Brittle Matrix Composites 8 (v. 8)

Proc. Int. Symp. ”BrittleMatrix Composites 8“ A.M. Brandt, ?!C. Li and I. H.Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

PLENARY I W T E D PAPER

HIGH PERFORMANCE CONCRETE: WHERE DO WE GO FROM HERE? Sidney MINDESS Department of Civil Engineering University of British Columbia 6250 Applied Science Lane Vancouver, British Columbia V6T 124, Canada e-mail: [email protected]

ABSTRACT We can now prepare concretes with an enormous range of compositions and properties. Indeed, we can now “tailor-make” concretes for a wide variety of applications. These advances have been driven by the need to both renew our infrastructure, and to ensure that the concrete industry remains sustainable in the face of environmental pressures. These advances, however, are not without cost, both economic and technological. In this discussion, possible future developments in the field of high performance concrete will be described.

INTRODUCTION Modem concretes constitute a family of Portland cement based materials with an enormous range of compositions and properties. Certain members of this family are referred to as high performance concretes, that is, concretes that meet “special combinations of performance and uniformity requirements that cannot always be achieved routinely using conventional constituents and normal mixing, placing, and curing practices” [ 11. Stated more simply, high performance concrete is simply concrete that is better in one or more respects than the concrete that we usually make. In its commentary on the above definition, the American Concrete Institute goes on to say [I] that “A high performance concrete is a concrete in which certain characteristics are developed for a particular application and environment. Examples of characteristics that may be considered critical for an application are: Ease of placement, Compaction without segregation, Early age strength, Long-term mechanical properties, Permeability, Density, Heat of hydration, Toughness, Volume stability Long life in severe environments.”

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Sidney MINDESS

It should be noted that high performance concrete is not the same thing as high strength concrete; while many high performance concretes also possess high strengths, it is perfectly possible to have high performance concretes with quite ordinary strengths, and high strength concretes that do not perform at all well. Of course, the term “high performance” is highly time-dependant; even the most ordinary of today’s concretes would have seemed high performance to Joseph Aspdin in 1824. Indeed, it is worth contrasting the concrete technology of today with that of 100 years ago. In 1901 [2]: Portland cement is thought to consist of C3S and C3A (or maybe C2A); it is referred to as a “fusible calcium silico-aluminate.” Cement costs about $2.00 - $3.00 per barrel (about $10.60 - $15.90 per ton). Chemical tests of Portland cement are not thought to be of much value, though “it is not impossible ...... that chemical tests may yet play a more important role in cement testing, especially if the method of analysis can be made more simple and rapid.” 0 Effect of wlc ratio on strength was not understood. 0 No commonly used mix design criteria for concrete. 0 No workability tests, just vague descriptions (e.g., “the mortar was wet enough to quake like liver under moderate ramming”). Compressive strength measured typically on cubes ranging in size from 6 to 12 inches, with the load applied over all or just part of the surface. No admixtures, either chemical or mineral. 0 Cost in place: $4.50 - $6.50 per cubic yard. ($5.90 - $8.50 per m3) “Concrete-steel” combinations (k,reinforced concrete) was still not in common use; simplified design methods were available for reinforced concrete beams, but not for other structural members. Today, the picture has changed completely: 0 We understand the chemistry of cement and of the hydration reactions. 0 We have well-developed cement standards, both physical and chemical. 0 We can “tailor-make” cementitious materials for particular applications. 0 Concrete mixes are designed largely on the basis of w/c ratio. 0 There are well-developed tests for concrete. 0 We can (though we often do not) design for durability. 0 We can reliably make concretes with strengths ranging from 1 MPa to 600 m a . 0 A large variety of both chemical and mineral admixtures are available. 0 Fibre reinforced concretes are coming into frequent use. 0 There are well-developed design procedures for reinforced concrete, which is now the most widely-used construction material in the world. Given these tremendous advances, the universal use of concrete, and the wealth of published information on cement and concrete, 0 Why do there continue to be large numbers of construction problems? Why is concrete perceived to be so environmentallyunfriendly? Why are there questions about the long-term viability of the cement and concrete industries?

High performance concrete: where do we go from here?

17

In what follows, some future developments of high performance concrete, which might help to alleviate some of these problems, will be discussed.

SUSTAINABILITY By far the biggest issue facing the concrete industry is its sustainability. There are two basic problems. 1. The cement and concrete industries are perceived to be environmentally unfriendly: C02 emissions from the production of cement; diminishing sources of good aggregates near major population centres combined with a public unwillingness to see new aggregate sources developed; and the problems of disposing of concrete from demolished structures. 2. Too much of the concrete that we produce suffers from a lack of durability. While we know how to produce concrete that will be durable under a given set of conditions, all too often this is not done in practice, due to a combination of (false) economic considerations, and lack of understanding on the part of the structural designers and the construction industry. It is in these areas that high performance concretes can play their most significant role.

“Green” concrete Probably the easiest way of reducing COz emissions is to reduce the amount of cement produced, by the replacement of cement with other cementitious (waste) materials, such as fly ash, blast furnace slag, silica fume, finely ground limestone fillers, and so on. The use of fly ash, at replacement levels of up to 50%, is now well-known [3. 41, as is the use of ground granulated blast furnace slag. For high strength concretes, silica fume is commonly used, and other finely divided forms of reactive silica (metakaolinite, rice husk ash) are other possible cement replacements. More recently, there has been the development of ternary blends of cement, silica fume and other mineral fillers [ S ] , and even more complex cement blends are being studied. Odler [6] has provided a description of a very large number of special cement systems, many of which would reduce C02 emissions. For instance, Gebauer et al. [7] have developed a cement that shows close to zero COZ emissions during its production, and provides a concrete with low water demand, low heat of hydration, ands excellent durability and chemical resistance. In the context of reduced greenhouse gas emissions and often lower energy costs, such cements and concretes are truly high performance. Another part of sustainability is the efficient use of aggregates which are currently considered to be “marginal” [8]. Such aggregates, even though they fall outside of normal aggregate specifications, may be perfectly acceptable for at least some particular applications. They may also be beneficiated so that the resulting concrete properties are not compromised. Finally, here must also be much greater use made recycled concrete aggregates [9]. This would greatly alleviate both the need for new aggregate sources, and the problem of disposing of concrete demolition wastes. While concretes made with recycled aggregates tend to have inferior properties to those of the original concrete [lo-1 11, they are nonetheless suitable for a great many uses. It should be remembered that in the Unite states, over 40% of the concrete produced is used in house foundations, which does not place a high demand on concrete strength, and there are many other applications in which “ordinary” concretes are perfectly adequate. Again, concretes made with recycled aggregates or marginal aggregates should be considered high performance in the context of resource utilization.

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Sidney MINDESS

Durability If the concrete industry is to remain sustainable, the concretes that we produce must be durable. In general, we know how to make concretes that will be durable in most environments. We know how to protect against sulfate attack, alkali-aggregate reactions, other forms of chemical attack, freeze-thaw cycling, salt scaling, corrosion of steel, and so on. Indeed, we can “tailor-make” concrete for almost any situation. However, most of our understanding is on the materials level. The basic approach is to make the concrete as impermeable as possible by using a low w/c ratio. (This, in turn, generally leads to higher strength concretes, a fact that may not be appreciated by the structural engineer, who often does not make use of this “extra” strength in the design process, leading to less efficient structures). There is much less of an appreciation of how a complete structure can be made durable. While the concrete as a material can be made to be durable, one characteristic of reinforced concrete structures is that they will inevitably crack somewhere, due to some combination of plastic and drylng shrinkage, creep, structural loads, and environmental exposure. We may well end up with the case of sections of highly impermeable concrete separated by wide cracks! As has been shown by Wang et al. [12], strength, permeability and cracking resistance have to be considered together in concrete design in order to achieve durable concrete and concrete structures. We thus require a much closer cooperation between structural designers, contractors and materials engineers, but all too often this does not occur (at least in part because we do not produce enough engineers who have a real understanding of concrete science and technology). However, such cooperation seems unlikely to take place until considerations of sustainability become paramount in the cement and concrete industries [13]. It should be noted that the economic costs of this lack of understanding and cooperation are enormous. It has been estimated that in the USA alone, an investment of $1.6 trillion would be required over the next five years for repair and retrofit of the existing infrastructure; the corresponding costs for Asia are estimated at $2 trillion. One novel approach is that taken under the auspices of the ISIS Canada Research Network Intelligent Sensing for Innovative Structures. Since corrosion of steel reinforcement is the most pervasive durability problem today, they have been advocating the use of GFRP or CFRP reinforcing bars in bridge decks, to replace conventional steel reinforcement. Several “steel-free’’ bridges have been constructed in Canada, and appear to be performing well in service. Perhaps of greater importance, they have pioneered the use of remote sensing technologies to obtain real-time data on the performance of these bridges, so that any necessary repairs can be carried out in a timely fashion. The incorporation of fibre optic or other sensors in new construction, combined with a systematic monitoring program, would be a significant step towards a rational management of our civil engineering infrastructure. CRACKING

As indicated above, we are now quite capable of making concrete itself highly impermeable, through a combination of chemical admixtures, mineral additions, low wlc ratios, and proper placing and curing techniques. However, this is not the same as making impermeable structures or structural elements. For this, apart from getting the design details right, it is necessary with today’s technology to use hybrid systems, containing fairly high volumes of fibres andor textile meshes. For instance, Naaman et al. [14] have described a number of such systems, containing discontinuous steel or PVA fibres combined with 2D or 3D continuous meshes made of Kevlar or steel. They were able to obtain flexural strengths of up

High performance concrete: where do we go from here?

19

to 100 MPa, and also excellent crack control (finer cracks), with a total volume fraction of reinforcement of less than 3%. Despite its apparently high initial costs, this technology needs to be developed further. Another approach to minimizing the crack widths in concrete is the development of ECC (Engineered Cementitious Composites) pioneered by Li [ 15-17]. ECC is a fibre reinforced cementitious composite, containing typically about 2% fibres by volume. Using a micromechanics-based approach to the mix design, involving careful matching of the matrix strength and the fibre pull-out strength, it has been possible to achieve ductility values of up to 3% in direct tension. This material can be placed in many ways - by ordinary casting techniques, as self-consolidating concrete, and by shotcreting. Because of its ductility, and the fact that it keeps crack widths small (Fig. l), this material too can lead to more durable structures and better sustainability,even though the initial costs again can b substantial. 1W

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Fig. 1. Typical tensile stress-strain curve and crack width development of ECC [ 171. More generally, it has been suggested [18] that High Performance Fibre Reinforced Cementitious Composites (HPFRCC) might increasingly be used as the matrix in reinforced concrete structures. While these materials (which contain 2-3% of fibres by volume) are very expensive in terms of first cost, if we factor in life-cycle costs, as well as the social and environmental costs of rehabilitation and replacement, their use should become feasible. Indeed, a family of such materials, with very high fibre contents, very high strengths, and very high durability are now beginning to appear. The common features of these materials are very low water:binder ratios, the use of silica fume and superplasticizers, high fibre contents, severe limitations on te maximum aggregate size (often less than 1 or 2 mm), careful control of the particle size distribution of all of the solid materials in the mix, and tight quality control in their production, placement and curing. Not surprisingly, these materials are very expensive, though they are beginning to find a place in certain specialized applications. Some examples of these materials are: DUCTAL@:This material consist of fine aggregate (<2mm), crushed quartz, silica fume, and of course cement, water and superplasticizer, and up to 2% by volume of fibres. When the fibres are steel, compressive strengths of about 150 to 180 MPa may be achieved, with

20

Sidney MINDESS

flexural strengths of about 32 MPa. When polypropylene fibres are used instead, these strengths are reduced by about 25%. BSI@-CERACEMwas recently used to construct the toll gate roofs for the new Millau viaduct in the south of France [19]. With about 2.5% of steel fibres it achieved a compressive strength of 165 MPa, and a tensile strength of 8.8 MPa. It was also self-consolidating. CEMTECmduse.le@'has much higher cement and fiber contents than the two previous materials, though the underlying principles remain the same. This material can achieve flexural strengths of 60 MPa, and has a very low permeability [20].

Typical mix proportions for two of these materials are given in Table 1.

Portland cement Silica fume Crushed quartz Sand Water Fibers Superplasticizer

710 230 210 1020 140 40 - 160" 10

1050.1 268.1 514.3 180.3 858b 44

'Either steel or polypropylene fibers (13x 0.20mm) bA mixture of three different geometries of steel fibers CRC (Compact Reinforced Composite): This material is made with a very low waterhinder ratio ( 4 . 1 6 or less) and contains from 2 to 6% steel fibres, with matrix compressive strengths ranging from 140 to 400 MPa. It is different fiom the materials mentioned above in that it is combined with closely spaced conventional steel reinforcement. It is used primarily for precast elements, but can also be use in cast-in-place construction [21].

It should be noted that while all of these materials are promoted largely on the basis of their compressive strength (since this is still the principal obsession of structural engineers), it is their high impermeability and durability that is probably of greater importance in the long run. In spite of their high initial costs, this family of materials will become more important as sustainabilityconsiderationsare embraced by he industry. One major impediment to the more extensive use of fibres in structural applications is that most building codes and design manuals, particularly in North America, are strength-based. It is therefore very difficult to incorporate FRC into these codes, since the fibres have little effect on strength; their contribution is to the post-cracking behaviour (or toughness) of the material or the structure. While rational methods for the incorporation of FRC into structural design have been developed in Europe [22-231, these are still not in common use. Much more work must be done in the near future to permit a more extensive use of fibres in truly structural applications, as a complement to conventional reinforcement.

High performance concrete: where do we go from here?

21

PERFORMANCEvs. PRESCRIPTION One final problem to be overcome if we are to be able to develop innovative materials and construction processes is our current reliance on prescriptive specifications for concrete. In most cases, it is assumed that prescribing the waterhinder ratio (andor the compressive strength), and perhaps the type of cement, is sufficient to assure durable concrete. However, it is clear that this does not always work, judging from the amount of deteriorated concrete that we see around us. In any event, we know that the waterhinder ratio is not the only thing that controls permeability; the proper use of supplementary cementing materials can lead to much greater reductions in permeability than merely reducing the waterbinder ratio. Similarly, at the same waterhinder ratio, one can produce concretes with very different strengths, particularly when dealing with modem, high performance concretes. As has been pointed out by OLeary and Lemay [24], the principal problems with the current prescriptive specifications are that they: Do not necessarily address the intended performance of the concrete; Do not provide an incentive for the producer to maintain strict quality control; Prevent the producer from using newer or better raw materials, or adjusting the mix to account for changed environmental conditions. In other words, they tend to inhibit the production of the highest quality (and often the most economical) concrete for a particular application. If we were to move towards performance specifications, the focus would be primarily on the functional requirements of the concrete. This would permit much more flexibility in the choice of materials, and should in the long run lead to better concretes, and a better use of the available resources. Probably the highest hurdle to be overcome in the move to performance specifications (apart from the inertia of a large and conservative industry) is the lack of effective test methods that can be used as reliable early predictors of the concrete quality. While much work in this area is being camed out in Europe, North America lags far behind. However, it will be difficult to put the innovations discussed above into general practice without a major paradigm shift as to how we specify concrete.

CONCLUSIONS In principle, we know how to make good concrete for almost any application. However, faced with increasingly Severe environmental and economic pressures, the industry is at a crossroads. If it does not find effective ways of introducing the new generation of high performance concretes that continue to be developed (at least in the laboratory) into the marketplace, the long term viability of the industry will be in jeopardy. It is up to all of us to continue to press €or innovation and change in what is today the most widely used material in the world. REFERENCES 1. Russell, H.G., ACI defines high-performance concrete. Concrete International, 21, 1999, pp. 56-57. 2. Sabin, L.C., Cement and Concrete. Archibald Constable and Co., London, 1905. 3. Malhotra, V.M., High-performance high-volume fly ash concrete. Concrete International, 24,2002, p ~45-49. .

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4. Malhotra, V.M. and Mehta, P.K., High Performance, High-Volume Fly Ash Concrete. Supplementary Cementing Materials for Sustainable Development, Inc., Ottawa, Canada, 2002. 5. Nehdi, M. and Mindess, S., Microfiller partial substitution for cement. In: “Materials Science of Concrete VI”, J. Skalny and S. Mindess eds., The American Ceramic Society, Westerville, Ohio 2001, pp. 5 13-574. 6. Odler, I., Special Inorganic Cements. E & FN Spon, London, 2000,395 pp. 7. Gebauer, J., KO, S.-C., Lerat, A. and Roumain, J.-C., Experience with a new cement for special applications. In “Non-Traditional Cement & Concrete 11”, V. Bilek and Z. Kersner, eds. Bmo University of Technology, Czech Republic, 2005, pp. 277-283. 8. Alexander, M.G. and Mindess, S., Aggregates in Concrete. Taylor & Francis, Oxford, England, 2 0 0 5 , 4 3 5 ~ ~ . 9. Hansen, T.C. and Lauritzen, E.K., Concrete waste in a global perspective. In: “Recycling Concrete and other Materials for Sustainable Development”, T. Liu and C. Meyer eds., SP2 19, American Concrete Institute, Farmington Hills, MI, 2004, pp. 11-34. 10. Buttler, A.M. and Machado, E. F., Jr., Properties of concrete with recycled coarse aggregates. In: “Quality of Concrete Structures and Recent Advances in Concrete Materials and Testing”, P. Helene, E.P. Figueiredo, T.C. Holland and R. Bittencourt eds., SP-229, American Concrete Institute, Farmington Hills, MI, 2005, pp. 497-510. 11. Kasai, Y., Recent trends in recycling concrete waste and use of recycled aggregate concrete in Japan. In: “Recycling Concrete and other Materials for Sustainable Development”, T. Liu and C. Meyer eds., SP-219, American Concrete Institute, Farmington Hills, MI, 2004, pp. 11-34. 12. Wang, K., Igusa, T. and Shah, S.P., Permeability of concrete - relationships to its mix proportion, microstructure and microcracks. In: “Materials Science of Concrete: The Sidney Diamond Symposium,” M. Cohen, S. Mindess and J. Skalny eds., The American Ceramic Society, Westerville, Ohio, 1998, pp. 45-54. 13. Bentur, A., Mindess, S. and Katz, A., Future of concrete: vision and challenges. In: “Concrete Technology”, J. Skalny, S. Mindess and A. Boyd eds., The American Ceramic Society, Westerville, Ohio, 2006, pp. 135-157. 14. Naaman, A.E., Wongtanakitcharoen, T. and Likhitruangsilp, V., High performance hybrid composites for thin cementitious products: the next generation. In “High-Performance Cement-Based Concrete Composites”, J.J. Biemacki, S.P. Shah, N. Lakshmanan and S. Gopalakrishnan, eds., The American Ceramic Society, Westerville, Ohio, 2005, pp. 55-72. 15. Li, V.C., On engineered cementitious composites (ECC) - A review of the material and its applications. Joumal of Advanced Concrete Technology, 1,2003, pp. 215-230. 16. Li, V.C., Wang, S. and Wu, C., Tensile strain-hardening behavior of PVA-ECC. ACI Materials Joumal, 98,2001, pp. 483-492. 17. Li, V.C., Engineered cementitious composites. In Construction Materials, Proceedings of ConMat ’05 and Mindess Symposium, N. Banthia, T. Uomoto, A. Bentur and S.P. Shah, eds. University of British Columbia, Vancouver, 2005, CD-ROM. 18. Li, V.C. and Stang, H., Elevating FRC material ductility to infi-astructure durability. In “Fibre-Reinforced Concrete, BEFIB 2004, M. di Prisco, R. Felicetti and G.A. Plizzari, eds. RILEM Proceedings PRO 39, RILEM Publications, Bagneux, France, 2004, Vol. 1, pp. 171186. 19. Thibaux, T., Hajar, Z., Simon, A. and Chanut, S., Construction of an ultra-highperformance fibre-reinforced concrete thin-shelled structure over the Millau viaduct toll gates. In “Fibre-Reinforced Concrete, BEFIB 2004, M. di Prisco, R. Felicetti and G.A. Plizzari, eds.

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RILEM Proceedings PRO 39, RILEM Publications, Bagneux, France, 2004, Vol. 2, pp. 11831192. 20. Parent, E., and Rossi, P., A new multi-scale cement composite for civil engineering and building construction fields. In “Advances in Concrete through Science and Engineering, RILEM Publications, Bagneux, France, 2004, CD-ROM Paper No. 14, Hybrid-Fiber Session. 21. Aarup, B., CRC - a special fibre reinforced high performance concrete. In “Advances in Concrete through Science and Engineering, RILEM Publications, Bagneux, France, 2004, CD-ROM Paper No. 13, Hybrid-Fiber Session. 22. Vandewalle, L., Test and design methods €or steel fibre reinforced concrete based on the o-&-relation. In “Some Aspects of Design and Application of High Performance Cement Based Materials”, A. M. Brandt, ed. AMAS Lecture Notes 18, Institute of Fundamental Technological Research, Polish Academy of Sciences, Warsaw, 2004, pp. 135-190. 23. Stang, H., Mechanics of FRC materials and structures. In “Some Aspects of Design and Application of High Performance Cement Based Materials”, A. M. Brandt, ed. AMAS Lecture Notes 18, Institute of Fundamental Technological Research, Polish Academy of Sciences, Warsaw, 2004, pp. 83-133. 24. O’Leary, J. and Lemay, L., The P2P initiative. In: “Concrete Technology”, J. Skalny, S. Mindess and A. Boyd eds. The American Ceramic Society, Westerville, Ohio, 2006, pp. 163170.

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Pubt., Warsaw 2006

ON CONNECTIVITY OF POROSITY IN MODEL CEMENT PASTE Piet STROEVEN, Huisu CHEN, Martijn STROEVEN Faculty of Civil Engineering and Geosciences, Delft University of Technology 2628 CN, Delft, The Netherlands e-mail: [email protected]

ABSTRACT Fresh model cement mixtures were simulated by SPACE,which is based on a dynamic mixing algorithm, because of the configuration-sensitivity of depercolation. Mixtures had the same w/c ratio and particle size distribution. Thereupon, they were hydrated by HYMOSTRUC 3D system. Boundary conditions were varied, rendering possible assessment of the influences of aggregate grains in concrete on the depercolation process accompanying maturation. Serial sectioning and image analysis was applied for that purpose. Simulation results revealed an increase in total porosity and in connected fraction of capillary pores due to the existence of aggregate grains. The various stages in the depercolation process were found related to total porosity, and to image resolution, but were also governed by the spatial distribution of capillary pores, and thus of the cement particles in the fresh state. This configuration-sensitivity declines during the depercolation process. Hence, simulation of hydrated cement pastes on the basis of HYMOSTRUC 3D simulations of the fresh particle state would generally lead to biased information, whereby the degree of bias will decline during the process. Thus, generation of approximate information on the correct value of the depercolation threshold can be expected only.

Keywords Cement paste, percolation, porosity, boundary, image resolution, modelling approach

INTRODUCTION Connectivity of pore space exerts significant influences on particularly the durability performance characteristics of concrete. Pore connectivity is declining during the pore de-percolation process that accompanies the cement maturation in time. The porosity level at which the capillary pore network is completely de-percolated, so that pores do not connect outer surfaces of specimens, is defined as the de-percolation threshold. Computer simulation makes it possible nowadays to study this phenomenon. Various approaches can be mentioned [ 1-31. Image resolution is recognized to significantly influence the de-percolation threshold of capillary porosity [ 1,3-51. Research work generally focused on pore percolation in bulk paste. Detailed information on the de-percolation process and its limiting stage of full de-percolation in cement paste with dispersed aggregate grains surrounded by ITZs is missing. So, this paper will extend the scope to this relevant issue, because it is found that neighbouring aggregate grains will influence the de-percolation process in the cement pocket [6, 71. Moreover, it will critically analyze the simulation methodology of the model pastes on which studies are based. Volume fraction of aggregate varies from 60% to 80% in ordinary concrete [8]. SEM observations by Diamond, et al., revealed average nearest surface spacing between neighbouring aggregates in concrete to be in the 75 pm to 100 pm range [9]. Nearest surface

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Piet STROEVEN, Huisu CHEN, Martijn STROEVEN

spacing between neighbouring aggregates has been demonstrated analytically fluctuating (mainly) between 0.1 pm and 200 pm in normal concrete [lo]. This sets the scene for simulation approaches pursuing quantitative assessment of the influence of aggregate spacing on the de-percolation process. MODELING METHODOLOGY

Generation of fresh model pastes can be achieved following different strategies. The particle packing system in the fresh state is simulated in most existing cases as schematization of reality by what is called a physical model. This is not the only possibility, but a presentation of other approaches is outside the scope of this paper [ll]. However, even the category of physical models encompassestwo different methodological strategies, (implicitly) aiming to schematize reality on different levels. The two systems have distinctly different operational capabilities, however. The commonly employed RSA system is based on a Random Sequential particle Addition procedure, whereby particle overlap during the random generation process results in rejection followed by re-generation. The HYMOSTRUC 3D system, which is (only) used for hydrating the fresh cement mixtures, is an example of RSA systems. They provide only relatively low-density mixtures in which the particle dispersion does not properly schematize reality; this is simply not part of their operational capabilities. The second category encompassesconcurrent algorithm-based systems, either of static [121 or of dynamic character. The SPACE system, resorting under the last category [13], solves the particle overlap problem in a dynamic mixing stage. Particle mixtures in the full range of densities can be generated, whereby the operational capabilities of such systems are extended to providing proper schematization of reality as to the particle dispersion characteristics. The proper schematization of reality as to material composition does not constitute a very demanding requirement for the physical simulation system. Systems based on RSA or concurrent algorithms can both be employed. However, when proper schematization of reality is extended to material configuration (particle dispersion as to location and size), only concurrent algorithm-based systems can fulfil the requirements; it is part of their operational capabilities. Since the interest in this paper is in the pore de-percolation process it can be expected that pore configuration aspects will be highly relevant. The fresh model cement paste cubes were therefore generated with the SPACE system in which the cement particles were modelled as spheres. Hydration of the cement pastes was accomplished by means of the HYMOSTRUC 3D system, although SPACE also encompasses a hydration stage. However, for assessment of pore connectivity use was made of serial sectioning procedures implemented in HYMOSTRUC 3D by Ye [11. EXPERIMENTALAND MATERIALS

Three boundary conditions of the container were implemented. The first is denoted “periodic boundaries” (abbreviated as P) (shown in Fig. 1 and Fig. 2(a)), and used to simulate bulk paste. In the second case, “rigid boundaries” (denoted as C) were employed for all container walls. This is to simulate cement paste in the pockets formed by the aggregate skeleton. The last type is called “semi-rigid boundaries” (S). This implies all container surfacesto be periodic, but a rigid plane is inserted parallel to z-axis in the middle of the container, and is used to simulate the case

On connectivity ofpomsity in model cement paste

27

in which two surfaces of cubic cement paste are rigid (due to neighbouring aggregate particles), but other four sides (in x- and y-direction) of cement paste are periodic boundaries. Fig. 2 presents a schematic diagram of the three boundary conditions. To analyze the microstructure of cement paste with semi-rigid boundaries, the approach of periodic replication and coordinates translation (along f x, f y and f z directions) is used to obtain a model similar to Fig. 1 (c). After that, a single cubic model cement paste will be obtained with the same size as the original but just located between two rigid planes (in Fig. 3). Particle size distribution (PSD) in Fig. 4 and initial w/c ratio of 0.30 were similar in the various cement mixtures. Size of the representative volume element (RVE)for composition is three to five times the maximum grain size [ 14-16]. Diameter range of cement was set between 1-2Opm. Initial side length of cubic container was l O O p n , which is reduced after dynamic mixing to 92pm. Hence, this can be considered representative for compositional homogeneity only. Other relevant parameters of model structures are given in Table 1 . It can be expected that the de-percolation process will start as a highly configuration-sensitivephenomenon, and will reveal sensitivity to particle (and hence, pore) dispersion to decline during the hydrationinduced pore evolution process.

r

I

I

I

i

I

I

I

I

I

(a) 2D case: 8 additional periodic structures.

(b) 2D case: replication of particles in periodic units.

(c) 3D case: 26 additional periodic structures.

Fig. 1. Diagram for 2D and 3D periodic structures.

(a) Periodic boundaries ("P") (b) Rigid boundaries ("C")

(c) Semi-rigid boundary ("S")

Fig. 2. Fresh model cement paste structures produced with different boundary conditions.

Piet STROEVEN, Huisu CHEN, Martijn STROEVEN

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Table 1. Characterization of fresh model cement Code P300, C300, S300

Size [pm3] 92x92~92

Specific Surface Area [m2/kgl 300

Diameter

[WI 1-20

No. of Particles 31406

Table 2. Mineral composition of model cements

530

Fresh model stcucture

63.0

13.0

8.43

J

9.0

Model element for

hydration and microstructural analysis

Fig. 3. Selected fresh model cement paste structure with semi-rigid boundary condition.

e

Fig. 4. Particle size distribution of model cements (Blaine specific surface area of 300m2/kg).

29

On connectivity ofporosity in model cementpaste

Table 2 gives the mineral compositionof model cements. Curing temperature is 20 "C. The desired degree of hydration (DOH), defined as the amount of cement that has reacted at time t divided by the total amount of cement at time PO, should be defined before the simulation process starts [l, 171. The evolution of the degree of hydration (DOH) with time for model pastes is shown in Fig. 5. The hydrated model cement cubes are subdivided in serial sections, yielding successive images of pore and solid phases. What is displayed in the images depends on image resolution. Slice thickness governs the micro-structural level of pore space that can be analyzed for connectivity. Hence, image resolution and slice thickness should be adjusted to each other for economic purposes. In the present case, image resolution and slice thickness were set as 0.1 pdpixel in all model structures. Area fractions of solid and pore phases (identifiedby colour coding) are determined in each slice by image analysis, yielding upon multiplication with slice thickness volume fraction of phases of interest. Additionally, making use of partial overlap of slices allows assessment of pore (or solid phase) connectivity, and the ratio of connected to the total porosity. These algorithms are similar as described. 0.8

0.6

r

t

/

03

0

1

10

100

1000

lo000

Time @ours)

Fig. 5. Degree of hydration of cement paste as function of hydration time in [I].

RESULTS Effects imposed by category of simulation system Fig. 6 presents the de-percolation process that one single paste (Blaine number 157, w/c=0.3) undergoes during hydration simulated by HYMOSTRUC 3D [ 181. In one case, the fresh state was also simulated by the RSA-based HYMOSTRUC 3D system. Contrary, the fresh state of the cement paste was generated by concurrent algorithm-based SPACE system in the other cases, whereby observations pertain to different resolution levels. Following the curves starting at the upper right-hand corner, the three SPACE curves reflect a gradual decline of connected porosity with total porosity during the hydration. The too regular particle dispersion in RSA systems results in a far more sudden de-percolation of pore space. Hence, the practical situation of porous concrete of which part of the pores is still connected cannot be properly simulated by RSA-based systems as the phenomenon is of configuration-sensitive nature. The diminishing distance between the SPACE curves and the HYMOSTRUC 3D one is a reflection of the declining configuration- sensitivity during the de-percolation process. So, the de-percolation threshold seems quite configuration-insensitive,with the practical consequence that different systems can be expected leading to comparable estimates, as they indeed do [1,4, 181.

30

Piet STROEVEN, Huisu CHEN, Martijn STROEVEN

-

1.o

J

.g 0.8 u)

-e f

a 0.6

0.4

n

B

g

0.2

c

8 Total porosity (96)

0.0 0.0

0.1

0.3 Porosity (-)

0.2

0.4

0.5

Fig. 6. Evolution of pore de-percolation during hydration (versus porosity), at the left for model cement pastes of which the fresh state was generated by HYMOSTRUC 3D and by SPACE [181, and at the right according to Garboczi and Bentz [4]. Influence of resolution is also depicted (loo3,2003 and 4003correspond to a resolution of 1,0.5 and 0.25 pdpixel, respectively). Influences exerted by boundary conditions Fig. 7 shows the visualized 3D images of model pore structures at ultimate DOH for the three differentboundary conditions, P, C, and S. Differences are obvious. Almost all of the boundary pores in C300 are inter-connected. The boundary pores were shown also inter-connected for S300.The " wall effect" in the fresh state is known leading to relatively inferior cement packing densities in a thin layer near to the aggregate grain's surface [19]. Hydration products of neighbouring cement particles will tend to fill these deficiencies, and local hydration rates will be higher, as revealed by computer simulation [6, 201. The inert aggregate itself cannot contribute to filling the free space between aggregates and hydrated cement particles near its surface. Thus boundary pores may have more chance than pores in bulk to be mutually connected. The total porosity at ultimate degree of hydration for " P300", C 3 0 0 and "S300" are 1.24%, 2.92% and 2.22%, respectively. The connected fraction of porosity, which is defined as the ratio of connected to total porosity, is zero for "P300", and 0.857 for "C300". For "S300", this connected pore fraction is zero in z-axis direction, but 0.482 in x- and y-direction. Since the w/c ratio is the same for all pastes, initial porosity must have been equal. However, different boundary conditions led to the large differences in total porosity and in pore connectivity of pastes at ultimate DOH. The differences in pore connectivity of "P300" and "C30O"seem to indicate that this is due to spatial distribution of pores. The pores tend to accumulate in the boundary region and form inter-connectedpaths due to the existence of rigid boundary. It is the inter-connectedboundary porosity that led to higher connected kaction of pores. Even when all pores were located in the boundary region and porosity would be zero in the interior, connected fraction of pores could still be high when boundary pores were largely interconnected. Hence, the de-percolation threshold of pores is governed by the spatial distribution of pores. The differences in connected porosity in x-, y-, and z-direction of "S300" also provides similar information about the influence of boundary conditions. To verify this viewpoint, a boundary

31

On connectivity ofporosity in model cement paste

layer of gradually increasing thickness was removed on all six sides (similar to the peeling process) from "C300" at ultimate degree of hydration, whereupon the total porosity and pore connectivity of the remaining cubic core were analyzed (shown in Fig. 8). For instance, when the thickness of the removed shell is 0.4 pm, correspondingto a cube with linear dimension of 91.2 pm, the porosity in this cube is 2%, and the connected fraction of porosity is 20% (see Fig. 8). Results in Fig. 8 demonstrate pore connectivity restricted to 0.5 pm thick surface layer. Hence, connected fraction of porosity of the inner core is declining steeply with core size, despite only moderate decline in total porosity of the core.

(a) P300 (periodic boundary) (b) C300 (rigid boundary) (Total Porosity=l.24%) (Total Porosity=2.92%)

(c) S300 (Semi-rigid boundary) (Total Porosity=2.22%)

Fig. 7. Porosity for different boundary conditions at ultimate degree of hydration.

4'

(a)

I

.

'

'0

"

'

'

0

a4 OB 13 L6 Thickness of the removed layer (bm)

OA

0.0

1.2

1.6

Thickness of the removed layer b m )

(b)

Fig. 8. Total capillary porosity and connected pore fraction in inner core of model paste. The three-dimensional information on the spatial distribution of porosity and of the connected fraction of porosity, strikingly reflected by Fig. 8, is obtained by a laborious procedure of serial sectioning. In [20, 2 11, the two-dimensional descriptors area fraction, AA, and mean &ee spacing, 1, are used for this purpose. Fig. 9 demonstrates that this simpler, two-dimensional andfar more economic approach can provide similar type of information

32

Piet STROEVEN, Huisu CHEN, Martijn STROEVEN

1

Hydrated for 100 h

Hydrated for 100 hours

-k

0.8 -

8

0.4

0.6 -

-

0.4 d

0.2-

0.2

07

De-percolation during hydration Proceeding similarly, the influence of DOH on the total and connected porosity was further investigated in all these cases. Capillarypores at the boundary are connected at ultimate DOH in x- and y- directions for “S300, so, only evolution in pore connectivity along the z-axis is investigated in this section. The results are shown in Fig. 10. Differences in total as well as connected fraction of porosity between the three model pastes are small at relatively low DOH. The influences of boundary conditions become apparent only after a certain DOH. Nevertheless, connected fraction of porosity of “C300” exceeds that of the other two model paste at any value oft. De-percolation occurs at a total porosity of 3.8%-4.2% for “S300”, and of 2.3%-2.6% for ”P300” (close to Elam, et al., [22]). No de-percolation occurs for model paste “C300”; even at ultimate DOH, the connected fraction is still larger than 80%. Thus, the de-percolation threshold of capillary porosity is not only related to the total porosity of material and to image resolution, but also to the spatial distribution of capillary pores.

I j

’

II

I I

II

a

an

41

Total porosty (%)

51

Degree of hydration

Degree of hydration

Fig. 10. Evolution of pore structure of model paste generated with different boundary conditions. The solid line stands for periodic boundaries (P300), the dashed line for rigid boundaries (C300), and the pointed line for semi-rigid boundaries (S300), respectively.

On connectivity of porosity in model cement paste

33

CONCLUSIONS This contribution investigated by computer simulation the influence of boundary conditions on the de-percolation process of capillary pores in maturing concrete. To do so, the operational capabilities of the available physical computer simulation system had to be evaluated. The following conclusions can be drawn: Only usage of concurrent algorithm-based physical computer simulation systems (like SPACE) for the generation of the fresh cement particle system can guarantee reliability of estimates for successive stages in the pore de-percolation process that accompanies hydration of cement paste. Operational capabilities of RSA-based systems (like HYMOSTRU 3D) for the same purpose are restricted to compositional properties. It seems that the de-percolation threshold has low configuration-sensitivity, so that RSA-based systems can provide approximate information on this final stage of pore de-percolation only. The aggregate in matured concrete not only affects the total porosity in the cement pockets between aggregate grains, but exerts a significant influence on pore connectivity as well. The connected fraction of porosity is concentrated in part of the ITZ layer in contact with the aggregate grains with a thickness of less than 1j.m. The connected fraction of porosity therefore reduces significantly inside the ITZ at increasing distances from the aggregate grain’s surface. The degree of de-percolation at any stage of the pore space evolution process during hydration of concrete is governed by total porosity and by image resolution, but also influenced by the spatial distribution of capillary pores, and thus of the cement particles in the fresh state. The pore dispersion influence is smallest at the end of the de-percolation process.

