three seams are indicated by numbers 1–3, with 1 being the upper seam and 3 being the lowermost. Table 1 indicates that Seam 1 is generally enriched in ash, sulfur, silica, and alumina, but depleted in nitrogen, oxygen, calcium, and sodium relative to Seams 2 and 3. Chemical fractionation analyses in Table 2 indicate generally high proportions of extractable (i.e., soluble in H2O and NH4OAc) calcium, magnesium, sodium, and sulfur in all of the samples. The proportions of extractable sodium consistently range from 70% to 90% of

Predicting Ash Behavior in Conventional Power Systems

713

the total sodium present. The extractable constituents are generally assumed to be organically associated in the coal. The predominant minerals in the coal seams, as indicated in Table 2, are pyrite quartz kaolinite and mixed clays ([Na,K,Ca,Mg, The CCSEM mineral analysis method could account for only about one-third of the inorganic constituents (i.e., total ash) in the samples because of a significant component of very fine-grained minerals. Consequently, a significant amount of the inorganic constituents of the coals occurs as submicron minerals and as organically bound ions dispersed throughout the organic coal matrix. The chemical fractionation and CCSEM analysis results were used to quantify the proportions of the total ash for each sample derived from the pyrometamorphism of discrete supermicron in diameter) minerals, submicron in diameter) minerals, and organically bound ions in the coal. Figure 2 shows that a larger proportion of the Seam 1 coal ash in a given drill core is derived from discrete supermicron minerals relative to Seams 2 and 3. The enrichment of submicron minerals and organically bound inorganics in Seams 2 and 3 is possibly a reflection of downward groundwater movement, which manifests in the boiler as greater fouling propensity.

4.2. Deposit Viscosity Modeling A thermochemical equilibrium and slag viscosity model, termed Phase Ordering and Equilibrium Evaluation (PHOEBE), was used to predict the amounts and viscosities of melt phases that could deposit in the lignite boiler as a function of temperature and coal ash composition. The model uses proximate ultimate and ash chemistry data as input and calculates solid, liquid, and gas phases based on Gibbs free energy minimization. Liquid-phase viscosity is then predicted using a modified Urbain equation.

714

C. J. Zygarlicke

Typical PHOEBE modeling results for the three seams in Cores 1734 and 1745, respectively, are presented in Figs. 3 and 4. Viscosity and proportion of liquid phase are indicators of fouling and slagging propensity, with lower viscosity and greater proportion of the liquid phase indicative of a more molten, “stickier” ash. Seam 2 and 3 coal ashes are predicted to generally produce lower-viscosity slags relative to Seam 1 coal ash for a given temperature. Differences in the viscosity–temperature distributions of the coal ashes may have a profound effect on slagging and fouling propensity. For example, the predicted viscosity of 1734-3 is relatively high over the range of FEGT encountered at the power station (i.e., boiler design FEGT of 1,950°F to measured FEGT of 2,150°F [1,065° to 1,180°C]), but the viscosity of 1745-3 is relatively low throughout the FEGT range. Along with the liquid-phase viscosity values, the amounts of liquid phase at different temperatures may give an indication of deposit sintering potential. Coal from Core 1745 shows a greater abundance on average in all three seams of lower-viscosity material available for sintering compared to coal from Core 1734. Although the calculated viscosities have not been correlated to actual slagging and fouling behavior, it appears that perceived coal quality may be strongly related to operating conditions, especially the FEGT.

4.3. Advanced Indices for Predicting Slagging and Fouling Propensity Advanced slagging, high-temperature fouling (HTF), and low-temperature fouling (LTF) indices were recently developed at the EERC and incorporated into a computer program termed the Predictive Coal Quality Effects Screening Tool (PCQUEST) (Zygarlicke and others, 1996). The indices are expressed numerically as whole numbers ranging from 1 to 100. General classifications of low, medium, and high propensity are

Predicting Ash Behavior in Conventional Power Systems

715

assigned to the following specific ranges of index values: 1–33, 34–66, and 67–100, respectively. A greater value corresponds to an increase in severity or adverse effect for a given

index. The slagging index applies to the amorphous deposits that form primarily on furnace walls and generally consist of aluminosilicate melts with various amounts of basic (fluxing) elements. The HTF index pertains to the partially fused, silicate-based type of deposit formation that generally occurs in the secondary superheater and reheater regions of a coal-fired utility boiler. The LTF index applies to the alkali- and sulfatebonded deposits that occur in the primary superheater and economizer regions of a boiler. The indices are most useful for comparing the relative slagging and fouling propensities of two or more coals or for determining optimum coal-blending ratios. The PCQUEST slagging and fouling indices presented in Fig. 3 were calculated

from the detailed coal analysis results. The index values for Seam 1 coals in a given drill core are generally lower than the corresponding values for Seams 2 and 3. Coals from

716

C. J. Zygarlicke

the 1745 drill core, located at the eastern end of Orange pit (Fig. 1), are predicted to have

the greatest overall propensities for slagging and fouling. Although the PCQUEST slagging and fouling indices are useful for assessing coal quality and developing coal-blending strategies, it is not economically feasible to obtain all of the analysis parameters, such as CCSEM input data, required to calculate the index

values for a large number of coal samples. Consequently, it was imperative to associate the PCQUEST slagging and fouling indices with one or more standard ASTM analysis parameters (Table 1). It was expected that the coal in the mine would be generally similar in mineral composition and physical properties, although the proportions of specific minerals and elements would show significant variation. The correlations between total ash

Predicting Ash Behavior in Conventional Power Systems

717

content and several of the inorganic analysis parameters determined with advanced analysis methods suggested that the development of such a relationship was feasible. The PCQUEST indices, when used on a suite of generally similar coals, such as in a lignite strip mine, tend to be dominated by a few terms in the rather complex index algorithms. These dominant terms, in turn, are associated with, or tend to trend with, the standard ASTM analysis parameters. The PCQUEST indices reported in Fig. 5 were constrained to a range of 1–100 by nonlinear scaling of raw index values that can range from 0 to infinity. The scaling parameters were originally defined to facilitate the comparison of coals with very different compositional characteristics and, hence, slagging and fouling propensities. The current scaling system, however, is not optimized to detect subtle differences in the slagging and fouling propensities of coals with similar compositional characteristics. Further, at high index values (>80), differences tend to be masked by the nonlinear scaling. Consequently, the unsealed index values were found to be more useful for comparing the coals considered in this study and for identifying correlations to standard ASTM analysis parameters. Comparisons of the unsealed PCQUEST index values with standard coal analysis results revealed useful working relationships. Figure 6 shows a well-defined correlation between (HTF index) and %/total ash %. This term is, in fact, used in the index calculation. The correlation indicates that the other terms involved in calculating the HTF index are either insignificant or are scaling similarly. A new scale was then devised for the fouling index according to 0.0–1.0 = low, 1.0–1.5 = moderate, and >1.5 = high predicted potential for adverse deposition. In short, the higher the value is, the greater the predicted propensity for adverse fouling. Similar relationships were developed for slagging deposition in the radiant section of the boiler and lowtemperature calcium sulfate-based ash deposition in the cooler region of the convective pass. After a reasonable fouling and slagging index relationship was found for rather

718

C. J. Zygarlicke

simple coal properties, the mining company’s database of core and pit sample analyses was used to draw a detailed isopach map of the mine seam plan area based on fouling propensity. Previously, mine personnel found that isopach contour maps of coal quality parameters, such as Na % and ash % were useful in interpreting and predicting trends in coal quality, so ample data were available. Calculations were performed using a commercial SURFER ® mapping package with default kriging options used to interpolate values between core data locations [SURFER ® , 1995]. Figure 7 shows the fouling map for Seam 3. Hatch marks on contour lines point “downhill” toward lower values. The areas are marked for Cores 1734 on the west end of the Orange Pit and 1745 on the east end. Coal at 1734 shows lower (low-moderate fouling) values for fouling propensity on the map compared to 1745 (severe fouling), which corresponds to the initial PCQUEST prediction and predicted viscosity data. The lignite mining company now has a strategic mine-planning system that is based on more than just sulfur, ash, and sodium values and includes ash deposition potential. Blending can now be accomplished for mitigating fouling by blending low- and highfouling lignite.

4.4. Comparison of Coal Quality to Boiler Operation Daily coal analyses coming into the plant were used to calculate fouling indices, which were then compared to boiler performance. As mentioned earlier, boiler performance was quantified using inspection sheets designed to document the severity of boiler tube fouling and cyclone burner operation. These were designed to provide a shift-byshift record of boiler cleanliness for the furnace walls, the secondary superheat region, and the reheat region with consistent evaluation criteria. The inspection program was carried out for about 6 months for this project. Additional plant operating information

was logged by the plant system and provided by plant engineers.

Predicting Ash Behavior in Conventional Power Systems

719

The reheater and superheater regions, where most of the troublesome fouling characteristically occurs in this boiler, showed progressive deterioration with time, as shown for the superheater in Fig. 8. Note that an increase in superheater ratings indicates increased fouling. The major trend noted is a general progressive deterioration in boiler cleanliness as the outage date is approached, which is essentially the end of the curve in Fig. 8. Although there is a fair amount of “noise” in the observations because of individual perception and shift-by-shift changes in furnace cleanliness, there appear to be

four individual “step” increases in boiler fouling, followed by stable periods, before the progressive deterioration leading to the outage began. These are indicated by the horizontal lines of Fig. 8. Also marked on Fig. 8 are periods of influx of very high-fouling coal as predicted by the fouling indices. It appears that episodes of an influx of very highfouling coal can cause the boiler to degrade in cleanliness without the ability to fully recover; thus the boiler continues to accumulate ash deposits, which eventually causes an outage for cleaning. Ash deposits were collected from the boiler reheater and analyzed for composition and liquid-phase viscosity. The SEMPC analysis of the deposit indicates that the majority of the deposit consisted of calcium or sodium aluminosilicates, iron—calcium–sodium sulfates, and calcium or magnesium silicates (Table 3). The bulk ash chemistry of the deposits compared to that of the feed coal shows an enrichment in calcium, sodium, iron, and especially a tremendous reaction of organically bound Ca, Na, and Mg with clays, quartz, and sulfur.

720

C. J. Zygarlicke

VISCALC, a program developed at the EERC to calculate point-by-point viscosities from SEMPC data based on a modified form of the Urbain equation, was used to estimate viscosity distributions with temperature for the deposit. The results are plotted for temperatures from 1,800° (980°C) to 2,300°F (1,260°C) in Fig. 9. The deposit shows generally low viscosities across the temperature range of interest. Particularly striking is the percentage of material with viscosity below 250 poise as a function of temperature: 1,700°F (925°C)—2%, 1,800°F (980°C—5%, 1,900°F (1,035°C)—11%, 2,000°F

Predicting Ash Behavior in Conventional Power Systems

721

(1,090°C)—21%, 2,100°F (1,150°C)—34%, and 2,200°F (1,200°C)—44%. This indicates that relatively smal, 1100°–200°F (40–90°C) changes in FEGT would produce dramatic increases in the amount of sticky, low-viscosity material in the deposit. The graph is in agreement with PHOEBE viscosity plots in Figs. 3 and 4, which show that very low viscosity of liquid phases occurs (up to 70% of material less than poise 3) in a temperature region that corresponds closely to the measured FEGT for this boiler of about 2,050°F (1,100°C). The design FEGT for the boiler was only 1,950°F (1,070°C), which means that the current mode of operation of the boiler drives liquid-phase viscosities down to values that are highly conducive to ash sticking and deposit sintering. Plant operators are now using this information about the lignite ash sensitivity to temperature excursions by monitoring FEGT more effectively and experimenting with flue gas recirculation, attemperation sprays, and air balancing to reduce FEGT. Daily ash-fouling prediction, using the modified PCQUEST model, also gives the operators an early warning for “bad” coal in flukes.