REFERENCES 1. Ye, G, Experimental Study & Numerical Simulation of the Development of the Microstructure and Permeability of Cementitious Materials. PhD Thesis, Delft University of Technology, Delft, 2003, pp 186 2. Garboczi, E.J., Bentz, D.P., Modelling of the microstructure and transport properties of concrete. Construction and Building Materials, 10, 1996, pp 293-300 3. Navi, P., Pignat, C., Simulation of cement hydration and the connectivity of the capillary pore space. Advanced Cement based Materials, 4, 1996, pp 58-67 4. Garboczi, E.J., Bentz, D.P., The effect of statistical fluctuation, finite size error, and digital resolution on the phase percolation and transport properties of the NIST cement hydration model. Cement and Concrete Research, 3 1,2001, pp 1501- 1514 5. Ye, G., Percolation of capillary pore in hardening cement pastes. Cement and Concrete Research, 35,2005, in press 6. Chen, H., Ye, G.., Stroeven, P., Computer simulation of structure of hydrated cement paste enclosed by interfacial transition zone in concrete. In: International Conference on Durability of High Performance Concrete and Final Workshop of CONLIFE, Setzer, M.J. and Palecki, S. (Eds.), AEDIFICATIO Publishers, Freiburg, 2004, pp 133-144 7. Chen, H., Stroeven, P., Ye, G., Influence of aggregate surface spacing on the microstructure of fresh and hardened interfacial transition zone. Proceedings of 3rd International Conference on Construction Materials: Performance, Innovations and Structural Implications (ConMat’O5) Vancouver, Canada, August 22-24,2005. 8. Nawy, E. G . ,Fundamentals of High Performance Concrete. 2nd Edition. John Wiley & Sons,

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Piet STROEVEN, Huisu CHEN, Martijn STROEVEN

New York, 2001, pp 14 9. Diamond, S., Mindess, S., Lovell, J., On the spacing between aggregate grains in concrete and the dimension of the aureole de transition. In: Liaisons Pastes de CimentlMat6riaux Associgtes, RILEM, Toulouse, 1982, C42-C46. 10. Chen, H., Stroeven, P., Stroeven, M., Nearest surface spacing between neighbouring aggregate particles in concrete: theoretical solution. In: International Conference on Advances in Concrete Composites and Structures, Chennai, India, January 6-8,2005, in press 11. Stroeven, P., Hu, J., Modelling in Concrete Technology - Balancing between Science and Alchemy? Cement and Concrete Research, 2005, submitted for publication. 12. Williams, S.R., Philipse, A.P., Random packings of spheres and sphero-cylinders simulated by mechanical contraction, Physical Review E, 67,051301,2003, pp 1-9 13. Stroeven, M., Discrete Numerical Modelling of Composite Materials. Delft University of Technology, Delft, 1999, pp 224 14. Stroeven, P., Some Aspects of the Micromechanics of Concrete. Delft University of Technology, Delft, 1973, pp 329 15. Brown, C.B., Minimum volumes to ensure homogeneity in certain conglomerates. Journal of Franklin Institute, 279, 1965, pp 189-199 16. Hashin, Z., Analysis of composite materials: a survey. Journal of Applied Mechanics, 50, 1983, pp 481-505 17. Breugel, K. van, Simulation of Hydration and Formation of Structure in Hardening CementBased Materials. Delft University of Technology, Delft, 1997, pp 305 18. Chen, H., Numerical Modelling on ITZ Microstructure and Its Influence on the Effective Elastic Property and Diffusivity of Concrete. PhD Thesis, Delft University of Technology, 2006, to be published 19. Stroeven, P., Stroeven, M., Reconstruction by SPACE of the Interfacial Transition Zone. Cement and Concrete Composites, 23,200 1, pp 189-200 20. Hu, J., Stroeven, P., Properties of the Interfacial Transition Zone in Model Concrete, Interface Science, 12,2004, pp 389-397 21. Hu, J., Chen, H., Stroeven, P., Concrete porosimetry: aspects of feasibility, reliability and economy, Cement Concrete Composites, submitted for publication, 2005 22. Elam, W.T., Kerstein, A.R., Rehr, J.J., Critical properties of the void percolation problem for spheres. Phys. Rev. Lett., 52, 1984, pp 1516-1519

Proc. Int. Symp. 'Brittle Matrix Composites 8" A.M. Brandt, YC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

MOISTURE DEPENDENCE OF PORE SIZE AND SPECIFICSURFACE AREA OF HARDENED CEMENT PASTE DETERMINEDWITH SAXS AND INVERSE GAS CHROMATOGRAPHY Jiirgen ADOLPHS POROTEC GmbH Niederhofheimer Strasse 55a, D-6519 HofXeidTs., Germany juergen. [email protected]

ABSTRACT The main constituent of hardened cement paste is the calcium silicate hydrate (CSH). Pore size and specific surfae areas change depending on relative humidity. New applications with two different techniques support these findings. A new interpretation of small angle X-ray scattering (SAXS) data correlates almost perfectly the fractal information (deduced from the linear Porod regime) with the pore sizes determined with Mercury intrusion porosimetry (MIP). In particular the moisture dependent pore size changes from MIP [2] are now shown independently by the noninvasive SAXS at least below 60% rh. Another issue is the specific surface area, that varies depending on relative humidity as previously reported from SAXS data. This is now supported by another technique - the inverse gas chromatography (iGC). In the iGC technique pulses of organic probe molecules are injected in almost infinite dilution into a carrier gas of controlled humidity and the retention time of the probe molecules is measured. From the retention times various thermodynamic properties can be determined, like the dispersive part of the surface energy [3]. However, it is also possible to gain the adsorption isotherms although limited to very low relative pressures due to the almost infinite dilution. Applying the ESW (Excess Surface Work Sorption Model) [4], allows to recover the isotherm so far that it is possible to determine the monolayer (with ESW or BET) and accordingly the specific surface area.

Keywords SAXS, Mercury intrusion porosimetry, inverse gas chromatography, excess surface work, isotherm, pore Size, relative humidity, specific surface area INTRODUCTION The main constituent of hardened cement paste either in normal concretes and even in UHPC is the calcium silicate hydrate (CSH) [l, 21. The nature of this phase is yet not completely understood, in particular its interaction with water vapor. CSH is characterized by its nanoporous structure, therefore all properties in particular transport, durability and strength must be regarded on the nanometer level. On this level a macroscopic physical description is no longer appropriate and the interaction with ambient vapours and gases cannot be regarded as a simple bulk reaction [2]. The ambient relative humidity changes in a nonlinear way the physic0 - chemical and mechanical properties of hardened cement paste. This includes the water uptake at cyclic wetting and drylng with the typical overall hysteresis, compressive strength, swelling and shrinkage, density, surface energies, transport properties and resulting chemical reactions like carbonation and acid - base properties [2,3]. In particular the pore size varies dramatically depending on relative humidity, this was measured with Merucry intrusion

36

Jiirgen ADOLPHS

porosimetry (MIP).In order to complete the above picture two new results will be presented. A new interpretation of SAXS data correlates almost perfectly the fractal information (deduced from the linear Porod regime) with the pore sizes determined with MIP. In particular the moisture dependent pore size changes from ME' [4] are now shown independently by the noninvasive SAXS at least below 60% rh. From inverse gas chromatography not only surface energies can be received [3], but also information about changes of the specific surface area depending on relative humidity. SAXS EXPERIMENT

The small angle X-ray scattering (SAXS) experiments were conducted at the Argonne National Lab APS DND-CAT Beam Sector 5ID. The variable intensity of the X-Ray source was fixed to 8 KeV corresponding to a wavelength of 0,155 nm. A pinhole with 1 pm diameter filtered the parasitic scattering in front of the sample. The X-rays entered then an evacuated tank with Kapton windows before they are detected by a two dimensional 1024x1024 pixel MAR-CCD detector (MAR Corp. Evanston, ILL. USA) with a diameter of 0.1399 m and a pixel resolution of 64.449 p d i x e l . The distance between detector and cement paste samples was 2.45 m and between the MCM-36 zeolites 0.75 m. The various relative humidities in the sample holder were generated by two methods. Either the hcp sample were preconditioned above saturated salt solutions or via an attached vapor generator where air is flowing over a temperature controlled water / ice bath and the humidity is measured with a dew point analyzer (EG&G 911). The inlet and outlet windows of the sample holder were sealed with Kapton tape while the preconditioned specimen were put on the inner sticky side of the tape. The vapor generator could generate relative humidities between 26.2% rh and 82% rh. The ambient temperature was 20.9 "C. The exposure time was set to 1 sec with a 10 times sampling rate. EVALUATION OF SAXS DATA FOR PORE SIZE DETERMINATION

The SAXS scattering patterns of the hardened cement paste measured with the 1024 x 1024 pixel MAR detector were circular while the center beam was stopped with a lead absorber. Each measurement was averaged over the diameter line of 1024 pixels with a width of 20 pixels. These values were normalized to the according beam intensity, which decreased with time after every ring filling of the synchrotron. This varying synchrotron beam intensity was first time measured at APS. The same procedure was applied for the empty sample cells with the same exposure times for all employed relative humidities. The intensities I are normalized to the maximum intensity of the according synchrotron ring filling at time tl. A background correction is done with the empty cell measurement at time tz and same relative humidity:

This intensity is plotted as a function of the scattering vector Q (Figure 1). Q is determined by the pixel distance and the tank length:

with d = pixel . 6,44.105 [m] the radial distance on the detector, tank length I T ~ A[m],

Moisture dependence ofpore size and specific surface area of hardened cement paste determined...

37

and wave length h =1,55.10-"m. From a double logarithmic plot of the intensity Z as a function of the scattering vector Q the Porod constants are determined (figure l), in particular the Porod exponent n and the Porod factor k.

k = Z(Q).Q"

(3)

The Porod factor k can be correlated with the specific surface area [5-71. In this work the Porod exponent n was not fixed for the plots as often done with values of 4 (applied for a 2dimensional detector, point geometry) or 3 (detector with slit geometry). Since this Porod exponent is related to the fractal geometry [8, 91 it is kept variable. However, the origin of this fractality can only be found in the porosity, as Winslow [9] mentioned. In the present case of hardened cement paste it would be more adequate to speak about the nano porosity of the mesopores in the pore radius range of 50 nm and smaller. By an empirical calibration procedure without any modeling assumption different to [7], it is possible deduce a reasonable pore size from the Porod exponent. For the calibration the SAXS data of a MCM 36 zeolite with a similar lamellar morphology like hardened cement paste is used. This MCM 36 zeolite is well investigated with a pore half width of 1.5 nm. Further the pore size of a dry hardened cement paste (OPC, w/c = 0.4, r = 25nm) measured with Mercury intrusion porosimtery [4] was taken. The sorption isotherms of both materials, in particular the water vapor sorption exhibit the typical triangular shape of slit like or lamellar pores. The pore radius r from the SAXS Porod exponent n was then computed as r = -10.235 . n - 3.5152 nm.

(4)

RESULTS AND DISCUSSION OF SAXS OF HARDENED CEMENT PASTE MEASURED AT VARIOUS RELATIVE HUMIDITIES At least not the absolute pore size values were of interest, in this work the more interesting goal was to investigate the behaviour with respect to relative humidity. In figure 2 the pore radii from the drylng branch (water vapor desorption) are presented. Indeed the observed changes of the pore size deduced from the Porod exponent are similar to those from the Mercury intrusion porosimetry as published in [lo]. Differences, in particular at the higher relative humidities may result from the slightly different wlc ratio (MIP 0.4 and SAXS 0.3). But even in the region above 60% rh the changes in the shape of the curve are similar, although not so pronounced for SAXS as for MIP.

Jiirgen ADOLPHS

38

33% rh y=1.46E-07~~~~~~'~ R2 = 9.99E-01 1E+Ol

0% rh 1E+01

5

-

d

5

IE-O1

-

4-

1E-01

d 1E-03

1E-03 1E-05 0.01

y = 3,330E-07~-~ 451E+Ci R2= 9.288E-01

-l

1E-05 0.01

0.10

Q [l/A]

-

1-

1

-

1E+01

5

1E-01

d 1E-03

~~

~

-

1E-01

d 1E-03

~1 1

1E-05 0.01

v H 54% rh

y = 1.408E-07~-~.~~~~'" RZ= 9.778E-01

1E+01

5

-

y =2 . 3 7 1 E - 0 6 ~ ~ . ~ ~ ~ ~ + ~ R2 = 9.989E-01

1-

I

1E-05 0.01

0.10

0.10

Q [l/] 75% rh

I

0.10

Q [l/A] 44% rh

'd

-I-

I

Q [l/A] 86% rh

y=3.058E-05~~~.~~~'~ R2 = 9.972E-01

y= R2 = 9.995E-01

1E+01

5

1E-01

-

1E-05 0.01

1E-03 1E-05

0.0 Q [l/A]

Q [I/A]

I

I I

91% rh

y=

R2 = 9.973E-01

IE+OI

1E-02

5

1E-03

- 1E-04 L

5

1E-01

-

1E-03

d

1E-05 0.01

0.10

Q [VA]

I

I

i

94%

,+,

y =4 . 0 5 0 E - 0 8 ~ - ~ ~ ~ ~ ~ + ~ R2 = 9.961E-01

j

1

1E-05 0.0

0.1 a 1414

Figure 1. Plot of the scattering vector Q as a function of the intensity Z and determination of the Porod constants (factor and exponent).

Moisture dependence ofpore size and specific surface area of hardened cement paste determined...

1 -+MIP H20 Desorption hcp w/c=0.4

39

SAXS H20 Desorption hcp w/c=0.3

+

35

5 1

0

20

40

60

80

100

rel. Feuchte [%I

Figure 2. Pore size dsitribution for drylng from M P [4]and SAXS calibrated with MCM-36 pore width 1.5nm and 25nm from MIP at O%rh.

IGC EXPERIMENT Inverse gas chromatography (iGC) is a gas phase technique, where contrary to normal gas chromatography (GC) the solid phase is the unknown component to be tested with known vapours. While the aim of normal GC is the analysis of the single components of a gas mixture via their different capabilities of adsorpion on the solid phase in meters long curled columns, iGC is a technique where surface properties like surface energies or basic-acid properties are determined. Here the column in use (sample cell) is a straight 6mm and only 30 cm long glass tube treated with DMCS (dimethyldichlorosilane) with inner diameters of 2, 3 or 4 mm. The employed organic vapors are automatically injected via a loop into a carrier gas (Helium) at an almost infinite dilution to ensure a solely vertical interaction between the surface (sorbent) and the vapor molecules (sorptives). From the retention times measured with a TCD (thermal conductive detector) and a FID (flammable ionisation detector) it is possible to determine the various surface and interaction energies. The employed instrument was an IGC 2000 (SMS Ltd, London), which can provide controlled relative humidity of the carrier gas between O%rh and 90%rh. Further the iGC instrument has 9 reservoir flasks to work automatically with different organic liquids resp. vapors. The dead space time is measured with methane gas. In [3] iGC measurements were taken to deduce the surface energies. In particular it could be shown that there is a linear relationship between shrinkage and decrease in the dispersive part of the surface energy. This was supporting the idea of the Munich model regarding the effect of adsorbed water molecules reducing the surface energy of the solid hardened cement paste and causing shrinkage resp. swelling below 40% rh. In this contribution it will be shown how specific surface areas can be computed from iGC experiments. The hardened cement paste (hcp) specimens were produced from white cement with a w/c = 0.4 and were stored under saturated Ca(OH)* solution for 1.5 years before the iGC experiments. Shortly before the experiment the hcp samples were removed from the storage and crushed in a wet state to particles of ca. lmm size. About 1 gram was then filled into the iGC columns (30 cm long, 3mm inner diameter) and fixed at both ends with silanised glass wool. A special device was used for compacting the sample, so that still a gas flow was

40

Jiirgen ADOLPHS

possible through the column. The sampling procedure took 10 minutes, after this the column was set vertically into the separate iGC oven. The columns oven temperature was set to 30 “C f 0.1 “C. Decane, Nonane, Octane, Heptane and Hexane (HPLC quality) were employed in 4% concentration with a flow rate of 10 ml/min to investigate the dispersive surface energies. In addition with same conditions the polar probe molecules Triethylamine, Ethyl Acetate, Chloroform and 1,4-Dioxane were used for the determination of the specific polar interaction and the acid-base properties. The Helium carrier gas was conditioned at 85%rh for the fist test and then in a desorption (drymg) series to 70%, 50%, 30%, 20% and less than 0.1% (which will be called 0%) relative humidity. IGC THEORY

In an iGC pulse experiment an injection of an adsorbate is made into a humidity controlled carrier gas, usually helium, which is then transported through the GC to the column containing the solid material under investigation. Here the injected probe molecules are entirely adsorbed and after the so called retention time, they are eluted by the carrier gas. The time and intensity of the eluted probe molecules is then measured by the detectors (FID, TCD). The intensity is corrected by the baseline resulting from the carrier gas and by the “dead time” of the instrument. In the infinite dilution range it is assumed that preferably at locations with the highest interaction potentials probe molecules are adsorbed and no multilayer adsorption occurs. Ideally the adsorption isotherm would be linear (Henry type). In such an ideal case peaks are symmetrical (Gaussian) and the retention volume can be calculated from the retention time at the peak maximum (figure 4 lee). However, in reality a tailing is usually observed, and then the retention time is taken from the centre of mass of the peak, rather than the maximum height. The corrected net retention volume VNis given by

with T the column temperature, F the exit flow rate at 1 atm and 273.15 K, tR is the retention time for the adsorbing probe and to is the mobile phase hold-up time (dead time). ‘y” is the James-Martin correction, which corrects the retention time for the pressure drop in the column bed.

If the peak of the retention time shows a tailing, a method called Elution by Characteristic Point (ECP) which is part of the instrument software can be applied to deduce an isotherm up to the maximum concentration resp. relative pressure (fugure 4 right). The idea behind is that this retention time peak contains all the smaller peaks of lower concentration. By an iterative numerical method a series of retention volumes is computed and finally transformed into an isotherm. The retention time is directly related to the first derivation of the amount adsorbed and the peak height is related to the partial pressure.

Moisture dependence ofpore size and specific surface area of hardened cement paste determined ...

~

~~

1

_

IGC RetentionTime of Hardened Cement Paste White Cement wbO.4

I

5

0

10

Ret-time [min]

15

41

_

isotherm of Hexane on HCP at 50%RH from iGC RetentionTime with ECP-Method

0

0.001

0.002

0.003

PIP0

L_-__.--_-

Figure 4. Example of iGC Hexane retention time and transformation with the ECP method into an isotherm. Notice: The maximumplpo = 0.003 and can not be used for monolayer and specific surface area calculations according BET. Because the isotherm yield with the ECP method reaches only up to extreme low relative pressures (e.g. Hexane at 50% rh p/po = 0.003, figure 4), these small pieces of isotherm are I S 0 9277 to determine specific surface useless for the application of BET (0.05

The Gibbs equation describes the sorption r as the ratio of surface free energy in chemical potential Ap:

(7)

xvand change

AfZer integration we yield: Y,~

0

Jdy,,= - Jr.dAp=Yd+Yb

(9)

AP

s= solid, I= liquid, F vapour.

The term TAp is a new thermodynamic function called excess surface work @. We can define that the Excess Surface Work is the sum of Surface Free Energy and Isothermal Isobaric Work of Sorption. In this work T is replaced by the adsorbed volume Vah. A transformation from an isotherm into an ESW plot, which is fundamental and not based on a model, yields at least one minimum (figure 6). From the definition, we see that the maximum of interaction is described at this point, and therefore it may be related to the formation of an adsorbed monolayer. Then it is possible like in BET standard analysis to compute a specific surface

Jiirgen ADOLPHS

42

area. A comparison is given for over 300 isotherms in [lo]. The depth of the ESW minimum gives as a rule the loss of degree of freedom of an adsorbed molecule, and therefore it is suitable to characterize the sorption energies [14]. A solution at the minimum is

@,in

d a =a p r + r a p= o

(10) 3 ln(lAp1)= --

r

I)

ln(lApo = 0

rmO"0

Due to the exponential form it can be referred to the structural component of an isotherm, thus it is called the structural model of ESW (figure 5). The constants are explained as the , and the onset potential for adsorption A h . This model is superior to monolayer capacity r BET, since the isotherms of nonporous systems can be described over the entire range of relative pressures with the two parameters r , and A h , resp. from the ESW plot [101. ESW Linear Plot of Hexane on HCP at 50%RH

P-

5

4= -

2.1 2 1.9 1.6 1.7 1.6 1.5

0

0.0005

0.001

0.0915

0.002

V h (mug)

Figure 5. Example of iGC Hexane injection at 50% rh. Application of the structural model, the linear part of the Excess Surface Work (ESW). From the slope a monolayer capacity can be determined and a sorption energy from the intercept. The marked points are from iGC ECP. From the constants a complete isotherm can be computed.

ESWMinimum Plot Hexane on HCP at 509CRH 0

0.01

0.02

0.03

0.04

0.05

0 -0.5 1

-1.5

L

3

-2

-2.5 -3

-3.5

Figure 6. Example of iGC Hexane injection at SO%rh. Plot of the Excess Surface Work (ESW) in units of RT. Notice: This is fundamental and another modelless way of displaying an isotherm. The marked points are from the siotherm in figure 5. The minimum is related to the highest interaction of adsorbed molecules and surface, and therefore defined as a monolayer capacity. The loss of energy of -3 RT is typical for physisorption.

Moisture dependence ofpore size and specijic surface area of hardened cement paste determined ...

43

RESULTS AND DISCUSSSION OF THE IGC EXPERIMENTS

In table 1 the specific surface areas determined from the Excess Surface Work method ESW are summarized for the Hexane, Heptane and Octane iGC experiments. The value for Octane at O%rh was taken from a DVS (Dynamic Vapor Sorption, dynamic gravimetric sorption device) measurement. The absolute values differ due to the size and occupied molecular surface area of the alcanes. Table 1. Specific surface areas determined with ESW Mol. Area Hexane 51.5 A’

rh

Heptane 57.3 A*

Octane 63 A’

m2/g

SHeptane m2/g

Soctane m2/g

28.37 28.22 24.10 13.42 9.11 15.05

10.03 10.62

6.41 *

5.29 4.83 6.60

1.07 1.55 4.80

SHexane

In figure 7 the surface areas at the various relative humidities of Hexane and Heptane are related to the corresponding specific surface area at 0% rh. Both graphs almost coincide and show a distinct minimum around 70% rh with a decrease in surface area by almost 70%. Similar results were received with SAXS by Volk et a1 [5] and Bier [15], as well as from the author with SANS [16], all showing a minimum of specific surface area in the intermediate range of ambient relative humidity. 1.2

0.2

,

I

t 0

20

40

60

RH (%)

80

100

I

Figure 7. Relative change of the specific surface areas as a h c t i o n of relative humidity from iGC experiments with Hexane and Heptane. Taking the interpretation of Beddoe et a1 [8] regarding the fractal dimension measured with S A X S , the decrease in surface area is explained by smoothing the hcp surface with adsorbed water molecules. On the other hand these adsorbed water molecules decrease the surface tension as measured with iGC [3] and cause swelling. From this observation one may expect an increase of surface area, not a continous decrease. Further, in all the investigations after passing the minimum an increase of specific surface area at higher relative humidities was

44

Jiirgen ADOLPHS

observed. This nonlinear behaviour is hardly to understand, in particular special effects occur at 30% rh when CSH gelpores are filled with water [2].

ACKNOWLEDGEMTS The author acknowledges the fundings of the DFG Ad 14411-3 (German Research Foundation). Further he is grateful to Prof. Dr. R. Kohler and B. Rimer IAF Reutlingen and Dr. F. Thielmann SMS Ltd, London for helpful discussions. Portions of this work (SAXS) were performed at the DuPont-Northwestern-Dow Collaborative Access Team (DND-CAT) Synchrotron Research Center located at sector 5 of the Advanced Photon Source. DND-CAT is supported by the E.I. DuPont de Nemours & Co., The Dow Chemical Company, the U.S. National Science Foundation through Grant DMR-9304725, and the State of Illinois through the Department of Commerce and the Board of Higher Education Grant IBHE HECA NWU 96. Use of the Advanced Photon Source was supported b y the U.S. Department of Energy, Basic Energy Sciences, Office of Science, under Contract No. W-3 1-109-Eng-38.

REFERENCES 1. 2. 3. 4. 5. 6. 7. 8. 9. 10. 11. 12. 13.

14. 15. 16.

Adolphs, J., Schreiber, A., Microstructural Characterisation of Ultra-High Performance Concrete. Proc. Intern. Symp. on Ultra High Performance Concrete (eds. M. Schmidt, E. Fehling, C. Geisenhansltike),Kassel Germany 2004, pp 265-272 Adolphs, J., The Role of Nanopores in Concrete, Proc. Intern. RJLEM Workshop 'Frost Resistance of Concrete' (eds. Setzer, Auberg, Keck) Essen Germany 2002, pp. 45 - 52 Adolphs, J., Surface Energies of Hardened Cement Paste, Materials and Structures 38, 2005, 443-448 Adolphs, J., Heine, P., Setzer, M.J., Changes in Pore structure and Mercury Contact Angle of Hardened Cement Paste Depending on Relative Humidity. Materials and Structures 35, 2002, pp.477486 Volkl, J.J., Beddoe, R.E., Setzer, M.J., Cem. Concr. Res. 17, 1987, pp. 81 Mittelbach, P., Porod, G., Zur Rontgenkleinwinkelstreuung kolloider Systeme, KolloidZeitschrift - Zeitschrift fiir Polymere, Band 202, Heft 1, 1965, pp. 40-49 Radlinski, A.P., Mastalen, M., Hinde, A.L., Hainbucher, M., Rauch, H., Baron, M., Lin, J.S., Fan, L., Thiyagarajan, P., Application of SAXS and SANS in evaluation of porosity, pore size distribution and surface area of caoal, Intern. J. of Coal Geology 59,2004, pp. 245-271 Beddoe, R.E., Lang, K., Effect of Moisture on Fractal Dimension and Specific Surface of Hardened Cement Paste by Small-Angle-X-Ray Scattering. Cement & Concrete Res. 24, No.4, 1994, pp. 605-612 Winslow, D.N.: The Fractal Nature of the Surface of Cement Paste. Cement & Concrete Res. 15, NO.4, 1985, pp. 817 - 824 Adolphs, J., Setzer, M. J., A Model to Describe Adsorption Isotherms, J.Colloid Interface Sci. 180, 1996 pp. 70-76 Churaev, N.V., Setzer, M.J., Adolphs, J., Influence of Surface Wettability on Adsorption Isotherms of Water Vapor, J. Colloid Interface Sci. 197, 1998, pp. 327-333 Adolphs, J., Setzer, M.J., Description of Gas Adsorption Isotherms on Porous and Dispersed Systems with the Excess Surface Work Model, J. Colloid Interface Sci. 207, 1998, pp. 349-354 Adolphs, J., "Modeling of Gas Adsorption at Porous and Dispersed Surfaces", Encyclopedia of Surface and Colloid Science, 2002, pp. 3472-3482 Adolphs, J., Setzer, M. J., Energetic Classificationof Adsorption Isotherms, J. Colloid Interface Sci.184, 1996, pp. 443-448 Bier, T., Dissertation, TU Karlsruhe, 1988, p. 93 Adolphs, J., DFG Report Ad144/1 SANS Measurement on HCP, in German (2002)

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, YC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and WoodheadPubl., Warsaw 2006

THE EFFECTS OF MICROSTRUCTURAL FEATURES OF MORTARS ON THE LASER CLEANING PROCESS Poologanathan SANJEEVAN', Agnieszka J. KLEMM', Piotr KLEMh4* 'School of Built and Natural Environment, Glasgow Caledonian University, Cowcaddens Road Glasgow G4 OBA UK e-mail:[email protected] 'Technical University of Lo&, Poland

ABSTRACT The paper presents a part of the larger study on microstructural features of mortars and their effect on laser cleaning process. The paper focuses on presenting the results of investigation on the influence of surface roughness, porosity and moisture content of mortars on the removal of graffiti. ND: YAG laser was used for laser cleaning. The properties of this laser are as follow: wavelength (X): 1.06pm, energy: 500 mJ, pulse duration: 10 ps. The relationship between laser fluence and number of pulses required for the laser cleaning can be divided in two zones which are effective zone and ineffective zone. There is good linear relationship between number of pulses required for laser cleaning and the laser fluence in the effective zone. Surface roughness, porosity and moisture content have strong influence on the laser cleaning process. The effect becomes however negligible for high level of laser fluence. The number of pulses required for the laser cleaning is smaller for smooth, less porous and wet surfaces.

Key words Graffiti, Laser cleaning, Moisture content, Porosity, Surface roughness INTRODUCTION Over the last couple of decades graffiti has become a great problem particularly in the large agglomerations [3]. While it is relatively easy to remove graffiti from plastic, ceramic tile, metal, etc. it is very difficult to completely remove graffiti from cementitious composites and natural stone. The traditional cleaning techniques are of abrasive nature and may lead to the serious damage of the deeper layers of substrate. They may lead to the serious damage of the deeper layers of masonry. This unavoidable damage of surface is caused mainly by the direct contact of the cleaning agents with the surface and the difficulty in identifylng the border-line between dirt (graffiti) and the base material. Laser cleaning as a non-contact method appears to be a suitable method for removing surface contaminants from porous materials like masonry. Although graffiti removal by laser has been commercialised some years ago there are no records of the comprehensive analysis relating microstructure, geometrical characteristics of surface and moisture content of cementitious composites and their effect on laser cleaning process [ 1-61. Highly developed surfaces of cement based materials and the presence of water

Poologanathan SANJEEVAN. Agnieszka J. KLEMM, Piotr KLEMM

46

significantly complicate the mechanism of interaction between laser beam radiation and a base material. The paper focuses on presenting the results of investigation on the influence of surface roughness, porosity and moisture content of mortars on the removal of graffiti. The attempt is also made to present potential problems and side effects associated with laser application, such as pop-out, cracking, cratering and/or glazing. LASER CLEANING PROCESS

The laser irradiation is expected to be a self-limiting process. During the first stage of laser beam interaction photons are being absorbed by contaminants (dirt) which are then heated up and finally evaporate (ablation). Deeper layers of material can safely reflect further incidencing photons by diffusion. The mechanism of this process is based on the difference in the monochromatic reflection (absorption) of photons by dirt and background (masonry). Improper selection of laser parameters may however result in a severe damage of surface. The ablation process can have different forms - photo-thermal, photo-mechanical and photochemical or their combination. In the photo-thermal process (Fig.la), the graffiti is removed by vaporization. If the laser radiation is applied in long pulses or continuously, the heat has time to conduct away into the bulk material. When power increases, the material will melt, boil and vaporize [7]. When the shorter laser pulse is applied with high power density, photo-mechanical process occurs (Fig.1b). Here, the temperature rise occurs rapidly at the surface. This is accompanied by sudden thermal expansion of the heated material and subsequent generation of stresses and strains within the material. These ultrasonic waves can stress the material beyond its elasticity and material is ejected from the surface at high power densities laser. Sudden vaporization and ejection of surface material forms plasma (Fig 2). In the case of photochemical process, breaking of direct covalent bond of graffiti occurs (Fig. lc). The term photo-chemical is used if the photons are sufficiently energetic to break the covalent chemical bonds directly upon absorption without heating. The products of this chemical reaction have a larger volume than the original sample and this sudden volume increase expels the material. Figure 1

The effects of microstructural features of mortars on the laser cleaning process

47

Figure 2. Plasma patch Theoretically, deeper layers of substrate should not be affected during cleaning process. Since the temperature rise on the substrate occurs within a very short time (nanoseconds) and the thermal conductivity of cementitious materials is low, the sudden temperature rise should not affect deeper parts of cementitious material. The potential problems however may appear when water is present.

MATERIALS, MIXES AND CURING CONDITIONS The mix proportions of the samples are shown in Table 1, below. The study was aiming to establish existence of the relationships between laser parameters, microstructure of composites, their roughness and moisture content and the effectiveness of cleaning process. For these reasons all tested samples differ in their microstructure (plain mortar - 8 and air entrained mortar - 6), surface roughness (smooth samples A @a= 2.28-2.49prn)and rough D (Ra= 15.58-17.89,urn) and moisture content (WET and DRY). WET samples were fully saturated with water during 24 hours immersion. DRY samples were in equilibrium with surrounding air in a laboratory. Table 1 . The notations and compositions of samples.

Cement Sand

1 1

I 1

Water Conplasi AE38O

0.4 0.013

I

1 1

1 1

1 1

1

1

1

1

1

0.4

0.4

0.4

0.4

0.4

0.4

0.4

0.013

0.013

0.013

0.0

0.0

0.0

0.0

The mortar specimen was cured in the prismatic moulds for one day. Then these were cured in the laboratory conditions for 6 to 7 months under a polythene cover before being painted with a black paint. Roughness of samples has been measured by a stylus device. After the predetermined period of time samples have been subjected to the cleaning process. ~ 1 ps)0 was used. For this purpose ND: YAG laser (h=1.06pm, E=500 d,

Poologanathan SANJEEVAN, Agnieszka J. KLEMM, Piotr KLEMM

48

EFFECTS OF NUMBER OF PULSES ON THE LASER CLEANING Figures 3(a)-(h) present the relationships between the fluence and the number of pulses required for the cleaning. All mixes contained the same proportions of cement to sand (1:1) and the constant waterlcement ratio 0.4. Samples 8 refer to plain mortar and samples 6 refer to air entrained mortar (0.013%).

LBMvm

"1.

Figure 3. The relationship between the fluence and the number of pulses required for the cleaning. The above results can be summarised and presented in a form of graph (see Fig 6)

The effects of microstructural features of mortars on the laser cleaning process

49

FS

Fluence(Jcrn2)

Figure 4. Summary of the relationship between the numbers of pulses required for the laser cleaning and the laser fluence. The relationship between number of pulse required for laser cleaning and laser fluence is as follow: C = m*Fs+ Ns =

Ns when F>=Fs

-m*F+m*Fs+ Ns when FCFs Where N is a number of pulses, F is laser fluence, Fs is the fluence at the saturation point, Ns is the number of pulses at the saturation point, m is the gradient of the linear graph in the effective zone and C is the y intercept of the graph. The relationship between laser fluence and number of pulses required for the laser cleaning can be divided in two zones which are effective zone and ineffective zone (Fig 4). In the effective zone, the relationship between fluence and number of pulse required for the laser cleaning is linear. At the same time the number of pulses required for the laser cleaning at the ineffective zone is almost constant even when fluence increases. It reveals that at the high fluence, the thickness of removed paint per pulse is constant even when fluence increases. The existence of these two zones might be associated with two different processes taking place here. In the effective zone and ineffective zone laser cleaning process might be photothermal ablation and photo-mechanical ablation respectively. There is also a possibility of a change from photo-thermal to photo-mechanical process with the increase of laser fluence. The number of pulses required for the laser cleaning at the saturation for different samples is between 8 to 10 (Fig 3 (a)+) ), while the number of pulses in ineffective zone is almost constant for different the samples. It shows that at the high beam intensity the effect of sample characteristics (microstructure, moisture content and surface roughness) on the laser cleaning is insignificant. The fluence at the saturation point for different samples varies between 2.78 (Jcm-’) to 6.25 (Jcm-’) (Fig 3(a)-(h)).

THE EFFECTS OF SURFACE ROUGHNESS OF THE SUBSTRATE ON THE LASER CLEANING Figures 5 show that the number of pulses required for the laser cleaning is higher for the rough surface than a smooth surface (compare Fig 5 (a) and (b)). In the Figures 5 (c) and (d), the number of pulses required for the smooth and rough samples are almost the same.

Poologanathan SANJEEVAN, Agnieszka J . KLEMM, Piotr KLEMM

50

Nevertheless, number of pulses required for the laser cleaning from the rough surface is higher than from the smooth surface. When high laser intensity was applied, the effect of surface roughness is not that much prominent. However it is clear, when the laser intensity was low (Fig 5(a) and (b)). Fig 5(c) and (d) show the effect of surface roughness on the laser cleaning of highly porous sample. In this case, the number of pulses required for the cleaning is almost the same for both rough and smooth surfaces. It is because the highly porous samples absorb more laser irradiation (more than enough to remove the graffiti) than low porosity samples. Furthermore, The adhesion of paint and surface and the paint increases in the case of highly porous samples, thus effects of surface roughness is not clear in this case (Fig 5(c) and (d)).

0

5

10

15

20

15

30

Focal Imm (mm)

Figure 5. The effects of surface roughness of the substrate on the laser cleaning. When average roughness increases (more than wave length of laser), absorption increases, because of the multiple reflections. In the case of smooth surface, where average roughness is less than the wavelength of laser, absorptivity is fairly low. Further adhesion also increases with the roughness (&). However the laser cleaning process is influenced by adhesion of the graffiti and the absorption of laser radiation for the particular laser power. When adhesion increases, laser cleaning will be difficult. At the same time, laser cleaning will be easy when absorption increases for constant laser power. Both events happen at the same time in opposite directions. Therefore, the effect of surface roughness on the laser cleaning might be a combination of the above two events. However, the adhesion between paint and substrate has stronger effect than the absorption of laser radiation. (Fig 5(a) and (b)). In the case of crater formation and pop-outs, the surface roughness seems to be increased when compare to the initial roughness of cementitious material. At the same time, the surface roughness of the substrate decreases in glazing. Figure 6 shows the laser cleaned areas of different samples.

51

The effeets of microstructural features of mortars on the laser cleaningprocess

I

Flusnes

Number

11

OtpU(seS

-

12

9

I

5 44 Jcm.’

10

10

10

11

12

Figure 6. Laser cleaned areas Visually identified cleaned areas are surrounded by approximately 0.3 mm thick rings of burned paint. This can be attributed to lack of laser fluence in the outer part of the laser beam (TEMoo). Figure 10 illustrates distribution of the laser beam intensity with in the laser beam. Figure 7

T

Intensity

II

I, -.:

:-cleanedLaser cleaned area areaBurned paint

\\

,

Intensity. Intensitv. .. which is required for the removal of paint

,

radius

THE EFFECTS OF MICROSTRUCTURE OF THE SUBSTRATE ON THE LASER CLEANING Two sets of samples - plain and air entrained mortar have been tested. Total porosities of these have been determined with application of the Mercury Intrusion Porosimetry and were correspondingly 12.5% and 49.63%. At the same time Total Water Absorption values (water accessible porosity) were proved to be comparable 11.35% and 1 1.03%. This implies that the majority if not all of the air entrained pores are free from water and therefore able to facilitate water vapour dissipation produced as a result of sadden increase in temperature. No pop-outs should occur on the surface therefore. Table 4 shows the effects of microstructure of samples on laser cleaning. If the substrate is highly porous, the laser cleaning requires higher fluence than the less porous substrate. Figures 8(a)-(d) show the relationship between number of pulses required for the laser cleaning and the focal length of the laser beam. When the high laser fluence is applied, the effect of porosity is not that much prominent. However, it becomes clear for the low intensity of laser radiation (Fig 8(a) and 8(b)) and for the smooth samples. Furthermore, it appears that the effects of surface roughness suppress the effects of porosity.

52

Poologanathan SANJEEVAN, Agnieszka .I. KLEMM, Piotr KLEMM

b) -1

I

a.

Figure 8. The effects of porosity of the sample on the laser cleaning.

THE EFFECTS OF MOISTURE CONTENTS OF THE SUBSTRATE ON THE LASER CLEANING The effects of moisture content of the substrate are shown in Fig 9(a) to Fig 9(d). The number of pulses required for the laser cleaning is smaller for the wet sample than dry sample. It shows that wetness of the sample positively affects the laser cleaning process.

Figure 9. The effect of moisture content of the sample on the laser cleaning

The effects of microstructuralfeatures of mortars on the laser cleaning process

53

When high laser fluence was applied, the effect of moisture content is not that much prominent. However it is clear, when the laser intensity was low (Fig 9(a) and (b)). In case of smooth sample, the effects of moisture content are clear. It shows that, effect of surface roughness have more influence than moisture content. Thus, effects of moisture content on the laser cleaning are not that much clear in the case of rough sample.

SUMMARY Highly developed surfaces of cement-based materials significantly complicate the mechanism of interaction between laser beam radiation and a base material. The presence of water in pore system adds even further complication to the process. A clear need is therefore perceived to identify relationships between laser parameters and material characteristics on both macro and micro scale. Based on experimental investigation up to date some preliminary observation and conclusions can be formulated. 0

The relationship between laser fluence and number of pulses required for the laser cleaning can be divided in two zones which are effective zone and ineffective zone. There is good linear relationship between number of pulses required for laser cleaning and the laser fluence in the effective zone. Until certain beam intensity, number of pulses required for cleaning is constant. Ns when D = F s

N = { -m*F+ m*Fs+ Ns when F
54

PoologanathanSANJEEVAN. Agnieszka J. KLEMM, Piotr KLEMM

REFERENCES 1. A.Coste1, I. Garcia-Moreno, C.Gomez, O.Caballero, R.Sastre, Cleaning graffiti on urban building by use of second and third harmonic wavelength of a NdYAG laser :a comparative study, Applied surface science, ~01202,2003,pp 86-99. 2. Eric C. Crivella, Joyce Freiwald, David A. Freiwald, Laser Surface Cleaning, F2 Associates Inc, New Mexico. 3. I Matsui, K Nagai, N Yuasa,Y Ishigami, Removing graffiti on concrete surface by laser, Nihon University, Taisei corporation Japan, Proceeding of the international conference held at the University of Dundee, Scotland, UK, 2002. 4. Katherine Liu and Elsa Garmire, Paint removal using lasers, applied optics, Optical society of America, Vol34, No.21, 1995. 5. Kazimiea Rozniakowski, Piotr Klemm, Agnieszka J. Klemm, Some experimental result of laser beam interaction with surface layer of brick, Building and Environment, Vol36,2001, pp 485-491. 6. Lin.LI, W.M.Steen, P.J.Modem and J.T.Spencer, Laser removal of surface embedded contaminations odin building structures, laser Materials Processing and Machining, SPIE, Vol2246,1994. 7. Martin Cooper, Laser cleaning in conservation, an introduction, UK, 1998.

Proc. Int. Symp. “Brittle Matrix Composites 8“ A.M. Brandt, r C . Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

FLY ASH BASED GEOCEMENTS: GENESIS OF MICROSTRUCTURE AND PROPERTIES AT HYDRATION-DEHYJlRATION PROCESS Pave1 KRIVENKO*, Georgiy KOVALCHUK*, Angel PALOMO**, and Ana FERNANDEZ-JIMENEZ** *V.D. Glukhovskiy State Scientific Research Institute on Binders and Materials (Kyiv, Ukraine), e-mail: [email protected] **Eduardo Torroja Institute CSIC (Madrid, Spain), e-mail: [email protected] ABSTRACT Phase composition of fly ash based geocements after treatment at a low temperature is represented by zeolitelike products in amorphous or crystalline state which are capable to re-crystallize into stable feldspar-like aluminosilicates without destroying the underlying skeleton. It may afford a pathway worth exploring for the production of non-fired quasi-ceramic materials.

Keywords Alkaline cements, Alkaline activation, Curing conditions, Feldspars, Fly ash, Heat-resistant materials, Promoters of zeolite synthesis, Zeolites INTRODUCTION Alkaline cements are becoming an interesting set of products with high potential to help the cement industry to face the problem of limitation of CO2 emissions. These types of cements are still not very well known although they have been developed since 1957. They have a set of excellent service properties: high early mechanical strength development, fire and acid resistance, etc. Some experiences at industrial level have confirmed the behaviour of these cements as good building materials [ 1-51. Fly ash based Geocements, also called ‘Alkali activated fly ash’, a material produced by mixing this industrial by-product with an alkaline solution and then curing the mix, are suggested to be one of the most promising alkaline cements. The term “Geocement” reflects peculiarities of the structure formation process of such materials being similar in a general way to that of the geological transformations of aluminosilicate materials in the nature: the both processes consist in transformation of initial aluminosilicate glass to zeolite-like phases at low temperatures (below 2OO0C), and then to anhydrous aluminosilicate minerals once being conditioned at high temperatures [ 6 ] .In accordance to that, fly ash based Geocements are belonged to the first class of alkaline cements [3], in which the calcium compounds are present in small quantities, thus playing a supplementary role. Unique microstructure of fly ash based geocements predetermines a set of useful properties such as high mechanical strength (over 100 MPa after 8 hours of curing at 9SoC), good adhesion, ability to bond hazardous and radioactive wastes, acid resistance etc. Since zeolitelike structures and characteristics are extremely varied, close regulation of the structural process is the key to obtaining high-performance, durable materials. Some time ago the influence of

56

Pave1 KRIVENKO, Georgiy KOVALCHUK, Angel PALOMO, h a FERNANDEZ-JIMhNEZ

NazO/AlzO3 and SiOz/A1~03molar ratios of the starting system on the properties of the final product was already investigated by different authors [7, 81. Nowadays it is well known that the mechanical strength development of the material improves by increasing the starting SiOz/A1~03ratio [8, 91. It is due to the fact that the number of strong Si-0-Si bonds increases in the final product. It is also known that the amount of Alz03 in the reactant system plays an important role in the kinetics control of the process [9]. One of the most perspective areas of application of fly ash based geocements is development of high-temperature materials. As new materials have been mainstreamed into the construction industry, new cementitious materials are expected to feature a heat- and fire resistance. However, traditional cements have a lot of serious short-comings restricting their application in modem high-temperature materials. Thus, modified Portland cements have low residual strength after firing; alumina cements are expensive and have significant variations of mechanical strength; application of pure water glass is associated with serious technological problems; phosphate binders harden quickly and are toxic and expensive [lo]. As it was shown before [6, 10, 111, some properly designed geocements demonstrate excellent thermomechanical properties conditioned by the ability of some zeolites to re-crystallize into stable feldspar-like aluminosilicates without destroying the underlying skeleton. It may afford a pathway worth exploring for the production of non-fired quasi-ceramic materials. There would appear to be a need for such products, in light of the growing demand for high temperature-resistant masonries and fireproof structures ensuing fi-om the catastrophic consequences of recent fires in tunnels, high-rises and so on. However, some peculiarities of phase transformation in the whole temperature range are still studied insufficiently The investigation described in this paper was aimed to achieve a better understanding of the genesis of microstructure of fly ash based geocements in the N~zO-A~ZO~-S~OZ-HZO system within hydratioddehydration process during heating. In particular, role of mix composition and promising additives called “zeolite precursors” [ 121 were studied within the general process of alkaline activation of fly ashes and their heating up to 1200°C. This study is based on preliminary investigations led using the same methodology and raw materials [7, 131. EXPERIMENTAL A typical class F fly ash from a Spanish power plant was used (chemical composition is shown in Table l), with 78.86% of particles smaller than 45 pm, and Blaine specific surface of 202 mz/kg. Water glass (soluble sodium silicate) with a silicate modulus of 3.35 and a density of 1350 kg/m3 (8.2% NazO, 27% SiOz and 64,8% HzO) supplied by Panreac S.A., was used as the main alkaline activator. In order to obtain the necessary NazO/Alz03 ratio of the alkaline solution, waterglass was adjusted by additions of NaOH pellets of 98% purity.

ash

A1203 Fez03 CaO MgO SO3 NazO KzO LR.** L.O.I. Total SiOz 54‘42 2.19 99.44 26.42 7.01 3.21 1.79 0.01 0.59 3.02 0.78 (45.05)*

** I.R. - insoluble residue, L.O.I. - loss on ignition. Calculation of initial mixtures compositions was carried out according to a set of prefixed total ratios established between main oxides (Table 2). Watedsolid ratio (W/S) was always kept constant = 0.21. There were 3 representative compositions (# 1, 2 and 8) selected based upon the preliminary study [7].

Fly ash based geocements: genesis of microstructure and properties at hydration-dehydrationprocess

Sample Identification 1 1c 2 8

Molar ratio of mixture NazO/A1203 Si02/Al203 1.o 3.5 3.5 1.o 4.0 1.o 4.0 0.5

57

Sodium perchlorate Water/Solid ratio (W/S) additive, % 0.2 1 0 2.5 0.21 0. 0.21 0 0.21

Additionally, in order to detect the role of zeolite promoters, sodium perchlorate additive (NaClOd.H*O, 2.5% by dry mass) was introduced into the first composition (# 1C). This additive was preliminary founded to improve crystallization of a hydroxysodalite-type phase. "Hot moulding" procedure was used to prevent quick setting [6, 111. Moulds once filled with pastes were covered with a plastic film to prevent water evaporation. Then they were kept into the oven at the temperature of 95°C for 8 hours. This type of curing was successfully used in a previous part of this investigation [7, 131. In fact it can be easily applied in real industrial conditions [6, 111. Mechanical strength of paste prisms (sized 1 ~ 1 x 6cm) was determined 24 hours after curing process. Thermo-mechanical properties (absolute and residual compressive strength, thermal shrinkage) were tested after firing up to 105...1200°C with a heating rate of 15O"C/hour and with a subsequent isothermal conditioning during 4 hours. Mineral composition and microstructural characteristics of materials were studied by using XRD and DTA. X-ray diffraction patterns of powered samples were recorded in a Philips PW 1730 diffractometer with CuKa radiation. The tests were run in the 20 range of 5"-60" with a scanning rate of 2"/min, deliverance slit of lo, anti-scatter slit of 1" and receiving slit of 0.01 mm. DTA patterns were recorded in a MOM derivatograph with an alumina used as an internal standard. The samples were heated up to 1000°C with a heating rate of 10"C/min.

RESULTS AND DISCUSSION Mechanical strength evolution Figure 1 shows an evolution of mechanical strength and thermal shrinkage depending on curing temperature. Residual strength and thermal shrinkage values (in percent) were related to that of the samples cured at 105°C. It should be mentioned that the compositions # 1, 1C and especially 2 began funding since lOOO"C, whereas less alkaline composition # 8 remained workable almost up to 1200°C. When compared to previous investigations, an increased W/C ratio used in this study (0.2 1 vs 0.18) in order to introduce an additive. It resulted in decreased strength values after basic curing (42-85 MPa vs 64-102 MPa [7, 131). Additional dry curing at 105°C made for increasing strength for all compositions excepting the #2 which dropped in strength. During subsequent heating, the compositions #1, 1C and 2 demonstrated decreasing in strength until 400"C, then a slight increasing in strength took place in the range of 400.. .800"C, followed by a new strength drop after firing at 1000°C. Less alkaline composition #8, whereas demonstrated the same behaviour, had a displacement in a temperature of strength changes of about 400°C: thus, it was slightly decreasing in strength until 800"C, after that a little strengthening took place. It is to be noted that the residual strength of fly ash based geocements after firing at 800°C was within 44123% which is much higher than that of the OPC (-30%). The additive of sodium perchlorate initially gave worse strength values; however, it gave better residual strength results due to higher variations depending on the temperature of heating.

58

Pave1 KRIVENKO, Georgiy KOVALCHUK, Angel PALOMO, h a FERNANDEZ-JIMhNEZ

h

120

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o Cunng temperature, C 180 T---

a)

I ;

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200 4M) 600 curing temperature. c

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lo00

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Curing temperature. C

c) Fig. 1. Absolute (a) and residual compressivestrength (b) and thermal shrinkage (c) vs mix composition and curing temperature Thermal shrinkage of all the compositions slightly increased until 400.. .6OO0C.Further heating affected different behaviour depending on mix composition. Thus, the compositions #2 and 8 demonstrated some expansion instead of shrinkage caused by an increased SiOz/A1203 ratio. The lowest shrinkage variations within 105...8OO0C (it is essential for high-temperature materials) were fixed for the composition #lC (modified with the sodium perchlorate), however, the shrinkage of this composition increased sharply after 1000°C. The same composition demonstrated the lowest thermal cracking, too. Although the above presented data reflects structural changes in a geocement paste during heating, they do not clearly describe the behaviour of a binder in a real concrete exposed to high temperatures because an important interaction between a matrix and an aggregate takes place in real systems affecting an amount of a fusion phase appeared and a thermal shrinkage. Thus, further investigations are to be held using a heat-resistant filler [6, 10, 113.

X-Ray Diffraction Fig. 2 presents the XRD patterns of the systems studied at different temperatures. The data of microstructural changes depending on the composition and temperature are also assumed in the table 3.

Fly ash based geocements: genesis of microstructure and properties at hydration-dehydrationprocess

59

0

fly a s h I

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C)

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Fig. 2. XRD patterns of the compositions # 1 (a), 1C (b), 2 (c) and 8 (d). Fly ash crystalline impurities: Q - quartz, M - mullite, G - gematite. New formations: * and # - hydroxysodalite or sodalite, n - nepheline, a - albite, m - magnetite.

Pave1 k N V E N K 0 , Georgiy K O V&CHUK* Angel PALOMO, Ana FERNANDEZ-JIMENEZ

60

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--- completely sintered phase---

Quartz, Nepheline

Nepheline, Albite, Quartz Albite (max), Quartz

After low-temperature curing, the low-silica composition #1 had an amorphous structure with some hydroxysodalite-type crystalline inclusions. The amount of this phase grew with increasing temperature and reached maximum at 20OOC, and then began decreasing. At 6OO0C, only weak peaks of hydroxysodalite were fixed, whereas an anhydrous feldspathoid nepheline (Na2O.Al20~2Si02)began synthesizing at this temperature. The amount of this phase grew intensively with increasing temperature, and reached a maximum at 1000°C. Formation of acrystoballite, a meta-stable crystalline silica, was also fixed at 600°C. Although sodium perclorate-modified compositions demonstrated generally similar genesis of a microstructure (especially regarding to the synthesis of nepheline and a-crystoballite), there were important differences. First of all, an introduction of these zeolite promoter resulted in 3.. .4 times higher amount of the hydroxysodalite-type phase synthesized after low-temperature curing (compare this phase's peaks with the reference quartz's peak from the composition of fly ash for both mixtures, fig. 2, a, b). However, there are some doubts regarding the nature of this phase: since the promoter contained the ions of Cl-, it is thought that a pure sodalite might be crystallized instead of hydroxysodalite traditionally appeared in these systems [6-111. The amount of this phase tended to growth with increasing temperature (up to 600.. .800°C), but a significant structure changes took place between 400 and 600°C: all the main peaks moved right for about 0.004-0.008 nm (fig. 2a and table 4) becoming equal to that of the phase fixed in the composition #1 (the one without the promoter). It is thought that this structural change corresponds to the transformation sodalite-thydroxysodalite with a C1 released. The above mentioned decreasing of interplain distance in the crystalline structure is thus probably conditioned by a difference between ionic radii of Cl- and OH- (0.181 nm vs 0.132 nm, correspondingly). In any case, in contrast to the pure composition 1, the hydroxysodalite-type phase formed in the composition 1C remained stable up to 800.. ,100O"C.

Main XRD peaks

1C (105 ...400°C) 0.639 0.370 0.262 0.214

1C (600.. .1000°C) 0.63 1 0.364 0.258 0.210

1 (average) 0.634 0.366 0.258 0.21 1

Fly ash based geocements: genesis of microstructure and properties at hydration-dehydrationprocess

61

The composition #2, with a waterglass used as main alkaline constituent, had an amorphous structure described earlier [6-9, 131, and remained that type of structure until 600°C. Only the weak peaks of hydroxysodalite were fixed at 400°C. Since 800"C, the structure began changing into highly-crystalline one based on nepheline which reached a maximum content at 1000°C. The composition #8, which differed from the #2 only in the lower alkalinity, demonstrated the amorphous structure within 100...600°C. The difference began revealing from 800°C: nepheline crystallized less intensively at 800 and 1000°C. However, formation of a feldspar albite (Na20.AI203.6Si02) was fixed at lOOO"C, with the same intensity than that of the nepheline. The composition #8 was the only one remained stable until 1200"C, so that the XRD study showed in intensive synthesis of albite at this temperature. The synthesis of albite, a less Al-rich phase than nepheline, was because of lower rate of alkaline activation of fly ash which is the only source of alumina (since silica was initially present from waterglass in the both compositions 2 and 8). Quartz impurity from the composition of fly ash tended to dissolve slowly within 100...600"C, and disappeared at 800°C in all the high-alkaline compositions (#l, 1C and 2), thus giving an additional source of silica for high-temperature transformations. However, in less alkaline composition #8, it remained unreactive until lOOO"C, and disappeared only at 1200°C. The mullite impurity, in turn,remained unreactive in all the cases. DTA studies Thermal analysis of fly ash based geocements is complicated due to a lot of the effects related to the initial fly ash. All the compositions demonstrated similar behaviour in general (Fig. 3): endothermic peak at 160...2OO0C (dehydration), narrow exothermic one at 355.. .370"C (recrystallization of iron oxides from fly ash), wide exothermic one within 450.. .700°C (burning of coal residuums). However, there were some details depending on the composition. First of all, heating of the samples during DTA procedure were going on until a liquid sintered phase was appeared. The DTA results lay down in accordance with the preliminary studies: the highest thermal stability was reached at the lowest alkali content (composition #8). More alkaline compositions #1 and 2 demonstrated approx. equal sintering temperature, whereas modification of the first one with a promoter resulted in increased thermal stability. Dehydration of the amorphous matrix took place within the similar temperature range for all the compositions (endothermic peaks of 170.. .200"C at DTA curves and acceleration of a weight loss of 150...170°C at DTG curves). However, more crystalline compositions 1 and 1C (which contained hydroxysodalite) showed another deep peak on DTG curves at 325°C which might be related to the dehydration of the hydroxysodalite phase. The composition #I had the above mentioned peak more intensive, whereas the modified one (IC) had an additional endothermic peak on the DTG curve at 680°C. It might c o n f m the hypothesis mentioned above: in the sodium perchlorate-modified composition #1C, some amount of sodalite appears instead of the hydroxysodalite synthesized in the pure composition #l . During heating, the sodalite releases C1 and transforms into hydroxysodalite. Since the DTA method is dynamic (contrast to the XRD method), the temperature of this transformation measured by DTA (680°C) was higher than that measured by XRD (within 400.. .6OO0C). The compositions #1C and 8 were the more thermally stable, it allowed fixing exothermic peaks at 820 and 910°C correspondingly, which referred to the formation of anhydrous aluminosilicates (probably nepheline). Weight losses (given near TG curves on the fig. 3) of fly ash based geocements consist of bounded water and coal residuums from the composition of fly ash. Since the content of fly ash and, accordingly, coal residuums were approx. equal for all the compositions, a difference in weight losses may give some information on the amount of bounded water which is indirect index of the zeolite-like products formed (i.e. N-A-S-H gel and crystalline zeolites).

62

Pave1 KRIVENKO, Georgiy KOVALCHUK, Angel PALOMO, Ana FERhCANDEZ-JIMENEZ

1c 760

W

8

780

I

-2,Cm=12.0%

I

20 ,ylOO200 300 400 500 600 700 A

800

900 If

Fig. 3. DTA patterns of the studied fly ash based geocements Depending on the composition, weight losses were increased in the direction of: #8 (10.3%) < #2 (12.0%) < #1 (14.0%). Taking into account an initial alkalinity used in these compositions, it is to be concluded that the higher initial alkalinity used, the higher amount of hydration products formed and, accordingly, the higher rate of alkaline transformation of fly ash is achieved. Compositions #1 and 1C had approx. equal weight loss. It means that despite different quantity of crystalline products was formed, the level of alkaline activation of fly ash was equal, it’s to say, when using the zeolite promoter, the only difference consisted in higher crystallinity of the hydration products, whereas the overall quantity of hydration products was equal. Thus, the role of zeolite promoter is to support the zeolite crystallization in the later studies of structure formation.

Fly ash based geocements: genesis of microstructure and properties at hydration-dehydrationprocess

63

EXPERIENCE OF INDUSTRIAL APPLICATION OF HEAT-RESISTANT FLY ASH BASED GEOCEMENTS Once applied to the technology of non-autoclaved cellular concretes, fly ash based geocements allowed obtaining fly ash based heat-resistant aerated concretes with a maximal use temperature of 800°C using fly ash- based geocements and an aluminium powder as a gas producing agent [ 101. These materials were intended for thermal insulation of hightemperature equipment such as furnaces, conduits, chimneys, boilers, fireplaces, etc. The development of process parameters and mix proportions allowed to produce the heat-resistant aerated concretes of density varying between 300-1200 kg/m3, compressive strength up to 16 MPa, thermal resistance up to 34 cycles in the air, residual strength after firng at 800°C of 75.. .537 % and thermal shrinkage of 0.94.. .1.97 %. The manufacture of this material does not require high temperatures, and thus its manufacture is cheaper, more energy saving and environmentally friendly in comparison with the manufacture of conventional lightweight refractory materials. In order to launch a commercial scale application of this material as heat insulation of a furnace lining at the glassware plant (Kiev, Ukraine), the geocement-based heat-resistant adhesive was developed. Using this adhesive, the materials may be glued at ambient or increased temperatures and the adhesive maintains its high adhesion after heating up to 1000°C and cooling. Thus, lightweight (600 kg/m3) heat-resistant tiles 4 0 x 4 0 ~ 4cm in size made from the elaborated aerated concrete were glued directly to a hot furnace surface (up to 700°C) without interruption of the furnace operation. After six months in service, no destruction of the heat resistant aerated concrete products was visualized [6, 10, 1I].

CONCLUSIONS After firing, the zeolite-like products in geocement structures re-crystallize into anhydrous aluminosilicates.Genesis of microstructure may be represented as follows: 1. Formation of zeolite-type structures (crystalline or semi-crystalline). 2. Intensification of its synthesis (up to 400.. .600"C). 3. Amorphization or no (depending on the composition). 4. Crystallization of stable anhydrous phases (nepheline, albite or cristobalite, depending on mix composition and curing temperature), since 600.. .800"C. 5. Melting (sintering) leading to destruction. Fly ash based geocements generally increase in strength after drymg at 150"C, but loose in strength after dehydratiodamorphization up to the point of formation of new crystalline structure based on anhydrous alkaline aluminosilicates (nephelien or albite, 600.. .8OO"C depending on the composition). Formation of above-mentioned structure results in slight increasing in strength (until destruction by melting begins), but also results in increased volume changes due to appearance of liquid phase. Depending on the composition, the lesser alkali content, the higher thermo-mechanical properties. Introduction of zeolite precursor makes for highly crystalline structure formed which makes the composition more resistant to the loads during heating resulting in higher residual strength and lower thermal shrinkage (before 800"C, when a sintering process begins). However, it is thought that a pure sodalite forms instead of hydroxysodalite wihin the temperature range of up to 400°C. After 600"C, it seems to loose C1 and transform into conventional hydroxysodalite without destruction of a framework. Thus, a sodium perchlorate-modified composition demonstrated the highest residual strength and the lowest thermal shrinkage at 800°C.

64

Pave1 KRWENKO, Georgiy KOVALCHUK, Angel PALOMO, Ana FERNANDEZ-JIMENEZ

The residual strength of fly ash based geocements after firing at 800°C was within 44-123% (or 37.. .52 MPa) which is much higher than that of the OPC (-30%). Maximal heat resistance (1200°C) was fixed for the less alkaline composition. Small-scale industrial application of heat resistant aerated concrete based on these materials in the furnace lining at the glassware plant codirmed economical effectiveness of application of fly ash based Geocements as hightemperature composite materials. ACKNOWLEDGEMENTS The authors wish to thank the NATO Scientific Committee for the postdoctoral grant, under its Science Fellowships Programme (2003), associated with this study. REFERENCES 1. Glukhovskiy, V.D., Soil Silicates. Gosstroy publsh, Kiev, 1959. (in Russian) 2. Krivenko, P.V., Alkaline Cements. In: Proc. First Intern. Conf. "Alkaline Cements and Concretes", Kyiv (Ukraine), 1994, pp. 11-129. 3. Krivenko, P.V., Alkaline Cements: Terminology, Classification, Aspects of Durability. In: Proceed. lothICCC, Gbteborg (Sweden), 1997, pp. 4ivO464ivO50. 4. Rostovskaya, G., Illyin, V., and Brodko, O., The Investigation of Service Properties of the Slag Alkaline Concretes. In: Proceed. Intern. Symposium "Non-traditional Cement&Concrete", Brno (Czech Republic), 2002, pp. 5 10-523. 5. Shi, C., Roy, D., Krivenko, P., Alkali-Activated Cements and Concretes. Taylor & Francis publish, 2006. 6. Krivenko, P.V., Kovalchuk, G.Yu., Directed Synthesis of Alkaline Aluminosilicate Minerals in a Geocement Matrix. Special issue of J.Mat.Sci., 2006 (in publish). 7. Kovalchuk, G., Palomo, A., Fernhdez-JimBnez, A., Alkali Activated Fly Ash. Mechanical strength development as a function of the composition of the starting system. Mater. Construcc. (submitted for publication). 8. Fernindez-Jimdnez, A., Palomo, A., Composition and Microstructure of alkali activated fly ash mortars. Effect of the activator. Cem. Concr. Res. 35, pp. 1984-1992, (2005). 9. Fernhdez-JimBnez, A., Palomo, A., Sobrados, I., S a m , J., The role played by the reactive alumina content in the alkaline activation of fly ashes. Microporous and Mesoporous Materials, 9 1, pp. 111- 119, (2006) 10. Krivenko, P.V., Kovalchuk, G.Yu., Kovalchuk, O.Yu., Heat-Resistant Cellular Concretes Based on Alkaline Cements. In: Use of Foamed Concrete in Construction: Proc. Intern. Conf. "Global Construction: Ultimate Concrete Opportunities", Dundee (UK), 2005, pp. 97-104. 11. Krivenko, P.V., Kovalchuk, G.Yu., Heat-resistant fly ash based Geocements. In: Proc. "Geopolymer-2002" Conference, Melbourne (Australia), 2002. 12. Kumar, A., Bhaumik, A., Ahedi, R.K., Ganapathy, S., Promoter-Induced Enhancement of the Crystallization Rate of Zeolites and Related Molecular Sieves. Nature, Vol. 381, 1996, pp. 298-300. 13. Kovalchuk, G., Palomo, A., Fernhndez-Jimenez, A., Alkali Activated Fly Ash. Part 11: Mechanical and Microstructural Development as a Function of the Thermal Curing Conditions. Fuel, 2006 (in publish).

Pmc. Int. Symp. “Brittle Matrix Composites 8“ A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

INFLUENCE OF WATER-SOLUBLE POLYMERS ON THE MICROSTRUCTURE OF CEMENT MORTARS Elke KNAPEN, Dionys VAN GEMERT K.U.Leuven, Departement Burgerlijke Bouwkunde Kasteelpark Arenberg 40,300 1 Heverlee, Belgium, email: [email protected]

ABSTRACT The influence of small amounts of water-soluble polymers (polyvinyl alcohol-acetate, methylcellulose and hydroxyethyl cellulose) on the microstructurebuilding and on the cement hydration reactions is studied. Thermal analyses, canied out at several time intervals, show a retarded hydration during the first 24 hours of hydration for the cement pastes modified with hydroxyethyl cellulose. After 2 days, this delay is eliminated. Although evaporation of water is prevented, polymer modified mortars show a slightly higher amount of bound water than unmodified mortars. Nevertheless, the Ca(OH)2content is decreased by polymer modification. SEM investigation is used to characterize the morphology of the hydrate crystals. A large crystal growth is noticed at the air void surfaces of the modified mortars where the presence of water-soluble polymers is expected because of their strong affinity to the gas-water phase. Water-soluble polymer molecules are supplied on a molecular scale, improving the approach of the relatively large cement grains by the polymers. Therefore, film formation on the hydrate crystals proceeds more easily and uniformly with dissolved polymers than with polymer dispersions. In mortars modified with methylcellulose, undistorted layers of Ca(OH)2 crystals are formed. Polymers are found to form additional bridges between those Ca(OH)2crystals.

Keywords Polymer-modified cement mortars, water-soluble polymers, microstructure. INTRODUCTION Usually, polymer-modified cement concrete or mortar is prepared by mixing polymer dispersions or redispersible polymer powders with the ffesh mixture. The surface active agents, added to allow emulsification and stabilization of the dispersion during storage, hinder the cement hydration and the polymer film formation, proceeding from the dispersion [l]. When water-soluble polymers are added instead of polymer dispersions, polymer molecules are supplied on a molecular scale, which eases the polymer film formation and the approximation of the cement grains. In the absence of surface active agents, water-soluble polymers tend to require a lower proportion in order to be comparably effective as polymer dispersions [2]. The addition of very small amounts of water-soluble polymers results in an improvement of the durability and the adhesion strength of the cementitious materials, which makes them appropriate as repair materials [2-41. The effect of the addition of water-soluble polymers is twofold: the cement hydration reactions may be influenced and polymer film formation may take place. In previous work, mechanical strength tests showed a strong indication of polymer film formation in mortars modified with water-soluble polymers, even at low polymer-cement ratios (p/c=l%) [5]. In this paper, the influence of water-soluble polymers on the microstructure of cement mortars is

Elke KNAPEN, Dionys Van GEMERT

66

discussed. Cement hydration and polymer film formation are studied by means of SEM investigation and thermal analysis. SAMPLE PREPARATION

An ordinary Portland cement (CEM I 52.5 N) and CEN-Standardsand DIN EN 196-1 are used. Different types of polymers are added to the fresh mortar mixtures: a polyvinyl alcohol (PVAA) which is a 87-89% hydrolyzed polyvinyl acetate and two cellulose derivatives, methylcellulose (MC) or hydroxyethyl cellulose (HEC). For the SEM investigation, mortar beams (40*40*160mm) are prepared with a polymercement ratio (PIC)of 1%, a water-cement ratio (w/c) of 0.45 and a sand-cement (s/c) ratio of 3. The polymer powders are first dissolved in the mixing water before adding to the sand and cement in the mixer. The mortar beams are covered for two days before demoulding. Standard curing implies storage, after demoulding, in a moist room for 5 days (20"C, 85-90% RH), followed by a dry curing of 21 days (20"C, 63% RH). For the thermal analyses, cement pastes with a polymer-cement ratio of 1% and a watercement ratio of 0.45 are prepared. The polymer powders are first dissolved in the mixing water before adding to the cement in the mixer. The pastes are stored in closed recipients until the moment of testing (24h, 2, 7, 28 and 90 days). Prior to thermal analyses, the water which has not participated yet in the hydration reactions should be removed. After being crushed to fine powder, the cement pastes are vacuum dried for 2 hours. The samples are introduced without prefreezing in an Alpha 1-2 LD Martin Christ type freeze-dryer. Round bottom flasks containing the samples are connected to a vacuum device consisting of a vacuum chamber that allows applying a vacuum of 2.5~1O-~rnbar. Extracted water is continuously collected in an ice condenser at a temperature of -62°C. THERMAL ANALYSIS

The thermal analyses of the cement pastes are carried out on a Netzsch STA 409 PC, a simultaneous Thermogravimetry (TGA) and Differential Scanning Calorimetry (DSC) system. The samples are heated from room temperature to 1000°C with a heating rate of 10"C/min in a N2-atmosphere (60ml/min) after 24h, 2, 7, 28 and 90 days of hydration. The TGA-signal is used to calculate the weight loss during heating and to estimate the Ca(OH)2 content. Amount of bound water Assuming that all water which has not yet participated in the hydration reactions is removed by vacuum drying [6], the total weight loss on ignition is a measure of the amount of bound water during hydration and therefore of the degree of hydration. The amount of bound water is calculated according to formula (1). This weight loss is corrected for the weight loss due to the decomposition of the polymers (Lpol) and for the loss on ignition of the cement itself (Lcem).The results are shown in Figure 1.

After 24 hours, the hydration of the HEC-modified cement paste is seriously retarded. Nevertheless, after 2 days of hydration, this delay has almost completely disappeared. After

Infruence of water-soluble polymers on the microstructure of cement mortars

67

90 days, the pastes modified with HEC show the highest amount of bound water. Although the evaporation of water is prevented in the closed bottles during hydration, the polymer modified cement pastes show a higher amount of bound water than the reference mortar. This is probably due to a better dispersion of the cement particles in the mixing water. Because water-soluble polymers show a high water retaining capacity, the amount of bound water after standard curing (2 days wet, 5 days at 93% RH, 21 days at 63% RH) will be studied in future research projects.

35%

C

2

m 5%-

0

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40

30

70

60

80

90

100

Hydration time [days] , ,"P, ,

~~

~

~~

~

, :/. -=-MCl% +HECI% +REF; Figure 1: Bound water ["! of ignited weight] for unmodified cement pastes (REF) and pastes modified with 1% PVAA, 1% MC and 1% HEC after 24h, 2,7,28 and 90 days of hydration

Ca(OH)* content The weight loss in the temperature interval around 450°C is mainly due to the dehydration of Ca(OH)2. The Ca(0H)z content is calculated according to formula (2).

WLCa(OH)2 is the weight loss during the dehydration of Ca(OH)2 as percentage of the ignited weight and determined according to the method described by Taylor [7]. This takes into account the continuous loss due to the gradual decomposition of the C-S-H phase and the hydrated aluminate phase. MWCa(OH)z and MWHI0 represent the molecular weight of Ca(OH)2 and water. The Ca(0H)z-content as weight percentage of the ignited weight is presented in Figure 2. Similar to the amount of bound water, the formation of Ca(OH)2 is delayed for the pastes modified with 1% of HEC. Contrary to the amount of bound water, the amount of Ca(OH)Z, being the weak phase in mortars, is slightly lower for polymer modified pastes than for the unmodified paste. SEM investigation is carried out to study the morphology of the Ca(OH)2 crystals in the presence of polymers.

68

Eke mAPEN, Dionys Van GEMERT

18% 1

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/p+pVAA1%

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Figure 2: Ca(OH)2 content [% of ignited weight] for unmodified pastes (REF) and pastes modified with 1% PVAA, 1% MC and 1% HEC after 24h, 2,7,28,90 days of hydration

SEM INVESTIGATION Fresh fracture surfaces of mortar beams are investigated using a Philips XL 30 FEG and a JEOL JSM-6400 Scanning Electron Microscope (SEM) in the SEI mode. Samples are coated by evaporation with gold.

Microstructure at air voids Special attention was paid to the air void surfaces in the mortar specimens. At these surfaces, the presence of water-soluble polymers is expected because of their strong affinity to the gaswater phase. The polymers serve as surface-active agents that are initially dissolved in the mixing water. During mechanical mixing, they are attached to the air void interface and start to stabilize the entrained air voids in the fresh mixture. So, an enrichment of polymer at the interface between air void and wet cement paste may be detected [8], depending on the surface activity of the polymer. Nevertheless, visual detection of the polymer films is difficult. The polymers are added in very low amounts (only 1% of cement weight), there is a possible integration of the polymers in the cement matrix at a submicrometer scale and the polymer films are highly sensitive to damage by the electron beam. The air void surfaces of unmodified mortar beams have a smooth surface at the micrometer scale without any well-developed morphology of Ca(0H)z crystals [9]. This is illustrated in Figure 3. The microstructure of cement hydrates formed at the air void surfaces is strongly influenced by polymer modification, as shown in Figure . PVAA-modified and HEC-modified mortars show a similar microstructure at the air void surfaces. Needle-like CS-H covers the wall, disturbing the smooth surface that was found in unmodified mortars. The presence of polymer particles or polymer films is expected at the air void surfaces, although polymer film formation is not detected. Water molecules are bounded to the polymer particles until sufficient forces are exerted, e.g., by cement hydration. In this aqueous environment, crystal formation is possible.

Influence of water-soluble polymers on the microstructure of cement mortars

69

Figure 3: Air void (left) in unmodified mortar without any large crystals at the surface (right)

Figure 4:Needle-like C-S-H at the air void surface of PVAA-modified (left) and HECmodified (right) mortars

Figure 5: Abundant efflorescence of Ca(OH)*crystals at the air void surface of MC-modified mortars

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The air void surface of the MC-modified mortar (Figure 5 ) is characterized by an abundant efflorescence of Ca(OH)2 crystals, surrounded by needle-like C-S-H. The plate-like Ca(OH)2 crystals are formed parallel towards each other and they are almost undistorted. Methylcellulose has a high swelling capacity and can swell to forty times its dry volume in water [lo]. The presence of methylcellulose at the air void surfaces results in a very aqueous environment which promotes crystal formation. Ca(OH)2 crystals It has been shown by several authors [ll-131 that the presence of organics can significantly influence the morphology of Ca(OH)2 in cement matrices depending on the type of polymer, the brand of polymer, polymer-cement ratio or a combination of all these factors. In the absence of polymers, some crystals of Ca(0H)z are weak and unable to withstand the stresses generated during early hydration periods when the rearrangement of hydrates takes place in a limited space. However, the structure of Ca(OH)2 produced in the presence of polymer dispersions is modified to the extent that the crystals become capable of withstanding the stresses generated during early hydration [9]. The addition of polymer dispersions is reported to result in a layered deposition of Ca(OH)2 crystals arranged in stack without deformations [l]. The polymer particles act as a kind of a bonding agent between the different layers, increasing the interparticle bonding which produce more and better structure [13]. Although a lot of research is carried out on the influence of the addition of polymer dispersions on the microstructure, less is known about the effect of water-soluble polymers on crystal formation [14]. In Figure 6, the influence of small additions of methylcellulose on Ca(OH)2 formation is shown. A stack of layered and undistorted Ca(OH)2 crystals is formed inside an air void, as was also found in mortars modified with polymer dispersions [ 1,9]. In Figure 6, polymer bridges are detected between the Ca(OH)2 layers, which act as an additional bond, gluing the layers together. Because Ca(OH)2 crystals are the weak part in the binder matrix and the surfaces of those crystals form preferred cleavage sites, the strengthening by polymer bridges may improve the overall strength of the binder matrix. CONCLUSION

Unmodified mortar beams do not show any well-developed morphology of Ca(OH)2 crystals at the air void surfaces. Modification with small amounts of water-soluble polymers results in the formation of needle-like C-S-H and Ca(0H)Z crystals at those surfaces. Especially, mortars modified with 1% MC are characterized by an abundant efflorescence of Ca(OH)2 crystals. Between the closely packed layers of Ca(OH)2, polymer bridges are detected, forming additional bridges between the crystals. Thermal analyses show a smaller amount of Ca(OH)2 in cement pastes modified with water-soluble polymers, but a higher total amount of bound water when compared to the unmodified pastes. ACKNOWLEDGEMENT

The grant offered by the Institute for the Promotion of Innovation by Science and Technology in Flanders (IWT-Vlaanderen)is gratefully acknowledged.

Influence of water-soluble polymers on the microstructure of cement mortars

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Figure 6: Polymer bridges connecting layers of Ca(OH)Z crystals in an air void of MCmodified mortar

REFERENCES 1. Beeldens, A., Influence of polymer modification on the behaviour of concrete under severe conditions. PhD dissertation, Faculty of Engineering, Katholieke Universiteit Leuven 2002 2. Chung, D.D.L., Review: use of polymers for cement-based structural materials. Journal of Materials Science, 39,2004, pp 2973-2978 3. Hayakawa, K. and Soshiroda, T., Effects of cellulose ether on bond between matrix and aggregate in concrete, Adhesion between polymers and concrete. In: Proceedings of international symposium organized by Rilem Technical Committee 52 “Adhesion between polymers and concrete”, Ed. by H.R. Sasse, 1986, pp 22-31 4. Kim, J.-H. and Robertson, R.E., Effects of polyvinyl alcohol on aggregate-paste bond strength and the interfacial transition zone. Advanced Cement Based Materials, 8, 1998, pp 66-76 5. Knapen, E., Van Gemert, D., Water-soluble polymers for the modification of cement mortars. In: Proceedings of ISPIC International Symposium Polymers in Concrete, J. Barroso de Aguiar, S. Jalali, A. Camties, R.M. Ferreira eds. Guimarges 2-4 April 2006, pp 1-12 6. Knapen, E., Cizer, O., Van Balen, K., Van Gemert D., Comparison of solvent exchange and vacuum drying techniques to remove free water from early age cement-based materials. Submitted to 2nd International RILEM Symposium on Advances in Concrete through Science and Engineering, Quebec 11-13 September 2006 7. Taylor, H.F.W., Cement Chemistry. Academic Press Limited, London 1990

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8. Jenni, A., Holzer, L., Zurbriggen, R. and Herwegh, M., Influence of polymers on microstructure and adhesive strength of cementitious tile adhesive mortars. Cement and Concrete Research, 35(1), 2005, pp 35-50 9. Afiidi, M.U.K., Ohama, Y., Iqbal, M.Z., Demura, K., Morphology of Ca(OH)2 in polymermodified mortars and effect of freezing an thawing action on its stability. Cement & Concrete Composites, 12, 1990, pp 163-173 10. Bikales, N.M., Encyclopedia of polymer science and technology, vol 14, ed. N.M. Bikales, John Wiley & Sons (197 1). 11. Barker, A.P., Structural and mechanical characterization of calcium hydroxide in set cement and the influence of various additives. World cement,15, 1984, pp 25-31 12. Berger, R.L., McGregor, J.D., Effect of temperature and water solid-ratio on growth of Ca(OH)2 crystals formed during hydration of Ca3SiOs. Journal of American Ceramic Society, 56(2), 1973, pp 73-79 13. Afi-idi, M.U.K., Ohama, Y., Iqbal, M.Z., Demura, K., Behaviour of Ca(OH)2 in polymer modified mortars. The International Journal of Cement Composites and Lightweight Concrete, 11(4), 1989, pp 235-244 14. Mikhail, R.S., Shater, M., Al-Mad, T.M., Studies on premix water-soluble polymer cement pastes - I. Cement and Concrete Research, 13, 1983, pp 207-215

Proc. Int. Symp. '%BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

STATIC AND DYNAMIC RESPONSE OF CELLULOSE FIBERGYPSUM-BOARD WALL ELEMENTS

H. W. REINHARDT, R. FINN, S. AICHER Materials Testing Institute, Otto-Graf Institute, University of Stuttgart, Germany, e-mail: [email protected] ABSTRACT Tests of wall panel structures with fiber gypsum board sheathing are presented. First, static and cyclic tests of gypsum board to wood are described, second, panel tests with static, cyclic and seismic loading are carried out and discussed. The panels show a medium ductility. The presented results suggest that the design according to German standards is rather conservative. It is proposed to increase the allowable stresses to a higher level.