5. CONCLUSIONS In summary, the use of thermoequilibrium modeling and advanced fouling indices has proven to be beneficial in helping a minemouth utility to improve performance. In this case, the mine used advanced indices to derive a fouling index contour map to plan blending schemes that would provide the plant with a more consistent blend of lignite that minimizes fouling potential. The plant also benefitted from this scheme by being able to correlate operation parameters such as quantified boiler ash deposition monitoring, sootblower control, gas attemperation, flue gas recirculation, and fuel/oxygen ratios to help extend the operation of the boiler between cleaning outages. This case is more

extreme than most plants encounter, since the coal is very high in sodium and ash deposition and frequent boiler cleaning is inevitable. However, by gaining even just 2–3 extra weeks between cleaning shutdowns, the plant saves in costs. The plant also gains an indirect means of forewarning by analyzing the coal at the mine level in order to plan for lower FEGTs or additional sootblowing when a known high-fouling stream of coal is coming into the boiler.

C. J. Zygarlicke

722

6. REFERENCES Casuccio, G.S., Gruelich, F.A., Hamburg, G., Muggins, F.E., Nissen, D.A., and Vleeskens, J.M. (1990). “Coal Mineral Analysis: A Check on Interlaboratory Agreement.” Scanning Microscopy, 4(2), 227–236.

Galbreath, K.C., Brekke, D.W., and Folkedahl, B.C. (1992). “Automated Analytical Scanning Electron Microscopy and Image Analysis Methods for Characterizing the Inorganic Phases in Coal and Coal Combustion Products.” Prepr. Pap.—Am. Chem. Soc., Div. Fuel Chem., 37(3), 1170–1176.

Galbreath, K.C., Zygarlicke, C.J., Casuccio, G.S., Moore, T., Gottlieb, P., Agron-Olshina, N., Huffman, G., Shah, A., Yang, N., Vleeskens, J., and Hamburg, G. (1996). “Collaborative Study of Quantitative Coal Mineral Analysis Using Computer-Controlled Scanning Electron Microscopy.” Fuel, 75(4), 424–430. Jones, M.L., Kalmanovitch, D.P., Steadman, E.N., Zygarlicke, C.J., and Benson, S.A. (1989). “Application of SEM Techniques to the Characterization of Coal and Coal Ash Products.” In Advances In Coal Spec-

troscopy, Meuzelaar, H., ed., Plenum Press, New York. O’Keefe, C.A., and Erickson, T.A. (1994). “Quantitative XRF Analysis of Coal after Successive Leachings.” Advances in X-Ray Analysis, 37, 735–739. SURFER ® for Windows Version 6 Users Guide. (1995). Golden Software, Inc., Golden, CO. Zygarlicke, C.J., and Steadman, E.N. (1990). “Advanced SEM Techniques to Characterize Coal Minerals.” Scanning Microscopy, 4(3), 579–590.

Zygarlicke, C.J., Galbreath, K.C., McCollor, D.M., and Toman, D.L. (1996). “Development of Fireside Performance Indices for Coal-Fired Utility Boilers.” In Application of Advanced Technology to Ash-Related

Problems in Boilers, Baxter, L., DeSollar, R., eds., Plenum Press, New York, pp. 617–636.

THERMODYNAMIC MODELLING OF THE SYSTEM TO CHARACTERISE COAL ASH SLAGS Evgueni Jak 1 , Sergei Degterov2, Arthur D. Pelton2, Jim Happ3,

and Peter C. Hayes1 1

Cooperative Research Centre for Black Coal Utilisation, Department of Mining, Minerals and Materials Engineering,

The University of Queensland, St Lucia, Queensland, 4072, Australia Centre for Research in Computational Thermochemistry, Ecole Polytechnique de Montreal, P.O.Box 6079, Station Downtown, Montreal, Quebec H3C 3A7, Canada 3 Rio Tinto Research and Technology Development, 1 Research Avenue, Bundoora, Vic, 3083, Australia 2

1. INTRODUCTION In slagging coal gasifiers the mineral matter is present in the reactor in the form of liquid or partially liquid slag. Whilst slagging gasifier flame temperatures are high and most ash constituents are melted completely, the majority of modern coal gasifier designs have water-cooled tube walls. The coal ash slags in contact with these cooled surfaces are completely solidified. The thickness of the slag layer on the walls of the gasifier and the ease of removal of the slag from the reactor will then be determined by the flow characteristics of the slag under the prevailing operating conditions. The flow characteristics of the slag are in turn related to the ash composition, operating temperature and oxygen partial pressure. The design and operation of the plant should be such that small temperature decreases do not lead to large increases in the fraction of solids in the slag exiting the process, and that blockages to the taphole are avoided. In pulverised-fuel boilers, incorrect plant design or operation, or departures from design coal specification can lead to the formation and build-up of ash deposits on the

heat transfer surfaces of boilers causing operating problems. Variations in load can also disturb the temperature profile in the reactor, and initiate slag build-up. The build-up can only proceed in presence of partially molten ash particles or deposits (no liquid phase— no deposit). The behaviour of coal ash in these and other applications is mainly related to the Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

723

E. Jak et al.

724

melting characteristics of the ash. Fundamental knowledge of the melting behaviour of coal mineral matter/flux mixture is essential for design and operation of these power generation technologies. Even for the processes in which the whole system is far from equilibrium, like ash deposit formation, the analysis of local equilibria is beneficial to characterisation of the whole process. Prediction of phase equilibria of the coal ash following combustion is a longstanding problem for users of various coal-based power generation technologies. The industry still uses empirical “ash fusion tests”, such as the cone deformation test, and empirical slagging indices to estimate and compare the relative behaviours of coal ashes. Scientific approaches to the prediction of these properties are, however, starting to produce results which will have a significant impact on the industry. Improvements in chemical thermodynamic models of oxide systems and computational methods now make it possible to predict the phase equilibrium conditions of the coal ash slag systems. The purposes of this paper are to demonstrate a) how the latest advances in chemical thermodynamic modelling can be used to predict the solid, liquid and gaseous phases formed at equilibrium from the mineral matter present in coal combustion systems, and, b) how this information can be provided in forms which are useful to engineering

and marketing practice. The current study characterises the phase equilibria and thermodynamic properties in the five-component system by computer-aided thermodynamic modelling using the computer package FACT [1]. These components represent the major constituents of the coal mineral matter in coals targeted for IGCC usage.

2. DEVELOPMENT OF THE THERMODYNAMIC MODEL The development of a thermodynamic model of a chemical system involves bringing together all thermodynamic and phase equilibrium data available for the system. These data are then evaluated simultaneously to obtain one set of model equations for the Gibbs energies of all phases as functions of temperature and composition. In this simultaneous data evaluation process conflicts between the data sets resulting from inaccurate or inconsistent measurement can be resolved, resulting in an “optimised” model of the system. In this way, all the data are rendered self-consistent and consistent with

thermodynamic principles. From the model equations all of the thermodynamic properties and the phase diagrams within the system can be back-calculated. For the molten slag phase a modified quasi-chemical model has been used. A description of the model and its application to various systems has been reported previously [2–19]. The polynomial and sublattice models were used for the solid solutions in this study. The five-component system contains ten binary, ten ternary and five quaternary sub-systems. The normal way of building up the thermodynamic database is to start from binary systems, then proceed to ternary and so on. Nevertheless, this advance from low to highorder systems has to be modified for two reasons. First, the systems with iron almost always contain both ferric and ferrous iron, so, strictly speaking, there are no experimental data for the sub-systems containing only FeO or Secondly, for a number

Thermodynamic Modelling of the System

to Characterise Coal Ash Slags

725

of low-order systems, there are insufficient experimental data to characterise the thermodynamic properties of all the phases. In that case experimental data for higher-order systems are used to select parameters of the models for the low-order systems. Thermodynamic optimisations with the FACT package of many of the sub-systems including

have been reported previously [3–6, 20]. Alumina was incorporated into the FACT thermodynamic database for the iron—containing systems in the course of the present study [21, 22]. Thermodynamic optimisations were performed for the following systems Thermodynamic and phase equilibrium data on these multi-component systems were used to derive parameters of the models describing the thermodynamic properties of all phases. As a result, a self-consistent database has been developed for the five component system using the FACT computer package. Having completed the optimisation of the system as part of the present study, analysis of the model shows that approximately 95% of all experimental liquidus points in the are reproduced to within Those points having larger disagreement in liquidus temperature are in composition regions where the liquidus surface is very steep, so that the maximum disagreement in composition is approximately 3mol%.

3. FRACTION OF SOLID PHASES AT TEMPERATURES

BELOW LIQUIDUS The information inherent in the thermodynamic model can now be used in many different ways. The combustion engineer may wish to select or evaluate particular coals for use in a slagging gasifier. This is a particular issue for Australian black coals in which the liquidus temperatures are relatively high and flux additions may have to be made or coal blends utilised. The slags which are exiting a slagging gasifier can contain solid as well as liquid phases. The slag composition and operating temperature should be such that a small temperature decrease in the reactor or a small variation in ash chemistry does not lead to a large increase in the fraction of solids in the slag. This is an important consideration because the viscosities of solid/liquid slurries can increase dramatically with increasing solids content, particularly beyond 10–20% solids [23]. The slag should at all times be maintained under conditions where it is sufficiently fluid to be readily tapped from the reactor. The proportion of solids present in a slag system is not only a function of a temperature below the liquidus (undercooling), but of the shape of the phase diagram and the slag composition. This is illustrated in Fig. 1 where a hypothetical eutectic system AB is presented. In Figure 1 for bulk composition decreasing the temperature to 25 degrees Celsius below the liquidus will result, at equilibrium, in the precipitation of 70 weight % solids. In contrast, for the composition a 25 degrees Celsius drop of the temperature below the liquidus will lead to the precipitation of only 10 weight % solids. For compositions and the slag system will be completely solid 25 degrees Celsius below the liquidus. It is clear from these examples that each composition has a unique value of the fraction of solid phases at equilibrium at a fixed undercooling below the liquidus.

726

E. Jak et al.

4. MELTING BEHAVIOUR OF COAL ASH AT TEMPERATURES BELOW THE LIQUIDUS OF COAL ASH FLUXED WITH CAO AND FEO Although the proportion of solids can readily be derived from the phase diagram for binary systems, deriving the same information for a complex multi-component system is not possible without a complete thermodynamic model of the system. For the case of a coal ash of composition 63.1% 27.7% 6.1% FeO, 3.1% CaO under the reducing conditions expected in a coal gasifier, at metallic iron saturation, Fig. 2 and 3 demonstrate how the liquidus and proportions of solids present in the system vary as a function of FeO and CaO flux additions respectively. These data are predicted using the optimised thermodynamic model described above, in conjunction with the FACT computer package. In these figures thick solid lines show calculated liquidus temperatures as a function of FeO or CaO flux additions. The series of thinner lines marked with squares, triangles, circles and crosses present calculated proportions of solids at equilibrium with liquid slag as a function of FeO or CaO flux additions at temperatures 25, 50, 75 and 100 degrees Celsius below liquidus respectively. For the FeO flux addition the liquidus temperatures decrease continuously with flux addition level up to 20g per l00g initial ash (Fig. 2). The proportion of solids (mullite) for a given undercooling decreases continuously with the FeO flux addition (Fig. 2). In the case of CaO flux addition, to the initial coal ash the liquidus temperatures decrease continuously as flux is added up to levels of approximately l0g per l00g initial ash (Fig. 3) in the mullite primary field. As the CaO flux level is further increased above l0g per l00g of initial ash, the liquidus increases as it enters the anorthite primary field. This change in the crystalline phase forming below the liquidus has a parallel influence on the proportions of solids as a function of the CaO flux addition. For all undercoolings up to 100 degrees Celsius the proportions of solids decrease with the CaO flux addition to compositions just short of reaching the anorthite primary

Thermodynamic Modelling of the System

to Characterise Coal Ash Slags

727

field. A dramatic rise in the proportion of solids present at equilibrium at undercoolings of 25 to 100 degrees Celsius occurs before the composition reaches the point marking the boundary between the mullite and anorthite phase fields (between approximately 7.5 and

l0g CaO per l00g of ash). This is due to the co-precipitation of mullite and anorthite with is not readily predicted from examination of the positions of the primary phase fields. The sensitivity of the system increases with increasing undercooling, at 100 degrees Celsius undercooling for example, a relatively small change in fluxing level from 7.5g and 10g CaO leads to an increase from 6% to 26% solids!

728

E. Jak et al.

For slags containing over 10 g CaO flux per l00g of the initial ash the proportion of solids at a given undercooling below liquidus is almost independent of slag composition.

5. OPTIMUM FLUXING OR BLENDING STRATEGIES 5.1. Optimum Fluxing or Blending Strategies for Complete Melting at Known Operating Temperatures The selection of the proportion of fluxes or coals for fluxing or blending is difficult in multi-component systems without information on how the liquidus changes as a function of composition. This information can now be obtained using the FACT computer package [21, 22]. For example, a projection of the liquidus surface on to the pseudo-ternary section FeO-CaOat a weight ratio of 2.2 in equilibrium with metallic iron is shown in this Fig. 4. This pseudo-ternary section construction can be used for the prediction of liquidus temperatures of the slag with a given ratio as a function of CaO and FeO content. Pure calcium oxide CaO and pure iron oxide FeO are represented by two corners of the section and the sum of and at the proportions corresponding to a given ash composition, in this case = 2.2, is placed in the third corner. Typical Australian black coal compositions are indicated by the shaded area in Fig. 4. This pseudo-ternary section construction has been selected so that the slag composition remains on the plane of the section with addition of CaO and FeO. The amount of CaO or FeO required to achieve a certain slag chemistry can be easily read from this diagram by means of the lever rule [21, 22]. This diagram was calculated by using the FACT package with the optimised model for the system outlined above. The metallic iron activity was set equal to unity for the calculations. This corresponds to the reducing conditions which are expected in a gasifier. Almost all iron at metallic iron saturation is present in the state (on average less than 5mol % of all iron in the liquid phase is in the ferric state). The remaining small amount of was converted to that is, the projection is made from the oxygen corner onto the plane FeO-CaOThe “weight of FeO” is actually where and are the weights of the components in a phase, and 0.900 is the ratio of the molecular weights of 2FeO and Liquidus isotherms, marked on Fig. 4, represent all compositions in this section having a given liquidus temperature. The composition areas designated as the primary phase fields (delineated by the thicker full lines—boundary lines) indicate the first solid to form from the liquid slag when cooling from above the liquidus. The arrows on the boundary lines show directions of falling liquidus temperatures. It should be noted that since the solid compounds formed on crystallisation, have compositions which do not lie in the plane of the pseudo-ternary section, the boundary lines and isotherms cannot be used to trace the crystallisation path of the slag in these cases. That is, in general, as solidification proceeds, the composition of the remaining liquid will move out of the plane of the section. For the same reason the “lever rule” does not apply. This pseudo-ternary section has mullite anorthite gehlenite spinel wustite ((Fe,Ca)O), di-calcium silicate tri-calcium silicate and lime ((Ca,Fe)O) primary fields.

Thermodynamic Modelling of the System

to Characterise Coal Ash Slags

729

5.2. Phase Equilibria Below the Liquidus on the Pseudo-ternary Sections Using the optimised thermodynamic model for the system system outlined above, it is possible to predict the proportions of solid phases at temperatures below the liquidus. This information can be presented on the pseudoternary sections. Data for the section at weight ratio of 2.2 in equilibrium with metallic iron for undercooling of 25 degrees Celsius is shown in Fig. 5. The thin

solid lines represent compositions for which the proportion of solids at temperatures 25 degrees Celsius below the liquidus is the same. At 25 degrees below liquidus the unfluxed ash slag system contains less than 5 weight percent solids. The proportion of solids present at 25 degrees below the liquidus decreases as CaO or FeO fluxes are added to the ash. For slags with a weight ratio of 2.2 containing up to 40% total (CaO + FeO) flux at an undercooling of 25 degrees Celsius most of the slags contain less than 10% solids. This value is exceeded in some areas of the anorthite primary field. Data for the section at weight ratio of 2.2 in equilibrium with metallic iron for undercooling of 50 degrees Celsius is shown in Fig. 6. For undercoolings of 50

730

E. Jak et al.

degrees Celsius with up to 40% total (CaO + FeO) flux most of the slags contain less than 20wt% solids. The influence of the anorthite primary field on the slag characteristics is now becoming clear. Not only is the liquidus temperatures raised in the anorthite region of this section relative to surrounding compositions area, but the proportions of solids for a given undercooling are also at higher levels.

Because of the complex solidification path the model is required to predict the com position of the remaining liquid in the system where partial crystallisation has occurred. Knowledge of both the proportion of solids and the composition of the remaining liquid phase is necessary in order to predict the viscosities of these solid/liquid mixtures.

CONCLUSIONS The sensitivity of the coal ash slag system in slagging gasifiers to temperature excursions below liquidus is important in the design and operation of such systems. An opti-

Thermodynamic Modelling of the System

to Characterise Coal Ash Slags

731

mised model for the system has been constructed for use in conjunction with the FACT thermodynamic computer package. The model has been used to predict the liquidus temperatures and the proportions of solid phases at temperatures below liquidus over a wide range of compositions. Calculations of this type can be used to predict optimum fluxing strategies, and to improve operability and availability of slagging coal gasifiers.

ACKNOWLEDGMENTS The authors wish to acknowledge the financial support provided by the Cooperative Research Centre (CRC) Black Coal Utilisation, which is funded in part by the Cooperative Research Centres Program of the Commonwealth Government of Australia.

E. Jak et al.

732

REFERENCES 1. Bale C.W., Pelton A.D. and Thompson W.T. (1996). “Facility for the Analysis of Chemical Thermodynamics” (FACT), Ecole Polytechnique, Montreal, Canada. 2. Blander M. and Pelton A.D.(1984). “Analysis and prediction of the thermodynamic properties of multicomponent silicates.” Proceedings of the Second International Symposium on Metallurgical Slags and Fluxes, TMS-AIME, Warrendale, PA, 295–304. 3. Blander M. and Pelton A.D. (1987). “Thermodynamic analysis of binary liquid silicates and prediction of ternary solution properties by modified quasi-chemical equations.” Geochimica et Cosmochimica Acta, 51, 85–95. 4. Blander M., Pelton A.D. and Eriksson G.(1992). “Analysis and predictions of the thermodynamic properties and phase diagrams of silicate systems.” Proceedings of the Fourth International Symposium on Metallurgical Slags and Fluxes, Iron Steel Institute of Japan, Tokyo, 56–60. 5. Eriksson G. and Pelton A.D. (1993). “Critical evaluation and optimisation of the thermodynamic properties and phase diagrams of the and systems.” Metallurgical Transactions B, 24B, 795–805. 6. Eriksson G., Wu P., Blander M. and Pelton A.D. (1994). “Critical evaluation and optimisation of the thermodynamic properties and phase diagrams of the and systems.” Canadian Metallurgical Quarterly, 33(1), 13–21.

7. Pelton A.D. and Blander M. (1984). “Computer-assisted analysis of the thermodynamic properties and phase diagrams of slags.” Proceedings of the Second International Symposium on Metallurgical Slags and Fluxes, TMS-AIME, Warrendale, PA, 281–94. 8. Pelton A.D. and Blander M. (1986). “Thermodynamic analysis of ordered liquid solutions by a modified quasi-chemical approach—application to silicate slags.” Metallurgical Transactions B, 17B, 805–15. 9. Pelton A.D. and Blander M. (1988). “A least squares optimisation technique for the analysis of thermodynamic data in ordered liquids.” Calphad, 12(1). 97–108. 10. Pelton A.D. and Eriksson G.(1988). “A thermodynamic database computing system for multicomponent glasses.” Advances in the Fusion of Glasses, American Ceramic Society, Westerville, OH, 27.1–29.1. 1 1 . Pelton A.D., Thompson W.T., Bale C.W. and Eriksson G. (1988). “Phase equilibrium calculations in multicomponent systems.” Advances in Phase Transitions, Embury J.D. and Purdy G.R., eds., Pergamon Press, New York, NY, 52–67. 12. Pelton A.D., Eriksson G. and Blander M. (1989). “Quasi-chemical model for thermodynamic properties of multicomponent slags.” Proceedings of the Third International Symposium on Metallurgical Slags and Fluxes, The Institute of Metals, London, 66–69. 13. Pelton A.D., Eriksson G. and Blander M. (1992). “Modelling and database development for the thermodynamic properties and sulfide capacities of slags.” Proceedings of the Fourth International Symposium on Metallurgical Slags and Fluxes, Iron Steel Institute of Japan, Tokyo, 79–84. 14. Wu P. (1992). “Optimisation and calculation of thermodynamic properties and phase diagrams of multicomponent oxide systems.” Ph.D. Thesis, Ecole Polytechnique, Montreal. 15. Wu P., Eriksson G., Pelton A.D. and Blander M. (1993). “Prediction of the thermodynamic properties and phase diagrams of silicate systems—evaluation of the system.” Journal of Iron and Steel Institute, Japan, 33, 25 34. 16. Jak E., Liu N., Wu P., Pelton A., Lee H.G. and Hayes PC. (1994). “Phase equilibria in the system PbOZnO-SiO2.” Proceedings of the 6th Aus. I.M.M. Extractive Metallurgy Conference, Brisbane, 253–259. 17. Jak E., Degterov S., Pelton A.D. and Hayes PC. (1997). “Thermodynamic modelling of the system PbOfor lead/zinc smelting.” Proceedings of the Fifth International Symposium on Metallurgical Slags and Fluxes, Iron and Steel Society, AIME, Sydney, 621–628. 18. Jak E., Degterov S., Hayes PC. and Pelton A.D. (1997). “Thermodynamic optimisation of the systems CaO-Pb-O and PbO-CaO-SiO2.” Canadian Metallurgical Quarterly, accepted June 1997, in press.

Thermodynamic Modelling of the System

to Characterise Coal Ash Slags

733

19. Jak E., Degterov S., Wu P., Hayes P.C. and Pelton A.D. (1997). ”Thermodynamic optimisation of the systems PbO-ZnO, and Metallurgical Transactions B, 28B, 1011– 1018. 20. Pelton A.D. (1994). “Thermodynamically optimised phase diagrams of oxide systems.” Final report, submitted to “Phase Diagrams far Ceramists”. 21. Jak E., Degterov S., Hayes P.C. and Pelton A.D. (1997). “Thermodynamic modelling of the system

to predict the flux requirements for coal ash slag.” Fuel, 77(1/2), 77–84. 22. Jak E., Degterov S., Pelton A.D. and Hayes P.C. (1996). “Predicting flux requirements for coal ash slags.” Proceedings of the Workshop on Impact of Coal Quality on Thermal Coal Utilisation, CRC Black Coal Utilisation, Brisbane, Australia, paper 16. 23. Jinescu V.V. (1974). “The rheology of suspensions.” International Chemical Engineering, 14 (3), 397–420.

This page intentionally left blank.