Keywords gypsum, cellulose fiber, wood, static and cyclic loading, seismic resistance INTRODUCTION Cellulose fiber gypsum is a short fiber composite board material typically manufactured with thicknesses of 10 to 18 mm. The fibers with a length of about 0.05 to 2 mm (mean value 0.2 mm) supply a volume fraction of about 15 to 20%. The fiber orientation is mainly in-plane random. The material incorporates several advantages versus classical gypsum wall board. The cellulose fibre gypsum boards are characterised by quasi-ductile behaviour, despite the fact, that the components cellulose fibers and gypsum are brittle. A noticable strain softening behaviour and a high damage tolerance were determined in tests with necked fiber gypsumboard specimens [l]. The material has a high stiffness to strength ratio being pronounced in comparison with other sheathing materials. Cellulose fiber gypsum boards are used as sheathing of timber frame structures, with load carrying and bracing function. Several investigations were performed in the last years to describe and model the inplane fracture softening properties of the material using notched and unnotched bending specimens [2-4]. There are several studies concerning the cyclic and seismic behaviour of wall elements with a realistic combination of internal forces and moments, comprising axial force in addition to shear force and bending moment [5-71. Oriented strand boards and plywood were the sheathing materials of the investigated wall elements. Wall elements with cellulose fiber gypsum boards, which differ observably from the other sheathing materials in the fracture characteristics (brittle/ductile characteristics), were investigated with a combination of vertical and horizontal action in the presented paper. The force-displacement relationship of doweled cellulose fiber gypsum-board joints is quasi-plastic. This is effected by three contributions: i) plastic deformations and damage in the timber, ii) plastic hinge formation in the fastener and iii) plasticity and damagekoftening in the sheathing panel. The plastic characteristics comprising ductility of doweled joints were investigated in [8]. The behaviour of joints determines the ductile performance of wall ele-

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ments sheathed with cellulose fiber gypsum boards. The aims of the presented investigations were to determine the ductility and its dependency on load history and to understand the damage evolution of shear walls. The German earthquake design code DIN 4149:2005 [9], which is implemented just now, reduces the allowable shear force of wall elements sheathed with fiber gypsum boards considerably. The limitation of design shear forces carried by fiber gypsum boards to 10% of the total shear force in ductility class 2 and 3 implicates, that only in ductility class 1 shear walls can be designed with two-sided gypsum board sheathing without other sheathing materials. The behaviour factor q for determination of design spectral acceleration, describing the ratio of the response of the completely elastic system to the system with plastiddamaging zones, is maximum 1.5. The test results presented in this paper indicate a higher ductility. SCOPE AND TESTING PROGRAM Fibre gypsum board to timber joints An important influencing factor for the mechanical behaviour of timber frame structures is the behaviour of the joints, the connections of the wall panels (and floor panels respectively) on the one hand and the sheathing to timber stud joints on the other hand. Extensive test series were performed, presented in extracts of a statically and a cyclically loaded fastener, a statically and a cyclically loaded wall element, and a seismic test in this paper. The plastic deformation capacity of the fibre gypsum board to wood connections was quantified in the first test series. Typical dowel-type fasteners used in timber frame structures were investigated, namely staples Haubold KG 750 CNK (shank length 50 mm, crown width 11.3 mm, wire diameter 1.53 mm) and smooth nails (length 65 mm, diameter 2.8 mm). The experimental setup was designed so that the deformation behaviour and the load carrying capacity of a fastener in a wall element could be observed. The loading rate of the displacement controlled tests was 0.167 d s , so that the induced inertial forces are negligible. Cyclic test on joint The protocol of the reversed cyclic loading was elaborated on the base of the monotonic static test following EN 12512. Except for the start of the test, three full cycles are repeated with the same displacement amplitude, chosen as a predefined multiple of the yield slip v,, obtained from the static test (Fig. 1). At the end of the third cycle, the amplitude is increased to the next displacement level. The failure criteria are either failure or a decrease of load carrying capacity to 0.8Fu or a displacement of 30 mm. The displacement between the top lumber and the fibre gypsum board, which equals the deformation of the doweled joint, was measured with a resistive displacement transducer.

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Fig. 1: Loading protocol of the quasi-static cyclic test following EN 12512 Static and cyclic wall element test An essential aspect of the wall element test was to simulate load conditions corresponding to the loads in real structures. It is obvious that in multi-storey timber frame buildings the highest horizontal force in the frst vibration mode acts on the ground floor shear walls. In wood panel constructions with three to four storeys, which is the maximum number of storeys in Germany, the first vibration mode is excited most (by typical earthquakes) in the case of structures being regular in plan and elevation. The highest vertical load on a storey, originating from gravity, acts in the ground floor. At materials characterised by a compression strength being significantly higher than the tension strength as fibre gypsum board the combination of the bending moment, coming from horizontal inertial force, with a compressive axial force, coming from vertical load, yields an improvement of the stress state and thus a higher strength. This effect was taken into account by the applied load conditions, conforming to the ground floor of a three-storey residential building. The structure of the wall panel is depicted in Fig. 2. The foundation of the wall panel was made of a reinforced concrete beam. At both ends, the wall element was connected with the reinforced concrete beam by a tension anchor. In order to determine the contributions of the different joints separately, the tension anchors of the wall elements were designed considerably stiff. Y

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Fig. 2: Structure of the two grid wall panel, fasteners: staples Haubold KG 750 CNK (length 50 mm, width 11.3 mm, wire diameter 1.53 mm)

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tension anchors concrete beam

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Fig. 3: Test setup and instrumentation of cyclic test (wall panel installed upside down) The test setup of the quasi-static cyclic tests is shown in Fig. 3. The vertical load was applied by a lever mechanism with concrete masses and kept constant during the test. The vertical load was transfered as concentrated loads at the studs, whereas the forces at the end studs were half as high as the forces at the internal studs. The wall element was loaded horizontally by means of servo hydraulic actuators with displacement control. The load protocol was elaborated following EN 12512. The shear deformation of the sheathing as well as the displacements at its edges were measured by LVDTs (A1 - A6). The rocking displacement at the supports was controlled by the sensors D1 - D3. The displacement difference between the sheathing and the timber frame was obtained from the LVDTs G1 and G2. The storey drift (Hl) was used as control factor.

Seismic wall element test Fig. 4 shows the test setup of the dynamic tests. The construction of the wall panels and the anchorage was the same as in the cyclic test. As the lateral and the torsional stiffness of the investigated wall panels is significantly smaller than the longitudinal stiffness, a structure consisting of two wall panels stiffened by a ceiling panel and two braces was designed. The reinforced concrete beams were connected to the shake table rigidly. The test models were stiffened by diagonal bracing with steel rods in lateral direction. The realised bracing of the structure allowed the observation of damage evolution also on the inner side of the wall panels. The connection of the bracing to the ceiling panel provided the possibility for free deformation of the wall panel in longitudinal direction. A mass of totally 10.0 tons was installed on the ceiling panel, which equals a vertical load of 19.6 kN/m on each wall panel. An ambient vibration test preceded the earthquake test to determine the initial natural frequencies. The natural frequencies of longitudinal, lateral, and torsional mode were measured. They served for comparison with the eigenvalues evaluated analytically. The initial stiffness was calculated from the natural frequencies. In the seismic test, the structure was excited with ground acceleration by the shake table. As wall panels act essentially in one direction, the structure was investigated for horizontal acceleration in longitudinal direction. The time history of the Friuli earthquake from 1976, E-W direction, recorded in Tolmezzo, Northern Italy, was applied. This earthquake is representative for the Alps, situated in Central Europe (Fig. 5). After the seismic test, the natural frequencies were measured again.

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Fig. 4: Test setup with instrumentation, steel ingots connected rigidly to the ceiling panel, and diagonal bracing 4 3 3

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TEST RESULTS AND DISCUSSION Static and cyclic test of fibre gypsum board to timber joint Statically loaded fibre gypsum board to timber joints with staples show two failure mechanisms. More frequently, pull-out of the staples occurs, whereby a plastic hinge develops in the staple legs. The plastic hinge moves towards the staple tip with increasing deformation. Furthermore, pull-through of the staples through the sheathing was obtained. The typical quasiductile, strain-softening load-deformationbehaviour is shown in Fig. 6.

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Fig. 6: Load-displacement curve of statically loaded fibre gypsum board to timber joint with staple Haubold KG 750 CNK (length 50 mm, width 1 1.3 mm, wire diameter 1.53 mm) Cyclic test on joint The load-displacement curve of the cyclic test is depicted for one staple in Fig. 7a. An apparent force emerges when the fastener is bent back, arising from its plastic deformation. At part of the way at reversed loading, the force is nearly constant, due to the cavity in the wood and the cellulose fiber gypsum board, caused in the preceding cycle, in which the fastener has no contact with the bearing. Additionally to the two failure mechanisms described before, low cycle fatigue of the staples emerges. A convex hysteresis loop provides better damping behaviour during an earthquake than a concave one, which is characterised as pinching effect. The pinching effect is expressed most at very stiff fasteners (which deform only elastic), the investigated staples show a medium pinching effect. The envelope curves of the hysteresis loops for the lSt, 2ndand 31d cycle are depicted in Fig. 7b. The comparison of the Figs. 7b and 6 shows, that the envelope curve of the cyclic test (1" cycle) and the load-displacement curve at static load are not identical. The envelope curves are not symmetric to the point of origin which significantly results from the strength reducing damage process in the connection at the initial load cycle drawn in the first quadrant.

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Fig. 7: a) Force-displacement relationship of one staple at cyclic load (Haubold KG 750 CNK), b) Envelope curves The strength impairment defines the reduction in load when attaining an imposed joint displacement from the first to the second cycle and from the first to the third cycle of same amplitude. Fig. 8 depicts the strength impairments F$ / Fll and F3' / F1' for the first quadrant in Fig. 7a as depending on displacement. The ductility at ultimate limit state, being the ability of the joint to undergo large amplitude displacements in the non-lineadplastic range without substantial reduction of strength, is defined by the ratio of ultimate displacement to yield displacement D = vu / vy. The ultimate displacement vu?given as the displacement with a strength reduction from the first to the third cycle of F3' / FI' = 0.8, equals 5.8 mm. Thus, the ductility is 7.4 at a yield displacement of 0.78 mm.

Hans Wolf REINHARDT, R. FIMV. S . AICHER

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Fig. 8: Strength impairment of cyclically loaded wood to fibre gypsum board joint with staple Haubold KG 750 CNK

Static and cyclic wall element test Fig. 9 depicts the horizontal force versus the storey drift of the two grid wall panel under static loading. The storey drift is the global displacement of the wall element, see Fig. 3. At an ultimate load of F,, = 127 kN, the fracture characteristic was relatively brittle with a slightly non-linear force-displacement curve. The wall element collapsed due to shear failure of the fibre gypsum board sheathing. The value of the secant stiffness at 0,4F,, was 12100 kN/m. The yield displacement was 6,O mm determined with the 0,1-0,4-0,9-F, method [lo]. Together with an ultimate displacement of 203 mm, this gives a static ductility of 3.4.

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Fig. 9: Load-storey drift curve of two grid wall panel under static load Comparing Fig. 9 with Fig. 10 one can see a completely different behaviour. While static loading yields a brittle failure cyclic loading leads to quasi-ductile behaviour. The deformation behaviour of wall panels composed of a timber frame and sheathing on both sides is characterised by a rotation of the sheathing relative to the timber frame in addition to the shear deformation of the sheathing. The shear distortion was determined directly from the

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measured diagonal displacements, provided that no bending deformation develops in the sheathing. The shear modulus of 1680 N/mm2 is obtained as the slope in the shear stress-shear distortion diagram. The equality of this value with the shear modulus determined in material tests on plain fibre gypsum board confirms, that the bending deformation of the sheathing is marginal and that the shear distortion and fastener deformation dominate the global deformation behaviour. At cyclic loading, larger displacements than at static loading were obtained, shown exemplarily for the storey drift in Fig. 10. The damage behaviour, being different to the former test, was characterised by staple pull-out and, at larger displacements, by low-cycle fatigue of a part of the staples. The staple pull-out was coupled with an uplift of the fibre gypsum board sheathing. The comparison of the displacements measured by the LVDTs A4 (showing small non-linearity) and GUG2 (giving a force-displacement curve being geometrically similar to Fig. lo), indicates that the plastic deformation of the fasteners, as investigated in the joint tests, causes the quasi-ductile behaviour of the wall element at cyclic loading. The envelope curve in Fig. 11 is more non-linear than the static load-displacement curve. Fig. 12 depicts the strength impairment versus the storey drift. An ultimate displacement of 24.4 mm was determined and a ductility of 4.1 was obtained. The ultimate load was 110 kN, being 13% less than in the static test. 120

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In the ambient vibration test preceding the shake table tests, an initial natural frequency of 7.6 Hz was measured in longitudinal direction. This served as verification for the two analyses of natural frequencies which were done on the basis of preceding test series, first performed with single structural members and second with whole wall elements. The first analysis was a finite element computation with the experimentally gained stiffness of the components. On the condition of fixed bearings at the bottom the calculated fundamental natural frequency in longitudinal direction was 7.9 Hz. The vibration mode belonging to fundamental natural frequency in plane model is given in Fig. 13. The global behaviour of the wall panels was modelled as a single degree of freedom system in the second analysis. The whole wall panel is represented by a spring, the stiffness of which was determined in a static wall panel test. The calculation yielded a natural frequency of 7.8 Hz. The analytical results and the measured initial natural frequency were in good agreement.

Fig. 13: 1'' vibration mode, boundary condition of fixed bearing at bottom rib, related natural frequency fi = 7,9 Hz

Static and dynamic response of cellulose fiber gvpsum-board wall elements

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The time history of acceleration response at the top of the wall panel in the direction of longitudinal degree of freedom is shown in Fig. 14. The curves of excitation performed by the shake table and the response of the structure at the ceiling level are given. It can be seen that the structure follows the excitation with low phase shift providing an indication of high stiffness. The inertia force at maximum acceleration of 3.8 m/s2 loaded the wall panel in the elastic range, no plastic deformation of a structural member occured. There was no visible damage of the staples connecting sheathing with timber frame after Friuli earthquake. Sound emissions could be observed beginning at the earthquake test indicating the exceedance of friction between structural members. After the earthquake test, the total damage of the shear walls was investigated by measurement of the fimdamental natural frequency, giving 6.8 Hz. The stiffness of one wall element decreased by 20% from 1 1 400 kN/m to 9 060 kN/m. The stiffness of the investigated wall panels was comparatively high yielding high natural frequencies. The amplification can be regarded as relatively small. This partly results from the relatively high stiffness but also it provides an indication of damping being present already at small deformations. The slip at the connections allows friction between the structural elements causing damping before plastic deformation and damage occurs.

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CONCLUSIONS

Wall panel structures with fibre gypsum board sheathing are characterised by medium ductility. Caused by the different deformation mechanisms, the ductility depends to some extent on the type of loading; at cyclic loading the ductility is by 10% higher than at static loading. The limitation of the design shear force carried by fiber gypsum boards to 10% of the total shear force in the German earthquake code DIN 4149:2005 is not justified. The ductility characteristics would allow a behaviour factor of at least 2.5. The presence of vertical load on the wall element has a significant influence on the carrying capacity for horizontal forces. According to the literature [ 5 ] , it has also influence on stiffness (increasing) and ductility (slightly decreasing). The presented results suggest that the design shear resistance of the tested wall elements is conservative. The design shear resistance according to the new German timber design code DIN 1052:2004 is determined to 42 kN, whereby the positive influence of the vertical load increasing the design shear resistance is not considered. The comparison of the ultimate horizontal force with the design shear resistance gives a load factor of 3.0 for the static test and 2.6 even for the cyclic test. REFERENCES

1. Aicher, S., Finn, R.: Fracture characterisation of cellulose fiber gypsum composite subject to inplane tension loading. Otto-Graf-Journal15 (2004), pp. 91 - 102 2. Aicher, S., Klock, W.: Fracture modeling of wood fiber gypsum board. Proceed. Int. Conf. on Wood and Wood Fiber Composites, Stuttgart, 2000, pp. 469 - 480, University of Stuttgart 3. Aicher, S., Klock, W., Reinhardt, H.W.: Fracture properties of wood fiber gypsum boards from size effect. J. Eng. Mechanics, ASCE, 132 (2004), (7), pp. 730 - 738 4. Klock, W., Aicher, S.: Size effect in paper fiber-reinforced gypsum panels under inplane bending. Wood and Fiber Science, 37 (2005), (3), pp. 403 - 412 5. Dean, K., Shenton Ill, H. W.: Experimental investigation of the effect of vertical load on the capacity of wood shear walls. J. Structural Eng., ASCE, 131 (2005), (7), pp. 1104 - 1113 6. DujiE, B., iarnik, R.: Influence of vertical load on lateral resistance of timber-framed walls. Int. Council for research and innovation in building and construction, CIB-W18 Proc. Meeting 35, Kyoto 2002, paper CIB-W18/35-15-4 7. Durham, J., Lam, F., Prion, G. L.: Seismic resistance of wood shear walls with large OSB panels. J. Structural Eng., ASCE, 127 (2001), (12), pp. 1460 - 1466 8. Reinhardt, H.W., Aicher, S., Finn, R. (2005): Energy dissipation and fracture softening of doweled wood-fiber gypsum board joints: the potential for seismically loaded shear walls. Proceed. 31d Int. Conf. on Construction Materials: Performance, Innovations and Structural Implications, Vancouver, p. 193 and chapter 1.14 on CD (10 pages), University of British Columbia 9. DIN 4149:2005 Buildings in german earthquake areas - Design loads, analysis and structural design of buildings. DIN Deutsches Institut ftir Normung, Berlin 10. Kawai, N.: Pseudo-dynamic test on shear walls. Proc. Fifth World Conference on Timber Engineering, EPF Lausanne, Montreux, Switzerland (1998)

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

EFFECT OF MATERIAL NON-LINEARITY ON THE FLEXURAL RESPONSE OF FIBER REINFORCED CONCRETE Chote SORANAKOM, Barzin MOBASHER, Saurabh BANSAL Department of Civil and Environmental Engineering Arizona State University Tempe, Arizona, 85287-5306 e-mail: [email protected] ABSTRACT This paper demonstrates the use of a closed-form derivation of moment-curvature diagram and crack localization rules in simulating the flexural response of textile reinforced cement composites. Effect of non-linear material response obtained fiom uniaxial tension tests on the simulation of the flexural beam test results were studied in order to establish a rational approach to balance the discrepancy between experimental data. The study reveals that the material nonlinearity can be adequately addressed by modifying two intrinsic material parameters: the frst cracking strain and initial tensile modulus. The direct use of uniaxial tension response to simulate flexural behavior underpredicted the response of textile cement composites. Modifications to the first cracking strain in the original tension model are recommended to yield better fit to the flexural experimental data.

Keywords: cement, tension, stress-strain response bending moment, fiber reinforced, composite flexural members INTRODUCTION The tension weak nature of concrete and cement products is a major problem that prohibits the development of many potential products. Over the years, various techniques have been developed to improve tensile capacity, increase fracture toughness, and minimize crack width of these brittle materials. An early improvement, steel fiber reinforced concrete SFRC [l], adds short steel fibers into the mix. The random steel fibers reinforce the matrix to hold cracks from opening. With this method, the post peak response is significantly improved; yet, it marginally increases the peak strength. Using the same approach, but at much higher fiber volume fractions than standard SFRC, slurry infiltrated concrete, SIMCON [ 1,2] and Ductal [4] have exhibited significant enhancement in tensile strength and post peak response. In contrast to the discrete reinforcement, the use of continuous fiber such as ferrocement FRC [5] provide a more efficient way to increase tensile capacity. The superior performance comes from the fact that smaller wire mesh has more contact area per unit volume to transfer more force into the matrix. The transverse wire also acts as a mechanical anchorage adding more resistance to pullout force. Furthermore, smaller fiber to fiber spacing minimizes flaw sizes, reduces stress concentration and leads to overall increase in strength. Besides using steel as reinforcement, other materials such as polypropylene (PP) [6] and alkali resistant (AR)glass have also shown improvement in tensile capacity and ductility. More recently, various textile reinforced cements (TRCs) [7-91 have also exhibited distributed cracking mechanisms and strain hardening behavior [lo].

Chote SORAh!4KOM, Barzin MOBASHER. Saurabh BANSAL

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A closed form solution of moment curvature diagram for general cement based composites and the algorithm to predict load deflection response of four point bending test [l]. The material models are based on parameterized uniaxial constitutive response of cement based composites. Instead of using a conventional iterative strain compatibility approach, the moment-curvature diagram according to the level of applied tensile strain at the bottom of the beam is expressed explicitly in dimensionless and parameterized forms. By this mean, the shape of the moment-curvature diagram is separated from the specimen size and material strength. The obtained dimensional moment-curvaturerelationship can be used as a sectional property of a beam element in plastic analysis to solve a variety of structural loading cases. Tensile response obtained from uniaxial test underestimates the flexural experimental results due to the material nonlinearity and variations in the stress distribution between the test methods [ 111. The main objective of this paper is to seek a rational way to account for the material non-linearity effects for simulation of a beam under four point bending test. CLOSED FORM SOLUTION FOR MOMENT CURVATURE DIAGRAM The derived closed form solutions for moment curvature diagram derived earlier are briefly explained in this section [l 11. The derivations utilized the homogenization concept which assume that the cement matrix is uniformly reinforced such that it acts as homogenized material with two distinct uniaxial models for compression and tension as shown in Fig. 1. The compression model shown in Fig. l(a) is defined by a parabolic function as:

where ECois the initial compressive modulus, K is the rate of compressive softening and &,, is the ultimate compressive strain. Figure 1(b) shows a trilinear tension model, which consists of three straight lines: the first line ascending from zero to the bend over point (BOP), the second line ascending from the BOP level to the ultimate tensile strength, and the last one descending from the ultimate tensile strength to zero stress. The tensile stress strain relationship can be expressed for three zones of tensile strain:

0, =

where Et0 is the initial tensile modulus, c; is the first cracking strain, E,I is the secondary modulus after cracking, 4 is the strain at the maximum tensile strength and &?isthe terminate strain at zero stress.

Effect of material non-linearity on the jlexural response offiber reinforced concrete

87

fi arbitrary terminating

E~.=

I

I I

I

I I I I

I

GO

Figure 1. Material model for a homogenized cement based composite. 2fl 1 Note that the compression model has relationshipsE,, = - and K = Ec

2ECO

h

Moment

Figure 2. Moment curvature diagram and crack localized rules. The inset shows the schematic of the four point bending test (only left half of model shown). To obtain non-dimensional forms of moment-curvature diagram, the material parameters used in Eq. (1) & (2) are replaced by the two intrinsic parameters; Et0 and and four normalized parameters 3: 77 a1 and az, which are defined as:

Chote SORANAKOM, Barzin MOBASHER, Saurabh BANSAL

88

In this approach, various characteristic of materials can be modeled by changing the normalized parameters. For example, equal compressive and tensile modulus can be modeled by setting y= 1 or setting y> 1 for material with compressive modulus greater than its tensile modulus. Linearly elastic compression response is defined by setting parameter K to 0 (no compressive strength drop). For the tension model, parameter q represents the reduced modulus of the composite after the first cracking of the matrix ( 0 < q 5 2 ), to define a transition from elastic behavior to a perfectly plastic material. To create a complete moment curvature diagram, the normalized tensile strain at the tension side of the beam p = &bor/EtO is imposed incrementally from stage I to N,in which stage I (0 < ,8 5 I), stage I1 (2 < /3 5 al),stage 111 (a1< ,8 5 az)and stage IV (B > az).An arbitrary normalized ultimate tensile strain ,8& = EruJEto can also be imposed to terminate the computations at a specified strain level. The equations for generating moment curvature diagram in each stage of normalized applied tensile strain are given by Eq. (4)&(5) and sub equation listed in Table 1 . 1 M , =Mi'Mo M,, = 6 bd'E,,~,, (4)

where subscript i refers to stages of normalized tensile strain ( i = 1,.4), Mis the moment, M' is dimensionless moment MO is dimensional scaling factor that accounts for geometry and strength at first cracking. Similarly, 4is the curvature, 4' is the dimensionless curvature and q$~ is the scaling factor at fust cracking strain. Both Mi' and 4 ' in Table 1 are expressed in term of normalized neutral axis k, which can be determined from the solution to the non-linear internal equilibrium of force, which is written in dimensionless form as:

zF= 4 k 3+ F2kZ+F,k+ F, (k-1)' The three roots of Eq. ( 6 ) can be expressed as: A(2/3)+ B - 2F2A(I/3) k(')= 64

+ B + 4F2A('/3)+ 1 2 4A ( ' / ~ )

+C

1

where A=36E;;F2F,-lO8F0&' -8F,3&,/4434 -F,'F;' -18F,F2F,F, +27F,'<' +4F,F:4 B =-1244

-+ 4F;

C = 1 2 i & 4 4 -4i&Fz

and F/ Q=O, 1, 2, 3) are normalized force coefficient given in Table 2 . When generating the moment-curvature diagram, the initial k value (0 < k < 2 ) at zero normalized tensile strain is

Effect of material non-linearity on the jlexural response offiber reinforced concrete

89

taken as 0.5 and the next solution k at subsequent application of tensile strain is obtained by selecting the root closest to the previous one.

CRACK LOCALIZATION RULES Figure 2 presents the schematic representation of moment curvature diagram with crack localization rules for loading and unloading directions. The inset shows a symmetric one-half model of four point bending test. The crack is assumed to start and localize in the mid-zone while the outside zone undergoes unloading as the material softens. The length of the localized zone is defined by a parameter c and spacing S=L/3, where L is a clear span. Cracks in brittle material localize into small region and lead to abrupt failure response (c -0) while cracks in ductile material are more distributed covering larger regions and exhibit more gradual failure response (c -1). In fiber reinforced cement composites, a default value c=0.5 was used since according to the experimental observations the cracks distribute throughout the mid-zone section. In order to use moment-curvature resposne (Fig. 2) in determining the curvature along the beam, the diagram is divided into two portions: an ascending curve from 0 to M,, and a descending curve from M,, to M&,. The first portion is used to determine the curvature of the section subjected to loading direction. After the specimen is loaded past the maximum, material becomes soften and the section in localized zone is unloaded following to the second portion of the moment diagram. For the sections in non-localized zone, the uncracked section is unloaded elastically while the section that experience some damage will unloaded with a quasi-linear recovery path expressed as:

where 44 and hij., represent the previous moment curvature state and 4, and M/ are the current state. The unloading factor is 0 < 5 < 1; 5 = 0 indicates no curvature recovery while 5 = I is unloading with initial stiffness (the default and used in the present study). E and I represent the elastic modulus and the uncracked moment of inertia.

ALGORITHM TO PREDICT LOAD DEFLECTION OF FOUR POINT BENDING TEST The simulation of load deflection response of a four point bending test is summarized as follows: 1) For a given section and material property, normalized tensile strain at the bottom of complete section p is imposed incrementally from 0 to the ultimate tensile strain moment curvature diagram is generated by Eq. (4) - (7) and procedures given in Table 1 and 2. 2) The maximum load for a half beam model is determined using equilibrium P,, = M,,/S and failure load Pfai/= Mfai~/S. Note that the total load are 2P,, and 2Pfoi1. 3) Arrays of incremental load steps from 0 to ZP,, and unloading steps from ZP,, to 2Pfai1 are generated and at each step, the moment distribution across the length is computed from static equilibrium. The corresponding curvature is determined from the moment curvature diagram and crack localization rules. 4) The deflection at mid-span is obtained by application of moment-area method between the support and mid-span. 5 ) Steps 4 and 5 are repeated for all load steps to obtain a complete load deflection response.

a.

90

Chote SORANAKOM, Barzin MOBASHER. Saurabh BANSAL

PARAMETRIC STUDY OF FLEXURAL AND TENSILE STRESSES

It is well accepted that tensile strength obtained fkom uniaxial test is lower than flexural strength. This is due to the nature of loading in uniaxial tension such that the entire volume of specimen is uniformly subjected to stress, whereas in a flexural test, only the small mid section of the tensile fiber is exposed to the high stress levels. Uniform tension stress in a brittle material results in a higher probability of stress concentration at natural flaws than a triangular bending stress. Use of tensile data in simulation of flexural response underestimates the experimental data [ 111. Model predictions can be resolved to match the experiments by increasing the tensile capacity primarily through modifymg the first cracking strain in the trilinear tension model. This will increase all the stresses and provides a simple approach to account for the material non-linearity. Alternatively, one can efficiently adjust two intrinsic material parameters: the first cracking strain and initial tensile modulus, which form the scaling factor for the material strength in Eq. (4) - (5). The modification to the first cracking strain will linearly change both the moment and curvature while the modification to the modulus will linearly affect the moment. The load-deflection response according to the mathematical model has the same shape as the moment-curvature diagram and can be scaled by the first cracking strain and initial tensile modulus. By adjusting these two parameters, the best fit of the load deflection and the experimental can be found. Figure 3 shows the effect of modifylng the first cracking strain and initial tensile modulus. An increase in the first cracking strain, & , directly affects the strain at ultimate strength Et1=alEato, strain at termination Q=CQ&, stress at bend over point oa = EE~O and ultimate strength om = E& + E ( E -~&) ~ by the same incremental amount as shown in Fig. 3(a). From Fig. 3(b), the increase in initial tensile modulus Et0 increases stress at bend over point and ultimate strength while strain Etl and & remain unchanged. Figure 3(c) shows that modification to first cracking strain & affects the magnitude of curvature and moment while Fig. 3(d) shows the initial tensile modulus E,o affects only the magnitude of moment. The corresponding load deflection response Fig. 3(e)&(f) follows the shape of moment curvature shown in Fig. 3(c)&(d) such that increase in first cracking leads to increases in both load and deflection (the response shift to the upper right) while increase in initial tensile modulus increases load capacity only (the response shift upward). Based on this parametric study, the underpredicted load deflection response can be resolved to fit the experimental results by adjusting either first cracking strain alone or both first cracking strain and initial tensile modulus. These modifications to the original trilinear tension response in order to fit the flexural response result in generation of modified constitutive response for the flexural case.

91

Effect of material non-linearity on the flexural response offiber reinforced concrete

2

Tensile Strain, mmlmm 20000h

I

I

I

Curvature, l/mm

Curvature, l/mm

Deflection, mm

Deflection, mm

I

Figure 3. Parametric study on the first cracking strain and initial tensile modulus; (a) Trilinear (b) Trilinear model (vary Eta), (c) Moment curvature (vary GO),(d) Moment model (vary curvature (vary &), (e) Load deflection (vary &to) and ( f ) Load deflection (vary Eta)

Chote SORANAKOM, Barzin MOBASHER, Saurabh BANSAL

92

PREDICTION OF LOAD DEFLECTION RESPONSE Textile reinforced cement composites were manufactured using a cement paste with a water to cement ratio of 0.45, and Alkali resistant glass fabrics manufactured by Saint-Gobain Technical Fabrics Inc. [12, 131. The grid size was 25.4x25.4 mm with 2 yams in each of the longitudinal and transverse directions. Each yam consisted of 1579 filaments, each 19 microns in diameter. Two layers of fabric were placed at the top and bottom of the specimens to provide reinforcement in each direction: VL =VT = 0.70%. The simulation employed three sets of specimens at 0, 14 and 28 days accelerated aging. Uniaxial tension and four point bending tests were conducted to correlate the tension and bending responses. Tension specimens were approximately 10x65~200mm. Flexural specimens for the four point bending test were 10x75~356mm with a clear span of 254 mm. The compressive modulus ECofor textile reinforced cement was estimated by back calculation fiom the initial slope m of the flexural load deflection: E =--23 mL3 (9) co 54 bh3 The parameter y can be calculated by (3). Parameter K for the compression model was obtained by equating the relationship between the standard Hognestad’s parabola [ 141 and the parabolic representation of the compression model described in Fig. l(a):

One can obtain the relationship GOand K:

For these simulations the compressive strength& ’ of cement paste was assumed to be 45 MPa and ultimate compressive strain &u was assumed at 0.012. The initial tensile modulus E,o was obtained directly fiom initial response of the uniaxial test and parameters 7,ajand a2 are obtained by fitting trilinear tension model to uniaxial tension test results. The assumed material parameters on compression fiber affect the prediction marginally since the flexural response is strongly dominated by the tensile behavior. Table 3 lists the complete set of parameters for all three simulations. Figure 4 (a), (c) and (e) show the average tension responses for three samples (0, 14, 28 day accelerated aging) along with the fitted trilinear tension model. Modifications to the intrinsic material parameters (k and E&&o) to account for material non-linearity effects are also shown. Figure 4 (b), (d) and ( f ) show the direct use of the tensile responses and the modification models to predict flexural responses. Without any modification, the simulation for samples accelerated aging 0, 14, and 28 days underpredicts the peak load of experimental results by 20.8%, 37.4% and 40.1%, respectively. To measure how well the model and algorithm predict the experiment result, a standard error of estimate s is used, which is defined by:

Effect of material non-linearity on the flexural response ofjber reinforced concrete

93

where, yi andy,,,d are the experimental and predicted load at the same deflection point and the toal number of deflection points n = 100 from 0 to the deflection at the load in post peak drop to 0.75 of the maximum load is used in the calculation. With this statistic measure, the fitted trilinear model yield standard error of 52.4, 83.5 and 73.5 N, respectively. To compensate for the underprediction due to material nonlinearity, an increase in first cracking strain with the values shown in Table 3 and drawn in Fig.4 (a), (c) and (e) result in better predictions. The standard error reduced to 35.1, 3 1.4 and 2 1.9 respectively. An even better fit can be obtained by modifying two parameters &0&Eto as shown the value listed in Table 3 and drawn in Fig. 4(a), (c) and (e), the standard error is further reduced to 24.5,20.5 and 21.5 N, respectively. CONCLUSION

The closed form solution for moment curvature diagram provides an alternative approach to the conventional iterative strain compatibility approach for obtaining moment curvature diagram. With crack localization rules, the curvature in localized and non-localized zones can be used in calculation of mid span deflection. The simulations of three samples subjected to accelerated aging at 0, 14 and 28 days confirm that the direct use of uniaxial tensile response underpredicts the flexural response due to material nonlinearity effect. A parametric study of two intrinsic material parameters reveals that the first cracking strain can be used to scale the predicted load and deflection while the initial tensile modulus can be used to scale only the predicted load. The underestimated response of three samples can be resolved to match to the experimental result by adjusting these two parameters.

Chote SORANAKOM, Barzin MOBASHER, Saurabh BANSAL

94

0 0 OExperiment

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0.02 0.04 0.06 0.08 0.1 Tensile Strain, mmlmm

10 20 30 Mid Span Deflection, l/mm

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0

10 20 30 Mid Span Deflection, llmm

Figure 4: Simulation of textile reinforced cement at three ageing; (a) Trilinear tension model at 0 day, (b) Predicted load deflection at 0 day, (c) Trilinear model at 14 days, (d) Predicted load deflection at 14 days, (e) Trilinear tension model at 28 days and ( f ) Predicted load deflection at 28 days

Effect of material non-linearity on the flexural response offiber reinforced concrete

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Chote SORANiKOM, Barzin MOBASHER, Saurabh BANSAL

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Table 2. Coefficients in equilibrium of force for each stage of applied tensile strain diagram Stage Dimensionless net force F'i and the dimensional factor (Fo = bdE,,~,,)

+3y-

F ', = -P( 2 y K p E o 6

F

l2

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I

= -(9 P

6

3)

-3y)

= - - 3p 2

F l o = $

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= -{3q

6P 1 F I 2 = -{97

68

I1

1 F ', = -{9q 6P 1 F lo= -{37 6P -

F', =-

111

(-p2+ 2 p - 1 ) ( p 2- 2 p + 1 ) ( - p 2 + 2 p - 1)-

6 p + 2 Y P 3 E 0 K + 3 y p 2+ 31

3 y p 2 + 18p - 9 } 18p + 9 }

- 2p + 1 ) + 6 p - 3 }

(p'

3q (-a; - 2pa,a2+ 2pa2- p2+ ga2+ a,- a2+ alp2)

[

I I

9q(a2-q2a2-a, -alp2+2pa,a2+a;' +p2-2pa2)

F ' , =-

+9(-p2-a;'+al-a2)+3@2(al -a2)+18Pa2 F', = -

I

6P( a,-a2) +3( a; + @'a2+ a2+ 4 +p2- @'a,)+ 2 & ~ ( a2-a,)- Spa2

1

9q(-a;'+2pa2+alp2+a;2a2-p2-2~a,a2-a2+a,)

6p(al-a2) +9(-a,+p' +a;'+a2)-18pa2

F ' , = - {13 ( y p 2 -q -a2qa,+1-a,+a,g-a2+qa,)+2y~3&O~} 6P 1 F l2 = - -{3 (Val - a , + a2q- a,- q + 1 + yp2 - 3a,a,q)} 2P 3

Iv

F', =--(-l-qal+q+a2+al+a,a,q-a,q) F

2P 1

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= - -(-1-

2P

qa, + q

+ a2+ a , + ala2q- a,q )

14 days accelerated aging (fit experiment) 14 days accelerated aging (modified strain) 14 days accelerated aging (modified strain&stiffness) 28 days accelerated aging (fit experiment) 28 days accelerated aging (modified strain) 28 days accelerated aging (modified strain&stiffness)

0 days accelerated aging (fit experiment) 0 days accelerated aging (modified strain) 0 days accelerated aging (modified

Comuosite Suecimens Section

(-1

254

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254

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3,000

3,667

0.00089 0.00060 3,661 3750

0.00087 0.00086

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2,554

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0.00092 0.00070

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Material Parameters

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8,944

8,746

8,146

12949

11,024

11,024

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1,579

7,579

45.0

45.0

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98

Chote SORANAKOM, Barzin MOBASHER, Saurabh BANSAL

REFERENCES Umekawa, S., and Nakazawa, K., “OnMechanical Properties of Stainless Steel Fiber and Fiber Reinforced Stainless-SN-PB Alloy Composite,” J Jap Inst Metals, V.34, No.2, 1970, pp. 222-227 2. Krstulovic-Opara, N., and Malak, S., “Tensile Behavior of Slurry Infiltrated Mat Concrete (SIMCON),” ACZMaterial J, V.94, No.1, Jan.-Feb. 1997, pp. 39-46 3. Bayasi, Z., and Zeng, J., “Flexural Behavior of Slurry Infiltrated Mat Concrete (SIMCON), J Materials Civil Engineering, V.9, No.4, Nov. 1997, pp. 194-199 4. Rossi, P., “High performance multimodal fiber reinforced cement composites HPMFRCC: the LCPC experience,” ACI Muter J, V.94, No.6, Nov.-Dec. 1997, pp. 478-483 5 . Naaman, A.E., and Shah, S.P., “Tensile Test of Ferrocement,” J Amer Concrete Inst, V.68, No.9, Sept. 1971, pp. 693-698 6 . Naaman, A.E., and Shah, S.P., and Throne, J.L., “Some Developments in Polypropylene Fiber for Concrete. Publication SP,” American Concrete Institute, 1984, pp. 375-96 7. Swamy, R.N., and Hussin, M.W., “Woven polypropylene fabrics-an alternative to asbestos for thin sheet application: Fibre reinforced cement and concretes, recent developments,” EZsevier Science, 1989, pp.99- 100 8. Peled, A., Bentur, A., and Yankelevsky, D., “Effects of Woven Fabric Geometry on the Bonding Performance of Cementitious Composites: Mechanical Performance,” Adv Cement Based Materials, V.7, No.1, Jan. 1998, pp. 20-27 9. Peled, A., and Mobasher, B., “Pultmded Fabric-Cement Composites,” ACI Materials J; V. 102, No. 1, Jan.-Feb. 2005, pp. 15-23 10. Peled, A., Bentur, A., and Yankelevsky, D., “Flexural Performance of Cementitious Composites Reinforced with Woven Fabrics,” J Materials in Civil Engineering, V. 1 1, No.4, NOV.1999, pp. 325-330 11. Soranakom, C., and Mobasher, B., “Closed-Form Moment Curvature Expressions For Homogenized Fiber Reinforced Concrete,” paper submitted to ACI Journal, Apr. 2006 12. Aldea Corina-Maria, Barzin Mobasher, Nora Singla, “Cement-Based Matrix-Grid System For Masonry Rehabilitation,” ACI Special Publications, in review, 2006 13. B. Mobasher, N. Singla, C. Aldea, C. Saranakom, “Development of Fabric Reinforced Cement Composites for Repair and Retrofit Applications, Textile Reinforced Concrete (TRC) - German/International Experience symposium sponsored by the ACI Committee 549, ACI Special Publications, in review, 2006 14. Hognestad, E., “A Study of Combined Bending and Axial Load in Reinforced Concrete Members,” University of Illinois Engineering Experimental Station, Bulletin Series No. 399, Nov, 1951,128 pp 1.

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

THE EFFECT OF DURABILITY ON THE DESIGN OF SELF-BEARING SANDWICH PANELS WITH CEMENTITIOUS COMPOSITE FACES Heidi CWPERS, Petra VAN ITTERBEECK and Jan WASTIELS Department of Mechanics of Materials and Constructions Vrije Universiteit Brussel Pleinlaan 2, B-1050 Brussel, Belgium e-mail: [email protected]

ABSTRACT In this paper the effect of durability modelling on the design of a textile reinforced concrete (TRC) building element is studied. The effect of long-term evolution of the strength is discussed for self-bearing sandwich elements with TRC faces, used as wall cladding elements. In this paper, focus is put on design according to ultimate limit state. Several strength loss models are calibrated on accelerated ageing experiments. These calibrated models are then used to make long term predictions for the sandwich panels and results are discussed.

Keywords Durability, textile reinforced concrete INTRODUCTION Textile reinforced concrete (TRC) composites with glass fibres as reinforcement are interesting load-bearing materials for construction purposes. They are advantageous when used in freeform architecture and in lightweight construction elements (e.g. faces in sandwich panels) andor when a certain capacity of energy dissipation is required (e.g. impact resistant construction elements or earthquake resistant buildings). Development of AR-glass fibres, coatings and low-alkaline matrices led to a prolonged lifetime (durability) of glass fibre reinforced concretes. However, although the durability of this material combination is seriously improved, loss of material performance is yet to be incorporated into the global design concept of building elements with textile reinforced cementitious composites with glass fibres as reinforcement. In this paper three durability models are discussed. All three models are based on the idea that loss of strength of the glass fibres in TRC is mainly determined by the deepening of flaws in the glass fibres, combined with the classical Griffith relationship between the free energy of a cracked body under stress and maximum crack size. In this paper all discussed data are retrieved on AR-glass fibre reinforced cement (grc) specimens immersed in water. The three presented models will be discussed based on experimental data found in literature (on two series of specimens with Ordinary Portland Cement (OPC)) and own data (for the third series of specimens, an Inorganic Phosphate Cement (PC) is used). Once the presented models are calibrated based on these three series of data, they will be used as a predictive tool for the study of the design of sandwich elements with TRC faces in ultimate limit state. For

100

Heidi CUYPERS, Petra Van ITTERBEECK, Jan WASTIELS

illustration purposes, it will be assumed that the self-bearing sandwich panels will be used as wall-cladding elements, exposed to Belgian weathering conditions. SANDWICH DESIGN

The design procedure of the discussed sandwich panels is based on Eurocode 1 [l] and the preliminary European Recommendations for Sandwich Panels (Part I, Design) [2]. According to these documents it should be verified for a structure or structural element that following equation is satisfied at both the ultimate limit state and the serviceability limit state.

F = calculated value ofan action (and EyfF = FD= design value of action) R = value of resistance at relevant limit state (and Wy,= RD = design value of resistance) yf= relevant load factor (safety factor) ym = relevant material factor (safety factor) For self-bearing sandwich panels, the variable actions that should be taken into account are: wind, temperature and snow (for roof panels only). For this paper, the loads that are taken into account effectively are self-weight (as a permanent load), the wind and temperature load, since the panels are considered to be wall-cladding panels without any other structural components applying extra loading. The temperature at the inside face may be taken as 20°C in winter and 25°C is summer. The temperature of the outside face has a minimum winter value of -20°C and a summer value of +80°C (which is the worst case scenario when a dark colour is applied at the outside surface). Long-term effects (creep) are neglected in this paper, but should also be introduced for real design applications. The ultimate limit state corresponds to the maximum load carrying capacity and is characterised in this paper by the following failure modes (individually or in combination): (a) failure of one of or both the outer skins, (b) wrinkling of the outer skin in compression with consequential failure, (c) shear failure of the core. For verification of the ultimate limit state, the variable action or the action effect, corresponding to the variable action, which has the largest action effect ( Q ~ J )is, multiplied by yf and all other effects (Qk,i with i>l) by V0,iyf. The load combination to be considered for ultimate limit state is thus defined by following expression:

Gk = characteristic value of own weight Q k , l = characteristic value of

main variable load of other variable loads .= i combination values variable loads, other then main load

Qk,i = characteristic values ~ 0

According to the preliminary European Recommendations for Sandwich Panels (Part I, Design), the combination values to be considered are: ~0 = 0.5 both for temperature and wind. The load factors yf are 1.5 for variable actions and 1.35 for permanent actions in ultimate limit state. Wind effects are most pronounced at the Belgian coast. It is thus assumed here for all

The effect of durability on the design of self-bearing sandwich panels with cementitious compositefaces

101

case studies that the building is situated at the coast. The maximum height of the building, on which the sandwich panels are placed, is chosen to be 30m. In this case a typical characteristic wind load in Belgium would have a value of 1635N/mz(for a lifetime of 5Oyears). The design values of material or product properties can be formulated as follows:

The material safety factors y,,are set to 1.25 for the core strength in shear, 1.5 for the failure strength TRC faces and 1.25 for the wrinkling failure of the face in compression. q is in general a mean value of conversion to include scale effects or effects of the environment (under moisture and temperature). This factor will be used in this paper to express the loss of strength of the TRC faces during a chosen lifetime of a construction. DURABILITY MODELS

Several mechanisms have been identified as being sources of loss of strength of glass fibre reinforced concrete in general. Chemical attack of the fibres and embrittlement, resulting from the penetration of the fibre bundle space by cement hydration products are generally identified as being the main driving mechanisms for loss of strength in grc ([3], [4]). Durability models predicting the evolution of the strength of grc usually have a strong (semi-) empirical background. Early durability models [5], stated that if the rate of loss of strength of glass fibres is related to some chemical reaction and if the time taken for the strength to fall to any given value ot can be regarded as an inverse measure of the rate of strength loss, that one can assume that an Arrhenius type of relationship will exist between the time (t) taken for the strength to fall to any given value otand the temperature can be written as follows:

with k a constant for a given material combination (matrix and fibres). This relationship was determined from tests on SIC (strand in cement) specimens. It was proven [5] on series of SIC specimens that there indeed exists a good linearity between loglo(t) and 1/T over a temperature range of 20°C to 80°C. This is an indication that there is one main chemical reaction or corrosion mechanism controlling the strength changes. More recent models explore the idea that loss of strength indeed occurs due to one corrosion mechanism, but that this mechanism should be combined with the classical Griffith relationship between the free energy of a cracked body under stress and crack size ( [ 6 ] ,[7]). Due to the growth of the size of the flaws (nano-defects) with time (accelerated in the presence of moisture and pH), the failure stress will decrease as follows:

with

- ft the tensile strength at time t - K1, critical stress intensity factor (Mode I) - A shape factor

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Heidi CUYPERS, Petra Van ITTERBEECK, Jan WASTIELS

- a critical flaw size (= largest flaw depth in fibre) From measurement of the evolution of single fibres in simulated pore solutions, it was indeed noticed that non-uniform corrosion of the fibre surface occurs [8]. This idea is confirmed by the analysis of the statistical distribution of fibre strength with ageing [9] and by evaluation of the roughness of the surface of fibres in simulated pore solutions. If it is assumed that the reaction between the glass and the environment is simply controlled by the reaction kinetics and that the flaw depth (some tens of nm) is much smaller than the fibre diameter (usually around 14 or 24pm) we can write that [lo]: a = a, + k,t

(6)

with

- kl a rate coefficient, depending on temperature - a the current size of the crack - the initial size of the crack - t time of ageing When equation ( 5 ) and equation (6) are combined, one can find: 1

(7)

For the variation of kl with temperature, again an Arrhenius relationship will be used, since it indicates the rate of one chemical reaction:

- kc, I + Ink,,, ln(k,) = T in which k11 and klz are constants, determining the relationship between the rate of chemical degradation and the temperature (in degrees K). The model based on equations (6) to (8) is the first model, which will be further discussed in this paper. This model will be further referred to as 'kinetics model'. Several authors noticed a gradual transition from the initial fall in strength of grc under accelerated ageing towards a smoother - more gradual -evolution of strength in grc ([S], [l 11). The same evolution has been measured on series of single filaments stored in simulated pore solutions [S]. It is suggested by some authors (e.g. [12]) that pitting of the fibre surface, resulting fiom the removal of silicon atoms increases the relative zirconium content on the surface, thus leading to a progressively reducing area of the surface for further attack. After some time the surface might be covered by exposed zirconium ions or a zirconia-rich reaction product. In order to be able to model possible slowing down or acceleration of chemical attack, one can describe a non-linear progression of the flaw depth [131: a=a,+k,t"

(9)

When n > 0 the reaction accelerates, when n < 0 the reaction slows down with time. Assuming that slowing down of a reaction is an indication that chemical attack becomes diffusion-controlled after some time ([8], [ll]), one can write a relationship, indicating that

The effect of durability on the design of self-bearing sandwich panels with cementitious composite faces

103

the rate of degradation is initially determined by the kinetics and becomes diffusion controlled in a later stage. The flaw size growth X can thus be expressed as follows: d -x 1 dt 1 -+-

x

with:

a=a,+x The remaining strength can again be calculated from equation (5). The formulation of kl is the one expressed in equation (8). Also for k2 an Arrhenius type of relationship is assumed: ln(k,)=-+Ink,,, - kc,, T

(12)

in which b11 and k12are constants, determining the relationship between the rate of chemical degradation and the temperature (in degrees K). In the following sections, this model will be referred to as ‘combined model’. Apart from the two proposed models, a third option will be discussed. There is a possibility that chemical attack of fibres in TRC will be diffusion-controlled from the start. In this case the flaw size growth is expressed as follows:

The model, combining equation (13) with equations (7), (1 1) and (12), will be referred to as ‘diffusion-controlled model’ or ‘diffusion model’ in this paper. MATERIALS In this paper, the three presented models will be used to fit results coming from data published in literature, combined with own data on grc with OPC and PC.An overview of the discussed experimental series is listed in table 1. The terminology long or short into the names of the series refers to the maximum time of ageing (where series based on data from 200 days of testing or less are considered as short-term data). Table 1: series of data used to calibrate presented models ~

Series (OPC-short) (OPC-long) (IPC-short)

Reference Matrix [ 141 Portland cement [51 Portland cement [ 151 Inorganic Phosphate Cement

Temperatures 38,65, 80 4,19,35,50,60,80 50°C

Max ageing time 200 days 8600 days 90 days

Heidi CUYPERS, Petra Van IlTEmEECK, Jan WASTIELS

104

RESULTS Constants hi, are determined for per series of specimens tested at one specific temperature (for a specific material combination). A program was written in order to determine those values of hij leading to a lowest discrepancy between the experimental data and theoretical predictions. This optimisation is based on minimisation of a least square discrepancy factor:

S,,

(t) is a measured remaining strength of a grc specimen after ageing time t

Sheo(t) is a measured remaining strength of a grc specimen according to the model used

The resulting lowest LSC for the three discussed models on the (OPC-short) specimens is depicted in figure 1. -7

0.08 ,

A

0

3 0.06

OSo2 0,Ol

I

04 0

O 0 20

40

60

I

I

80 100 temperature (“C)

Figure 1 . Least square fit coefficient for discussed models for series tested at different temperatures (OPC-short series, indicated as OPCII in [ 141) As can be seen from figure 1 all models show rather similar accuracy to fit the degradation behaviour as a function of time for the different temperatures. Figure 2 shows the least squares coeficient for the data series that was recorded on an OPC grc for 8600 days [5] at various temperatures.

0 combined model

Adiffusion model

40

60

I 80 100 temperature (‘C)

Figure 2. Least square fit coefficient for discussed models for series tested at different temperatures (OPC-long series, grc specimens in [5])

The effect of durability on the design of self-bearingsandwich panels with cementitious composite faces

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For the second series of tests, the ‘kinetic model’ shows larger discrepancies, while the ‘combined model’ and the’ diffusion model’ show better coincidence with the experimental results. Based on figures 1 and 2 one would conclude that the ‘combined model’ is the most convenient model to predict both short time and long time degradation of the discussed grc series. However, in order to really verify the soundness of the three models, the Arrhenius relationship between the inverse of temperature and the logarithm of the degradation rates kl and k2 should be checked. These relationships are plotted in figure 3. Figures 3a, 3b and 3c show the Arrhenius relationships for the ‘kinetic model’, the ‘combined model’ and the ‘diffusion controlled model’ for the (OPC-short) series. Figures 3d, 3e and 3f show the Arrhenius relationships for the ‘kinetic model’, the ‘combined model’ and the ‘diffusion controlled model’ for the (OPC-long) series. Figures 3a, 3b, 3c show that for the (OPC-short) series all presented models show a good linear fit between 1/T and the logarithm of the rates. For the (OPC-long) series this conclusion is still valid for the ‘kinetic model’ and the diffusion model (see figures 3d and 30. However, when both the evolutions of kl and k2 are to be determined as a function of temperature for the combined model (figure 3e), no real clear evolution of kl can be found. It seems that trying to calibrate this model for the combination of both effects is an awkward task. Determination of the discrepancy between the theoretical curves and experimental results gave similar discrepancies for all models for the combined AR-glass with IPC. Since only one series of experiments was used (at 5OoC), these results are not presented in a separate figure. Using the determined Arrhenius constants for all models on (OPC-short) and (OPC-long) series, a prediction of the residual strength of these series under climatic conditions in Belgium (assuming an average temperature of 93°C and constant high humidity) is proposed here. In table 2, the predicted residual strength after a lifetime of 50 years at 9.8”C according to the three models is given. Since the results obtained on the IPC series were only determined at one fixed temperature, the Arrhenius constants cannot be determined. However, the rate coefficients at 50°C can be determined. The prediction of the long term remaining strength of the IPC specimens after 50 years at 50°C is also given in table 2. In order to have a at least some type of comparable result, predictions for all models on the (OPC-long) series at 50°C for 50 years are also depicted.

Model Kinetic Diffusion Combined

OPC-short at 93°C 0.24 0.42 0.30

OPC-long at 9.8”C 0.46 0.54 0.99

P C at 50°C 0.16 0.89 0.16

OPC-long at 50°C 0.03 0.15 0.14

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Heidi CUYPERSl Petra Van ITTERBEECK, Jan WASTIELS

lil 0,0032

0,0034 -2

-4

-3

E-l

“I

R2= 0,999

Fp = 0.9921

-5 (3a)

In(k1)

In(k2)

12

i

I01

Oln(k2)

+In(kl)

T 5(

10

8

R = 0,9941

0

0,003

0,0032

in p 10 Y

=

1\

R2=0,9976

oc-v----0,003

(3b)

10

Y

0,0028

2

0,0032

1il

0,0034

0,0025

0,003

(3c)

0,c (3

Figure 3: determination Arrhenius constants for three models on two series of data on OPC composites

As can be seen fiom table 2, the determination of the long term remaining strength is rather similar for all three models when the (OPC-short) series was used for calibration. Remaining strength values of 24%to 42% are predicted. When the resulting Arrhenius constants are used, as determined on the (OPC-long) series, one can see that the ‘kinetic model’ and the ‘difision model’ both give similar predictions. However, the combined model gives a result that can hardly be considered logical. Probably the reason for this large discrepancy can be found in the fact that it seems to be difficult to find values of b11,b12,b z l , kc22 in order to obtain proper Arrhenius relationships for kl and k2 (see also figure (3e)). Concerning the P C series, the ‘combined model’ and the ‘kinetic model’ predict very limited remaining strength. However, the ‘combined model’ predicts a similar strength for a high alkaline matrix as for a neutral matrix, which again seems not logical. The ‘diffusion model’ however predicts a much higher remaining strength for the IPC than for the OPC. In this stage of research it would be precarious to draw conclusions about which prediction seems to be the most logical for IPC: the ‘kinetic model’ or the’ diffusion model’. Since the

The effect ofdurability on the design ofself-bearing sandwich panels with cementitious composite faces

107

studied type of matrix differs considerably from more common mixtures, one cannot prefer one model above another yet. Extra long-term data are thus needed. Finally, the obtained remainihg strength values from table 2 are used as mean values of conversion (11) in equation (3) in order to study the effect of ageing in design of ultimate limit state for sandwich panels for TRC faces. For illustration purposes, a sandwich with a span of 3m, a width of lm, a PUR foam core with a density of 45kg/m2 is discussed. It is assumed that the shear strength of the core is 0.1MPa. The strength of glass filaments in a bundle, after implementation on a TRC matrix, is usually much lower than the initial average strength of single filaments. For this paper it is assumed that the reference value of the strength is about 1000 MPa and that the Weibull modulus is 9. In this case the characteristic value of the strength is 720 MPa. When a material coefficient of 1.5 is applied, this leads to a design value of 480 MPa of the fibre strength. If a fibre volume fraction of 10% is introduced under the form of random mats, this means the design value of the TRC is around 16MPa. The necessary thickness of the core is based on the shear strength criterion and is thus 50mm. The necessary thickness of the faces can be calculated and are depicted in table 3.

Model Kinetic Diffusion Combined

TRC face thickness (mm) OPC-short at 93°C 11 7 9

TRC face thickness (mm) OPC-long at 93°C 7 6 3

CONCLUSIONS In this paper three models to predict the evolution of the strength of grc as function of time and temperature were discussed: a kinetic model, a diffusion-controlled model and a combined model. These models were used for long-term predictions of TRC, after they were calibrated on existing data-series under accelerated ageing. From these results, long-term predictions under average Belgian weathering conditions were made. For a relative short-time accelerated ageing data series with classical OPC, the kinetic model seemed to fit the experimental evolution well. For the long-term series of data, the diffbion controlled model seemed to fit the behaviour better. Although it would sound logical to use a model, combining both mechanisms, it was shown that such a combined model should be calibrated with care. ACKNOWLEDGEMENT Funding by the Flemish Fund for Scientific Research (FWO) of the post-doctoral research of the first author is acknowledged. REFERENCES 1. Eurocode 1. Actions On Structures 2. Preliminary European Recommendations for Sandwich Panels, ECCS, TC7, 1991 3. Marikunte, S., Aldea, C. ,Shah, S.P,. Durability of Glass Fiber Reinforced Cement Composites: Effect of Silica Fume and Metakaolin, Adv Cem Bas Mat, 5, 1997, pp 100-108

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4. Larner, L.J., Speakman, K. Majumdar, A.J., Chemical Interactions between Glass Fibres and Cement, Journal of Non-Crystalline Solids, 20, 1976, pp 43-74 5. Litherland K.L., Oakly, D.R., Proctor, B.A., The use of Accelerated Ageing Procedures to Predict the Long Term Strength of GRC Composites, Cement and Concrete Research, 11, 1981, pp 455-466 6. Griffith, A.A.:,The Phenomena of Rupture and Flow in Solids, Philosophical Transactions of the Royal Society 22 1A, 1920, pp 163 7. Freiman, S.W., Fracture Mechanics of Glass, Elasticity and Strength in Glass 5 (Uhlmann, D.R. ;Kreidl, N.J. (Ed.)) Academic press, new York, 1980, pp 21-78 8. Durability modeling of glass fibre reinforcement in cementitious environment, J. Orlowsky, M. Raupach, H. Cuypers, J. Wastiels, Materials and Structures, 38,2004, pp 155-162 9. Gao, S.L., Mader, E., Abdkader, A., Offermann, P., Environmental resistance and mechanical performance of alkali-resistant glass fibres with surface sizings, J. of noncrystalline solids, 325, pp 230-241 10. F’umell, P., The Durability of Glass Fibre Reinforced Cements made with New Cementitious Matrices, PhD Thesis - Aston University UK, 1998 1 1. Orlowsky, J., Dauerhaftigkeit von AR-glabewehrungen in Betonen, PhD Thesis RWTH-Aachen, 2005 12 Lamer, L.J., Speakman K. and Majumdar, A.J., Chemical Interactions between Glass Fibres and Cement, Journal of non-crystalline solids, 20, 1976, pp 43-74 13 Purnell, P., Interpretation of Climatic Temperature Variations for Accelerated Ageing Models, Journal of Material Science, 39,2004, pp 1 13- 1 18 14 Purnell, P. Advances in Modelling GRC Durability, Proc. of CTRS2,29/9-1/10 2003, Dresden, pp145-160 15. Cuypers, H., Wastiels, J., Van Itterbeeck, P., De Bolster, E., Orlowsky, J., Raupach, M., Durability of Glass Fibre Reinforced Composites: experimental methods and results, Composites: Part A 37,2006, pp 207-215 16. EP 0 86 1 2 16 B 1, Inorganic Resin Compositions, ‘Their Preparation And Use Thereof

Proc. Int. Symp. "Brittle Matrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

DEBONDING PHENOMENA IN FRP - STRENGTHENED CONCRETE MEMBERS Renata KOTYNIA Department of Concrete Structures, Technical University of Lodz Al. Politechniki 6,90-924 Lodz, Poland e-mail: [email protected]

ABSTRACT Reinforced concrete (RC) members strengthened with fiber reinforced polymer composites (FRF') to enhance its flexural capacity can experience several failure modes due to premature debonding of the composites from the concrete surface. The crucial role of this technique plays the bond characteristic between the FRF' and concrete. The most commonly reported modes of failure are caused by the plate end debonding (ED), critical diagonal crack (CDC) debonding or flexural intermediate crack (IC) debonding. The classical failure modes due to concrete crushing (CC) and plate fiacture (PF) are possible as well. The paper focuses on identifying debonding mechanisms of external FRP in flexurally strengthened RC members. Generic flexural and shear deformation mechanisms and prediction models based on available official guidelines are discussed. Effects of different parameters such as compressive and tensile strength of concrete, axial stiffness, thickness and width of external FRF' plate are analyzed and discussed. The paper presents comparisons of simplified design models adopted by generic codes.

Keywords External strengthening, FRP strip, reinforced concrete member, bond mechanism, debonding INTRODUCTION The bonding of EB FRP (external bonded fiber reinforced polymer) plates to reinforced concrete structures has become very popular method of strengthening in the last several years. This popularity is attributed to the well-known advantages of the FRP composites such as: good corrosion resistance, weight and easy application. The major problem that accompanied the strengthening with the EBR (externally bonded reinforcement) is a stress transfer from concrete into FRP. Reinforced concrete members externally strengthened with fiber reinforced polymer composites fail prematurely due to loss of bond between FRP and concrete surface. The debonding initiates in the bond zone between concrete and EBR. Afterwards, local debonding propagates and the composite action is lost on the whole length of the FRP. Bond failure may occur at different interfaces between concrete and the FRP reinforcement. Due to usually higher tensile and shear strength of the adhesive than the tensile and shear strength of concrete, debonding failure occurs in the concrete. The most common debonding is observed in the concrete cover near the surface or along the embedded steel reinforcement. If there is the insufficient surface preparation during the FRP application, bond failure in the interface between concrete and adhesive or adhesive and FRP

110

Renata KOTYNA

may occur. The debonding failure modes observed in the tests are classified depending on the starting point of the debonding process defined by the high shear and normal interfacial stresses that exceed the tensile concrete strength. Many studies based on a simple shear test have been carried out in order to analyze the debonding behaviour. The aim of this paper is to discuss debonding modes, review and compare the existing bond models between concrete and FRP. DEBONDING MODES

The debonding mechanism means loss of composite action between concrete and FRP. Local debonding initiates when high interfacial shear and normal stresses exceed the concrete strength. Depending on the starting point of the debonding process stress concentration the following debonding modes can be classified into two main categories: the first in the anchorage zone (plate-end interfacial debonding and concrete cover separation) and the second in the flexural-shear or flexural distance (intermediate flexural shear crack-induced interfacial debonding and intermediate flexural crack-induced interfacial debonding). Debonding in the anchorage zone (end debonding - ED) In the anchorage zone two similar debonding failures may be observed. The premature failure due to concrete cover separation (Fig. la) initiates after formation of a crack at the plate end. The crack propagates to the level of the tension reinforcement and then progress along the ordinary steel reinforcement till the separation of the FRP plate end with the detached concrete cover (Fig. lb) [l]. The second type of debonding rarely observed in the anchorage zone refers to as plate end interfacial debonding (Fig. 2a). Debonding plane is located in a thin layer of concrete adjacent to adhesive interface (Fig. 2b) [2].

flexure w i t h shear 7”

LI

flexure

t

Figure 1 Concrete cover separation failure [11.

Figure 2 Plate end interfacial debonding [2]

Debonding phenomena in FRP - strengthened concrete members

111

Intermediate crack-induced debonding (IC) This mode of debonding initiates at a flexural or a mixed flexural-shear crack away from the plate ends. After the forming of a flexural crack in concrete, when the local tensile interfacial stresses between the FRP plate and concrete at the edge of the crack exceed the critical strength value in concrete, debonding initiates at the crack (Fig. 3a). Then debonding propagates towards one of the plate ends (Fig. 3b) [ 11.

u crack propagation

Figure 3 Intermediate crack-induced debonding [ 11 When debonding initiates in flexural-shear region (Fig. 4), the relative vertical displacement between two edges of the crack produces peeling stresses impact on the FRP plate. Three phases of the flexural-shear debonding are shown in Fig. 4a-c. In both cases of the intermediate crack debonding, a critical width of the crack decides about the initiation of the debonding process. b) phase II

a) phase I t

I

c) phaseIII

I

I

Figure 4 Gradual progress of flexural-shear debonding

DEBONDING STRENGTH MODELS Extensive research has been carried out on the bond behaviour of reinforced concrete members strengthened with externally bonded FRP plates in flexure. The results of the numerous tests indicated two major modes of debonding: plate end debonding and intermediate crack induced debonding (discussed earlier). These two different cases have been the bases of various bond models classified as empirical models (directly based on the test data) and fracture mechanics models. Different standards or guidelines dealing with EBR FRP strengthening in flexure provide simple design proposals to prevent debonding failure at the cut-off section and in the intermediate region.

112

Renata KOTYNIA

Plate end debonding models Several existing strength models for plate end debonding may be classified into three categories: shear capacity models, concrete tooth models and interfacial stresses models. Based on the detailed study of the plate debonding models by Smith and Teng [3], it was confirmed that the most safety but overly conservative predictions gives the Oehlers’ model [41. The Oehlers’ model [4] developed on RC beams strengthened with steel plates for general shear - bending regions is given as follows

where h&b,end and Vdb,,end are the plate end bending moment and shear force respectively at debonding and Mdt,,fis the calculated bending moment at the end of the plate given by (1) and vdb,.s is the shear capacity of the RC beam without any shear reinforcement.

where 1.4 - (d/2000)2 1.1. E, and E h are the modulus of elasticity of concrete and FRP respectively, is the cracked second moment of area of the plated section transformed to concrete, A, is the cylinder splitting tensile strength of concrete, t h is the FRP plate thickness, d and b, are the effective depth and the width of the section, respectively, ps= A & d is the tension steel reinforcement ratio, A ’ is the concrete cylinder compressive strength. Based on the Oehlers’ model (for steel plates) Smith and Teng [5] proposed the following simple debonding strength model

where q =1.4 and V, given by (4) or can be calculated based on national codes. The simple Smith and Teng’s model was widely calibrated on a large database of the RC beams strengthened with EB FRP plates. Another new proposal of limit shear force in order to avoid the cover delamination failure at the plate end has been recently presented by So and Harmon [6] given by (6). This model considers both the stiffness of the FRP reinforcement and also the stiffness of the steel reinforcement.

where A/Ef and AsEs are the stifhess of FRP and steel reinforcement, respectively, L, is a distance between the support and the FRP termination point, b and d are the width and the effective depth of the section,& ’ is the cylinder compressive strength of concrete, ps= AJbd is the tension steel reinforcement ratio. Harmon’s bond tests indicated that the adhesive

Debonding phenomena in FRP - strengthened concrete members

113

flexibility does not significantly affect the cover delamination failure load. The ratio of FRP stiffness to steel stiffness affects the failure load but with a nonlinear relationship [6].

Intermediate crack debonding models This model of debonding has been studied by limited number of research. Simple shear tests describe the most common debonding behaviour corresponding that due to flexural cracks. Hence, the simple strength models for FRP plate-concrete joints can be used to predict the intermediate flexural crack debonding failure. Many available studies about the intermediate crack debonding are based on the bond-slip model derived from direct shear bond tests (Fig. 5). In this approach there is a stress concentration near the crack. The interfacial slips occur on both sides of the flexural crack and the total amount of interfacial slip is equal to the width of the flexural crack.

Figure 5 Bond-slip model for the flexural intermediate crack debonding A review of existing models with a division on models that consider or not consider the effective bond length in calculation of the anchored load is provided below.

Empirical models Based on the test results of single double shear tests on the RC specimens externally bonded with FRP sheets, the following empirical expressions of the ultimate bond strength P,, at failure has been proposed by Tanaka [7] Pu

= bp L

(6.13 - In L),

L

[ml

(7)

by Hiroyuki and Wu [8]

and by Madea et al. [9]

P, = bpL, 110.2Eptp L,

=

6.13 0 0.580 InEptp 9

[ml

Eptp [GPa-mm]

(9) (10)

where L, is the effective bond length, Ep [MPa] and tp [mm] are the modulus of elasticity and thickness of the FRP plate, respectively.

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Renata KOTYNIA

Two different simple models proposed by Brosens and van Gemert [ 101 and Adhikary and Mutsuyoshi [ 111 considered the concrete surface tensile strength fClm and the concrete compressive strengthf, ’

Pu= 0.5 bpLA,,

(Brosens’s model)

(1 1)

Pu

(Aldhikary’s model)

(12)

=

bpL (0.25f,&n),

Khalifa at al. [12] used Madea’s model and proposed the inclusion of the concrete compressive strength factor

Izumo’s model [ 131 confirmed influence different type of fibres on the bond strength

P, = (3.8fc’Z~3+15.2)b,LEptp10-3 for carbon sheets

(14)

P, = (3.4fc’Z~3+69)bpLEptp10~3 for aramid sheets

(15)

Fracture mechanics models One of the first bond strength models between steel and concrete based on nonlinear fracture mechanics (NLFM) was investigated by Holzenkiimpfer [ 141. It was modified by Niedermeier [151 as follows [0.78b,,/w

if L 2 L,

m],

where P, is the bond strength Epfp[MPa-mm], L, is the effectivebond length [mm] and Gf is the fracture energy given by L,

=j$

wheref,, is the average surface tensile strength of the concrete determined in the pull-off test [MPa], cf is a constant determined in linear regression analysis using the results of double shear or similar tests, kp is a geometrical factor related to the width of the bonded plate bp [mm] and the width of the concrete member b, [mm]

Debonding phenomena in FRP - strengthened concrete members

115

2-b, / b , 1+bP/400‘ Neubauer and Rostasy [161 modified Holzenkampfer’s bond strength model for the FRP strengthened concrete

where L,

=E

The nonlinear fracture mechanics analysis by Taljsten [17] used to developed the bond strength formula

where E, and tc are the elastic modulus and thickness of the concrete member. Yuan and Wu [18] developed the bond strength between FRP and concrete based on linear elastic fracture mechanics (LEFM) and NLFM given by

Based on the results of different shear tests Yuan and Wu [18] performed the most real linearly ascending and then descending response shown in Fig. 6 . For this shear-slip model the bond strength was given by Yuan at al. [ 191

2.25 - b, / b,

=j12S+b,/400

Renata KOTYNA

116

u = - 1s i n - ' [ 0 . 9 9 / 7 ] ,

4

zma=1.5pwfc,, so =0.0195,L?wfc,,

s f =-2Gf

(29)

' m a

where&*is the concrete splitting tensile strength.

Figure 6 Bond-slip model by Yuan and Wu [ 181 Yang at al.'s model [20] considers the tensile strength of concrete and the constant value of the effective length L,=lOOmm.

where

(1

if L > L,

Teng and Chen [21,221 based their model on the NLFM analysis developed by Yuan and Wu [18].Chen and Teng's model predicts the bond strength and the effective bond length given by

for

l+b, /b,

L L L, (33)

Debonding phenomena in FRP - strengthened concrete members

117

Chen and Teng [21] modified their above expression to the ultimate strength design to the form given by

Teng et al. [23] presented a smeared crack approach for finite element simulation of intermediate crack-induced debonding. A design model, based on interfacial stress distributions defines the limiting FRP strain, +b given by = 0. 171kb(4.32- a)-

fc,

fi

(35)

where pW, a and Gfare given as follows

EXISTING DESIGN MODELS

ACI 440.2R Guidelines In order to avoid the intermediate crack debonding failure of the RC members strengthened with EBR FRP, the ACI 440.2R-02 [24] propose to limit strain in FRP to

where &fi is the FRP design rupture strain and the value of Km is given by &(l-&)

50.90 for E,t, I180,OOO N/mm (38)

10.90 for Eft, > 180,000 N/mm

where Ef and t/ are the tensile modulus and the total thickness of FRP, respectively. The FRP thickness is a total number of plies and the nominal ply thickness, ngI. The ACI 440.2R recommendations indicate that if the stiffness of the laminate increases the strain limitation becomes more severe. It is important to recognize that ACI does not include the effect of existing internal longitudinal or transverse steel, concrete strength, the properties of the adhesive layer bonding the FRP to the concrete or the width of the FRP laminate relative to the concrete width. Fib Bulletin 14 The fib Bulletin 14 [25] takes a design approach recommending a direct use of a shear stressslip relationship to predict the debonding failure. In the fib model the critical bond stress, and slip parameters are detemined from experimental analysis of the FRP system and substrate i

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Renata KOTYNIA

condition. The Bulletin 14 presents three approaches to assessing the potential for debonding modes. Approach 1 - FRP Tensile Force The maximum axial force in the FRP that may be anchored Nfirrrand the corresponding required anchorage length Lb-, are given by

where a i s a reduction factor to account for influence of inclined cracks on bond strength, a = 0.9 typically, a = 1.0 should be taken for beams having sufficient internal or external shear reinforcement and for slabs, k, is a factor accounting for concrete compaction, k, = 1.0 for FRP bonded to concrete faces cast against formwork, k, = 0.67 for FRP bonded to concrete faces not cast against formwork, b is the width of a beam soffit,f,, is the tensile strength of concrete, CI and c2 are empirical factors determined for CFRP to be 0.64 and 2.0, respectively.

21.0,

[b, b, inmm]

(41)

where b, is the FRP width. The maximum axial force in the FRP and the debonding FRP strain q d b are given by N , = EfEft,bf

(42)

Approach 2 - FRP bond stress The second fib approach involves determining the critical increase in tensile stress in the bonded FRP, transferred by bond stress, between adjacent concrete flexural cracks. This model requires the determination of a critical crack pattern and the corresponding bond stresses transferred to the FRP. This aspect of analysis is beyond the scope of the present discussion. However, the maximum stress a m and strain q d b that may be transferred are given by

L,,,

=..dm

Eft f em

9

[ml

E

wheref, ’ is the compressive strength of concrete, CI = 0.23 and c2 = 1.44 for CFRP.

(45)

Debonding phenomena in FRP - strengthened concrete members

119

In both fib approaches 1 and 2, the FRP capacity is reduced if the available bonded development length, L b < &,a. In cases were L b is less than L b m a (Fig. 7), the FRP capacity cha and the FRP strain limit E / b are reduced by the following factor

Lbmar

Lbmm

Figure 7 Anchorable tensile stress related to anchoring length [25] Approach 3 - Concrete bond strength The third fib approach comprises two steps. The first step involves verification of the end anchorage as in Approach 1. The second step involves verifying that the substrate concrete can transfer the expected shear stress developed across the FRP-concrete interface. The main assumption of this approach is that if the shear stress is maintained below the concrete bond shear strength, flexural cracks will not lead to debonding. JSCE Recommendations The Japanese Society of Civil Engineers Recommendations for Upgrading of Concrete Structures with use of Continuous Fiber Sheets [26] notes that the important contribution of the interfacial fracture energy between the bonded FRP and substrate concrete in determining the maximum stress and the FRP strain, prior to debonding are given by

where Gf is recommended to be taken as OSN/mm in the absence of experimental test data. Reported values of total interfacial fracture energy for CFRP strips bonded to the clean concrete substrate ranging from 0.44 to 0.55 N/mm. Concrete Society TR55 The Concrete Society Technical Report 55 [27] takes essentially the same approach to avoid FRP debonding as it is in the fib approach 1. The tensile bond capacity and corresponding FRP debonding train are given by Nf-

= 0 . 5 k b b f d m 9

where k b term is given by Equation (41).

&fdb =O.”b,/%

fc,

(48)

Renata KOTYNlA

120

Comparison of design FRP axial stiffness-strain relationship for design models and two Teng’s models [21, 231 is shown in Fig. 8. The following assumptions were taken for analysis: bp = 50mm, b, = 150mm,f,’= 40MPa,f,, = 2.5MPa, and G,= 0.5 N/mm. The ACI [24] and Teng’s [23] models give similar limitation of FRP debonding strain, particularly for the FRP axial stiffness E/f between 200 and 300 kN/mm. The Concrete Society TR55 [27] recommendation is similar to ISCE proposal [26] and slightly more conservative than that of Bulletin 14 [25]. For mitigating of flexural intermediate crack induced debonding the ACI recommendation and Teng et al.’s model [21] are generally non-conservative while they are compared with general fib [25] and Report TR55 [27] recommendations to avoid the intermediate crack debonding (with FRP limit debonding strain 6,0%05 @b 5 8,5%0).

I

50000 I00000 150000 200000 250000 300000 350000 400001 Ma1 stiffness E i t t [Nmml

Figure 8 Comparison of design models recommendation for predicting FRP debonding strain From analyzed design models only fib Approach 1, Report TR55 and both Teng’s models consider geometrical factor related to the width of the bonded plate bp and the width of the bonded member b,. This parameter has a big influence on the debonding strength confinned in the Teng’s model shown in Fig. 9. Only the ACT recommendation does not consider the concrete strength, but this effect on the debonding strength is rather small (see Fig. 9). 14

1

I

I

I

I

I

I

....- .. ,f

I

= 2.5MPa

50000 I00000 150000 200000 250000 300000 350000 4000C Axial stiffness E,?, [Nrnm]

Figure 9 Effect of factor

and concrete tension

on debonding FRP strain

Debonding phenomena in FRP - strengthened concrete members

121

CONCLUSIONS Based on the test results of the reinforced concrete members strengthened in flexure with externally bonded FRP plates two major modes of debonding were clearly observed: plate end debonding and intermediate crack induced debonding. In order to predict the debonding failure various bond models classified as empirical models (directly based on the shear test data) and fracture mechanical models have been proposed. Existing guidelines dealing with EBR FRP strengthening in flexure provide simple design proposals to prevent debonding failure at the cut-off section and in the intermediate region. Comparison of design FRP axial stiffness-strain relationships for design models indicated that the ACI recommendation and Teng’s model give similar limitation of FRP debonding strain, particularly for the FRP axial stiffness E j f between 200 and 300 kN/mm. These proposals correspond to intermediate crack debonding failure. Other recommendations are highly more conservative but they refer to the plate end debonding. For strengthening RC members with EBR FRP plates these two critical sections should be taken into account. Teng’s et al. model [21] is recommended for the intermediate debonding strength, while So and Harmon’s model [6] is recommended for the plate end debonding strength.