DEVELOPMENT OF A PREDICTION SCHEME FOR PULVERISED COAL-FIRED BOILER SLAGGING Jouni Pyykönen, Jorma Jokiniemi 1 , and Tommy Jacobson2 1

VTT Energy, Aerosol Technology Group P.O. Box 1401, 02044 VTT, FINLAND 2 IVO Technology Centre Rajatorpantie 8, 01019 IVO, FINLAND

1. INTRODUCTION All coals have a significant content of ash-forming inorganic material. The accumulation of deposits derived from the mineral matter can lead to substantial financial losses to an operator as a result of reduced thermal efficiency, reduced availability and high maintenance costs due to blockage, erosion and corrosion. Finnish power plants have traditionally imported coal feedstocks with low slagging propensities. However, there is an increasing tendency to use coals of inferior quality from several different sources on a shorter term basis. This has created a need to be able to cope with moderate or high slagging propensity coals. This paper describes a slagging prediction scheme with a focus on modelling the impact of furnace flow fields and burner arrangements. The emphasis of this paper is on the description of the modelling approaches and the model structure. The use of a slagging coal sometimes necessitates boiler modifications, such as burner arrangement retrofits. Based on operational experience, it is known that coals which cause severe slagging problems with one burner arrangement are quite manageable with another arrangement. The effects of the burner configurations and the associated flow patterns are further exemplified by the fact that the introduction of low-NOx firing systems has not exacerbated furnace slagging problems in spite of the local reducing conditions they have brought about. This is mainly due to lower local heat fluxes in the burner belt [Gibb, 1996]. There has been a considerable effort in the 80’s and 90’s in the development of models for the prediction of the slagging propensities of various coals based on their inorganic constituents [e.g. Helble et al., 1992; Baxter, 1993]. The development of models for the assessment of the impact of boiler specific factors, such as the flow and temperImpact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

735

736

J. Pyykönen et al.

ature fields, has received much less attention. However, in recent years, a number of computational fluid dynamics (CFD) based slagging prediction schemes have been reported. Richards et al. (1993) modelled slagging in an axisymmetric pilot-scale reactor using stochastic trajectories. The model considered ash particle impaction rates, sticking and timedependent evolution of the deposit layer properties. Boyd and Kent (1994) identified the role of burner swirl with particle tracking techniques. Mann et al. (1995) modelled the introduction of an air-curtain as a measure to combat wall-slagging in a front and back wall fired industrial furnace. A slag was assumed to form when the temperature of an impacting particle was above 1,370 °C. Lee et al. (1996) studied a single-burner ash deposition rig as a validation case and a full-scale front wall fired furnace. In the model, fly

ash particles were tracked until they reached the edge of the viscous sublayer. If a particle possessed a critical velocity to penetrate the viscous sublayer, it impacted onto the wall, otherwise it was considered to be trapped in the boundary layer. The predictions of the deposit compositions were found to be quite similar to bulk ash chemistry with some enrichment in iron in the initial deposit layers. Huang et al. (1996) concentrated on superheater tube deposit formation in a pilot scale furnace, which they modelled with CFD particle tracking techniques. All these modelling studies included comparisons with experimental data and qualitative agreement was reported in each case. The aim of the current research has been the creation of a CFD-based tool for industrial use in slagging prediction. More generally, we want to understand the phenomena involved in the formation of tenacious deposits and their connection to boiler design and operation. Simulations of the flow and temperature fields with the Ardemus CFD-based furnace modelling system [Kjäldman, 1993] have proved useful in the design of burner modifications. We have enhanced the modelling capabilities of the Ardemus system with the inclusion of a slagging prediction scheme. The model can indicate the regions that are vulnerable to slagging for various burner configurations, and thus assist

the design of the modifications. Novel features in our approach in relation to other models include a two-stage modelling strategy of the deposition by turbulent impaction applicable to relatively coarse computational grids, and a semi-mechanistic representation of coal mineral-to-ash transformation processes as a part of the CFD-based scheme. The model is restricted to slagging deposits in the radiant section of a boiler. The first results are related to submodel testing, partly carried out in conjunction with an industrial scale multi-burner furnace Ardemus simulation.

2. MODELLING SYSTEM 2.1. Overall Scheme The basis of our CFD studies is Ardemus, a Fluent-based pulverised coal fired boiler simulation system [Kjäldman, 1993], which includes sophisticated models of char

particle reactions relevant to pulverised coal combustion. Ash particles have some effect on the overall flow and temperature fields mainly due to their role in radiative heat trans-

fer. In the case of a slagging coal, the effects of temperatures and radiative properties of the slag layers are important, and cannot be neglected in the flow modelling. In Ardemus flow simulations, the absorption and the scattering of radiation due to ash particles are taken into account assuming typical pulverised coal fly ash properties. The temperature and radiative properties of the slag layers are simply given as boundary conditions, and slagging is modelled by post-processing Ardemus flow simulation results. For deposition modelling, we use a two-stage strategy. In the first stage, particle tra-

Development of A Prediction Scheme for Pulverised Coal-Fired Boiler Slagging

737

jectories are used for determining particle concentrations in the next-to-wall computational cells. In the second stage, deposition rates in these cells are determined based on

the predicted concentrations. Deposition rates, expressed in are obtained by multiplying the next-to-wall cell particle concentrations with deposition velocities expressed in m/s. Deposition velocities are, thus, determined independently of other transport mechanisms. Two cases are considered. The first is a superheater tube in a cross flow subject to deposition by inertial impaction, and the second is a furnace wall such as burner belt refractory wall or a furnace water wall. With this approach, very fine computational grids are not necessary for the boundary layer trajectory calculations. An exception to this rule is inertial impaction (as opposed to turbulent impaction) on furnace walls which is calculated solely based on trajectories. Thermophoresis is not modelled explicitly since the concentration of particles smaller than a few microns is low on a mass basis and the deposit initiation phase is not considered in the model. For larger particles, based on the results of Lind et al. (1997), thermophoresis seems to be less important than turbulent impaction. Slag layer growth rates not only depend on particle arrival rates, but also on sticking probabilities at the surface. The sticking of an ash particle on a heat transfer surface is a complex phenomenon, which depends on both the particle and the surface properties. The growth of the slag layer is further complicated by the fact that the sticking efficiencies vary as a function of the thickness of the layer. In addition, the shapes of the slag layers and the associated flow profiles are not constant. Our approach is to consider only the idealised cases of 1) a clean surface (with a thermophoretically formed initial layer) and 2) a surface which captures all impacting particles, such as a molten slag layer, and use them as indicators of the overall growth process. Gibb (1996) broke the slagging process down to three stages: deposit initiation, bulk ash deposition and deposit consolidation. The key to furnace slagging, according to Gibb, is the long term behaviour of the deposits, i.e. the consolidation process. To characterise the degree of consolidation we propose a sintering index, which gives an indication of the tenacity of the deposits, i.e. their resistance to sootblowing and gravitational shedding. Overall, the model gives three indices for slagging propensity assessment. Two assess the deposit layer growth rates at clean and sticky surfaces and the third provides the sintering index. In addition, heat flux to the wall is estimated as a part of the flow and combustion model. The heat flux is an important factor when considering the properties of the deposit layers. For general burner arrangement studies, we assume typical problematic coal fly ash properties. For coal specific studies, we can use either data from a fly ash computer controlled scanning electron microscopy (CCSEM) analysis or from a coal mineral composition CCSEM analysis as inputs to the calculation. The direct use of fly ash data eliminates the need to model the transformations that generate fly ash particles out of minerals in coal particles. The difficulty, in this case, is to obtain a representative fly ash sample. On the other hand, a mineral-to-ash transformation model makes it possible to use the CCSEM analysis of coal as the input to the calculations. This strategy also makes it possible to model the dynamic phenomena involved in the transformations, i.e. pyrite conversion and iron oxidation.

2.2. Ash Particles Trajectories The standard Ardemus furnace simulations provide char particle trajectories. When char burnout is complete, we continue the trajectories as fly ash particle trajectories. The

738

J. Pyykönen et al.

mean particle trajectory is calculated by solving the equation of motion for a single particle:

The trajectory calculations include stochastic velocity components calculated on the basis of the turbulence properties of the flow. We have used both the discrete and the continuous random walk models, which are a standard feature in Fluent [Fluent, 1997]. The solution procedure also keeps track of the temperature of the particle based on the following equation of heat transfer for a single particle:

The term representing radiative heat transfer is not important for particles smaller than in pulverised coal combustion applications [e.g. Kjäldman, 1993]. The term represents the heat addition due to combustion of the char particles. To maximise the amount of information retrieved from each trajectory calculation, a concentration variable C is attached to each trajectory. This represents the probability that the particle has not deposited onto furnace walls or superheater tubes. At the beginning, it is initialised to a value of one, and each time a trajectory passes a next-to-wall cell, a fraction representing the depletion in particle concentration by deposition is subtracted from it. The depletion is calculated based on the deposition velocity the particle residence time in the cell the area in the cell subject to deposition A dep and the cell volume Vol with the following equation:

If a trajectory hits a wall, a rebound is assumed, since trajectories are not used for estimating deposition velocities. The exception is the case of inertial impaction, which is discussed in the next section. Fly ash particles that contain considerable amounts of iron are heavier than other fly ash particles of equal size. To account for this, separate trajectories are computed for high-iron ash particles. With these methods, the number of trajectories required to obtain representative next-to-wall cell ash particle concentrations in industrial multi-burner boilers simulations is on the order of several tens of thousands.

2.3. Deposition Velocities The second step in the two-stage deposition modelling is to calculate the arrival rates, i.e. the deposition velocities, of different particle sizes onto heat transfer surfaces. Models of inertial impaction onto a tube in cross flow are a standard feature in slagging/fouling models. The distinction in our model is that parameters, such as gas velocity, are obtained directly from a CFD model. Wessel and Righi (1988) have derived impaction efficiency correlations for tubes in a cross flow based on numerical trajectory calculations. We have adopted their correlations following the example of a number of others [e.g. Allan et al., 1996]. This strategy only provides a rough approximation of the

Development of A Prediction Scheme for Pulverised Coal-Fired Boiler Slagging

739

actual impaction velocities. For instance, for tubes with a deposit layer on top of them, the impaction velocities are lower since the geometry is more streamlined [e.g. Yilmaz, S. and Cliffe, K., 1997]. The parameters for the correlation (particle density and size, free stream cross-flow gas velocity and viscosity) are obtained from standard Ardemus flow simulations supplemented with fly ash particle trajectory calculations. No special grid spacing for the effect of the superheater tubes is used. Instead, particle depletion is calculated by having a sink term for the particle trajectory concentration C in the superheater area. The locations, spacings and sizes of the superheater tubes are given as inputs to the computational scheme. Only the first superheater tube rows are considered in the present model. As deposition mechanisms onto furnace walls, we consider inertial impaction and turbulent impaction separately. Comparing the dimensions of the furnace with the dimensions of a superheater tube and the associated bends in the streamlines, one can infer that the cut size for inertial impaction onto furnace walls is much larger. Inertial impaction can be important for the larger particles of a typical pulverised coal fly ash size distribution when velocities are high and streamlines bend sharply as might occur in burner belts. The inertial impaction of ash/char particles onto furnace walls is estimated when their trajectories enter the next-to-wall cells. At that point, a particle trajectory is calculated without stochastic turbulent velocity components based on equation (1). If this trajectory reaches a furnace wall, the particle impacts inertially. Otherwise, a standard trajectory with turbulent velocity components is computed and used instead. The accuracy of the predictions depends on the near-wall grid spacing. Even relatively coarse grids can give predictions that are accurate enough for our purposes as long as they are able to capture the characteristics of the flows that impinge on furnace walls. A more difficult task is to model deposition by turbulent impaction onto furnace walls. To start with, there are no proper models for turbulent impaction in complicated geometries, and in the case of a slagging furnace wall, even the geometry is not well defined. The difficulties in the development of models of turbulent impaction arise from representing the coherent structures of turbulence and bursts in the turbulent boundary layer, especially in the strongly inhomogeneous buffer layer Kallio and Reeks (1989) have provided a sophisticated mechanistic representation of the phenomena leading to turbulent impaction. They simulated a number of stochastic trajectories in a random distribution of turbulent eddies. The properties and distribution of the eddies were taken from experimental observations. The use of their approach, however, is not appropriate here since it requires detailed data on the structure of turbulence in boundary layers, which is not available for furnace walls. Other mechanistic turbulent impaction models, such as those presented in Im and Chung (1983) and Jokiniemi et al. (1996), are based on a number of assumptions, the validity of which is questionable in furnace wall conditions. Instead of using a mechanistic model as a part of the slagging prediction scheme, we simply resort to an experimentally observed functional relationship between the dimensionless deposition velocity and the dimensionless particle relaxation time i.e. This relation has been observed to remain approximately constant at different flow rates in turbulent duct flows. The nondimensionalisation is made with respect to boundary layer properties (see the nomenclature for details). There is some scatter in the experimental data at lower relaxation times. Additional parameters, such as surface roughness and the ratio of particle to fluid density, have been used to account for this scatter via phenomena such as interception effects. A recent correlation form of the relation proposed by Muyshondt et al. (1996) was adopted for this study. The correlation