REFERENCES 1. Kotynia, R., Ductility and Load Capacity of Reinforced Concrete Members Strengthened with CFRP Strips. Ph.D. Dissertation Department of Civil Engineering, Architecture and Environmental Engineering, University of Lo&, Poland, Lodz 1999, (in Polish), pp 215 2. Kotynia, R., Kaminska, M.E., Ductility and failure mode of RC beams strengthened for flexure with CFRP. Report No. 13, Technical University of Lo&, 2003, pp 5 1 3. Smith, S.T., Teng, J.G., FRP-strengthened RC beams. I: review of debonding strength models. J. of Engineering Structures, 24(4), 2001, pp 385-395 4. Oehlers, D.J., Reinforced concrete beams with plates glued to their soffits. J. of Structural Engineering, ASCE, 118(8), 1992, pp 2023-2038 5. Smith, S.T, Teng, J.G., FRP-strengthened RC structures. II: assessment of debonding strength models. J. of Engineering Structures, 24(4), 2002, pp 397 4 1 7 6. So, M., Harmon, T.G., Cover delamination of R/C members with surface mounted FRP reinforcement. In review of ACI Structural Journal 7. Tanaka, T., Shear resisting mechanism of reinforced concrete beams with CFS as shear reinforcement. Graduation thesis, Hokkaido University, Japan, 1996 8. Hiroyuki, Y . , Wu, Z., Analysis of debonding fracture properties of CFS strengthened member subject to tension. Non-Metallic (FRP) Reinforcement for Concrete Structures., Proc., 3rd Int. Symp., Japan Concrete Institute, Sapporo, 1, 1997, pp 287-294 9. Maeda, T., Asano, Y., Sato, Y., Ueda, T., Kakuta, Y., A study on bond mechanism of carbon fiber sheet. Non-Metallic (FRP) Reinforcement for Concrete Structures., Proc., 3rdInt. Symp., Japan Concrete Institute, Sapporo, 1, 1997, pp 279-285. 10. Brosens, K., van Gemert, D., Anchoring stresses between concrete and carbon fiber reinforced laminates. Non-Metallic (FRP) Reinforcement for Concrete Structures., Proc., 3rd Int. Symp., Japan Concrete Institute, Sapporo, 1, 1997, pp 271-278 11. Adhikary, B. B., Mutsuyoshi, H., Study on the bond between concrete and externally bonded CFRP sheet. Proc., 6th Int. Symp. on Fiber Reinforced Polymer Reinforcement for Concrete Structures (FRPRCS-5), 1,2001,371-378

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12. Khalifa, A., Gold, W. J., Nanni, A., Aziz, A., Contribution of externally bonded FRP to shear capacity of RC flexural members. J. of Composites for Construction, ASCE, 2(4), 1998, pp 195-203 13. JCI., Technical report of technical committee on retrofit technology. Proc., Int. Symp. on Latest Achievement of Technology and Research on Retrofitting Concrete Structures, Japan, 2003, CI 14. Holzenktimpfer, O., Ingenieurmodelle des verbundes geklebter bewehrung fiir betonbauteile. Dissertation, TU Braunschweig (in German), 1994 15. Niedermeier, R., Stellungnahme zur Richtlinie fik das Verkleben von Betonbauteilen durch Ankleben von Stahllaschen-Entwurf M'drz 1996. Schreiben 1390 vom 30.10.1996 des Lehrstuhls fiir Massivbau, Technische Universitiit Miinchen, Munich, Germany, 1996 (in German) 16. Neubauer, U., Rostasy, F. S., Design aspects of concrete structures strengthened with externally bonded CFRP plates. Proc., 7th Int. Conf. on Struct. Faults and Repairs, ECS Publications, Edinburgh, Scotland, 2, 1997, pp 109-1 18 17. Taljsten, B., Strengthening of concrete prisms using the plate bonding technique. Int. J. Fract., 82, 1996, pp 253-266 18. Yuan, H., Wu, Z., Interfacial fracture theory in structures strengthened with composite of continuous fiber. Proc., Symp. of China and Japan: Science and Technology of 21" Century, Tokyo, 1999, pp 142-155 19. Yuan, H., Teng, J. G., Seracino, R., Wu, Z., Yao, J., Full-range behavior of FRP-toconcrete bonded joints. J. of Eng. Struct., 26(5), 2004, pp 553-565 20. Yang, Y. X., Yue, Q. R., and Hu, Y. C., Experimental study on bond performance between carbon fiber sheets and concrete. J. Build. Struct., 22(3), 2001, pp 36-42 (in Chinese) 21. Teng, J.G., Smith, S.T., Yao, J., Chen, J.F., Intermediate crack-induced debonding in RC beams and slabs. J. of Construct. Bldg. Mater., 2003;17(&7), pp 447-62 22. Teng J.G., Chen, J.F., Smith, S.T., Lam, L., FRP strengthened RC structures. UK, John Wiley and Sons; 2002,245 pp 23. Teng, J.G., Lu, X.Z., Ye, L.P. Jiang, J.J. Recent research on intermediate crack induced debonding in FRP strengthened beams. Proc., of the 4th Int. Conf. on Advance Composite Materials. for Bridges and Structures, Calgary 2004, (on CD) 24. American Concrete Institute, Guide for the design and construction of externally bonded FRP systems for strengthening concrete structures. ACI 440.2R-02, MI, USA.24., 2002. 25. Externally Bonded FRP Reinforcement for RC Structures. Technical Report,$b Bulletin no 14, Lusanne, Switzerland, 2001,130 pp 26. JSCE, Recommendations for the upgrading of concrete structures with use of continuous fiber sheets. J. of Concrete Engineering, Series 41, Japanese Society of Civil Engineers, Tokyo, 200 1,250 pp. (available in English on CD) 27. Design guidance for strengthening concrete structures using fiber composite materials. Technical Report no 55, Concrete Society, London, 2000,70 pp

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ.. Warsaw 2006

FRACTURE TOUGHNESS VARIATION OF PRESTRESSING STEELS BY BICARBONATE SOLUTIONS

J. SANCHEZ, J. FULLEA and C. ANDRADE Eduardo Torroja of Construction Science Institute Serrano Galvache, 4,28033 Madrid, Spain e-mail: [email protected] ABSTRACT The stress corrosion cracking process developing in metals is at present an unknown mechanism of deterioration. It is a surface process that implies a corrosion and stress synergy, but the most practical consequence is that stress corrosion cracking can modify the mechanical characteristics of the metal. Due to it leads into brittle faliures, it generally involves high level of uncertainty in the prediction. This research deals with steels for prestressed concrete and has the aim to show that the Fracture Toughness changes when the steel is susceptible to stress corrosion cracking, questioning the idea that the toughness is an intrinsic characteristic of the material. The reduction in the fiacture toughness of prestressing steels when they are in contact with aggressive media, involves that the material, for the same stress level, may reach a fracture having a lower crack size. That means the material becomes less damage tolerant, which implies that it is n e c e s s q to develop techniques able to detect defects of smaller size, as for example, small notch, pits or superficial cracks.

Keywords Stress corrosion cracking, fracture toughness, high strength steel and hydrogen INTRODUCTION Concrete has an alkaline pore solution (PH > 12.6) that guarantees the passivation of steel reinforcement in addition to be a physical barrier against the penetration of environmental aggressives. This protection can be maintained indefinitely until an aggressive element in enough concentration reaches the bar. The most common causes of corrosion are the carbonatacibn of the concrete cover, which produces a reduction of the pH of pore solution, and the penetration of chlorides, which induces pitting corrosion. A particular case of corrosion of the steel embedded in concrete is the Stress Corrosion Cracking (SCC), which can appear in prestressed structures. The SCC is produced by the simultaneous action and synergy of a mechanical tension and a corrosive media. Nucleated at the steel surface, the result is the appearance of microscopic cracks that are penetrating and inducing the brittle failure of the wire, due to a triaxial stress condition. The Fracture Toughness (KIc) is one of the most important parameters in Fracture Mechanics. Prestressed wires present high fracture toughness and, until now, this parameter has been considered as a characteristic of the materials. The fracture toughness is one of the fracture criteria [I]. This parameter is based on the knowledge of the stress ranges and displacement in the surroundings of the crack, that is to say, is based on the Stress Intensity

Javier SANCHEZ,Jod FULLEA, Carmen ANDRADE

124

Factor (KI) (Fig. 1). Therefore, the fracture takes place when the stress intensity factor reaches the condition: KI=KIC. There are a standard to measure the fracture toughness: ASTM E 399-90 (1997): "Standard Test for Method Plane-Strain Fracture Toughness of Metallic Materials". This standard provides details on the geometry of the specimens (Single Edge, Compact, Arc, etc.) and the minimum thickness based on the fracture toughness of the material and its elastic limit. This indicates that the fracture toughness varies with the thickness, decreases as increases the thickness until reaching a constant value from a big thickness [l]. In addition, the fracture toughness depends on the rate of the test and the temperature. Some authors [2] have shown the effect of the fatigue in the top of the crack. The cycles of load can produce the plasticity of the crack, which influences as well in the behaviour of the material.

X

Fig. 1. Stress intensity factor. The present work shows that the Fracture Toughness (KIc) of steel varies when it remains in the media susceptible to the corrosion. That is to say, during the process of Stress Corrosion Cracking (SCC) the fracture toughness diminishes. Until now, the fracture toughness has been considered like a constant of the material [5]. The reduction in the fracture toughness implies that the material, for a same tensional level, fractures with a defect much smaller. That is to say, the material becomes less tolerant to the damage, which implies that it is necessary to detect defects, like for example, small notches, superficial pits or cracks. In order to support this statement it is shown some stress corrosion cracking results of high strength steel in carbonated solutions. In these tests, instead of generating the crack by fatigue, it is generated by means of controlled electrochemical and mechanical conditions. After that, it is possible to estimate the fracture toughness in a simple test. The obtained results show decreases around 3040% of the fracture toughness with respect to the fatigue method value.

MATERIAL The material used in this study is a steel of eutectoide composition named parent pearlitic steel, whose composition is given in the following table.

125

Fracture toughness variation of prestressing steels by bicarbonate solutions

Table 1. Chemical composition of parent pearlitic steel (YO,.,). Cr Ni Mn P C Si 10.02 0.20 0.074 0.2 0.7 0.8

S 50.03

Parent pearlitic steel has treated thermally to a temperature about 250 "C during 15 minutes [3]. The purpose of this treatment is to increase the yield strength from 950 MPa of the raw material to 1273 MPa. The value of the fracture toughness for this material is K ~ c= 58 MPa m0.5[3]. The samples were mechanized to a diameter of 2.5 mm and a length of 13.2 mm. In this case it is not possible to obtain standardized geometries, and then it is necessary to test cylindrical samples. They have been prepared as shown in Fig. 2.

8 Fig. 2. Tested bar specimen (in mm). The mechanical properties of parent pearlitic steel are: Young module: E = 201 GPa Yield strength: oY= 1273 MPa Maximun stress: omax = 1870 MPa Strain: E = 13.1 %

TESTS A set of tests were carried out to localize the generation of single pit and avoid depassivation in the rest of the surface. After several trials, epoxi coating was used in order to avoid depassivation by generation of various pits. A notch artificially made to leave the steel surface in contact with the solution was used to reproduce a single pit. Actually, the more realistic conditions are based in the generation of a crack by electrochemical dissolution from a pit [4], which may represent better the reality than to generate the crack by fatigue. In the test method the mechanical and electrochemical parameters are combined and it is made up of the following stages: 1. Fixed potential test in the media: The specimen is immersed in a solution of sodium bicarbonate at constant temperature. A fixed potential is applied, during around 100 hours, simultaneously a data logger registers the current. The specimen is strained to 80% of its yield strength. The objective of this stage is to generate an anodic zone and control the crack growth. After this stage, the specimens are removed from the solution and dried.

126

Javier SANCHEZ, Jose FULLEA, Carmen ANDRADE

2. Slow strain rate test in air: SSRT is performed in air at a rate of 3*10-' s-' in order to determine the fracture toughness. It is possible to obtain the fracture toughness using the fracture mechanics calculations from the stress and crack size data.

3. Scanning Electron Microscope analysis (SEM): In order to examine the fracture surface is used a scanning electron microscope. From this fractographic analysis is possible to evaluate the size of the crack in the fracture surface and the existence or not of brittle zones. In addition it is possible to determine the reduction of area, the different zones of surface of fracture and formed oxides.

RESULTS

The Fig. 3 shows two different behaviours. In the representation on the lefi, it is possible to see an example of a test for a material without defects. In the right part it is shown the curve corresponding to a material that has a crack generated in bicarbonate solution. The fracture for first is completely ductile (Fig. 4) with the formation of micro-voids, whereas for the second case is completely brittle (Fig. 5).

Fig. 4. Ductile surface of fracture. The surface of fracture of one of the wires is shown in Fig. 5 . This type of fracture is characterized by a small area reduction and the fracture takes place in the same plane of the

Fracture toughness variation of prestressing steels by bicarbonate solutions

127

crack. In the image on the right, it can be observed the top of the crack and the mechanical fracture. This mechanical fracture is characterized by the appearance of brittle zones called clivage.

Table 1 shows the crack values reached in different stress corrosion cracking tests for 2,5 rnm diameter bars and their corresponding critical load in the tensile strain test.

Table 1. Size of the crack (a) and critical load (Pc). a (mm)

pc (W

0.60 0.76 0.88 0.92 0.98 1 .oo 1.08

7.152 7.256 6.922 7.000 7.1 16 7.213 6.699

DISCUSSION Due to limited size of the samples, prestressed steels cannot be prepared to obtain standardized specimens for testing fracture toughness of the material (ASTM E399-78) and therefore other approaches are necessary. For the case of a cylindrical geometry of the material, the calculation of the stress intensity factor and the criterion of fracture have been proposed by Elices, Astiz, and Valiente, A. [ 5 ] . The above mentioned authors have assumed that cracks along the whole perimeter of the specimen are formed and the superficial cracks have semi-ellipse shape (Fig. 6). In equation 1 is given the expression corresponding to the stress intensity factor for a superficial crack with semi-elliptical form.

Javier SANCHEZ,Jose FULLEA, Carmen ANDRADE

Fig. 6. Superficial crack in a wire.

Where:

KI is the stress intensity factor. u is the stress. a, b the semi axes of the elliptical crack. R is the radius. Cij are constants, see Table 2.

Table 2. Valor de 10s coeficientes C,,. Value of the coefficients

I i=O i=2 i=3 i=4

I

j=O 1.118 1.405 3.891 8.328

j=1 -0.179 5.902 -20.370 21.895

j=2 -0.339 -9.057 23.217 -36.992

j=3 0.130 3.032 -7.555 12.676

Table 3 shows the values of fracture toughness for the conditions defined in Table 1. Table 3. Fracture toughness of stress corrosion cracking specimens. a KIC(MPa mO.5 0.6 46.1 0.76 39.8 0.88 34.9 0.92 38.3 0.98 38.0 1 40.3 1.08 40.6 The average value of the fracture toughness for a crack generated by stress corrosion cracking is 39.7 MPa mO,’with a standard deviation of 3.4 MPa mO”. The tests made with specimens cracking by fatigue in air present a fracture toughness value of 58 MPa ma.’ [3], which implies a reduction of 3 1.5% in the fiacture toughness.

Fracture toughness variation of prestressing steels by bicarbonate solutions

129

CONCLUSIONS The principal conclusion of this research is the measurement of large reduction of the fracture toughness when the material is immersed in media inducing to stress corrosion cracking. Although these tests should be extended to other media in order to know how much this conclusion regarding prestressing steels can be generalized, it indicated a possible need to review the damage tolerance of prestressed structures in some contaminated atmospheres.

ACKNOWLEDGMENTS The authors wish to thank Ministerio de Foment0 of Spain for the funding the accomplishment of the project “Not destructive methods and strategies for the control of the corrosion in pretested steels”, to the Department of Science and Technology (MAP200303912), to the CSIC for the scholarship of investigation 13P and, specially, to Prof. Gustavo Guinea (UPM).

BIBLIOGRAPHY 1. Elices, M. (1996) “Mecanica de la Fractura. ” Escuela TBcnica Superior de Ingenieros de Caminos (UPM). Madrid. 2. J. Toribio and V. Kharin (2001) “Localizedplasticity near a crack tip in a strain hardening material subjected to mode Z loading“ Materials Science and Engineering A, Volumes 3 19321, Pag. 535-539. 3. Caballero, L., Fullea, J., Alonso, M. C. and Andrade, C. (2002) “EnvironmentallyAssisted Cracking of Pearlitic Steels in Simulated Carbonated Concrete Pore Solutions” 15” Int. Corrosion Congress, Granada. 4. Shchez, J., Fullea, J., Alonso, C. and Andrade, C. (2004) “Estudio de un Nuevo Mktodo de Fisuracibn por Via Electroquimica de Aceros de Alta Resistencia” Anales de Mecanica de la Fractura, Vol. 21, pp. 175-180. 5. Valiente, A. and Elices, M. (1998) “Premature Failure of Prestressed Steel Bars” Engineering Failure Analysis, Vol. 5, no 3, pp. 219-227.

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

SENSITWITY OF RAPID-SETTING SELF-CONSOLIDATINGCONCRETE (RSSCC) TO MIXTURE PRODUCTION VARIABLES Yogini Deshpande' and Jan Olek' School Of Civil Engineering, h d u e University, 550 Stadium Mall Drive West Lafayette, In 47907 [email protected], *[email protected]

'

ABSTRACT When dealing with the issue of repair of the infrastructure, especially bridge deck and concrete pavements, the desire to minimize the traffic delays and inconvenience to the traveling public often leads to the use of rapid hardening repair materials. Frequently, the repairs need to be performed in confined spaces where repair materials are placed around the existing or newly installed reinforcement. As a result, it is very desirable for the repair material to have high fluidity that can ensure good compaction and facilitate flow to tight spaces, preferably without the use of a vibrator. Also, typically such repair concretes are prepared in small (-25-30 L) batches using low-capacity mortar mixers. The existing literature on self-consolidating concrete (SCC) clearly indicates that its stability, in terms of flowability and segregation resistance, can be significantly influenced by the quantities as well as by physical and chemical properties of the component materials. This paper presents the results of laboratory investigation on the sensitivity of rapid-setting self-consolidating concrete (RSSCC) to material and production variables that included: aggregate gradation, aggregate moisture content and the type of the mixer. The maximum size of the aggregate used in production of RSSCC in this study was 9 mm and all mixtures were prepared using Type I11 Portland cement, silica h e , micro-fine fly ash, high-range water reducer, and an accelerator. Two types of mixers were used in this study: a 56 L-capacity rotary pan mixer and 26 L-capacity mortar paddle mixer. All mixtures were prepared using the same general proportions but the "as-mixed" aggregate moisture condition varied fiom dry (0% moisture) to twice the saturated surface dry (SSD) value. The aggregate gradation was also varied by using aggregates with different fineness modulus. It was observed that variation in aggregate moisture content and aggregate gradation resulted in noticeable changes in fresh concrete properties such as the slump flow, stability and V-funnel flow values. While changes in moisture content and gradation of aggregates had an impact on the early (6 h) compressive strength, the compressive strength at the end of 24 hours was not significantly affected.

Keywords Self-consolidating concrete, concrete repair, rapid-setting self-consolidating concrete, rehabilitation, mixing action, moisture content of aggregates, gradation of aggregates

INTRODUCTION Excessive service loads, reduced durability due to low quality of construction materials, improper construction practices and severe environmental conditions often lead to damage

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Yogini DESHPANDE, Jan OLEK

and deterioration of concrete structures. Damaged structures are typically repaired by replacing all, or part, of the deteriorated elements. The performance of the repair materials, especially cementitious based materials, depends upon factors such as: flowability, rate of strength gain, impermeability to water and chloride ions, shrinkage cracking resistance, freeze-thaw resistance and bond between old and new concrete [l]. Literature review suggests that self-consolidating concrete (SCC) is often more sensitive to any deviation from the target recipe (or from the mixing technique) than ordinary concrete. Some of the recent studies indicate that the main factors influencing the robustness of SCC include [2-51: a. type of mixing equipment b. total water content in the mixture as well as the amount of the free moisture supplied by the aggregate c. variations (within the specified limits) of aggregate grading curve Mixing intensity affects the viscosity of the SCC mixture and the amount of high range water reducer (HRWR) required for the same water to powder volume ratio [3, 61. Takada et al. [6] carried out a laboratory investigation to study the effect of mixer type on fresh properties of SCC using forced pan mixer (with four rotating paddles and two fixed paddles) as well as tilting drum mixer. They have related the requirement for low dosage of HRWR observed for tilting drum mixer and the high viscosity of the resulting mixtures to low mixing efficiency of this type of mixer and to its effects on the dispersion of the powder particles. While reviewing the study by Takada et al. [6], Emborg [4] commented that the fact that the dosage of HRWR is influenced by the mixer type has been well documented. He added, however, that an observation that a lower dosage of HRWR is required for drum mixer is new and warrants fiuther investigation. The same author also found that changes in aggregate gradation and moisture content had significant effect on the T50 flow properties and the L-box blocking ratio [4]. The coarser gradation resulted in lower TSOflow time values and higher blocking ratios as compared to the finer gradation. The natural moisture content of aggregate affects the mixing water content in two ways: a. if the natural moisture content of the aggregate is higher than that required for SSD condition, the amount of mixing water (or the free water) is reduced b. if the natural moisture content of the aggregate is lower than that required for SSD condition, the amount of mixing water is increased. In a study by Mori et al. [7] mixtures with 74 different types of aggregates and varying water absorption values were prepared. The authors concluded that the slump flow value tends to prominently decrease with an increase in natural moisture content of fine aggregate for mixtures with 0.35 water-cement ratio as opposed to 0.5 water-cement ratio. Higuchi [8] studied the effects of surface moisture of aggregates on concrete properties and the electric power consumed by the mixer. He observed that the 0-funnel time increased with an increase in the surface moisture content of sand. In addition, the viscosity of the mixtures also increased as did the electrical power consumption of the mixer. The mixer’s power consumption data were used by Nishizaki et al. [5] to adjust the composition of SCC which varied due to fluctuations in the moisture content of the fine aggregate. Power consumption data of every batch was collected and the SCC properties were controlled by adding water in the amount depending on the power consumption values. Since RSSCC would need to be prepared on site, it is essential to study closely the sensitivity of RSSCC to the above factors. The main objective of the present paper was to

Sensitivity of rapid-setting self-consolidatingconcrete (RSSCC) to mixture production variables

133

evaluate the sensitivity of RSSCC to mixture production variables including the effect of mixing equipment, the effect of variation in moisture condition of aggregates and the effect of variation in aggregate gradation. EXPERIMENTAL PROCEDURE MateriaIs The cementitious materials used in this study included Type I11 portland cement with fineness of 6210 cm2/g and tricalcuim aluminate (C3A) content of 10%. Densified silica fume (SF) with a specific gravity of 2.2 and micro-fine fly ash (MFA) with a specific gravity of 2.57 were also used as a part of the binder system. The chemical admixtures used to prepare the RSSCC included: polycarboxylate-based high range water reducer (HRWR), air entraining agent (AEA) conforming to ASTM C 260 [9] and non-chloride accelerator conforming to ASTM C 494 Type C [lo]. Three different sources of fine aggregate (Sand-1, Sand-2 and Sand-3) as well as two different sources of pea gravel with four different gradations (PG-1, PG-2, PG-3 and PG-4) were used to prepare the RSSCC mixtures. The selection of a particular aggregate source was a function of mixture design variables as described in the next section (Experimental Variables) of the paper. The specific gravities of Sand-1, Sand-2 and Sand-3 were 2.63, 2.70 and 2.65 respectively. The gradation curves of these sands are shown in Figure 1. Sand-I was the coarsest of the three sands with the fineness modulus (FM) of 4.14. Sand-3 was the finest with FM of 3.70. The FM of Sand-2 was 3.87. The water absorption values of the sands were 1.8, 1.85 and 1.5% for Sand-1, Sand-2 and Sand-3, respectively. The gradation curves of all the three sands fit between the upper and lower gradation limits given in the Indiana Department of Transportation (INDOT) Standard Specifications for # 23 sand [ 1I].

100

80

20

0 0

0

1

Sieve Size (mm)

10

100

Figure I : Gradation of various sands used in the study. The maximum diameter (Dmax)of all pea gravel aggregates was 9.5 mm. The specific gravities of PG-1, PG-3, PG-4 aggregates were all 2.70 as PG-3 and PG-4 were derived from PG-I source by changing the gradation to obtain either coarser blend (PG-4) or a finer blend

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Yogini DESHPANDE, Jan OLEK

(PG-3) as shown in Figure 2. The specific gravity of PG-2 aggregate was 2.68. The FM of the coarse aggregates was 5.45, 5.50, 5.67 and 6.05, respectively for PG-1, PG-2, PG-3 and PG-4 gradations. The water absorption of the pea gravels was 2.43 and 1.91% for PG-1 and PG-2 respectively whereas for PG-3 and PG-4 it was 2.51 and 2.64% respectively. 100

80 bo

.9 60 v1

*-PG-2 (FM-5.67)

s

I%

s 40 20

0 1

0.1

10

100

Sieve Size (mm) Figure 2: Gradation of different pea gravels used in the study. Experimental Variables To evaluate the influence of the variation in moisture content, the type of mixing equipment and the aggregate gradation on the properties of RSSCC two groups of mixtures (Group I and Group 11) were investigated as shown in Figure 3. A total of 18 different mixtures were prepared and tested during the study (in reality 19 mixtures were prepared but one of them (mortar mixer Sand-1 and PG-1) was common to both groups.

+ Research Variables

Variation in Moisture Content

Group I1 Variation in Gradation of

k Mortar Mixer

Figure 3: Schematic of experimental variables.

Sensitivity of rapid-setting self-consolidating concrete (RSSCC)to mixture production variables

135

Group I mixtures were prepared in two types of mixers: mortar mixer and pan mixer. For this group of mixtures, two water-cementitious material ratios (0.3 1 and 0.36) were used in the mortar mixer whereas one w/cm value (0.31) was used in the pan mixer. All Group I mixtures were prepared using Sand- 1 and PG- 1. In preparing Group I mixtures in the mortar mixer, the moisture content of both Sand-1 and PG-1 was varied from dry condition (0% moisture) to twice the moisture needed to achieve SSD condition, (2 x SSD), in steps of 0.5 x SSD. As a result, for each of the two wlcm values five different mixtures were prepared with aggregate moisture content of 0%, 0.5 x SSD, 1.0 x SSD, 1.5 X SSD and 2.0 x SSD, respectively. When preparing Group I mixtures in the pan mixer the aggregate moisture content used was O%, 1.O x SSD and 2.0 x SSD, thus resulting in three different mixtures. Table 1 gives the mixture proportions for Group I concretes. These proportions were developed assuming that all aggregates will be in SSD conditions and that the mixtures will have 6.5% of entrained air. The cementitious content in all mixtures was kept constant at 570 kglm’. For mixtures prepared with wlcm = 0.31 the design water content was 176 kg/m3, total volume of aggregate in the mixture was about 57% and the volume of fine aggregate as percentage of total aggregate volume was about 63%. For mixtures prepared with w/cm of 0.36, the design water content was 205 kg/m3, total aggregate volume was 54% and fine aggregate volume as percentage of total aggregate volume was 65%. In this study water to cementitious ratio is defined as the design amount of water divided by the total cementitious content assuming the aggregates to be in SSD condition. TheJi-ee water to cementitious ratio p e e w/cm) is defined as the ratio of actual water added to the mixture (accounting for the moisture condition of the aggregates) to the total cementitious content. When batching the mixtures, the water content of the different chemical admixtures used for preparing RSSCC was subtracted from the total (design amount) of water (assuming the aggregate in SSD condition). As a result, the water quantities given in Table 1 are lower than the design values discussed above. Table 1: Mixture proportions for Group I concretes

Materials

I

Cement Silica fume Micro fine flv ash Pea gravel Sand HRWR Air entraining agent Accelerator Water

I

Quantity for w/cm 0.31 (kg/m3) 485 48.5 36.5 58 1 928 10.5 0.17 43.3 134

I

Quantity for w/cm 0.36 (kg/m3) 485 48.5 36.5 510 923 8.8 0.22 43.3 164

I

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Yogini DESHPANDE, Jan OLEK

Table 2: Combination of aggregates used for Group I1 mixtures

*Note-Mixtureproportions as listed in Table 1for w/cm=0.31

In Group I1 mixtures the same general mixture proportions as those used for Group I mixtures with w/cm of 0.31 (with slight variations due to change in specific gravity of aggregates) were prepared. The mixture proportions adopted are shown in Table 3. All aggregates used in Group I1 mixtures were in SSD condition. As mentioned earlier, the water added as a part of the chemical admixtures was accounted for and subtracted fkom the total (design) water content. Table 3: Mixture proportions for Group I1 mixtures

I

Air entraining agent Accelerator Water

I I

0.17 43.3 134

I

0.17 43.3 134

I

0.17 43.3 134

I

Mixing Methodology The 28 L- capacity mortar mixer (MM) shown in Figure 4a is the type of mixer typically used on repair sites and was also used in this study. The mixing methodology adopted for mortar mixer depended on w/cm value of the mixture as shown below: wkm-0.31- mortar mixer Pea gravel + water required to bring the pea gravel to surface saturated condition (if pea gravel is not in SSD condition)+ mix for 30s+sand + AEA + cement + silica fume + MFA + % remaining water + % HRWR + accelerator (mixer stopped for 135 seconds)+ mix for 4 5 s j % remaining water + % HRWR+ mix for 225 seconds

Sensitivig of rapid-setting self-consolidating concrete (RSSCC) to mixture production variables

137

w/cm-036 - mortar mixer Pea gravel + water required to bring the pea gravel to surface saturated condition (if pea gravel is not in SSD condition) 3 mix for 30s+sand + AEA + cement + silica fume + MFA + ?4remaining water + ?4 HRWR + accelerator (mixer stopped for 135 seconds)+ mix for 45s+ ?4remaining water + ?4 HRWR+ mix for 180 seconds

~~

Figure 4a: Mortar mixer The 56 L-capacity pan mixer (PM) shown in Figure 4b is a counter-current type of mixer that was used in study. The mixing methodology used with this mixer was as shown below: w/cm-0.31- pan mixer Pea gravel + water required to bring the pea gravel to surface dry condition (SSD) (if pea gravel is not in SSD condition)+ mix for 30s+sand + AEA + cement + silica fiune + MFA + ?4 remaining water + !4 HRWR + accelerator (mixer stopped for 135 seconds)+ mix for 45.93 % remaining water + % HRWR+ mix for 330 seconds

In an attempt to get an insight into the effects of aggregate moisture content and gradation on mixing efficiency of a given mixer, variations in electrical current levels during the mixing process were monitored using two different ampprobes. The AC current clamp ampprobe manufactured by Fluke@was used for monitoring the current for mixtures with w/cm of 0.31. The Ohio Systems@probe was used for monitoring the current for mixtures with w/cm of 0.36. The primary difference between these ampprobes was their sensitivity. For the Fluke@probe the sensitivity factor was 100 mV = 20 A and for the Ohio Systems@ probe the sensitivity was 2.5 mV = 20 A. The measured variation in the current drawn by the mixer was converted to the power consumed by the mixer using the relationship below: Power = Voltage x Amperagex 0.85 x 0.86 where: 0.85 = Power Factor, 0.86 =Efficiency Factor Testing of Fresh and Hardened Concrete Properties The fresh concrete properties measured were slump flow, flow time for the concrete patty to flow a distance of 500 mm (T~o),visual stability index (VSI), V - m e 1 flow time and the passing ratio (using L-box test). Using the recommendations of ASTM C 1611 [12] for single operator precision, the acceptable value of slump flow was fixed at 5 25 mm of the value obtained for the mixture

138

Yogini DESHPANDE, Jan OLEK

having aggregates in SSD condition for both w/cm. The time taken by the slump flow patty to flow 500 mm is termed as T50 flow time and is expressed in seconds. For this study the acceptable value of T5o was fixed at & 2 of the value obtained for the mixture having aggregates in SSD condition for both the w/cm. This value is slightly above the repeatability value of 1.18s reported in the European Guidelines for SCC [131. The VSI is an index (as per ASTM C 1611 [ 121) to describe the distribution of coarse aggregate within the concrete mass, distribution of the mortar fraction along the perimeter of the slump flow patty and the bleeding characteristics of the slump flow patty. A VSI of 0, which indicates a stable nonsegregating and non-bleeding concrete, was the target index for this study. The V-funnel test and the L-box test are described in detail in other references [l, 141. As per the European Guidelines for SCC [13], the acceptable variation for repeatability for Vfunnel flow time of 8s is 2s whereas for 15s it is 4.4s. For this study, the average of these two values (t3.1 s) was considered as the acceptable deviation of V-funnel flow time from the value obtained for the mixture having aggregates in SSD condition for both w/cm. The acceptable value for the L-box test was t 0.05 of the value obtained for the mixture having aggregates in SSD condition for both wlcm. Both of these values are within the repeatability ranges reported by the European Guidelines for SCC [131. Rate of strength gain at 6, 8 and 24h was the only hardened concrete property measured in this study. The acceptable value for compressive strength was deviation of t 2 MPa from the strength obtained for mixture having aggregates in SSD condition. The adopted deviation constituted about 10 % of the ultimate strength at 6 h. RESULTS The effect of variation in aggregate moisture content and mixer type Presented in this section are the test results dealing with the influence of aggregate moisture content and type of mixing equipment on the fresh and hardened properties of Group I mixtures. The results for mixtures prepared in mortar mixer will be presented first, followed by the results of mixtures prepared in the pan mixer. Figure 5 illustrates the variations in slump flow for different aggregate conditions prepared in the mortar mixer at two different water-cementitious ratios. Table 4 provides additional test results for these mixtures including, VSI, L-box passing ratio and the air content. As seen in Figure 5, the slump flow values for w/cm = 0.31 mixtures exhibit variation from 673 mm for 2 x SSD condition to 787 mm for DRY condition of aggregates. For mixture with w/cm of 0.36, the slump flow variation was between 71 1 mm to 787 mm for the different aggregate moisture conditions. These results indicate that the reduction in the amount of actual mixing water added affects the slump flow to a larger extent for lower water to cementitious ratios than for the higher wlcm ratios. As discussed in the section on testing of fresh concrete properties, the acceptable deviation of slump flow value from that at SSD condition was t 25 mm. For w/cm of 0.3 1 the slump flow values for 2 x SSD condition and DRY condition of aggregates did not fall within the stipulated target range and were, respectively, 38 mm above and 76 mm below the SSD value (see Table 4).

Sensitivity of rapid-setting self-consolidating concrete (RSSCC) to mixture production variables

800

E

I ~

750

E v

139

Ow/c=O.36

I

I

I

W w/c=0.31

2xSSD

r-

1

1.5xSSD

SSD

1

0.5xSSD

DRY

Moisture condition of the aggregate Figure 5: Slump flow of mixtures mixed in mortar mixer

Table 4: Fresh concrete properties of Group I mixtures mixed in mortar mixer

2xSSD 1.5 x SSD SSD 0.5xSSD DRY

0.312 0.344 0.360 0.375 0.408

176 195 204 213 231

0 0 0 0 2

0.82 0.83 0.85 0.87 0.88

5.1 5.1 5.3

-5.2

25 0 0 -25 -5 1

Mixture with w/cm of 0.31 and aggregates at 2 x SSD condition was stiff in comparison to mixture with SSD aggregates whereas mixture with aggregates in dry condition had low degree of flowability. For w/cm = 0.36 the slump flow for all the aggregate conditions was within the stipulated target of 736 2 25, except for mixture with DRY aggregates. The VSI of all the mixtures with w/cm of 0.31 and 0.36 was zero except for those mixtures with aggregates in the dry conditions (see Table 4).

Yogini DESHPANDE, Jan OLEK

140

0.27

0.295

-8- T SO

(w/cm-0.3 1) MM

4 T so

(w/cm-036) MM

-A- T

(w/cm-0.31) PM

0.32 0.345 free w /cm

0.31

0.395

0.42

Figure 6: TSOflow time for Group I mixtures Figure 6 shows the Tso flow time values plotted versus thefiee w/cm. The Tso flow time values indicate a trend similar to that observed for the slump flow. The mixtures with 0.31 w/cm and wetter aggregates (i.e. 2 x SSD) have a high Tso flow time value (10.3 s) as compared to the flow time of mixtures in SSD condition (6.0 s). This indicates that the mixture with 2 x SSD aggregates was stiffer as compared to mixture with aggregates in SSD condition. The Tso flow time values do not vary significantly for mixtures with w/cm = 0.36. Figure 7: V-funnel flow time for Group I mixtures

50 40

$

1

-+w/cm-0.3 1 (MM) immediately +w/cm-0.3 1 (MM) (after 20

i

++w/cm-0.36 (MM) 20 ;*

lo

DRY

-1

0.21

r\

TI

0.295

0.32

0.345

0.37

0.395

0.42

free w/cm The V-funnel flow time values for mixtures made at different w/cm and in two different mixers are shown in Figure 7. The V-funnel flow time for mixtures made with w/cm of 0.3 1 and prepared in the mortar mixer was determined either immediately after mixing or 20 minutes after mixing. It can be seen (Figure 7) that the V-funnel flow time for mixture with

Sensitivity of rapid-setting self-consolidatingconcrete (RSSCC) to mixture production variables

141

dry aggregates increases from 9 s (when measured immediately after mixing) to 19 s (when measured 20 minutes later). For the same time interval, the increase is only 2 s (from 18 s to 20 s) for mixtures with aggregates in the SSD condition. The increase in the V-funnel flow value observed for DRY aggregates indicates that the aggregate started absorbing the water from the mixture. Mixtures with w/cm of 0.36 prepared with aggregate at different moisture conditions did not have large variations in the V - h n e l flow time as compared to mixtures with w/cm of 0.3 1 . The L-box passing ratio values (see Table 4) vary from 0.74 to 0.84 for mixtures with w/cm of 0.31 and from 0.82 to 0.88 for mixtures with w/cm of 0.36. For only one of the mixtures (2 x SSD, w/cm = 0.31) was the deviation from the L-box passing ratio value for SSD condition greater than 0.05. 7o

$ 60 h

W(w/~-O.31)6h

O(w/c-0.31)8h

61 (w/c-O.~ 1 ) 24h

-

70

( ~ / ~ - 0 . 36h 6) (wic-0.36) 24h

(w/c-0.36) 8h

”50

5

$40 W

$30 W

.-P 20 v1

g 10 g 0

u

2x SSD

1.5 x SSD 0.5 x DRY SSD SSD Aggregate Condition

0

V

2 x 1.5 x SSD 0.5 x DRY SSD SSD SSD Aggregate Condition

(4 (b) Figure 8: Compressive strength for mixtures (a) w/cm 0.3 1 and (b) w/cm 0.36 The rate of compressive strength development over 24 h for mixtures with w/cm of 0.3 1 and 0.36 is illustrated in Figures 8a, and 8b, respectively. The trend in the rate of strength development is similar for both w/cm. The compressive strength of mixtures with w/cm of 0.31 varies between 19.3 MPa (2 x SSD aggregate condition) to 7.6 MPa (DRY aggregate condition). The 6 h compressive strength of w/cm = 0.36 mixtures was 17.4 MPa. At 24 h all mixtures with w/cm = 0.3 1 had nearly the same compressive strength of 60 MPa while the compressive strength of the mixtures with w/cm = 0.36 showed slight decrease with an increase of the “free” water content in the mixture. Figures 9 and 10 illustrate the power consumption of the mortar mixer obtained during mixing of w/cm = 0.3 1 and 0.36 mixtures, respectively.

142

Yoghi DESHPANDE, Jan OLEK

0.3

0.2 h

First additionof water and HRWR

3

5 B

g

0.1

&

0.0 2

1

0

3

4

5

6

Time (minutes)

Figure 9: Power consumption curves for mixtures with w/cm = 0.3 1 mixed in mortar mixer As explained in the section on mixing methodology, two different ampprobes were used to measure current variations during mixing mixtures. Despite differences in the sensitivity of these probes, the trends in the power curves for both of these mixtures are similar.

90.0

-1

-2

x SSD

h

v1

Final addition of

c) c)

m

B

v

B

B

0

h

1

0

1

2

3

4

5

6

Time (minutes) Figure 10: Power consumption curves for mixtures with w/cm = 0.36 mixed in mortar mixer Comparing the power consumption data for mixtures made with w/cm = 0.31 obtained for varying aggregate moisture conditions (Figure 9), it can be seen that the power consumption is highest for mixture with aggregates in 2 x SSD condition and it is lowest for mixture made with DRY aggregates. All mixtures with w/cm = 0.31 show significant variation in the power consumption after addition of all the water and HRWR has taken place (see Figure 9). The power consumption-time curves obtained during the 3-5 minutes mixing period for mixtures with aggregates in 2 x SSD condition and SSD condition exhibit steeper slope than the same curves for mixtures with aggregates with 0.5 x SSD or 0 % moisture. All curves become relatively flat after about 4 minutes of mixing, indicating that mixture

Sensitivity of rapid-setting self-consolidating concrete (NSCC) to mixture production variables

143

components have been more or less uniformly distributed throughout the volume of the mix and thus implying the end of the mixing process [2]. The main conclusion that can be formed on the basis of these results is that as thefiee water to cementitious ratio decreases from 0.379 to 0.281 the time required for the mixtures to achieve uniform dispersion of components decreases. For mixtures with w/cm= 0.36 (see Figure 10) the time required to achieve uniform mixing is shorter (between 3.15 minutes for DRY aggregate condition to 4.15 minutes for 2 x SSD condition) than that required by mixtures with w/cm = 0.3 1. So far, only the results pertaining to the mortar mixer have been presented. The next section of the paper discusses the results obtained for the mixtures prepared in the pan mixer. The properties of these concretes (w/cm = 0.3 1) are given in Table 5. Table 5: Properties of concrete mixtures (w/cm = 0.3 1) made in pan mixer Aggregate Condition

2xSSD SSD DRY

w/cm

Slump flow (mm)

VSI

0.281 0.3 11 0.379

610 762 750

0 0 1

free

Air content (YO)

2.3 4 3.9

L-box passing ratio 0.65 0.71 0.75

Compressive Strength (MPa) 6h 24 h 15.3 56.8 17.6 60.1 9.2 56.2

As mentioned earlier (section on Experimental Variables) the mixtures mixed in pan mixer were prepared using aggregate with three different moisture conditions: 2 x SSD, SSD and DRY. The slump flow for these mixtures was between 610 mm to 750 mm and the difference in the slump flow value for mixture with 2 x SSD condition from that of mixture with aggregates in SSD condition was very large (152 mm). The VSI was zero for mixtures with aggregates in 2 x SSD and SSD conditions and the mixture with aggregate in DRY condition had the VSI value of 1 . The Tso flow time value of the pan mixtures was higher in comparison to the mixtures prepared in mortar mixer, irrespective of the w/cm (see Figure 6). Similarly, for all three moisture conditions, the V - W e 1 flow time values (see Figure 7) were higher for all mixtures mixed in the pan mixer. The L-box values varied from 0.65 to 0.75 and were lower in comparison to mixtures mixed in the mortar mixer (see Tables 4 and 5). First addition of water and HRWR

0.8 0.7 0.6 v

a

0.5 0.4

I

i\ I I

0.3

+SSD

(0.31)

-DRY (0.31) Final addition of remaining water and HRWR

Figure 11: Power consumption curves for mixtures mixed in pan mixer

144

Yogini DESHPANDE, Jan OLEK

Figure 11 shows the power consumption curves for mixtures mixed in pan mixer. Contrary to what was observed for mortar mixtures (Figure 9), these curves do not show large variations in power consumption values as different ingredients are added to the pan mixer (i.e., at the point when final addition of remaining water and HRWR has taken place at the end of 3.45 minutes). Though the curves imply that a stable state has been achieved after addition of all water and HRWR, in reality the mixture had not achieved homogeneity. When the mixer was stopped after 5 minutes of mixing the presence of undispersed cement particles and clumps was observed. It took almost three minutes of additional mixing time before the homogenous dispersion of all ingredients was observed (compare Figures 9 and 11). This difference between the degree of dispersion achieved in the mortar mixer and pan mixer is most likely due to the differences in the mixing action provided by these two mixers. The pan mixer has a vertical axis of rotation and consists of the rotating pan and rotating blades. The rotating action causes movement of the ingredients, which results in uniform mixing. However, due to a single axis of rotation, the mixture components have only one direction of movement. This results in low dispersion of all cementitious particles and reduced flowability. In the mortar mixer, the mixer has a horizontal drum with a rotating shaft to which two blades are attached. During the mixing action in the mortar mixer the concrete ingredients are subjected to dual actions - shear caused by the rotating blades and tumbling due to the free fall of mixture during turning of the paddles (see Figure 4a). Due to this dual mixing action, the cementitious particles are probably getting more dispersed and, as a result, mixtures exhibit higher flowability.