740

J. Pyykönen et al.

was obtained by measuring the penetrations of a number of different sized monodisperse

particles through pipes of various diameters and flow rates. They used the duct Reynolds to remove the scatter in experimental data at low dimensionless particle relaxation times resulting in a form The correlation is as follows:

number

where the values of the coefficients are For the duct Reynolds number, we have somewhat arbitrarily adopted the value of This value allows significant deposition for smaller particles, and thus, implicitly takes into account the effects of surface roughness and thermophoretic forces. The differences due to this “Reynolds number effect” are significant only at low particle relaxation times. At values of larger than 100 (which is out of the valid range for equation 4), the relationship used in this study is based on the results of Kallio and Reeks (1989). The whole correlation we have adopted is presented in Fig. 1. For turbulent impaction modelling, the key parameter required from an Ardemus simulation is the shear stress at the wall. Commercial CFD codes, such as Fluent, provide a wall functions option for the near-wall modelling. Wall functions are based on

an assumption of the shape of the velocity profile in the boundary layer, and they make it possible to obtain estimates of the shear stresses even on a relatively coarse grid. Nevertheless, Ardemus furnace grids are too coarse in the near-wall regions to capture the effect of the water wall tube structure and the degree of grid refinement required for their modelling is inappropriate for practical furnace studies. Moreover, if a slag layer builds up on the tube surface, the geometry of the wall changes with time. In line with the overall concept of considering only clean and sticky surfaces, we examine 1) an unslagged wall tube structure and 2) a flat slag layer. To account for the effect of the tube structure, a series of detailed flow simulations of a single water wall tube has been carried out. The flow fields have been studied at different tube orientations and flow velocities using periodic boundary conditions. We have tried a number of different grid configurations. In

Development of A Prediction Scheme for Pulverised Coal-Fired Boiler Slagging

741

one configuration, the boundary layer was solved numerically with the first grid point at

and used the two-layer zonal model of Fluent [Fluent, 1996] with the RNG turbulence model in the fully turbulent region and the Wolfstein one-equation model in the viscosity-affected near-wall region. This grid is presented in Fig. 2. The idea is to examine the overall flow pattern around a tube for various conditions and, for a first approximation, to derive correlations with which we can estimate the average “windward” side shear stress at the tube wall based on the shear stress obtained with a coarse grid using a flat wall approximation. Predictions of turbulent impaction must be regarded only as first estimates of deposition velocities. The main difficulty is that most studies on turbulent impaction, both experimental and computational, have only considered turbulent duct flow. Nevertheless, theoretical modelling studies have demonstrated that it is the buffer layer structure which determines turbulent impaction velocities and boundary layer theories hold fairly well at buffer layer distance. Therefore, the adopted correlation should give a reasonable indication of the degree of turbulent impaction for flat walls or for flows parallel to the water wall tubes. Furthermore, Yau and Young (1986) have used a somewhat similar approach for modelling turbulent impaction velocities on steam turbine blades and obtained results that are in good accordance with measurements. Here, the overall idea is not to model single tube slagging in a detailed manner, but to indicate spots that are vulnerable to slagging on a larger scale. The main uncertainty, in our case, arises from the representation of the effect of the tube structure in cases with a significant flow component perpendicular to the tubes. The use of the average shear stress as the basis for the turbulent impaction correlation is quite a rough approximation. We justify it by the experiences gained with a similar approach in the case of deposition in pipe bend flows [Kissane, 1996]. Still, in case of a flow perpendicular to the tubes, the effects of flow separation and reattachment on the structure of near-wall turbulence render the foundations of the average shear stress based approach rather weak.

2.4. Sticking Efficiency and Sintering Index In the case of a clean surface, the sticking efficiency of ash particles is estimated with the model of Walsh et al. (1992), which has been adopted also in the CFD-slagging

742

J. Pyykönen et al.

models of Richards et al. (1993), Lee et al. (1996) and Huang et al. (1996). The model is based on the concept of a critical viscosity for complete sticking. Overall, the sticking efficiency is calculated as follows:

A number of different values have been suggested for critical viscosity as discussed by Huang et al. (1996). We have adopted the value of Pa s following the example of Lee et al. (1996). This simple sticking model does not take into account the role of flow fields on sticking and re-entrainment is not considered. These are important phenomena when modelling the formation of deposit layer shape on a single tube. In addition, the

effects of deposit microstructure are neglected. Therefore, deposition rate predictions must be regarded as indices. In some cases, a deposit layer may build up and consolidate in a stagnant flow region due to the absence of forces that remove particles. For this case, however, the model would predict low deposition rates. This is a topic for future model development and a point to consider in the interpretation of model results. Large particles do not necessarily cool fast enough to achieve the wall temperature as they traverse the boundary layer. The temperature at which the viscosity is evaluated

for the calculation of the sticking efficiency is therefore a function of the particle size. Because of our two-stage modelling strategy for deposition, additional considerations are

necessary to estimate the temperatures of the particles at the point they reach the wall. We do this by solving the particle heat transfer equation (2) numerically along a separate

boundary layer trajectory using a particle speed obtained from the deposition models and assuming either a laminar or turbulent boundary layer temperature profile. In the case of inertial impaction on a superheater tube, the trajectory is calculated at the stagnation point assuming a Blasius profile for the laminar boundary layer temperature. The velocity of an impacting particle is obtained from the correlations of Wessel and Righi (1988). In the case of the turbulent boundary layer, the temperature profile is assumed to obey the logarithmic law when is greater than 20, the linear law when is less than five and an interpolation formula in between. The trajectory is assumed to end at the height of the surface roughness. For turbulent impaction, the deposition velocity of a single ash particle is not the same as the average deposition velocity obtained from the adopted correlation. Considering that half the time turbulent eddies are in a downsweep motion and half the time in an upsweep motion, we assume, as a lower bound approximation, that the deposition velocity of a single particle is twice the average deposition velocity. Most of the fly ash particles generated in pulverised coal combustion are glassy aluminium silicate particles. Their consolidation occurs by viscous flow sintering, which depends on temperatures, viscosities, sizes and surface tensions of individual ash particles. Compared with viscosity and size, surface tension only slightly varies between various ash particles. Therefore, we characterise the degree of consolidation by a sintering index, which is based on ash particle sizes and viscosities. We assume that we can characterise the role of viscosity in the sintering process in a manner analogous to the sticking process, i.e. we assume the existence of a critical sintering viscosity with maximum sintering efficiency. To get the contributions of different particle sizes to the overall sintering process we use particle surface areas as weighting factors. The overall scheme for the sintering index, based on N ash particles that have deposited and stuck on a wall cell, is as follows:

Development of A Prediction Scheme for Pulverised Coal-Fired Boiler Slagging

Sintering

743

(Sintering efficiencyi·

The viscosities of individual ash particles are evaluated at the wall or deposit layer temperature, which is given as an input in the present scheme. For meaningful localised data, this temperature should be estimated with a deposit layer heat transfer model. At

this stage, however, we have deferred all deposit layer modelling to future studies. For the critical sintering efficiency, we assume a value of Pa s as our starting point. An inverse square root dependence of the sintering efficiency on the ratio of actual to critical viscosity is tentatively assumed. The square root makes the relation less sensitive to the choice of the critical value. The sintering index bears some similarity to the silicate strength variables presented in Allan et al. (1996). The viscosities of the aluminium silicate based particles are evaluated with the model of Senior and Srinivasachar (1995), which is a modification of the well-known Urbain viscosity model. Other minerals are assumed to be either of very low or of very high viscosity based on their melting temperatures. For pyrite and its products we make the same assumption on the basis of the stage of the transformation process.

2.5. Mineral-to-Ash Transformation Model The slagging prediction scheme provides an option to study the behaviour of a certain coal based on a computer controlled scanning electron microscopy (CCSEM) mineral composition analysis. The phenomena involved in the transformations of coal minerals into ash particles are complicated, the CCSEM analysis and its interpretation are fraught with difficulties and the viscosity models are not very accurate. Therefore, it is not surprising that the success of the purely mechanistic representation of the transformation processes has been limited. Realising these difficulties, the overall approach adopted here for coal specific studies is somewhere between the mechanistic models and semi-empirical slagging indices. For instance, we only use CCSEM mineral identifications of pyrite, quartz and calcite. Otherwise, we directly use EDX elemental composition data. The coalescence behaviour and the viscosity of, for example, an illite mineral particle are assessed based on its elemental contents of K, Al and Si (and impurities) as opposed to using illite specific properties. Since we have limited our attention to bituminous coals, organically bound inorganic constituents in coal are not taken into account in the present model. As input for the deposition models, the mineral-to-ash transformation model provides ash particle size distributions and viscosity distributions as a function of particle size. An alternative to the mineral-to-ash transformation model for coal specific studies is the use of fly ash CCSEM data. The viscosity distribution of fly ash is also calculated using EDX elemental composition data. Our mineral-to-ash transformation model for included minerals is based on the assumption of a single ash particle per coal particle, which according to Sarofim and Helble (1994) is a good starting point for ash particle size distribution calculations. High volatility coals may form cenospheric char that fragments. For such coals, an estimate of the degree of fragmentation should be modelled separately. Included mineral particles of various sizes and compositions detected by the CCSEM are randomly distributed among coal particles of different sizes. In reality, however, the distribution is not always random

744

J. Pyykönen et al.

[Slater et al., 1995]. The size distribution of the injected coal particles is usually known reasonably well. All minerals are assumed to become completely assimilated with aluminium silicate minerals, with the exception of pyrite and quartz. For them we use a fluxing effectiveness between 0.5 and 1.0, depending on the amount of clay present. These assumptions are based on the results and reasoning of Gibb (1996) and Helble et al. (1990). The size of the mineral inclusion may also have an effect on fluxing effectiveness, but that is not modelled. Although the fragmentation of included minerals is not thought to be an important process, reactive extraneous minerals have been observed to fragment in experiments relevant to pulverised coal combustion. A constant fragmentation behaviour for pyrite and calcite particles is assumed. A rough correlation for the degree of fragmentation as a function of particle size has been estimated from the results of Helble et al. (1990). The roles of the kinetic phenomena of pyrite conversion and aluminium silicate iron oxidation can be modelled conveniently with CFD. Pyrite has been identified as a major source of wall slagging especially in the near-burner region [e.g. Bryers, 1996]. Under pulverised coal combustion conditions pyrite decomposes to pyrrhotite, which subsequently melts to form an Fe-O-S melt. Until the Fe-O-S melt droplet oxidises and crystallises as magnetite it remains sticky. The period in which the particle remains sticky depends on its size and its residence times in various temperatures and oxygen con-

centrations. We have modelled these transformations based on the model of Srinivasachar

and Boni (1989). As a simplification to their model, we do not consider the stages that lead to the formation of a Fe-O-S melt. Instead, we simply start the calculation once the temperature required for pyrrhotite melting (1,356K) is reached. The oxidation of the melt to magnetite is modelled as a diffusionally limited process. Both gas phase and liquid

phase oxygen diffusion are considered. The oxidation rate is given by the following equation:

The mass transfer coefficient includes the diffusional resistances and a conversion factor from moles of to mass The crystallisation is assumed to begin when the magnetite particle has supercooled to 1,600K. In the case of pyrite mineral inclusions embedded in coal particles, the oxidation can begin only after the inclusions are exposed by the receding char surface. Bool et al. (1995) have presented a detailed model of the mechanisms involved. As a simplification, we calculate the Fe-O-S melt surface area based on the amount of char in the burning coal particle assuming a homogeneous distribution of pyrite within the original coal particle. Char combustion data is obtained from Ardemus char combustion models. As the oxidation state of iron affects the viscosity of glassy aluminium silicate particles, we also model the kinetics of iron oxidation. Bool and Helble (1996) have shown that iron oxidation is limited by oxygen diffusion inside glassy ash particles. We have

adopted their model of the conversion of Fe(II) to Fe(III). The rate of change of the volume fraction of glassy ash that has reacted to Fe(III) is calculated as follows:

Development of A Prediction Scheme for Pulverised Coal-Fired Boiler Slagging

745

For the oxygen concentration we assume a value of 90% of the oxygen solubility in molten glass. This value gave the best agreement at the highest measured temperature of 1,400°C in the experimental and modelling studies of Bool and Helble (1996).

3. SUBMODEL TESTING A correlation for the average windward side shear stress enhancement due to wall tubes has been established with Fluent simulations and is presented in Fig. 3. At a flow

angle of 0°, the flow is parallel to the tubes. At a flow angle of 90° the flow is perpendicular. No significant velocity dependence was identified. The correlation should give reasonable estimates of average wall shear stresses. The overall effects of the inaccuracies in the shear stress enhancement correlation are, in any case, secondary to those due to the uncertainties in the applicability of our turbulent impaction model to this geometry. Figure 4 shows how the adopted turbulent impaction correlation converts to typical pulverised coal-fired furnace wall conditions at various shear stress levels. A representative clean wall temperature of 700 K and a particle density of were assumed. Furnace wall shear stresses in areas with major near-surface flows are typically between 0.01 Pa and 0.1 Pa. In the near-burner areas they may be even higher. At shear stress levels between 0.01 Pa and 0.1 Pa, turbulent impaction velocities range from 1 to 5cm/s for particles. For particles in these conditions, the deposition velocities are below 1 cm/s. The results show that, in general, as the particle size increases, the corresponding turbulent impaction velocity increases as well. Large particles do not precisely follow buffer layer turbulent eddies but can deviate from them and reach the wall. In areas of shear stress as high as 0.1 Pa, however, deposition velocities are larger for particles than for particles. The reason for this is that the particle size corresponding to the nondimensional particle relaxation time with maximum turbulent

746

J. Pyykönen et al.

impaction velocity is closer to than to Under high shear stress, a particle is so large that it does not react to all turbulent fluctuations in the boundary layer. Turbulent impaction velocities under the furnace wall conditions outlined above are presented as a function of particle size in Fig. 5. A common feature at different shear stress levels is that there is an increase in deposition velocities above a certain particle diameter. At shear stress levels between 0.01 Pa–0.1 Pa turbulent impaction velocities are considerably higher for particles larger than than for bulk mass mode fly ash particles (typically between ). Turbulent impaction velocity estimates at the lower particle sizes are not very reliable, since these are sensitive to surface roughness and the choice of the Reynolds number in the correlation (Eq. 4). In reality, these velocities could

Development of A Prediction Scheme for Pulverised Coal-Fired Boiler Slagging

747

be lower. In contrast, much higher velocities are not expected, since the present Reynolds number represents something close to the peak experimentally observed velocities. Overall, turbulent impaction velocities are much lower than those associated with inertial impaction onto a superheater tube. This means that inertial impaction onto the furnace walls cannot be neglected, and even relatively low inertial impaction velocities can lead to deposition rates that are comparable to deposition rates by turbulent impaction. Nonetheless, the overall result is that, in areas with major near-surface flows, particles larger than deposit preferentially over smaller bulk fly ash particles, be it via turbulent or inertial impaction. Considering particle masses, this tendency should lead to the wall deposit enrichment in heavier species, such as those of iron. The models of particle temperature evolution in the boundary layers were tested for inertial impaction onto a superheater tube and turbulent impaction onto a furnace wall. The particle density was assumed to be and the specific heat capacity to be l,000J/(kg K). These values are typical of aluminium silicates. In the case of inertial impaction, the free stream velocity was given values of l0m/s and 5 m/s and the calculations were made at the stagnation point. In the l0m/s case, Fig. 6 shows that, for particles smaller than the temperature at the impact is in the midrange between the free stream and the superheater tube surface temperature while, for particles larger than it is very close to the free stream temperature. Since the impaction efficiencies are low for particles smaller than most of the impacting particles have temperatures close to the free stream temperature. In the 5 m/s case, there are more particles that have impaction temperatures in the midrange between the free stream and the superheater tube surface temperature. In the case of turbulent impaction onto furnace walls, the calculations with similar temperatures produced temperature differences between the impacting particle and the surface on the order of only 1 K even at the relatively high shear stress level of 0.1 Pa. These small temperature differences are a result of low turbulent impaction velocities. A major assumption is that the deposition velocity of a single ash particle is twice the average deposition velocity. However, even deposition velocities that are much higher than the average would not produce significant differences between the particle temperature at the impact and the surface temperature.

748

J. Pyykönen et al.

The turbulent impaction model was tested in conjunction with an Ardemus simulation of an industrial front and back wall fired furnace. Figure 7. shows the deposition velocities by turbulent impaction for particle sizes of and A par-

ticle density of was assumed. The wall was assumed to be flat, i.e. the water wall tube structure correction was not applied in these studies. The results are preliminary in the sense that the computational grid has not yet been adapted to slagging studies. The grid is too sparse in the high shear stress burner area for the effective use of wall functions. Thus, the results are only indicative in that region. For particles, turbulent impaction velocities are low under an assumption of a flat surface. The largest deposition velocities due to turbulent impaction are in the near-burner areas. Large turbulent impaction velocities occur also at the furnace bottom. There is a high deposition velocity zone on the side wall as well, apparently due to flow impingement. The results, once again, indicate the trend that when the flow velocities are higher near the walls, they generate greater turbulence which more effectively transports particles to the walls. The more turbulent the flow is the smaller are the particles that are affected and, up to a point, the larger the particle deposition rates.

4. CONCLUSIONS An extensive CFD-based modelling system for the study of slagging in pulverised coal combustion boilers has been built. Ash particle size and viscosity distributions are given input values that are typical of problematic coals. For coal specific studies, these inputs can be estimated based on CCSEM analysis of fly ash or coal minerals. For the case of coal mineral analysis, the scheme includes a semi-mechanistic model of the mineral-to-ash conversion process. By post-processing Ardemus CFD boiler simulation data, ash particle trajectories within the furnace can be computed and important time dependent phenomena that affect the ash particle viscosity distribution can be studied. Once we have ash particle concentrations in the next-to-wall cells, deposition models and a sticking model are used to evaluate the growth rates of slag layers. A correlation-based model was adopted for turbulent impaction. A sintering index was introduced to give an

Development of A Prediction Scheme for Pulverised Coal-Fired Boiler Slagging

749

indication of the removability of the slag layers by sootblowing or gravitational shedding. In the future, a deposit layer heat transfer model should be added in the scheme to provide sintering index values at deposit layer temperatures. The testing of the various submodels is still in progress. At this stage, estimates of deposition velocities by turbulent impaction have been obtained. For a first approximation, we have constructed a simple correlation to account for the effect of the water wall tube structure on the wall shear stress. Due to the complex flow patterns in this geometry and the neglect of their effect on particle sticking, the estimates of turbulent impaction velocities are still well justified only in the case of a masoned wall, a flat wall where the tube structure is eliminated due to slag build-up or in the case of a flow parallel to the tubes. For these scenarios, trends are correctly predicted, though they may not be quantitatively accurate. The general trend is that flows impinging on walls generate turbulence in the boundary layer which results in turbulent impaction of particles. The greater the degree of turbulence the smaller are the particles that possess large enough fluctuating velocities to be transported to the wall. Results obtained from submodel testing indicate that turbulent impaction velocities are largest in near-burner areas. In other parts of the furnace, maximum turbulent impaction velocities are on the order of 1 to 2 cm/s, i.e. rather low compared with inertial impaction velocities onto superheater tubes. Overall without sticking considerations, particles larger than deposit onto furnace walls preferentially over smaller bulk mass mode fly ash particles, be it via turbulent or inertial impaction. When completed, the slagging prediction scheme can be a valuable tool in the evaluation of various burner arrangements and other furnace modifications. In addition, it

can be used for troubleshooting and for gaining an understanding of various slagging problems in operating boilers. Further work is still needed to validate the overall scheme against experimental data. The future development of the scheme will be directed towards incorporating industrial scale experimental data for better correlations in the model. Also, even if the scheme is not meant for routine coal selection processes, it can be used to identify critical parameters in different kinds of slagging problems. Based on such model studies, problem-specific CCSEM-based indices or other methods for coal slagging propensity ranking can be formulated.

5. ACKNOWLEDGEMENTS We acknowledge the funding of the Finnish Technology Development Centre TEKES via the research programme LIEKKI II. We acknowledge the contribution of Mr. Risto Paavilainen in the water wall tube simulations. In addition, we are grateful to Dr. David Brown for useful suggestions concerning this manuscript.

6. REFERENCES Allan, S., Erickson, T. and McCollor, D. (1996) “Modeling of ash deposition in the convective pass of a coal-fired boiler.” In: Baxter, L., DeSollar, R. (eds.) Application of Advanced Technology to Ash-Related Problems in Boilers, Conference Proceedings. Waterville Valley, USA July 16–22, 1995, pp. 451–470. Baxter, L. (1993) “Ash deposition during biomass and coal combustion: A mechanistic approach.” Biomass and Bioenergy, 4(2), pp. 85–102.

750

J. Pyykönen et al.

Bool, L. and Helble, J. (1996) “Iron oxidation state and its effect on ash particle stickiness.” In: Baxter, L., DeSollar, R. (eds.) Application of Advanced Technology to Ash-Related Problems in Boilers, Conference Proceedings. Waterville Valley, USA July 16–22, 1995, pp. 281–292. Bool, L., Peterson, T. and Wendt, J. (1995) “The partitioning of iron during the combustion of pulverized coal.” Combustion and Flame, 100, pp. 262–270. Boyd, R. and Kent, J (1994) “Comparisons of large scale boiler data with combustion model predictions.” Energy & Fuels, 8, pp. 124–130. Bryers, R. (19%) “Fireside slagging, fouling and high-temperature corrosion of heat transfer surfaces due to impurities in steam-raising fuels.” Prog. Energy Combust. Sci., 22, pp. 29–120.

Fluent Inc. (1996) Fluent user’s guide, Release 4.4. Gibb, W. (1996) “The UK collaborative research programme on slagging in pulverised coal-fired boilers: Summary of findings.” In: Baxter, L., DeSollar, R. (eds.) Application of Advanced Technology to Ash-Related Problems in Boilers, Conference Proceedings. Waterville Valley, USA July 16–22, 1995,

pp. 41–65. Helble, J., Srinivasachar, S. and Boni, A. (1990) “Factors influencing the transformation of minerals during

pulverized coal combustion.” Prog. Energy Combust. Sci., 16, pp. 267–279. Helble, J., Srinivasachar, S., Wilemski, G., Boni, A., Kang S., Sarofim, A., Graham, K., Beer, J., Peterson, T., Wendt, J., Gallagher, N., Bool, L., Huggins, F., Huffman, G., Shah, N. and Shah, A. (1992) Transformation of Inorganic Coal Constituents in Combustion Systems, Final report submitted to the U.S. DOE/PETC, November, Available NTIS Document DE/PC/90751–T15 (BE93013076). 3 volumes.