Effect of variation in aggregate gradation The effect of variation in aggregate gradation was studied for six (Group 11) mixtures. The properties of all mixtures in this group are compared with the properties of mixture prepared with Sand-1, PG-1 in SSD condition. Figure 12 shows the slump flow for Group I1 mixtures. It can be seen that none of the mixtures had a slump flow within the stipulated range of 7 11 25 mm.

Sand-1, Sand-1, Sand-1, Sand-1, PG-1 PG-2 PG-3 PG-4

Sand-2, Sand-3, PG-1 PG-1

Aggregate Gradation

Figure 12: Slump flow for Group I1 mixtures Table 6 lists the fresh concrete properties for Group I1 mixtures, including deviation of slump flow, VSI, L-box passing ratio and air-content values. The slump flow value increased by 64 mm when the PG-1 aggregate (FM4.45) was replaced by PG-2 aggregate (FM4.67). PG-1 and PG-2 aggregates differ in the amount of material passing sieve opening of 4.75

Sensitivity of rapid-setting self-consolidatingconcrete (RSSCC) to mixture production variables

145

mm, with nearly 70% being retained on higher sieve sizes for PG-2 material. The TSOand the V-funnel values are slightly lower for mixture with Sand-1 and PG-2 aggregates in comparison to Sand-1 and PG-1 mixture (see Figure 13). The 6 h compressive strength values for Sand-1 and PG-2 aggregates are slightly lower than the allowable deviation of (18.4 -2 MPa) (see Figure 14). Table 6: Fresh properties of Group I1 mixtures Deviation of

Mixture prepared with Sand-1 and PG-3 aggregates was slightly unstable with a VSI of 1 and slump flow of 51 mm higher than that of mixture containing Sand-1 and PG-1 aggregates. The gradations of PG-1 and PG-3 are comparable up to sieve size 2.36 mm above which PG-1 has finer particles (see Figure 2). These slight variations also reduce the power consumption for mixture containing Sand-1 and PG-3 aggregates as compared to Sand-1 and PG-1 mixture (see Figure 15). 25

I

20

I

'

b

T.

so (s) V-funnel (s)

I

Sand-1, Sand-1, Sand-1, Sand-1, Sand-2, Sand-3, PG-1 PG-2 PG-3 PG-4 PG-1 PG-I

Type of aggregate

Figure 13: T50 and V-hnnel flow time values for Group I1 mixtures The PG-4 aggregate (FM = 6.05) was much coarser than the PG-1 aggregate (FM=5.45) and the use of this aggregate resulted in significant reduction of the slump flow of Sand-1, PG-4 mixture compared to the slump flow of Sand-1, PG-1 mixture (see Figure 12). The coarser mixture also exhibited high T50 and V - W e 1 time values (see Figure 13). The 6 h

146

Yogini DESHPANDE, Jan OLEK

compressive strength of that mixture was very low in comparison with mixture containing Sand- 1, PG- 1 aggregates (see Figure 14). m6h

Sand-1, Sand-1, Sand-1, Sand-1, Sand-2, Sand-3, PG-1 PG-2 PG-3 PG-4 PG-1 PG-1 Aggregate Gradation

Figure 14: Compressive strength at 6 Bnd 24 h for Group I1 mixtures When the gradation of sand was changed by replacing Sand-I (FM4.14) with a finer Sand-2 (FM=3.87), the Sand-2 and PG-1 mixture exhibited tendency to segregate, as indicated by VSI = 2 (see Table 6 ) . This mixture also had lower 6 h compressive strength than the mixture containing Sand-1, PG-1 aggregates (see Figure 14). The power consumption curve of this mixture is comparable to the mixture containing Sand-I, PG-3 aggregates (see Figure 15). Mixture containing Sand-3 and PG-1 aggregates had the highest slump flow (837 mm) in this group of mixtures (see Figure 12). The Tso and V-funnel flow time values were also low (Figure 13) but with VSI = 2 this mixture also exhibited some amount of segregation. 0.3 j I

- A - Sand-I, PG-1

%-Sand-l,PG-2

First addition of water

0

1

2

3 4 Time (minutes)

5

6

Figure 15: Power consumption curves for mixtures Group I1 mixtures.

Sensitiviw of rapid-setting self-consolidatingconcrete (RSSCC) to mixture production variables

147

CONCLUSION 1. Variation in aggregate moisture content and aggregate gradation primarily affects the

fresh properties of RSSCC and compressive strength at 6 h. 2. Presence of excessive surface water on aggregate does not facilitate the flowability of RSSCC. Mixtures made with aggregates in 2 x SSD condition exhibited the least favorable flow properties. 3. An increase in total mixing water added (in cases where dry aggregate was used) results in reduction of flowability within 20 minutes of mixing. 4. Mixtures having w/cm of 0.36 were more robust and less sensitive to variations in aggregate moisture conditions than those made with w/cm of 0.3 1. 5. For mortar mixer, the power consumption curves provided useful information regarding the completeness of the mixing cycle. Prominent deviation in power consumption can be observed for mixtures made with very wet or dry aggregates. 6. Due to differences in the mixing action, mixtures prepared using mortar mixer exhibited more favorable rheological properties than mixtures made in pan mixers.

7. For the same mixture proportions, mixtures prepared in mortar mixers require shorter mixing time to achieve comparable fresh and hardened concrete properties than mixtures mixed in the pan mixer. 8. Reduction in fineness modulus of sand increases the flowability of the mixtures but also increases their tendency to segregate.

REFERENCES 1. Deshpande, Y . S., Development of Rapid-Setting Self-Consolidating Concrete for Infrastructure Repair, 2006, Doctoral Thesis, School of Civil Engineering Purdue University. 2. Chopin, D., De Larrard, L. and Cazacliu, B., Why do HPC and SCC require a longer mixing time?, Cement and Concrete Research, 34,2004, pp 2237-2243. 3. Deshpande, Y. S. and Olek, J., Effect Of Mixing Equipment And Mixing Sequence On Rapid -Setting Self-Consolidating Concrete, in Proceedings of The Second North American Conference on the Design and Use of Self-Consolidating Concrete (SCC) and the Fourth International RILEM Symposium on Self-Compacting Concrete, Shah, S. P., 2005, Hanley Wood Publication, pp 897-904. 4. Emborg, M., Final Report of Task 8.1, BRITE EURAM 2000, Proposal No. OBE96-3801, pp 1-65. 5. Nishizaki, T., Kamada, F., Chikamatsu, R. and Kawashima, H., Application of HighStrength Self-compacting Concrete to Prestressed Concrete Outer Tank for LNG Storage, in Proceedings of the First International Rilem Symposium on Self-compacting Concrete”, Skarendahl, A. and Peterson, O., 1999, RILEM, pp 629-638. 6. Takada, K., Pelova, G. I. and Walraven, J., Influence of Mixing Efficiency on the Mixture Proportion of General Purpose Self-Compacting Concrete in Proceedings of International Symposium on High-Performance and Reactive Powder Cements, Aitcin, P.-C., 1998, University of Sherbrooke, pp 19-39. 7. Mori, H., Tanigawa, Y., Wakabyashi, S. and Yoshikane, T., Effect of Characteristics of Aggregate on Properties of High-Fluidity Concrete, Transactions of the Japan Concrete Institute, 18, 1996, pp 53-60.

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8. Higuchi, M., State of the Art Report on Manufacturing of Self-compacting Concrete, in Proc. Int. Symposium Proceedings of the International Workshop on Self-Compacting Concrete, 1998, Japan Society of Civil Engineering, pp 360-367. 9. ASTM C 260, "Standard Specification for Air-Entraining Admixtures for Concrete", Annual Book of ASTM Standards, American Society for Testing and Materials, Vol. 04.02, 2002, pp 165-167. 10. ASTM C 494, "Standard Specification for Chemical Admixtures for Concrete", American Standards for Testing Materials (ASTM), Annual Book of ASTM Standards, American Society for Testing and Materials, Vol. 04.02,2002, pp 269-277. 11. Standard Specifications, Section 900 Materials Details, Indiana Department of Transportation (INDOT), 2006. 12. ASTM C 1611, I' Standard Test Method for Slump flow of Self-Consolidating Concrete" Annual Book of ASTM Standards, American Society for Testing and Materials, Vol. 04.02, 2005, pp 1-6. 13. The European Guidelines for Testing Fresh Self-compacting Concrete - Specification, Production and Use, The European Federation of Specialist Construction Chemicals and Concrete Systems, 2005, pp 1-68. 14. Deshpande, Y. S. and Olek, J., Development of Rapid-Setting Self-Consolidating Concrete (RSSCC) Using Small Size Aggregate, to be published in the Proceedings of the International Symposium on Advances in Concrete Through Science and Engineering, 2006, RILEM.

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

A DISCUSSION ON THE ESSENTIAL ISSUES FOR THE SUCCESSFUL PRODUCTION OF SELF-COMPACTING CONCRETE (SCC) Syed Ali RIZWAN, Thomas A. BIER and Katja DOMBROWSKI Institute for Ceramics, Glass and Construction Materials Technology Agricolastr. 17, TU Bergakademie Freiberg,09599,Germany Fax: 0049-373 1-39-2223, e-mail: [email protected]

ABSTRACT A number of papers on various aspects of SCC can be found in the literature but most of them are based on the

laboratory investigations using optimized conditions, complex terminologies and procedures. The results so obtained are not easy to comprehend and are unsuitable for engineers hying to make use of this wonderful technology of the decade. A laboratory SCC mix generally gives a different response when used at a readymixed concrete plant. This is due mainly to the differences in mixing regimes, procedures, climatic conditions and aggregate surface moisture estimation. The information provided in this paper is based on the experience gained from laboratory and plant mixes with subsequent field placements 150 meters below the ground level, in a local tunnel of a research and teaching mine, by means of pumping. The total horizontal pumping distance (above and below ground level) was also 150 meters. 50 liter laboratory mixes of combination type of SCC using natural aggregates with different gradings, cement types as well as mineral and chemical admixtures were made and tested followed by the tests on 1.5 or 3.0 m3 ready-mix plant batches with almost similar mixing regimes. A typical field batch consisted of 8 m3 of SCC. This paper addresses areas such as mix design, aggregate grading, system's water demand, flow and strength of SCC formulations. The results suggest that a thorough understanding of all relevant aspects is essential for successful SCC production. A simple procedure to determine the water demand of a typical SCC formulation is also suggested and is contained in this paper.

Keywords SCC, mix design, aggregate grading, system's water demand, superplasticizer, viscosity enhancing agents, flowability, strength and blockage

INTRODUCTION Any systematic study on high performance (HP) self-compacting cementitious systems (SCCS) must start with the pastes and from there reach mortars and concretes. The research work on HP SCP (self-compacting pastes) by the authors has already been published [ 1, 21 and a paper on HP SCM (self-compacting mortar) systems is also included in these proceedings. SCCS do not require any external compaction. The special features, characteristics, definitions and required tests for SCC are well documented [ 3 , 41 and are therefore not described here. The acceptable SCC compositions based on laboratory results may not be directly used for site placements as they need further adjustments at ready-mix concrete plants due to the difference in mixing regimes, aggregate surface moisture conditions, environmental and the SCC batch size differences. Therefore it is advisable to make plant trials before placements [5]. At some of the sites additional considerations due to differences in the environmental parameters may call for fuaher adjustments in SCC formulations. In case of an improperly

Syed Ali RIZWAN, Thomas A. BIER, Katja DOMBROWSKI

150

designed SCC formulation, the pumping process can result in the expulsion of air and the separation of water from other constituents thus causing problems leading to blockage at times. In this reported research significant differences between the existing environmental parameters in the laboratory, at the plant and the site were observed especially during summer months. For example the temperature and relative humidity in the laboratory and at the plant were 27"C-28% and 35"C-30% respectively while in the tunnel underground these were 8°C 85 % respectively. The mixing water temperature in the laboratory and at the plant was different and was not studied. But it is believed that it has a significant role in the flow response of SCC formulations. SCC MIX DESIGN CONCEPTS Various types of SCC are known including powder type, viscosity agent type and combination type differing mainly in the way the segregation resistance is achieved. In the powder type SCC a low water-powder ratio guarantees adequate segregation resistance while the same role is played by viscosity enhancing agent (WA) in the viscosity agent type of SCC. The combination type of SCC allows the production of a robust SCC due to the combination of slightly reduced powder with a VEA. This type is believed to have excellent segregation resistance and was therefore selected for site placements. Several SCC mix design approaches exist with each one being drastically different from the others. However continuous aggregate grading is preferentially used in SCC formulations as it requires less paste to fill the voids and has obvious advantages. SELECTION OF MIX PROPORTIONS

I Reference

I Volume of coarse aggregate w.r.t SCC 0.28-0.36 0.266-0.281

I Reference I Volume of fine

I aggregate y.r.t scc

I

volume (m3/m3> 0.256-0.327 48-5570 of total

0.27-0.36 0.313-0.359

Usually a sand content of about 50 70of the total aggregate mass has been found satisfactory for SCC [lo]. This highlights the importance of material passing 1 mm sieve for the stability of SCC mixes. The powder content and the water-cement ratio can also be selected considering the guidelines and desired SCC properties in the light of strengths given by local cements [5, 71. Suitable coarse aggregate content, fine aggregate content and powder content along with water-cement ratio can therefore be pre-selected along with the percent of other fillers etc. before workability tests can be started. Adjustments are then made till the desired test values are obtained. The other recommendations for the size fractions of aggregates including sand (0-2 mm) content in SCC passing the 1 mm sieve are given in Table 3.

A discussion on the essential issuesfor successful production of self-compacting concrete (SCC)

Reference Generally recommended [31 [ 111 CSA (A23.2-2A)

151

Fraction < lmm in % Sand (0-2 mm) Total aggregate 30-45 70 39 67 40-82 (General)

SHAPE AND GRADING OF AGGREGATES

In general the desirable aggregate shape requirements for both normal concrete and SCC are the same. These state that no more than 15 % of aggregates should be elongated. According to DIN EN 933-4 an aggregate is considered elongated if the ratio of length to the maximum thickness is more than 3. These elongated aggregates can cause increased internal friction, bigger voids and also pipe blockings at times. Such blockages can also be caused by bleeding, high coarse/fine aggregate ratio, and using pipes with different wear [12]. A high elongated aggregate content also needs a higher paste volume for SCC conveyance. . A random sample of 8/16 mm size fraction of aggregates when tested according to DIN EN 933-4 contained about 15 % of elongated material rendering it to be a boundary line aggregate. The importance of shape of coarse aggregate regarding flow and passing ability has also been highlighted in the literature [5]. Fig 1 shows the shape of 8/16 mm size fraction of aggregates.

The calculation of a system’s water demand (WD) is often the first step required in SCC design [5] necessitating a simple procedure for its evaluation. Complicated procedures using sophisticated equipments for determining the WD of system have been reported in the

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152

literature [8, 91. For the production of durable SCC mixtures it is important that the water content of the mixture does not exceed the water demand of the system by a big margin. The total WD of SCC system is the sum of individual water demands of the powder and the aggregate components. The WD of the powder component can be determined by taking cement and other selected fillers in the desired pre-selected proportions by mass and testing with the Vicat needle. For determining the water demand of various size fractions of coarse and fine aggregates, simple procedures outlined in ASTM C 127 and 128 can be followed. The results can then be added to get the system’s WD as shown in Table 4. Example The following amounts of materials were used in a typical SCC formulation. The water demand of 1 m3 of SCC was calculated according to the procedures and ASTM standards cited above. Table 4 Calculation of the water demand for SCC mix

EXPERIMENTAL Materials For this reported work CEM IYA-LL 32.5R (C 11) and a hard coal fly-ash (FAl) were selected for various SCC formulations. Table 5 gives the properties of the powders used.

Table 5 Properties of the powders used Powder CII FA1

Particle Size

tw)

16.90 26.59

BET Area (m2/g) 1.353 1.65

(dcc) 3.11 2.31

Alz03 NazO KzO

SOz

Fez03

MgO

Cao

18.74 51.44

2.23

1.38 2.51

58.9 4.78 4.03 26.13

5.55

1.25 1.23

1.01 7.09 2.63

s; 3.20

The Bogue’s potential parameters of CEM IYA-LL 32.5R are Cz S= 19.56, C3 S= 51.18, C3 A= 10.5 and C4AF= 6.87. Siliceous sand (0-2 mm) and natural gravel (2-8 mm and 8-16 mm fractions) was used for the SCC-mixes. Figure 2 shows the grading of aggregates.

1

153

A discussion on the essential issues for successful production of self-compacting concrete (SCC)

Grading Curve of Sand

Gradmg curves of -gates 1

100

100 80

60 40

//

20

0 0

0,063

0.125

0.25

Fig 2(a) Sand : 0-2 mm

0.5

1

0.063 0.125 0.25

2

Fig

0.5

I

2

4

8

16

2(b)

Contents (0-2:2-8:8-16 mm)-G1 (39:33:28), G2 (45:27:28), G4 (50:25:25), G6 (5:22.5:22.5)

The grading curves of aggregates used in this research project with a maximum aggregate size of 16 mm fall within the limits of German standards [13] DIN EN 206-1 and DIN 1045-2. MIXING PROCEDURE

The mixing sequence for the SCC ingredients both in the laboratory and at the plant consisted Of:

Mixing of dry constituents for 30 seconds Then adding water, superplasticizers and viscosity agent at the same time. Mixing the ingredients for another 90 seconds. The total mixing time was 2 minutes. This may not be the most efficient mixing regime but it had to be adjusted to approximate to the stringent plant mixing procedures which normally do not allow the use of optimal mixing schemes as recommended elsewhere [5]. Because of the low shear rate and small size of the laboratory sample, concrete was kept undisturbed for seven more minutes after two minutes of initial mixing and was thereafter given one minute of final mixing to insure full activation of the superplasticizer before the start of flow measurements. The other possible alternative could have been to mix the sample in one go with a higher mixing time (say up to 4 minutes or so). However the mixing sequence in the laboratory and at plant was kept the same. Two minutes of mixing was done at the plant with one minute of fast truck mixing before starting flow tests there. Same procedure was used before pumping concrete at site. To avoid thixotropic gelling, concrete was kept agitated on route to the site prior to placing. SCC FLOW TESTING

In the fresh state tests including slump spread (cone standing on narrow end), V-funnel time, L-box, J-ring (blocking ring) and air content were carried out in sequence. It took about 20-25 minutes with a three men party. The sample was mixed again for 5 seconds each before starting another test after the slump spread test. Following the plant mixing regime in the laboratory may not give sufficient time for the activation of superplasticizer (SP) because of low shear rates of the mixer and smaller SCC sample volume (usually about 50-70 liters or so). It is therefore suggested that slump test parameters should be measured after the activation of the superplasticizer otherwise their comparison with J-ring values would be inaccurate and unrealistic. In some research papers J-ring spread is shown to be greater than

154

Syed Ali RIZWAN, Thomas A. BIER, Katja DOMBROWSKI

that of slump spread [I41 which can not be easily justified. It would indicate the inadequate activation of SP andor the absence of a time frame to measure slump flow only after its complete stoppage as some cohesive concretes keep on creeping for some time. A final consistence adjustment with water is allowed [5] however experience showed that addition of even a small quantity of water after the chemical admixtures could significantly reduce the mix cohesion andor could yield slightly inaccurate results. It should therefore be avoided as far as possible. Figure 3 shows therelation between V-funnel time and the ratio of SPNEA for two aggregate gradings used with CEM 11. The powder and aggregate content used has been shown in Table 4 while aggregate gradings are given in Fig 2. The target slump spreads for G4 and G2 formulations were 66*1 and 7W1 cm respectively. Grading-Admixtures Response with CEMII 21

17

13 9

5

I

1.4

I

I I

I

1.5

1.6

1.7

I

1.8

I

I

1.9

2 ~

Figure 3 V-funnel time and PIS relation for two aggregate gradings with CEM I1 Although a funnel time of 6-11 seconds is considered essential for SCC, yet literature suggests a time between 10-20 seconds may still be good for practice [15].It appears that for G4 grading a slight incmse in PIS ratio drops the V- W e 1 time considerably. A higher flow target with this grading might have been attained near PIS range of 1.8-2 and the funnel time might have been even lesser than G2 grading. Fig 4 shows relations between slump and Jring spreads of some SCC mixes. It can be seen in Figure 4 that J-ring spreads were lesser than the corresponding slump spreads for all mixes by varying margins due to the elongated particles and the degree of obstruction offered by J-ring though other factors like slightly different plasticizer-stabilizer ratios etc. could also affect this response. A linear regression fit to the data is also presented in figure 4. A difference of 10 mm between slump and J-ring spreads is permitted [5] but the German literature [ 151 allows upto 50 mm and is considered to be more realistic as it seems to take care of rather elongated natural aggregates as well. It should be remembered that any increase in VEA content (for the same SP content) would reduce flow as some of SP is engaged by the VEA. The flow of SCC mixes seems to start at a typical PIS ratio for a given grading and other mix constituents.

155

A discussion on the essential issues for successful production of self-compacting concrete (SCC)

Slump spread and J-ring Spread 75

y = 1.0117~- 3.9436

8

70 65

*

E0 2 L

a

CII M

+ /

C . 1

Y

3

+ GradingG4

60 55

Slump Spread cm 50

62

64

66

68

70

72

74

Figure 4 Relation between slump and J-ring spreads of some of SCC mixes SEGREGATION RESISTANCE

The segregation resistance was visually estimated by longitudinally sawing of cylindrical concrete samples (approx. length: 1 m; diameter: approx. 100 mm). The uniform presence of aggregates of all sizes along the height simply indicated a good segregation resistance and adequate system’s viscosity. Perhaps the same could have been stated had the three equal portions of this cylinder been weighed and compared. COMPARISON OF LABORATORY, PLANT AND SITE RESULTS

The difference between the flow response of similar mixes measured in the laboratory, at the plant and within the tunnel site can be attributed to: Activation time of SP Continuous slow stirring en-route to site and waiting time at site before placement 0 Incorrect estimation of the aggregate surface moisture and its correction Environmental factors Pumping process. The other reasons could be the difference in the laboratory and plant batch sizes or in the shear energy imparted to the sample. It should be remembered that SCC is more sensitive than normal concrete to the variations in physical properties of its constituents and especially to changes in aggregate moisture content [ 5 ] . Pumping generally increases the flow and reduces air content of SCC at the discharge end. This can also be seen in Table 6 .

Syed Ali RIZWAN, Thomas A. BIER, Katja DOMBROWSKl

156

WORKABILITY RETENTION

The transportation time, waiting time and placement time of SCC may be significant requiring adequate slump retention. It also requires good scheduling and planning. Workability loss is small if SCC is agitated continuously. However a loss of around 12 % is generally expected over two hours [lo]. The experience showed good workability retention of the mixes over the total processing time. Literature reports that dispersant structure and the amount of applied shear energy influences the rate of dispersant depletion from the aqueous phase influencing workability retention [ 161. Reported depletion levels are higher in concrete systems compared to the paste systems and the dispersants adsorbed faster usually show lesser retention time. The resulting quantitative differences of similar formulations tested in the laboratory, at the plant or at the placement site are given in Table 6. Table 6 A typical difference in flow parameters as measured in laboratory, plant and tunnel site (mixture: CEMII/A-LL 32.5 R= 380 kg/m3,FA1=147 kg/m3,total aggregate content with G4 was 1625.3 kg/m3.MI and M2 were two dflerent measurements on the same day. LocationMeasurement No. Tests

I

"unnel

I

I

SPNEA Slump: T50 cm in s, spread in mm V-Funnel, time in s

2.511.05 7 690 13.4

2.511.05 3.8 735 6.43

211.3 8.5 650 12.8

211.3 1.9 785 4

M2 211.3 2.5 750 5.3

STRENGTH OF SCC MIXES

Samples were cured in moist air with 90% relative humidity for initial 24 hours. After demoulding those at this age these were put in water for seven days and thereafter were cured at 20°C and 67% relative humidity. The target strength at 28 days was 60 MPa and it was exceeded by all mixes. The margin was higher for the lower water-cement ratios. There was an excellent agreement between actual and fitted curves for G6-mixes with CEM 142.5 R within the water-cement range investigated. The average 28 days compressive strength of the tested formulations was 0.216 MPa, 0.175 MPa and 0.239 MPa per kg of cement for CEM I 42.5 R, CEM I1 /A-LL 32.5 R and CEM IUB-M 32.5 R respectively as against 0.14 reported elsewhere [7] which may be due to the lower cement content. Figure 5 shows 28 day compressive strength results obtained kom 100 mm cubes of some SCC laboratory formulations.

157

A discussion on the essential issues for successful production of self-compacting concrete (SCC)

Actaul and Fitted Strength-Water/Cementratio Response of SCC Mixes

I

0.4 0.45 0.5 0.55 0.6 Fig 5 28 day compressive strength versus w/c-ratio relation of SCC mixes

Table 7 shows the strength results of tunnel castings with two curing conditions. The samples were cast in the tunnel and half of them remained in the tunnel environment till the age of testing while the others were taken to the laboratory and were cured as stated above. The mixes Mix 1 (C II-FAl-P2.5-S1.05-G4) and Mix 2 (C 11- FAI-P2.0-S 1.3- G4) indicate in sequence the cement type, fly ash type, SP and VEA per cent dosages in terms of cement mass and aggregate grading. Table 7 Strength results of some tunnel castings with different curing conditions ~

7 Days strength, Curing Condition Tunnel Laboratory Tunnel

MPa

Flexural strength

I

90 days strength,

28 days strength, MPa

Compressive strength

Flexural strength

54 54 53

7.9 7.5 9.9

I

I

I

I

MPa

Compressive strength

Flexural strength

62

9.3

78

80

7.2 10.1

101

64

I

I

Compressive strength

78

It can be seen that flexural strength of tunnel cured samples was higher than that of laboratory cured ones. It may be due to the higher relative humidity in the tunnel. After observing all the available data it can be said that a little variation in P/S ratio does not change the strength significantly if all the other components remain the same. Moreover an increase of 5 % sand content (of the total aggregate content) does not reduce the SCC strength significantly if the total aggregate content is kept constant. The strength seems to be more influenced by the w/c ratio.

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Syed Ali RIZWAN, Thomas A. BIER, Katja DOMBROWSKI

CONCLUSIONS Considering all the data, the following conclusions are drawn from the study. 1. The determination of SCC system water demand is an essential starting step. It should be determined and compared with the total mixing water content. 2. Variation in the environmental factors and an incorrect estimation of aggregate surface moisture at plants are the main sources responsible for variation in the flow response of similar SCC mixes tested at the plant, site and in the laboratory. 3. For similar formulations lower PIS ratio (plasticizer-stabilizer ratio) slows down the flow times and entraps more air. 4. Other reasons for the different flow response of identical SCC mixes used at these locations can be attributed to the degree of SP activation, mixing water temperature, continuous slow stirring en-route to site, waiting at site and to the pumping process. The difference in batch sizes could also be one of the reasons. 5. A little variation in plasticizer or stabilizer contents does not change the SCC strength significantly if other constituents remain the same. An increase of 5 % sand (of the total aggregate content) does not reduce the SCC 6. strength significantly for a constant total aggregate content. 7. For the same formulation, flow times are reduced with increased cement content keeping the total powder content constant. 8. For a given grading and SCC mix ingredients the flow starts at a typical PIS ratio beyond which any increase does not improve the flow response very significantly. REFERENCES 1. 2. 3. 4. 5. 6. 7. 8.

9.

Rizwan, S.A. and Bier, T.A., Inclusion of Mineral Admixtures in Cement Pastes for High Performance Concrete. CD 7-004 Proc. 2”d International Conference on Concrete & Development, Tehran, Iran, May 2005, pp. 1-12. Rizwan, S.A. and Bier, T.A., Role of Mineral Admixtures in High Performance Cementitious Systems. Proc. of International Conference on “Concrete and Reinforced Concrete-DevelopmentTrends” 5-9 September 2005, Moscow, Russia. Khayat, K.H.; Assad, J. and Daczko, J., Comparison of Field-Oriented Test Methods to Assess Dynamic Stability of Self-Consolidating Concrete. ACI Materials Journal, Vol. 101, No .2, March-April 2004, pp. 168- 176 Assad,J.; Khayat, K.H. and Daczko, J., Evaluation of Static Stability of SelfConsolidating Concrete. ACI Materials Journal, Vol. 101, No. 3, May-June 2004, pp. 207-2 15 The European Guidelines for Self-compacting Concrete, May 2005. EFNARC, www.efnarc.org, pp. 1-63 JSCE.: Guide to Construction of High Flowing Concrete. Gihoudou Pub, Tokyo 1998 (In Japanese) Su, N.; Hsu, K-C. and Chai, H-W., A Simple Mix Design Method for SelfCompacting Concrete. Cement and Concrete Research 3 1 (2001), pp. 1799-1807. Marquardt, I.; Vala, J. and Diederichs, U., Optimization of Self-compacting Concrete Mixes. Proceedings of Second International Symposium on Self-compacting Concrete, Tokyo, 2001, pp. 295-302 Marquardt et al.: Proceedings of First North American Conference on the Design and Use of Self-Consolidating Concrete. ACBM, USA, November 12-13,2002,

A discussion on the essential issues for successful production of self-compacting concrete (SCC)

10. 11. 12. 13. 14. 15. 16.

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Brouwers, H. J. H. and Radix, H. J., Self-compacting Concrete: Theoretical and Experimental Study. Cement and Concrete Research, 35 (2005), pp. 21 16-2136 Ghezal, A. and Khayat, K. H., Optimizing Self-Consolidating Concrete with Lime Stone Filler by Using Statistical Factorial Design Methods. ACI Materials Journal, Vol. 99, No. 3, May-June 2002, pp. 264-272 Kaplan, D; de Larrad, F. and Sedran, T., Avoidance of Blockages in Concrete Pumping Process. ACI Materials Journal, Vol. 102, No.3, May-June 2005, pp. 183191 Readymix Baustoffgmppe: Baustofftechnische Daten. 18 Auflage, pp. 94., www .readymix.de Reinhardt, H. W and Stegmair, M., Influence of Heat Curing on The Pore Structure and Compressive Strength of Self-compacting Concrete (SCC), Cement and Concrete Research 36 2006 879-885. BrameshubeqW.,” Selbstverdichtender Beton”, Schriftenreihe SpezialBeton Band 5, verlag Bau+Technik(in German) ,67pp . Vickers, Jr., T.M., Famngton, S. A., Bury, J. R . and Brower, L. E., Influence of Dispersant Structure and Mixing Speed on Concrete Slump Retention. Cement and Concrete Research, 35 (2005), pp. 1882-1890.

Proc. Int. Symp. Y3rittle Matrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

PRELIMINARY OPTIMIZATION ANALYSIS OF TERNARY MIXTURES FOR BRIDGE DECKS Mateusz RADLINSKI", Jan OLEK'**,Haejin KIM', Tommy NANTUNG3 and Anthony ZANDER4 1 Purdue University, Schopl of Civil Engineering, 550 Stadium ** Mall Drive, West Lafayette, IN 47907, USA, e-mail: [email protected], [email protected] 2 University of Maryland, Civil and Environmental Engineering, 1173 Glenn L. Martin Hall, College Park, MD 20740, USA, email: [email protected] 31ndianaDepartment of Transportation, Research Division, 1205 Montgomery Road, West Lafayette, IN 47906, USA, email: [email protected] 41ndianaDepartment of Transportation, Materials and Test Division, 120 Shortridge Road, Indianapolis, IN 462 16, USA, email: [email protected] ABSTRACT The main objective of this study was to select the optimum ternary mixture design for high performance concrete (HPC) bridge decks. Three key concrete properties (rapid chloride permeability (RCP), free shrinkage and compressive strength) were evaluated for four different mix designs. The cementitious material used in each of these designs consisted of Type I portland cement and various percentages of fly ash (FA) and silica fume (SF). The second objective of this work was to investigate the influence of variability of materials sources on permeability and compressive strength test results of ternary mixes. Accordingly, the laboratory study was conducted in two series, consisting of 20 and 10 mixes each, incorporating different sources of cement, fly ash, and fine and coarse aggregate. Collected test results were subjected to statistical analysis and multiple linear regression models were developed for each of the investigated key properties (responses). As a last step, the multicriteria optimization analysis was performed. The values of response functions for each of the mix designs were back-calculated by using the design binder composition and target air content as input parameters into the prediction models obtained during the previous step. Subsequently, three objective functions were formulated: durability function (permeability and shrinkage), mechanical (compressive strength) and economical (unit cost). Normalized weighted sum method was employed as means of selection of optimum mix design. The major finding from the study was that although the selection of the optimum mix design depended on the values of weights assigned to the corresponding criteria, for any configuration of the weights the optimum mixture always contained 20 rather than 30% of fly ash. The change in material sources altered 56-day RCP results by approximately 30%, whereas the 28-day compressive strength appeared to be insensitive to change in materials source. Keywords Compressive strength, optimization, permeability, regression, shrinkage, ternary mixtures

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Mateusz RADLINSKI, Jan OLEK, Haejin KIM, Tommy NANTUNG, Anthony ZANDER

INTRODUCTION The term “high-performance concrete” (HPC) almost automatically implies incorporation of supplementary cementitious materials, such as fly ash, silica fume or slag [l]. When properly designed, produced, placed and cured, HPC will offer higher durability and therefore an increased service life compared to plain concrete mixtures. HPC is being increasingly utilized in such structures as bridge decks to control shrinkage and permeability and reduce the risk of premature deterioration [2-41. Based on an extensive body of research data available in the literature on concretes containing the pozzolanic materials, there is a general agreement that the best performing mixtures typically use ternary or quaternary cementitious blends rather than just a single mineral admixture [5-81. In particular, a combination of cement, fly ash and silica fume has been suggested to be very promising for bridge decks applications [9]. Although the optimum (and relatively wide) ranges have been established for incorporation of individual pozzolanic materials in binary cementitious mixtures, i.e. 20-40% for fly ash, and 5-10% for silica fume by mass of binder [ 1,lO-121, the issue of optimum proportions of these materials when used in multicomponent system is still unresolved. In order to adequately select the best, i.e. optimum mixture, one has to carry out a complete optimization process which involves selection of experimental variables, constraints, objective functions, and properly assigned weights. Furthermore, the optimization method and the method of selection of the optimum mixture need to be adequately chosen. Several approaches have been proposed, as far as concrete mixture optimization methodology is concerned, including factorial designs [ 13-171, mixture method, response surface method, Tagushi’s method [ 181, genetic algorithm [ 191 and artificial neural networks [20]. Among these response surface methodology (RSM) appears to be the most popular [21-241. This method was developed for relatively large and complex experimental matrices, which result in selective examination (production and testing) of considered mixture designs, followed by development of regression models which approximate the properties of all mixture designs included in the experimental program [25]. Although this is undeniably a great advantage from the standpoint of workload and time savings, it also implies that in order to draw any valuable conclusions about hardened concrete properties, the fresh properties of all mixtures need to be relatively similar. The current study was focused on the selection of the optimum mixture from the relatively narrow range of four ternary (cement + fly ash + silica fume) mixture designs, each of which exhibited different fresh concrete properties, i.e. slump and air content. Moreover, in order to determine the sensitivity of evaluated concrete properties to change in the source of raw materials, i.e. cement, fine and coarse aggregate, the study was conducted using two series of mixtures (Series 1 and Series 2), each consisting of 20 and 10 mixtures, respectively. The detailed description of mixture designs and summary of properties of materials used in Series 1 and 2 are provided in the following section. EXPERIMENTAL PROGRAM

Materials Ordinary ASTM C150 Type I portland cement from two different sources was used. One source of Class C fly ash, meeting requirement of ASTM 618, and one source of dry, densified silica &me was utilized in both series of mixtures. Table 1 contains the chemical composition and physical properties of all cementitious materials used in the study as provided on the plant certificates.

163

Preliminary optimization analysis of ternary mixtures for bridge decks

Table 1. Physical properties and chemical composition of cementitious materials (as reported by the suppliers) Description of test Specific gravity Fineness -retained on #325 mesh -Blaine’s surface area (cm2/g) Compressive strength tested on mortar cubes (MPa) - 1 day - 3 day - 7 day - 28 day Silicon Dioxide S O 2 (%) Aluminum oxide A1203(%) Femc oxide Fe203(%) Calcium oxide CaO (%) Magnesium oxide MgO (%) Sulphur trioxide SO3(YO) Loss on ignition (YO) Sodium oxide Na20 (%) Potassium oxide K20(%) Total alkali as sodium oxide Na20 ( Y ) Insoluble residue (%) Tricalcium silicate C3S (%) Dicalcium silicate C2S(YO) Tricalcium aluminate C3A (%) Tetracalcium aluminoferrite C4AF(YO)

3470

Cement (Series 2) 3.15 99 3620

Class C Fly ash (Series 1&2) 2.67 9.6 -

-

16.5 26.0 32.2 -

15.9 26.8 33.6 41.8

-

-

-

-

-

20.04 5.84 2.28 64.87 1.63 3.28 1.13 0.14 0.88 0.72 0.47 60 12 12 7

20.60 4.70 2.60 64.90 2.60 2.50 1.31

36.16 20.32 7.58 23.94 5.47 1.91 0.43

93.07 0.62 0.41 0.66 1.16 <0.01 2.71

-

-

-

0.53 0.34 65 10 8 8

1.35

0.67

Cement (Series 1) 3.15

Silica fume (Series 1&2) 2.20 -

-

-

-

-

-

-

-

-

-

Two different sources of Indiana #S crushed limestone, with the maximum aggregate size of 19 mm, and natural silicious Indiana #23 sand were used for each series of mixtures. The physical properties of fine and coarse aggregate used are given in Table 2, while gradations are shown in Figure 1. Table 2. Physical properties of aggregate

Specific gravity (SSD) Absorption (%) Aggregate correction factor for air content (%)

Coarse aggregate (Series 1) 2.69 1.4

Fine aggregate (Series 2) 2.62 1.2 0.05

Coarse aggregate (Series 1) 2.65 1.1

Fine aggregate (Series 2) 2.66 1.6 0.50

Three chemical admixtures, namely air entraining agent, polycarboxylate-based highrange water reducing admixture (HRWRA) and mid-range water reducinghet retarding admixture, were used in all mixtures. The dosages of the first two admixtures varied, and were adjusted individually for each mixture in order to reach the target values of the air content in the range 5.0+8.5% and slump in the range 100+190 mm. The mid-range water reducer was used to ensure retention of good workability of concrete for period of time sufficiently long to perform tests on fresh concrete and to cast the test specimens.