Huang, L., Norman, J., Pourkashanian, M. and Williams, A. (1996) “Prediction of ash deposition on superheater tubes from pulverised coal combustion.” Fuel, 75(3), pp. 271–279. Im, K. and Chung, P. (1983) “Particulate deposition from turbulent parallel streams.” AIChE J., 29(3), pp. 498–505. Jokiniemi, J., Pyykönen, J., Lyyränen, J., Mikkanen, P. and Kauppinen, E. (1996) “Modelling ash deposition during the combustion of low grade fuels.” In: Baxter, L., DeSollar, R. (eds.) Application of Advanced Technology to Ash-Related Problems in Boilers, Conference Proceedings. Waterville Valley, USA July

16 22, 1995, pp. 591 615. Kallio, G. and Reeks, W. (1989) “A numerical simulation of particle deposition in turbulent boundary layers.” Int. J. Multiphase Flow, 15, pp. 433 446. Kissane, M. (1996) “Inertial impaction in bends: Tests of simple models against data from large-scale experiments with turbulent flow.” Abstracts of the Fifteenth Annual Conference of American Aerosol Research Society, p. 463. Kjäldman, L. (1993) Numerical simulation of combustion and nitrogen pollutants in furnaces, Dissertation. VTT Publications 159. 132 p. Lee, F., Riley, G. and Lockwood, F. (1996) “Prediction of ash deposition in pulverised coal combustion systems.” In: Baxter, L., DeSollar, R. (eds.) Application of Advanced Technology to Ash-Related Problems in Boilers, Conference Proceedings. Waterville Valley, USA July 16–22, 1995, pp. 637 664. Lind, T., Kauppinen, E., Sfiris, G. and Maenhaut, W. (1997) “Fly ash deposition onto the convective heat

exchangers during CFBC of willow.” Proceedings of this conference. Mann, A., Moghtaderi, B. and Kent, J. (1995) “Computational modelling of a slag reduction strategy in a wallfired furnace.” Journal of the Institute of Energy, 68, pp. 193–198. Muyshondt, A., Anand, N. and McFarland, A. (1996) “Turbulent deposition of aerosol particles in large transport tubes.” Aerosol Science and Technology, 24, pp. 107–116. Richards, G., Slater, P. and Harb, J. (1993) “Simulation of ash deposit growth in a pulverised coal-tired pilot scale reactor.” Energy and Fuels, 7, pp. 774–781. Sarofim, A. and Helble, J. (1994) “Mechanisms of ash and deposit formation.” In: Williamson, J. and Wigley, F. (eds.) The impact of ash deposition on coal fired plants, Conference Proceedings, Solihull, England

June 20 25, 1993, pp. 567–582. Senior, C. and Srinivasachar, S. (1995) “Viscosity of ash particles in combustion systems for prediction of particle sticking.” Energy and Fuels, 9, pp. 277–283. Slater, P., Richards, G. and Harb, J. (1995) “Pyrite and illite associations in two eastern US bituminous coals.” Fuel Processing Technology, 44, pp. 55–69. Srinivasachar, S. and Boni, A. (1989) “A kinetic model for pyrite transformations in a combustion environment”, Fuel, 68, pp. 829–836.

Walsh, P., Sayre, A., Loehden, D., Monroe, D., Beér, J. and Sarofim, A. (1990) “Deposition of bituminous coal

Development of A Prediction Scheme for Pulverised Coal-Fired Boiler Slagging

751

ash on an isolated heat exchanger tube: Effects of coal properties on deposit growth.” Prog. Energy

Combust. Sci., 16, pp. 327–345. Wessel, R. and Righi, J. (1988) “Generalized correlations for inertial impaction of particles on a circular cylinder.” Aerosol Science and Technology, 9, pp. 29–60. Yau, K. and Young, J. (1987) “The deposition of fog droplets on steam turbine blades by turbulent diffusion.” Journal of Turbomachinery, 109, pp. 429–435. Yilmaz, S. and Cliffe, K. (1997) “Simulation of coal ash deposition on to a superheater tube.” Journal of the Institute of Energy, 70, pp. 17–23.

7. NOMENCLATURE

b C

g

R t

Vol

Surface area subject to deposition Fe-O-S melt area on particle surface particle surface area stoichiometric conversion coefficient Concentration probability variable Cunningham correction factor drag coefficient oxygen concentration particle specific heat capacity (J/kg · K) effective diffusivity of oxygen in glass particle diameter (m) mass concentration gravitational acceleration mass transfer and conversion coefficient bulk phase gas thermal conductivity (W/m · K) particle mass (kg) particle Nusselt number partial pressure (Pa) radiative heat transfer (W) universal gas constant (J/mol · K) duct Reynolds number time (s) residence time in a next-to-wall cell (s) gas temperature (K) particle temperature (K) bulk phase gas temperature (K) deposition velocity (m/s) dimensionless deposition gas velocity vector (m/s) particle velocity vector (m/s) friction Cell volume volume fraction of glass that has reacted to Fe(III) gas dynamic viscosity particle dynamic viscosity critical particle sintering viscosity critical particle sticking viscosity

752

J. Pyykönen et al.

gas kinematic viscosity concentration of Fe(II) in unreacted glass gas density particle density

particle relaxation wall shear stress (Pa) dimensionles particle relaxation

MODELLING THE INITIAL STRUCTURE OF ASH DEPOSITS AND STRUCTURE CHANGES DUE TO SINTERING Hamid R. Rezaei, Rajender P. Gupta, Terry F. Wall, S. Miyamae1,

and K. Makino 1

Department of Chemical Engineering The University of Newcastle Callaghan, 2308 Australia. 1 Ishikawajima-Harima Heavy Industries Co. Ltd. Japan

1. INTRODUCTION A number of deposit growth models have been developed in the literature. Visscher and Bolsterli (1972) presented random packing of equal and unequal spheres in two and three dimensions with an uni-directional gravitational force. Jodery and Tory (1979) describe a model that simulates the settling of rigid spheres from a dilute suspension into a randomly packed bed. In their model the spheres fall and roll till they occupy the lowest

possible level. During heat treatment, the production of a liquid phase plays a major role in densification. When the volume of liquid is sufficient to fill all the pores in the compacted powder, complete elimination of the porosity can occur by flow of the liquid into the pores under the influence of surface tension forces, resulting in a dense, pore-free, solid product. Densification is driven by the energy reduction brought about by the reduction in surface area of the porous body. Various attempts have been made to describe the phenomenon in terms of energy balance and deformation. These models show the particular importance of three variables, i.e., a geometrical factor, the particle size; a kinetic factor, the viscosity; and a thermodynamic factor, the surface tension. The effect of particle size distribution in the initial stage of sintering has been studied by Coble (1973). German and Lin (1994) used a constant heating rate sintering densification of bimodal alumina powder mixtures. He found at low sintering temperatures a bimodal powder mixture can result in high sintered densities. As the sintering temperature increases the smaller powder leads to a higher sintered density. The effect of particle size and temperature with varying time are considered for four coal ash samples obtained from a deposit probe test section and from a bag filter collection fly ash. The ash analyses for the samples are shown in Table 1. Impact of Mineral Impurities in Solid Fuel Combustion, edited by Gupta et al. Kluwer Academic / Plenum Publishers, New York, 1999.

753

754

H. R. Rezaei et al.

2. PREDICTION METHOD FOR INITIAL POROSITY A simple 2-dimensional model similar to that of Jodrey and Troy (1979) is presented here to estimate the initial porosity of particles. In this study it is assumed the particles are sticky so that particles stay where they impact. The porosity of a particulate ash deposit is predicted assuming the particles have only two dimensions, therefore the calculation predicts the two dimensional (2D) porosity. The position of a particle in x, y coordinate is determined at random. The initial heat transfer surface is a line of constant y. In the 2-D model, the spherical particles are represented by a circular disc. The particles fall from a fixed height (y) but at random x. The particles fall initially to the x-axis representing a clean heat transfer surface. As the deposit grows, the surface of the deposit also grows. The addition of each particle modifies the boundary, as shown in Fig. 1. There are other random numbers for the particle size distribution and angle of impact. These random numbers correspond to input data. The 2D porosity of particles is determined by selecting a rectangular area as shown on Fig. 2.

A prediction of average 2D porosity is obtained by taking the average of 40 runs. The average porosity of each run is calculated by varying the position of the rectangle across the deposit.

Modelling the Initial Structure of Ash Deposits and Structure Changes Due to Sintering

755

3. THEORETICAL MODELS FOR SINTERING The kinetics of densification or porosity removal in an ash deposit, which is undergoing vitrification, may be described by a curve shown in Fig. 3. The porosity decreases and the extent of shrinkage increases as a function of time.

Among the large number of models in the literature, it is possible to distinguish between two different approaches:

• one based on geometrical assumptions that related to an idealised system, i.e., assemblies of spherical, viscous particles whose dimensions and physical properties (surface tension and viscosity) remain constant during processing. A schematic diagram for sintering of two spheres is presented in Fig. 4. • the other based on a phenomenological description for the full span of the densification process. Frenkel (1945) was the first to propose a model describing the sintering of viscous materials from geometric assumptions. The model describes the rate of coalescence of

particles in terms of measurable parameters; viscosity, surface tension, and particle size. By equating the energy variation due to surface decrease to the energy dissipated by viscous flow, Frenkel’s model describes the first stage of vitrification by Eq. 1.

where are the time of sintering, surface tension and viscosity, respectively. According to this model, the volume shrinkage is proportional to the duration of the thermal treatment. The kinetic constant depends on the surface tension, viscosity, and powder grain size.

756

H. R. Rezaei et al.

Shrinkage level or degree of sintering is defined as the ratio of neck radius to the

radius of particles. The degree of sintering in terms of density is presented in following equation.

where of initial powder of sintered product of the interface (assumed to be circular)

of the spherical particles The initial density and density at time t can be expressed as a function of the density of the solid in following equations:

where the initial porosity of and is the porosity at time t. The porosity of sintered product at time t can be obtained from Eqs. (2) to (4).

The range of application of this model is limited to a shrinkage level of 0.3. Zagar (1975) described vitrification kinetics from phenomenological assumptions:

Modelling the Initial Structure of Ash Deposits and Structure Changes Due to Sintering

757

i) the sphere volume decreases with time, and ii) the sphere centers approach each other.

However, he did not take into account the solid volumes involved in these two phenomena, which contribute to the pore volume decrease. Assuming that during the period t the original radius of the particle does not change appreciably the ratio of the pore at time t to the initial pore volume is given by Eq. (6).

According to Frenkel (1945) for viscous flow,

equals

Thus we obtain

Ivensen (1970) investigated the change in pore volume during the sintering of powder compacts from phenomenological assumptions. This variation can be expressed

by a ratio, which can be found from the bulk densities as follows:

and the true density

The porosity at time t can be determined by the ratio of the pore volume at time 0 and t (Eq. 8) and Eqs. 3 and 4. The porosity at time t therefore can be obtained by Eq. (9).

The surface tension of an amorphous phase can be estimated from its chemical composition. Tillotson and Oppen (Jouenne, 1990) showed that the surface tension

can be expressed as a linear function of the composition

where is the surface tension corresponding to each oxide i and is its molar percentage. Table (2) gives the tension coefficient values for some oxides present in ash deposits. From this table, it is possible to approximate the surface tension versus temperature variation by subtracting or adding 0.004 N/m when the temperature decreases or increases by 100°C, respectively. Table (2) shows that for silicate ceramics the surface tension is weakly dependent on both the composition and temperature. The viscosity of a melt at any temperature (T) is given by Urbain et al. (1981) and presented in the following equation:

where T is the absolute temperature, and has unit of A and B are two parameters depending only on the composition of the melt. The A value can be obtained from Eq. (12).

758

H. R. Rezaei et al.

The parameter B is related the composition and temperature of the melt, which are presented by Urbain et al. (1981).

4. RESULTS The effects of particle size distribution and angle of fall of particles are considered in prediction of initial porosity. The effect of particle size distribution is considered when all particles fall vertically. When considering the effect of angle of fall on porosity, all particles are assigned same dimension. A typical deposit growth is shown in Fig. 5.

4.1. Effect of Particle Size Distribution on Initial Porosity In this discussion all particles drop vertically The average porosity was predicted from a number of runs. In the first instant, the porosity of single and double size particles is predicted. Figure 6 shows the average predicted porosity for single and double size particles is about 0.6.

In the second case the effect of particle size distribution on porosity is considered. The particle size distribution is classified into two groups. In group one the ratio of particles in the size distribution is less than 6, while this ratio is greater than 6 for group two. The particle size distribution for group 1 is presented in Fig. 7. In this case the maximum to minimum particle size is 17.5 to showing the ratio is less than 6. The average porosity which is predicted for group one is the same as that for the mono-size particles showing there is no effect of particle size distribution on average porosity in this group. In group 2A the maximum to minimum particles size are 30 to and this value for group 2B is 100 to as shown in Fig. 7. The average porosity is predicted to be

about 0.63 showing 5% increasing in porosity due to particle size ratio. The next prediction is based on the actual particle size distribution of a coal ash. The particle size

distribution of particulate coal ash is also presented in Fig. 7. This figure shows that 50% and 80% of the mass fraction are below 3 and respectively. The average porosity of this ash is also presented in Fig. 6. As the ratio of particle sizes is above 6, the average

porosity as expected is about 0.63.

Modelling the Initial Structure of Ash Deposits and Structure Changes Due to Sintering

759

4.2. Effect of Angle on Initial Porosity The effect of angle of fall on porosity is investigated for mono sized particles only. For mono size particles the average porosity is predicted for one direction, two directions and multiple directions. Figure 8 shows the average porosity for different angles of fall. For two directions of fall it is assumed that 50% of the particles fall vertically and 50% at another angle. No significant effects are obtained. When considering multiple angles of fall, the average porosity decreases to 0.57, showing a 5% reduction.

760

H. R. Rezaei et al.

4.3. Effects of Angle and Particle Size Distribution Together on Initial Porosity Finally, the porosity was determined for a particle size distribution (as in Fig. 6), where fine particles came vertically down and the largest particles fall at an angle of 30°

Modelling the Initial Structure of Ash Deposits and Structure Changes Due to Sintering

761

to the deposit surface. Intermediate size particles fell down at 40° to 70°. Figure 9 shows the porosity of ash deposits for several cases, such as:

1) mono size, single direction 2) mono size, and different angle of fall 3) particle size distribution, single direction

4) particle size distribution in different angles of fall Different angles of fall with mono size reduced the porosity, but when accompanied by a wider particle size distribution, the porosity increases. This suggested that the effect of angle of fall is less dominant in determination of the porosity for a sample than particle size distribution. The porosity of the deposit according to Fig. 9 is about 0.63.

4.4. Converting the Two Dimensional (2D) Porosity to Three Dimensional (3D) Porosity There is a simple approach to obtain the porosity in three dimensions. When the porosity in two dimensions is zero, the porosity in three dimensions is also zero. The porosity equal to one in two dimensions corresponds to one at three dimensions. The third point can be obtained by using Eq. 13 according to Fig. 10.

762

H. R. Rezaei et al.

Figure 11 presents the calibration curve for converting the 2D porosity to 3D porosity for single size particles. As shown in the figure the 3D porosity is higher than the 2D porosity. For example, for a 2D porosity of 0.62 the porosity in three dimensions is about 0.8.

4.5. Effect of Temperature, Particle Size and Viscosity on Sintering In this discussion the effect of temperature and particle size are considered on the sintering process. Each coal has a different composition, which is presented in Table 1. All ash particles are assigned the same composition in the calculation. The surface tension is obtained at each temperature. In this prediction the viscosity is estimated from Urban’s model. Figure 12 shows the viscosity of deposit and filter bag samples versus temperature. The initial porosity used in sintering model is 0.8 based on the ash deposition model in 3D.

Modelling the Initial Structure of Ash Deposits and Structure Changes Due to Sintering

763

Coal ashes C and D, respectively, have lower and higher silica and alumina contents compared with the other coal ashes. The ash fusion temperature for coal ashes A, B, C, and D is 1,140, 1,180, 1,220, and 1,580°C, respectively. The porosity of coal ash deposit A versus time for particle sizes of 10, 20, 50, 75, and

764

H. R. Rezaei et al.

is shown in Fig. 13 for a temperature of 950 K. The figure shows that sintering occurs faster for finer particle sizes. For example a decrease in porosity from 0.8 to 0.4 takes place in less than 15 and 80 minutes for particles of 10 and size, respetively. The reduction of porosity for coal ash C is presented in Fig. 14 for temperatures

of 700, 752, 750, 775, and 800 K. As shown in this figure, there is no sign of sintering below a temperature of 700 K. The reduction of porosity from 0.8 to 0.4 at 750 K occurs in 200 minutes, whereas at 800 K it takes about 25 minutes. The results presented in Fig. 15 show the effect of viscosity on the sintering temperature of coal ash samples A, B, and D. Among these samples, coal ash C has the lowest viscosity then A, B, and D, respectively. According to Fig. 15, there is no reduction in porosity for coal D at a temperature of 900 K. The reduction in porosity from 0.8 to 0.72 for coal ash A takes about 100 minutes, whereas this reduction takes more than 220 minutes for coal ash B. Figure 16 shows the porosity of coal ash samples before and after slow and rapid sintering. In slow sintering it is assumed that the reduction in porosity is 5% of the initial porosity during the four hours of the process. Rapid sintering corresponds to 50% reduction in porosity in one hour of sintering. The results are compared with those obtained from the thermal conductivity experiments. In a curve of thermal conductivity versus temperature, the sintering temperature is a point where the slope of the curve changes rapidly. The prediction of sintering temperature of ash deposit A corresponding to a 5% reduction in porosity is so close to the experiment. The difference in sintering temperature for ash deposits B and D obtained by prediction and experiment is about 50 to 100 °C, respectively. This difference for ash deposit C is about 200 °C.

Modelling the Initial Structure of Ash Deposits and Structure Changes Due to Sintering

765

766

H. R. Rezaei et al.

5. CONCLUSIONS The effect of the size distribution was to increase the predicted 2D porosity from 0.6 to 0.63, whereas the effect of varying angle of fall was to reduce the 2D porosity from 0.6 to 0.57. The effect of particle size distribution was more important than the effect of

angle of fall when both the effects were considered together. However, the anisotropy in the structure with varying angles of fall is clearly seen. For all practical particle size distributions, the initial porosity for a typical ash particle size distribution can be taken as , irrespective of angle of fall. Converting the 2D porosity to 3D porosity results in an increase in porosity of 19% for the porosity at 0.6. The initial porosity changes with temperature and time. The variation of porosity depends on particle size, and the chemical component determining the viscosity. Sintering happens at higher temperatures for coarser particles. The samples with higher viscosity need a higher temperature or longer time to sinter and then fuse. Among samples A, B, C and D, the variation of porosity with temperature was highest for sample C and was lowest for sample D. The results revealed that sintering occurs at temperatures above 800, 850, 650, and 1,000 K for coal ashes A, B, C, and D, respectively. Below these temperatures there is no significant effect of sintering on coal ash samples.

6. REFERENCES Coble, R. L., (1973), Effect of Particle Size Distribution in Initial-Stage Sintering, Journal of The American Ceramic Society, 56(9), 461–466. Frankel, J., (1945), J. Tech. Phys. Leningrad, 9, 385. German, R. M., and Paul, Lin, S. T, (1994), Constant-Heating Rate Sintering Densification of Bimodal Alumina Powder Mixtures, Journal of Materials Synthesis and Processing, 2(5), pp. 291–294. Ivensen, V. A., (1970), Powder Metall. (USSR), 4(20).

Jodrey, W. S., and Tory, E. M., (1979), Simulation of Random Packing of Spheres, Simulation, 1–12. Jouenne, C. A., (Ed.), (1990), Traite de Ceramiques et Materiaux Mineraux, 5th ed., Paris: Septima, 568. Urbain G., Cambier, F., Deletter, M., and Anseau, M. R., (1981), Viscosity of Silicate Melts, Trans. J. Br. Ceram. Soc., 80, 139–141.

Visscher, W. M., and Bolsterli, M., (1972), Random Packing of Equal and Unequal Spheres in Two and Three Dimensions, Nature, 239, 504–507. Zagar, L., (1975). Sci. Sintering, 7(1), 35–43.

INDEX Advanced systems, 45, 653 Additives, 55, 319 Alkali behaviour biomass combustion, 525, 595 fluidised bed, 525, 555 pf combustion, 10 Analytical techniques, 25–30, 134–144, 155, 171, 181, 260

CCSEM, 23–26, 137, 148 FTIR, 30

Standard Thermal Analysis, 181 Thermo Mechanical Analysis, 155, 261

X-Ray Diffraction, 37, 217 X-Ray Fluorescence, 30, 376 Ash adhesion, 195 agglomeration, 49, 147 amorphous, 30, 217

classification, 205 crystalline, 30, 217 formation, 4, 117, 584, 621, 743 mineral phases, 621

particle size, 541 shrinkage measurement techniques, 155 sticking efficiency, 195 synthetic, 184 Ash fusibilty Image Analysis, 171

Coal char particle combustion, 621 mineral matter in, 5, 105, 225, 389 trace elements, 609 Coal Cleaning, 645 Co-firing

biomass and coal, 247, 271, 367, 525, 569 ash deposition, 249, 367, 417 utility boilers in, 271, 367, 417 blast furnace gases, 285 burnout, 254 chlorine corrosion, 255

NOx emissions, 253 Combustion entrained flow, 525 fluidised bed, 13, 47 full scale tests, 358, 367, 383 pulverised coal, 9, 390

Computer-Controlled Scanning Electron Microscope deposits, 207, 278, 334, 375, 400, 534

fluidised bed application, 148 fly ash, 425, 441, 621 technique, 27–28, 137, 225, 237, 743 Corrosion chlorine, 126, 255, 497, 513, 583 co-fired combustion, 271 hydrochloric acid, 513

Standard Thermal Analysis, 181 temperature, 402, 433 Thermo Mechanical Analysis, 155

Deposit classification, 205 SEMPC, 148, 207, 334, 375, 378, 534

minerals in, 445 particle size, 545 Blending, 35, 297, 383, 471, 728

strength, 473 structure, 66–78, 206–208, 375, 753 superheater, 407 Deposition biomass-coal co-firing, 247, 272, 333, 525 full-scale experiments, 205, 358, 367, 383 mechanisms, 7, 121, 212, 322, 352 rate, 697, 735 surface temperature effect, 357

Cement kiln, 675 Chlorine, 126, 255, 497, 513, 583 Circulating Fluidised Bed, 41, 88, 319 Clean coal technologies, 33, 85

Fluidized bed agglomeration, 49–50, 260, 319, 333 ash coating, 309 bed materials, 56

Biomass

ash, 186, 341, 441, 541, 555 combustion, 595, 635 contaminated, 635 deposit, 341, 405, 635 heavy metals, 595

767

768

Fluidized bed (cont.) cement kiln, 675 circulating, 41, 87, 319 combustion, 36, 47, 309

Index

Phase equilibria, 581, 723–728 Pressurised fluidised bed, 41, 87, 90 development, 653 experience, 666

defluidisation, 49–52 entrained, 525 gasification, 51, 55 pressurised, 41, 87, 653, 663 Fouling, see also Slagging, 107, 124, 417, 541

calcium project, 11 FOULER, 17 PCQUEST, 16 sodium project, 10 Fuel cell, 42, 92

Gasification entrained flow, 14, 609

fluidised bed, 14, 51, 555 Hot gas filter, 195

Integrated Gasification Combined Cycle, 38 Low-rank coal combustion, 47, 309, 319 Minerals in coal, 5, 112, 389 analytical methods, 25 geological origin, 24 particle size, 230, 239 Minerals in oil-shale, 685 Mineral-mineral associations, 581 Mineral transformations, 4, 117, 584, 621, 743 Modelling ash behaviour, 4, 352, 697, 709 utility boilers, 389, 706, 735

Radio-nuclides, 635

SEMPC, 27, 148 deposits, 207, 334, 375, 378, 534 Sintering, 49, 699, 743, 755 Slagging, see also Fouling, 4, 122, 325, 383, 395, 455, 581

blending, 297 chemical composition, 4, 105

indices, 7, 233, 298, 458, 471, 714 operating conditions, 4, 107, 357 sulphur, 405 theoretical predictions, 405, 728, 735 Stickiness models, 195, 697, 735 Thermal conductivity ash deposits, 65, 405

in-situ measurements, 405 models, 65 Thermodynamic modelling blending, 723, 728 fluxing, 728 Trace element distribution, 595, 609

Vaporization of inorganic, 595, 609 Viscosity of ash, 8, 261 X-Ray diffraction, 27, 133

X-Ray fluorescence, 30, 376