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Mateusz RADLINSKI, Jan OLEK, Haejin KIM, Tommy NANTUNG, Anthony U N D E R

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9.5

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Sieve size (mm)

Figure 1. Gradations of fine (left) and coarse (right) agg. (Series 1 and Series 2 mixtures)

Mixture proportions Four ternary mixtures containing different combinations of fly ash (20% and 30% by mass of cementitious materials) and silica fume (5% and 7% by mass of cementitious materials) were evaluated in the study. The detailed proportions of these mixtures are given in Table 3. The cement content of all mixtures was the same (231.3 kg/m3) and they were all prepared at a constant water-cementitious materials ratio of 0.41. As a result, the paste content of these mixtures varied from 23.1% to 27.9% of their total volume. Table 3. Mixture proportions Mix designation Total cementitious materials (kg/m3) Paste (%) Cement (kg/m3) Fly ash (kg/m3) Silica fume (kg/m3) Fine agg. (SSD) (kg/m3) Series 1 Series 2 Coarse agg. (SSD) (kg/m3) Series I Series 2 Water (kg/m3) Air entraining agent ( d m 3 ) Series 1 Series 2 HRWRA (mum3) Series 1 Series 2 Mid-range water reducing/set retarding admixture ( d m ’ )

20FN5SF 308.3 23.1 231.3 61.7 15.3 746.3 (738.9) 11 15.7 (1 141.6) 126.5 50.3+100.6 40.2t50.6 804.6+3017.1 704.0+1609.1

20FN7SF 317.0 23.9 23 1.3 63.4 22.3 738.4 (719.7) 1104.0 (1 128.7) 129.9 51.7+82.6 41.3151.7 1033.7i1550.6 723.6s 1343.9

30FN5SF 355.9 26.9 23 1.3 106.8 17.8 705.9 (695.2) 1055.3 (1 072.1) 145.9 58.0s2.8 58.049.6 928.3+1624.6 812.3

30FN7SF 367.4 27.9 231.3 110.3 25.8 695.5 (688.0) 1039.8 (1 063.0) 150.7 59.9+107.7 59.9+71.9 958.511676.0 718.9

402.3

413.5

464.2

479.2

Testing procedures All mixtures were prepared in a laboratory pan mixer with a nominal capacity of 0.05 m3. Slump, air content and unit weight were measured immediately after completion of mixing following the relevant ASTM procedures. Compressive strength test was camed out after 28 days of standard moist curing on three 102x203 mm cylinders. In order to evaluate the resistance of concrete to chloride ion penetration, the rapid chloride permeability (RCP) test (AASHTO T277) was carried out on specimens moist cured for a period of 56 days. Despite

165

Prelimina y optimization analysis of terna y mixtures for bridge decks

its frequent criticism [26], this test method is believed to provide a useful comparative measure of concrete resistance to chloride penetration [6,7,27,28], which was the major objective of this study. The reported RCP results represent an average obtained from four 52 mm thick disks, two sampled from the upper part of the 102x203 mm cylinder (after discarding about 6.5 mm thick top layer) and two sampled from the bottom part of a cylinder (located about 114-165 mm from the top). The free shrinkage test was performed on three 7 6 ~ 7 6 ~ 2 mm 8 6 prismatic specimens following the AASHTO T160 procedure, except for the changes in moist curing method (moist room instead of lime-saturated water curing) and the change in the length of curing period which (7 days instead of 28 days). At the end of the curing period the specimens were removed from the moist room and subjected to drying at 23°C and 50% RH. The shrinkage was monitored using a length comparator over a period of 365 days, however for the purpose of discussion presented in this paper, only the average values of shrinkage after 56 days of drying have been reported. With the exception of the free shrinkage test which was only performed on Series 2 mixtures, all tests described above were conducted on mixtures in both series.

TEST RESULTS AND STATISTICAL ANALYSIS Figure 2 presents the relationship between slump and air content along with the corresponding target ranges for these properties. It can be seen that 8 out of 20 mixtures for Series 1 and 8 out of 10 mixtures for Series 2 fell into the target range of both slump and air content. It can also be noticed that, particularly for Series 1 mixtures, there exists a strong relationship between slump and air content. This is confirmed by relatively high coeficient of determination (R2 = 0.78) for Series 1 mixtures. That coefficient was reduced to 0.58 if both series were considered. The reason for this reduction is most likely related to the fact that for some of the Series 2 mixtures the dosage of the air entraining admixture or HRWRA was readjusted if the concrete exhibited too low slump or air content,

9.0

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8.0 7.0

3

6.0

5.0

3.0 0

20

40

60

80

100

120

140

160

180

Slump (mm)

Figure 2. Air content vs. slump (Series 1 and Series 2 mixtures)

200

220

240

260

166

Mateusz RADLINSKI, Jan OLEK, Haejin KIM, Tommy NANTUNG,Anthony ZANDER

In the statistical models developed for all three tested hardened concrete properties, i.e. 28-day compressive strength, 56-day rapid chloride permeability and 56-day free shrinkage, and discussed in the following section, only air content was taken into account as a variable representing the different fresh concrete properties. This was done to avoid serious multicollinearity effects, which would have resulted in obtaining erroneous models due to intercorrelationbetween the predictor variables despite the fact that an improved R2 value can be obtained [29]. Presented in Figure 3 are the results of 28-day compressive strength. It can be noted that, regardless of composition, most of the mixtures tested. achieved satisfactory level of compressive strength of about 40 MPa and higher, which is often perceived as a desired level of compressive strength for bridge deck concrete [30]. Similarly, almost all mixtures, irrespectively of the composition, exhibited good resistance to chloride-ion penetration (Figure 4), as indicated by the fact that only one mixture showed permeability higher than 1500 coulombs, value more than satisfactory for most of the bridge decks specifications. Likewise, for none of the mixtures was the 56-days free drying shrinkage (Figure 5) greater than -560 microstrains, which was roughly an expected value based on the of previous studies carried out on high-performance concretes of similar composition [3 11.

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3 4

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9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30

mix #

Figure 3. The 28-day compressive strength (Series 1 and Series 2 mixtures) In analyzing the data, it is evident that the scatter present in the test results, which is often as large between the replicate mixtures of one design as between the mix designs themselves, makes it virtually impossible to determine which mixture composition is actually the best, i.e. has the highest strength, lowest coulomb value, and lowest shrinkage. It appears that the only conclusion that can be drawn from visual examination of data presented in Figures 3-5 is that RCP results obtained within the same mixture composition for Series 2 are somewhat higher than those obtained for Series 1 mixtures (Figure 4). Therefore, for each of

167

Preliminary optimization analysis of ternary mixtures for bridge decks

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the properties mentioned above statistical analysis was performed and, by the least squares method, the following multiple regression models were developed: STR = 72.854 - 0.040CM - 0.258.AIR2

(Rz=0.86) (1)

RCP = 991.656 - 37.568.AIR + 18.761.FA- 124.346.SF + 341.780.Series# (R2=0.89) (2) SHR = 625.000 - 2.563CM - 4.083.AIR'

(R2=0.82) (3)

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Mateusz RADLINSH, Jan OLEK, Haejin KIM. Tommy NmTLWG, Anthony ZANDER

where: STR - 28-day compressive strength (MPa) RCP - 56-day rapid chloride permeability (coulombs) SHR - 56-day free shrinkage (microstrains) CM - mass of cementitious materials (kg/m3) AIR - fresh concrete air content (%) FA - amount of fly ash (YOby mass of total binder) SF - amount of silica h e (% by mass of total binder) The values of coefficients of determination obtained for the prediction models were not the only parameters used for selection of the best-fitting model. The t-test was also used to assess the statistical significance of predictor variables in each of the models, whereas variance inflation factor (VIF) [29] was computed to check for the presence of multicollinearity effect in the models. Furthermore, development of any models with reasonable value of R2 was only possible when several outlying results were removed from the data set. In the case of model for 28-day strength, the results for mixtures 3, 8 and 24 were excluded, for 56-day RCP the results for mixtures 6 and 8 were removed, and for 56-day shrinkage the results obtained for mixture 16 was discarded, as they appeared to contradict the expected trends. The developed regression models are very useful in terms of identifying which factors, and to what extent, affected the examined parameters. However, due to specific conditions under which the mixtures were prepared and tested, as well as because of variations in composition and properties of raw materials used, the models can be expected to yield only rough estimates of the anticipated concrete properties. As it can be seen, the increase in air content, which was found to be significant in all the developed models, appeared to reduce compressive strength and increase shrinkage, while it decreased RCP results. It is also apparent that the increasing amount of cementitious materials, or in other words the volume of paste, reduced compressive strength and increased shrinkage. The model developed for RCP contains further refinement, and implies that increase in fly ash content increases the coulomb values, as opposed to silica fume which results in lower RCP results when used in larger quantities. This is in agreement with findings from the earlier study as reported by Olek and Lu [16]. Finally, the change in source of materials appeared to alter rapid chloride permeability results by approximately 30%, while compressive strength was not found to be affected by this variable. In order to allow for actual comparison of results obtained from mixtures of different compositions, the values of response functions for evaluated concrete properties (as expressed by Eqs. 1-3) were back-calculated for each of the mixture design and series (if present in the model) and are given in Table 4. The air content used in this process was set at the target value of 6.5 %. In addition, Table 4 also provides an estimated unit costs of lm3 of concrete mixtures calculated using unit prices of the raw materials acquired from the local ready-mix concrete plant in Indiana. Table 4. Back-calculated values of response functions Criterion

Rz for model 20FN5SF 20FN7SF 30FN5SF 30FN7SF

28-day Compressive 56-day Rapid chloride strength (MPa) permeability (coulombs) Series 1&2 Series 1 Series 2 0.86 0.89 1185 51.5 843 51.2 594 936 49.9 1030 1372 1123 49.5 782

56-day Free shrinkage (ptrains) Series 2 0.82 -338 -360 -460 -489

Unit cost ($/m’) Series 2 96.4 103.7 97.5 106.1

Preliminary optimization analysis of ternaiy mixtures for bridge decks

169

OPTIMIZATION ANALYSIS As a last step, an ultimate goal of this study - optimization of mixture composition, was carried out. As it was mentioned previously, due to relatively narrow range of mixture composition of interest, a simple optimization problem has been formulated and all 4 mixture compositions were examined. Since ffee shrinkage measurements were not performed for Series 1 mixtures, the optimization analysis was conducted using only the results for Series 2 mixtures (Table 4). The decisive variables vector for this optimization problem was defined as follows:

x = [XI,

X*lT

(4)

where x1 and x~- percentage of FA (20% or 30%) and SF (5% or 7%), respectively The objective functions vector has been specified as follows:

where fl(x) - durability function (permeability and shrinkage) -to be minimized fz(x) - mechanical function (compressive strength) - to be maximized f3(x) - economical function (unit cost) - to be minimized Since the durability function consists of two components, it can be expressed as:

where: fl,l(x) - permeability - to be minimized fI,z(x) - shrinkage - to be minimized The presented optimization problem is a typical multi-objective problem which, due to conflicting objectives, often prevents simultaneous optimization unless special methods of single objective determination are employed. In this paper, the normalized weighted sum method has been employed to select the preferred (optimum) value of the objective functions and corresponding solution (mixture composition). The underlying principle of this method is calculation of value of function f, for each of the evaluated solutions, i.e. mixture compositions, using a formula which in general has the form:

where: J - number of objective functions w, - weight of the j-th objective function fj(x) - value of the j-th objective function for solution x In general, function f, is to be minimized. In order to properly account for fact that durability function fi consists of two sub-functions f1,l and f1,z (representing resistance to

170

Mateusz RADLINSn, Jan OLEK, Haejin kYM, Tommy NANTUNG, Anthony U N D E R

chloride ion penetration and shrinkage respectively), as well as for the negative values of shrinkage and the need to maximize the compressive strength, the actual formula used for calculations takes a form:

f, = W I .

The objective functions weights have been specified as follows: ~1 = [ w ~ Jw1,2IT , = [0.5,

w = [WI,~ 2~ ,

0.5IT 0.2IT

3 = [0.6,0.2, 1 ~

The choice of the weight coefficients assigned to each of the objective functions, however arbitrary, was driven by the intended application of optimized mixture for HPC bridge decks. As selected, the relative weights imply that permeability and shrinkage were equally important and comprised 60% in terms of the importance, while the strength and unit cost of mixture were assigned an importance 20% each. The calculated values of function f, for each mixture composition as well as their rank (where lowest = the best) are given in Table 5 . As it can be seen, for the selected optimization criteria and specified weight coefficients assigned to them, the optimum mixture is the one which contains 20% of fly ash and 7% silica fume by mass of binder, for which the minimized function f, is 0.44.Moreover, it appears that the mixture 20FN5SF which ranked as a second best mixture, is not very much worst than the optimum mixture. In addition, the performance of both mixtures containing 30% of fly ash seem to be poorer as compared to both mixtures containing only 20% of fly ash. Table 5. The calculated values of function f, and ranks of the examined mixture designs Mix designation fs Rank 20FN5SF 2OFN7SF 30FA/5SF 30FN7SF

0.46 0.44 0.59 0.57

2 1 4 3

Specific relationships between the evaluations of all the normalized and weighted objective functions are shown in Figure 6. To help with the interpretation of these graphs an example is provided using plot “a” that shows the relationship between the normalized and weighted objective sub-functions f1,l (permeability) y d f1,2 (shrinkage). Because both functions are being minimized, the ideal value Yidl (and corresponding solution) are determined by the x-coordinate of the point having the minimum value of wl,l.f1,l/maxf1,l(x) (mixture 20FN7SF) and y-coordinate of the point having the minimum value of wl,2.f1,2/maxf1,2(x) (mixture 20FN5SF). The preferred (optimum) evaluation ypl is Parameter yidl denotes an ideal evaluation of the objective sub-functions(permeability and shrinkage). Parameter Yid occurring in the plots “b” through “8’is an ideal evaluation of the objective functions in the entire optimizationproblem (durability,mechanical and economical)

-

171

Preliminary optimization analysis of ternary mixtures for bridge dech

determined using a distance method [32-341, and it constitutes a point which is closest (geometrically) to ideal point Yidl. Similarly, the anti-ideal evaluation (not shown in Figure 6) would be represented by the point having x-coordinate of the point having the maximum value of wl,l.fl,l/maxfl,l(x) (mixture 30FN5SF) and y-coordinate of the point having the maximum value of wl,2.f1,2/maxf1,2(x)(mixture 30FN7SF). Also, by analogy, the worst solution is mixture 30FN5SF being the closest to anti-ideal point. Analogical interpretation can be applied to the three other plots showing relationships between functions f1, f2 and f3 (Figures 6, plots “b-d”).

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Figure 6. Evaluations of objective functions for w1 = 0.6; w2 = 0.2; w3 = 0.2 It is interesting to note that mixtures 20FN5SFA and 20FN7SF are in all cases so-called non-dominated solutions (as indicated by non-dominated evaluations). That outcome implies

that at least one of the normalized weighted objective functions reaches the extreme value (minimum for durability and economical function and maximum for strength). Furthermore, as shown in plot “d”, mixture 20FN5SF is an ideal solution, i.e. it minimizes the unit cost and maximizes the compressive strength at the same time. The dominating tendency of the mixtures containing 20% of fly ash has been examined further by determining the optimum mixture composition at any value of weights w1, w2 and w3. Due to computational constraints, the values of weights W I J (corresponding to permeability) and w1,2 (corresponding to shrinkage) were kept constant at 0.5. The ranges in which the examined mixtures become the optimum solution have been illustrated in a form of diagram (Figure 7) in which each comer of a triangle denotes 100% weight of the respective objective function. Evidently, as the entire area of the diagram is taken by either mixture

172

Mateusz RADLINSK, Jan OLEK, Haejin KIM,Tommy N A N m G , Anthony ZANDER

20FAl5SF or 20FAl7SF, none of the two mixtures containing 30% of fly ash has become an optimum solution, regardless of what weights are applied to the evaluated properties.

fs=ECONOYlCAL -~ (Unit cost) ~

A

Proposed objective

fr=DURABILlM (Permeability8. Shrinkage)

fz=MECHANICAL (Compressive strength)

Figure 7. Diagram for optimum mixture selection for full range of objective functions weights CONCLUSIONS On the basis of the work presented in this paper, the following conclusions can be drawn: The proposed approach of statistical analysis of data prior to the actual optimization analysis helped in determining the parameters affecting each of the evaluated response functions and accounted for variability of concrete properties and test results. The developed regression models indicate that the increase in air content lowers compressive strength and shrinkage but at the same time it also reduces rapid chloride permeability. The change in material sources altered 56-day RCP results by approximately 30%, whereas the 28-day compressive strength appeared to be insensitive to change in materials source. For the predefined set of optimization criteria and corresponding weight coefficients, the optimum mixture contains 20% of fly ash and 7% of silica fume The selection of the optimum mix design depends on the values of weights assigned to the corresponding criteria, but for any configuration of the weights the optimum mixture always contains 20 % rather than 30% of fly ash. ACKNOLEDGEMENTS The authors wish to express their gratitude to Adam Rudy and Jason Taylor for assisting in preparation of concrete mixtures and casting of the test specimens.

Preliminary optimization analysis of ternary mixtures for bridge decks

173

REFERNCES 1. Neville, A. M., Properties of concrete, Fourth Edition, Pearson Education Ltd. 1995. 2. Boddy, A., Bentz, E., Thomas, M. D. A., and Hooton, R. D., An overview and sensitivity study of a multimechanistic chloride transport model, Cement and Concrete Research 29, Elsevier 1999, pp 827-837 3. Hearn, N., Effect of Shrinkage and Load-Induced Cracking on Water Permeability of Concrete, ACI Materials JournalMarch-April 1999, pp. 234-24 1 4. Weigel, J., Kapur J., and Khaleghi, B., Key factors to durable bridge deck slab made with high performance concrete in Washington State, ISHPC 2003 5. Toutanji, H., Delatte, N., Aggoun, S., Duval, R., and Danson, A., Effect of supplementary cementitious materials on the compressive strength and durability of short-term cured concrete, Cement and Concrete Research 34, Elsevier 2005, pp 3 11-319 6. Hooton, R.D., Titherington, M.P., Chloride resistance of high-performance concretes subjected to accelerated curing, Cement & Concrete Research 34, Elsevier 2004, pp 1561-1567. 7. Bleszynski, R., Hooton, R. D.,. Thomas, D. A, and Rogers, C. A., Durability of Ternary Blend Concrete with Silica Fume and Blast-Furnace Slag: Laboratory and Outdoor Exposure Site Studies, ACI Materials JournaVSeptember-October2002, pp 499-508 8. Mehta, P. K., Gj0rv, E. E., Properties of portland cement concrete containing fly ash and condensed silica fume, Cement and Concrete Research 12, Elsevier 1982, pp 587-595 9. BouzoubaP, N., Bilodeau, A., Sivasundaram, V., Fournier, B., and Golden, D. M., Development of Ternary Blends for High-performance Concrete, ACI Materials JournaVJanuary-February 2004, pp 19-29 10. Whiting, D., Dehviler, R., Silica Fume Concrete Bridge Decks, NCHW Report 410, TRB, National Research Council, Washington D.C. 1998 11. Nikam, V. S., and Tambvekar, V. Y., Effect of Different Supplementary Cementitious Material on the Microstructure and its Resistance Against Chloride Penetration of Concrete, 2003 ECI Conference on Advanced Materials for Construction of Bridges, Buildings, and other Structures 111. 12. Aitcin, P.-C., High-performance concrete, E&SN Spon, London 1998 13. Simon, M. J., Lagergren, E. S., Snyder, K. A., Concrete mixture optimization using statistical mixture design methods, International Symposium on High Performance Concrete, New Orleans, Louisiana, October 20-22 1997 14. Patel, R., Hossain, K. M. A., Shehata, M., Bouzoubab,, and Lachemi, M., Development of Statistical Models for Mixture Design of High-Volume Fly Ash Self-Consolidating Concrete, ACI Materials JournaVJuly-August 2004, pp 294-302 15. Ghezal, A., Khayat, K. H., Optimizing Self-Consolidating Concrete with Limestone Filler by using Statistical Factorial Design Methods, ACI Materials JournalMay-June 2002, pp 264-272 16. Olek, J. and Lu, A., "Optimization of Composition of HPC Concrete with Respect to Service Life and Chloride Transport Properties", Cement, Wapno, Beton (Cement, Lime, Concrete), Vol. WLXXI, No. 6, pp. 271-283 December 2004 17. Sonebi, M., Svermova, L., and Bartos, P. J. M., Factorial Design of Cement Slurries Containing Limestone Powder for Self-Consolidating Slurry-Infiltrated Fiber Concrete, ACI Materials JournalMarch-April 2004, pp 136-144 18. Lin, Y.H., Tyan, Y.Y., Chang, T.P., and Chang, C.Y., An assessment of optimal mixture for concrete made with recycled concrete aggregates, Cement & Concrete Research 34, Elsevier 2004, pp 1373-1480

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19. Lim, C.-H., Yoon, Y.-S., and Kim, J.-H., Genetic algorithm in mix proportioning of highperformance concrete, Cement and Concrete Research 34, Elsevier 2004, pp 409-420 20. Tao, J., Lin, T., and Lin X., An concrete mix proportion design algorithm based on artificial neural networks, Cement and Concrete Research (In Press), Elsevier 2006 21. Bayramov, F., Tasdemk, C., and Tasdemk, M. A., Optimisation of steel fibre reinforced concretes by means of statistical response surface method, Cement and Concrete Composites 26, Elsevier 2004, pp 665-675 22.Nehdi, M. L., Sumner, J., Optimization of ternary cementitious mortar blends using factorial experimental plans, Materials and Structures, Vol. 35, issue 235 (2002), pp 495-503 23. Abbasi, A. F., Ahmad, M., and Wasim, M., Optimization of Concrete Mix Proportioning Using Reduced Factorial Experimental Technique, ACI Materials JournaVJanuary-February 1987, pp 55-63 24. Muthukumar, M., Mohan, D., and Rajendran, M., Optimization of mix proportions of mineral aggregates using Box Behnken design of experiments, Cement and Concrete Research 25 (2003), Elsevier, pp 751-758 25. Ruiz, J. M., Rasmussen, R. O., and Simon, M., Performance-Based Concrete Mixture Design and Optimization, Proceedings of the Seventh International Symposium on the utilization of High-Stren@h/High-Performance Concrete, ACI, 2005, pp 97 1-989 26. Shi, C., Another look at the rapid chloride permeability test (ASTM C1202 or AASHTO T277), FHWA Report, (2003) 27. Aitcin, P.C., The durability characteristics of high performance concrete: a review, Cement & Concrete Composites, 25, Elsevier, pp 409-420 (2003) 28. Stanish, K. D., Hooton, R. D. and Thomas, D. A., A rapid migration test for evaluation of the chloride penetration resistance of high performance concrete, Proceedings of the PCVFHWA International Symposium on High Performance Concrete, Orlando Florida, (2000), pp 358-367 29. Kutner, M. H., Nachtsheim, C. J., Neter, J., and Li, W., Applied Linear Statistical Models, Fifth Edition, McGraw-Hill/Invin 2005 30. Xi, Y., Shing, B., and Xie, Z., Development of Optimal Concrete Mix Designs for Bridge Decks, Report No. CDOT-DTD-R-2001-11, Colorado Department of Transportation, June 200 1 31. Olek, J., Lu, A., Feng, X., and Magee, B., Performance-Related Specifications for Concrete Bridge Superstructures, Vol. 2, High-Performance Concrete, Report No FHWA/INDOT/JTRP-2001/08-11(2002) 32. Paczkowski, W. M., Selected Problems of the Evolutional Discrete Optimization (in Polish), Published by Wydawnictwo Uczehiane Politechniki Szczecinskiej, vol. 544, Szczecin 1999 33. Brandt, A. M., Marks, M., Optimization of the Material Structure and Composition of Cement Based Composites, Cement and Concrete Composites 18, Elsevier 1996, pp 271-279 34. Deb, K., Multi-Objective Optimization using Evolutionary Algorithms, John Wiley & Sons 2001

Proc. Int. Symp. 'Brittle Matrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

HIGH PERFORMANCE SELF-COMPACTING MORTARS CONTAINING POZZOLANIC POWDERS Syed Ali RIZWAN', Thomas. A BIER' and Muhammad Sharif NIZAMI' 'Institute for Ceramics, Glass and Construction Materials Technology Agricolastr. 17, TU Freiberg, 09599, Germany Fax: 0049-373 1-39-2223, e-mail: [email protected] 'Glass and Ceramics Division, PCSIR Laboratories, Lahore, Pakistan.

ABSTRACT A study on the role of powders in high performance (HP) self-compacting mortar (SCM) systems has been made using three different cements and five powders that included lime stone powder (LSP), class- F Fly-ash (FA) and three 20% mass replacement blends of FA with two rice-husk ashes (RHA- amorphous and crystalline) and silica fume(SF). Naturally occurring locally available sands of 0-2 mm size 100% (Sl) and 0-2mm 80% with 20% 2-4 mm size (S2) have been used in these powder type SCM systems. The targeted flow level of 31i 2 cm was obtained by using Melflux 2500L a poly-carboxylate ester (PCE) type liquid super-plasticizer(SP) having 30% total solids made by Degussa Germany. The parameters reported herein are flow, strength, early shrinkage and micro-structure. Contrary to the frequent literature reported use of LSP in self-compacting systems, the results of present investigation in terms of strength, flow and other parameters show that among all powders used, LSP gives the lowest strength, requires highest super-plasticizer (SP) content to meet flow target and shows highest early linear shrinkage with all the three cements used. 20% mass replacement of FA by RHA's and SF increases the strength further and hence improves the pore structure as well. By incorporating 20% 2-4 mm sand in 0-2 mm sand, the shrinkagekxpansion is slightly reduced when compared with that of 100% 0-2 mm sand for the exposure conditions investigated. The maximum pore size, however, was slightly increased. Simple strength quantification is proposed and results are reported for the used powders in such HP SCM systems.

Keywords High performance, self-compacting mortars, secondary raw materials, rice-husk ash, silica fume, lime stone powder, fly-ash, flow, shrinkage and microstructure

INTRODUCTION SCM systems find numerous applications while the published work on such systems is rather limited especially for those having pozzolanic powders which are invariably incorporated in HP/SCCS (self-compacting cementitious systems) to make a variety of improvements like reduction in water demand, increased flow and/or strength, reduced peaks of liberated heat, early shrinkage control and to improve the microstructure for enhanced durability. Essentials of SCCS including high fluidity and segregation resistance are in general contradictory in nature. Powders and SP have to be used imaginatively to achieve these contradictory requirements. This paper is a part of ongoing research on HP SCCS by the authors and a part on HP self-compacting paste systems has already been published [l, 21. In addition to this work another paper on self-compacting concrete by the authors is also included in these proceedings thereby completing the systematic study starting from self-compacting pastes and ending on self-compacting concrete. In Germany the combined use of FA with other

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Syed Ali RIZWM, Thomas A. BIER, Muhammad Sharif NIZAMI

admixtures like silica fume etc. in concrete seems to be generally unusual although in Scandinavian countries such combination of both materials had proven successful in practice [3] and literature also verifies such combinations [4]. Generally the cement based systems incorporating SF show greater liberated heat in comparison to the control system while incorporation of FA reduces the heat evolution [5]. Differential scanning calorimetry by the authors also showed and confirmed these points. Moreover the presence of SP also delays the heat peaks observed in the calorimetry of cement pastes (compared with those without it) and the amount of heat liberated depends on the cement type, its content and wlc ratio. Using CEM 142.5 R with SP increases the heat liberated while with CEM III/B 32.5N-NW/HS/NA it gets reduced at the same wlc ratios. At higher wlc ratios in SCCS, the liberated heat is also reduced possibly due to lesser cement content. EXPERIMENTAL Materials Three cements from Lafarge, CEM 142.5 R, CEM IVA-LL 32.5R (pr EN 197-1,6-20% lime stone) and CEM III/B 32.5 N-NW/HS/NA and five powders including LSP, FA and three of its 20% mass replacement blends with two rice-husk ashes (amorphous RHA and crystalline RHAP) and SF were used. Powder exclusively means mineral admixtures over and above cements. Table 1 lists their chemical compositions and physical properties.

Table 1 Physical properties and chemical compositions of powders

Table 1 presents the physical and chemical properties of all the powders used. Two types of natural local sands S1 and S2 consisted of 0-2 mm size and 80% 0-2 mm and 20% 2-4 mm size respectively with fineness moduli 2.39 and 2.78 were used. Figure 1 shows the particle

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shapes of the powders obtained by scanning electron microscopy. The particle sizes and shapes would help understand various aspects of HP SCM systems described in this paper.

SEM PICTURES OF POWDER PARTICLES

Fig: 1 (c) SEM of FA Darticles powder particles used. It is essential to know the shapes of powder particles for understanding their role in flow, strength, shrinkage and microstructure of HP SCM systems. LSP is seen to have very rough and irregular morphology which is responsible for its very high SP demand for meeting the flow target. SF and FA are round particles of varying sizes; RHA has internal porosity and irregular shape which results in high viscosity. RHAP is crystalline and gives earlier setting. The aggregate grading also plays an important role in SCCS and per cent material passing sieve 1 mm has a direct bearing on their segregation resistance. For fillers alone, the percent passing sieve 0.063 mm is very significant and should be more than 70% [6]. S1 and S2 had passing percentages at 1 mm sieve of 73.37 and 58.47 respectively which itself would suggest about the range required for SCM systems. Flow target was obtained by adjusting SP content. The water demands, setting times, flexural and compressive strengths of the cements (at water

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Syed Ali RZZWAYAN, Thomas A. BZER, Muhammad Sharif N Z W I

demands) are given in Table 2. The compressive strength seems to depend on the clinker contents of the cements. Table 2 Setting times and strengths of cements

ROLE OF POWDERS IN STRENGTH ENHANCEMENT

Inert and pozzolanic mineral admixtures modify the physical and chemical properties of mortars and concretes and the compressive strengths can be separated into fractions of strength related to physical and chemical effects of mineral admixtures. When mineral admixtures are added, three effects can be quantified including, dilution, heterogeneous nucleation (physical) and pozzolanic reaction (chemical) depending on the amount and solubility of amorphous silica. Heterogeneous nucleation is a physical process leading to a chemical activation of hydration of cement such that mineral admixture particles act as nucleation centers for the hydrates thus enhancing cement hydration. A smaller amount of powder has an optimum efficiency and results in a large increase in compressive strength while the use of large amount of powder has a smaller effect [7]. Therefore only 20% mass of FA was replaced with both types of RHA and SF. LSP particles being finer act as nucleation centers for hydrations products and also adsorb SP in its porous structure with bottle neck pores. Low purity of LSP in terms of CaC03 content will also adsorb more SP [8]. Incorporation of FA in cement based materials generally reduces the water demand, increases the setting times and reduces the early shrinkage due to the delayed hydration. Packing effect is dominant for FA systems during 3-28 days [9] and pozzolanic effect becomes more pronounced thereafter and that the pozzolanic reaction of FA decreases with increase in its particle size. Quantification of SF in concrete systems has shown that upto an age of 7 days, physical effects contribute to the compressive strength while beyond that chemical effects become significant [lo]. Increase in the strength of a cementitious system brought about by the inclusion of amorphous RHA in a replacement mode is due to its pore refinement effect, reduction of effective wlc ratio due to absorption of water in internal porosity of RHA particles, improvement of cement hydration and to the pozzolanic reaction between silica and Ca(0H)z [ 111. By virtue of its reduced pozzolanic activity, crystalline RHAF' showed lesser strength enhancement than amorphous RHA. FLOW OF SCM SYSTEMS

Literature suggests the indices of deformability and segregation resistance in terms of minislump cone spread and V-funnel time [12, 131. It has been proven that the relationships between slump spread and yield stress and those between V-funnel time and plastic viscosity are very close [ 141. CEM I1 is particularly useful for the improvement in corrosion resistance

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[15]. Mixing sequence is also very important for obtaining flow target with minimum SP content in all HP/SCCS. It consisted of 30 seconds of dry mixing of materials at slow speed in a 1OL Erisch mixer (900 rpm slow and 1800 rpm fast speed) followed by the addition of 80% of the total mixing water with 30 seconds slow and then 60 seconds fast speed mixing. The interior of the mixer was then cleaned and remaining 20% water along-with SP was added in a single dose for enhanced workability retention [16]. Three minutes of additional mixing at a fast rate was done. The total mixing time was five minutes which is higher than that for normal concrete [6, 171. CEM 11, CEM 111 and CEM I formulations were tested in order for a target mini-flow spread achieved by varying SP content. After testing flow, strength and volume stability of SCM with CEM I1 and CEM 111 formulations with both the sands, it was noticed that sand type S2 had a marginal effect on such properties of SCM systems. Therefore the SCM formulations with CEM I were tested with sand type S1 only. The mix proportions of all SCM systems were 1:1:2 (cement: powder: sand) by mass with water-cement ratio of 0.40 and water-powder ratio of 0.2. Figure 2 gives the V-funnel time and T25 cm time relation ship of various HP SCM systems having equal target flow which has been adjusted by SP.

Slump T25 cm Time and V-Funnel Time Resposne of HP SCM 254

*

6

'8

I

20

-

1

5

15

5

r A C1-LSP

1)

T25 cm ,s 10 A c2-LSP

15 &

c3-LSP

20 H C1-FA

Fig 2 T 25 cm time and V-Funnel Time of HP SCM systems Fig 2 shows that FA+ SF (80:20 by mass) powder gives the least V-funnel flow times (lowest viscosity). It means that flow especially through changing cross-sections depends very heavily on the shape of powder particles. Spherical particles of FA and SF of various sizes get easily adjusted in such sections by just rolling over each other without much internal friction. Highest funnel time was for FA+RHA combination showing increased viscosity due to the addition of 20% RHA in FA. Higher mini-slump cone spread times for FA+ RHA formulations were due to their increased internal friction. Tables 3, 4 and 5 give details of SP content required for different powders used in HP SCM systems with sand S1.

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Table 3 Flows of HP SCM systems with CEM 142.5 R Cement content

Series

SCM system

SSSRl 522 521 523 522

SSSR3 SSSR3A SSSR4

I 1

I

1 :1 :2(Cl:LSP:Sl) 1: l:Z(CI:FA:Sl) 1:0.8:0.2:2(Cl:FA:RHA:SI) 1:0.8:0.2:2(Cl:FAR:S) 1:0.8:0.2:2(Cl:FA:SF:Sl) I

8.27 2.90 5.43 3.02 3.68

I I

31.5

I I I

32.25 32.0 31.5

I

12.97 12.82 14.46

I I

I

I 1

23.77 25.00 17.74

I

Table 4 Flows of HP SCM systems with CEM II/A-LL 32.5R Cement content kg/m’

Series

SCM system

Plasticizer /cement ratio (%)

Cone Spread, cm

Vfunnel

T25cm sec

time, Sec

1:1:2(C2:LSP:S1) 1: 1:2(C2:FA:Sl) 31.88

SR3

SR3A SR4

1 1

521 521

I

I

1:0.8:0.2:2(C2:FA:RHAP:S) 1:0.8:0.2:2(C2:FA:SF:Sl)

3.28

Table 5 Flows of HP SCM systems with CEM I11 /I3 32.5 N-NW/HS/NA

I czI 1

i Series

Cement content

~~

ssR2 ssR3

SSR3A SSR4

517 516 518 517

SCM system

1: 1:2(C3:LSP:Sl) 1:1:2(C3:FA:S1) 1:0.8:0.2:2(C3:FA:RHA:Sl) 1:0.8:0.2:2(C3:FA:RHAP:S) 1:0.8:0.2:2(C3:FA:SF:Sl)

/cement ratio (%)

7.5 2.54 4.05 2.79 3.05

TE;m

spread,

I

31.37 33 32.25 33.0 32.5

I

13.08 8.36 14.2 9.05 9.8

I

$11

sec 18.91 14.92 28.5 18.00 13.02

The flow times can also be seen in Tables 3-5. All the HP SCM systems had same flow target. The reported flows and respective times are average of the two batches of each formulation. Determination of T25 cm mini slump cone spread time for pastes and mortars is suggested by the authors (ratio of 25 cm spread to diameter of cone bottom is 2.5) by drawing analogy from the ratio of the diameter of the Abrams cone to T50 cm spread ( 2.5) for self-compacting concrete.

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STRENGTH AND ITS QUANTIFICATIONFOR HP SCM SYSTEMS The casting, curing and strength determinations of HP SCM systems were made as per EN 196-1 of 1994. The mix proportions used were 1:1:2 (cement: powder: sand). W/C ratio and W/P ratios were 0.40 and 0.20 respectively. The strength results of these formulations had 95% statistical acceptance level. Structural and construction engineers are interested only in the strength increases given by various powders so that they could select suitable powders for the placements. They are not interested in the proportions of strength increases attributed to physical and chemical effects. The LSP is relatively inert [6] and it is proposed by the authors to consider it as a base line material for simple quantification of strength increments obtained with other reactive powders. For pozzolanic powders used with the three cements, the increase in compressive strength (w.r.t base line powder) is the minimum for FA+ 20% crystalline RHAP powder showing it to have a little short term pozzolanic activity. A progressive increase in compressive strength of SCM containing pozzolanic powders like FA, FA+ RHA (amorphous) and FA+SF powders (80:20 by mass) is obtained. Table 6 gives the strength results of SCM made with different powders. Table 6 Simple 28-day strength quantification of various powders with CEM I and CEM I1

HP SCM made with CEM 111 showed exactly similar trend and increase in compressive strength for respective formulations was +36.86, +43.46, +31.46 and +39.36 MPa respectively. It should be kept in mind that the amorphous RHA was imported from USA and had medium reactivity. By using optimized special incineration techniques, RHA with very high pozzolanic activity exceeding even that of SF can now be obtained [18]. RHAP was crystalline and was imported from Pakistan. As expected, it showed minimal pozzolanic activity amongst all the pozzolanic powders used but was still better than LSP in terms of strength and other properties.

EARLY VOLUME CHANGES AND DIMENSIONAL STABILITY Engineers are only interested in the total shrinkage value of a cement based system and not in respective contributions of various parallel operating shrinkage mechanisms. This is why that most codes specify only the maximum permissible total shrinkage values. Total shrinkage may be considered as the sum of shrinkages of various parallel operating mechanisms. In general shrinkage is caused by the internal consumption of water and by its evaporation [6].

Syed Ali RZZWm, Thomas A. BZER, Muhammad Sharif N Z W I

182

In this study a modified version of German classical “Schwindrinne” meaning shrinkage channel apparatus measuring 4 ~ 6 x 2 5cm was used for linear measurements at 2&loC and RH of 3 1&5% with specimen in uncovered and covered conditions. Fig 3 shows typical shrinkage plots of SCM with CEM I with various powders in three exposure conditions. Samples were covered by using a specially tailored thin plastic sheet and adhesive tapes. For the third case, covered samples were topped by an insulation sheet afiixed by using adhesive tapes. CI-LSP-3 E x p o m CI-FA-) Exponmr

OW I

0

-1wOW

-2000 00

CI-FA+SF-3 ExpoPuns

Cl.FA+RHA-3 Ecpoiurcr 0.00

250.00

0 -250.00

w

-5W.W -750 00

-1wO.W -1250.W -1500.00

-

0

0.00

-2000.00

Fig 3 Shrinkage response of powders wi CEM I under three exposure conditions.

The insulation toping fiuther reduced the shrinkages, or increased the expansions, by about 20% due to elimination of condensation. Therefore insulation was not tried with CEM I1 and CEM 111formulations. More air may be entrapped in case of finer powders which require high SP doses for a given target spread and a part of it may leave during self-compaction process resulting in pores. This could result in a higher shrinkage. Swelling (or reduction in shrinkage) can be expected when relative humidity of air in covered specimens is about 100% (fully saturated air). It is mentioned that the formation of Ca(OH)2, AFt and AFm crystals produces expansion [19, 201. This is especially true if powder particles have some internal porosity (like amorphous RHA) and are porous or hollow (like FA) so that they could accommodate water inside and give delayed setting due to high carbon content. Incorporation of FA in cementitious systems results in prolonging setting times, which get magnified with SP, and hence liberates reduced heat of hydration due to high carbon content [21]. Increased FA contents in cementitious systems results in an increased amount of absorbed calcium ions. This inhibits calcium ions concentration build-up giving delayed setting. When FA and RHA are combined the total carbon content of the systems is further increased resulting in even more delayed setting. It is stated that incorporation of RHA in high strength mass concretes reduces adiabatic temperature rise by about 10°C for 30 % RHA incorporation [22] and serves

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183

as a heat sink. The HP SCM system using FA showed expansion in the covered conditions. In addition to the reasons stated earlier it appears that the presence of FA increases the effective water content which in turn happens to be conducive to the growth of expansive hydration crystals. In uncovered condition water is internally consumed as well as gets evaporated leaving insufficient water for the growth of expansive species. The shrinkage profiles of the other cements were similar to those of CEM I except for different values which have been reported in Table 7. Table 7 First 24-hour linear shrinkage values of HP SCM systems with three cements

From Table 7, it is obvious that when moisture loss to the environment is prevented, total early shrinkage values are arrested to a great extent emphasizing the need of early covering and of preventing loss of moisture in actual SCCS placements. LSP in SCM systems shows highest early shrinkage in both conditions due to faster internal depletion of water. FA formulation showed expansion in the covered condition. MICROSTRUCTURE MIP was done, with the help of Autoscan 33 Porosimeter with a contact angle of 140' on 1 l0'C 24 hour oven dried specimens. Sand type S2 gave slightly higher maximum pore size than S1 in these formulations. LSP formulation with CEM LI gave the highest maximum pore size while SF gave the lowest. Fig 4 shows some typical pore-size - intruded mercury volume response of powders with CEM I1 and S 1.

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Syed Ali RIZWAN, Thomas A. BIER, Muhammad Sharif N I W I

CEMII-FA at three ages

CEM II-LSP at three ages 0.45 I 0.4 0.35 0.3 0.25 0.2 0.15

I

I

I

I

0.45 0.4 0.35

I

g $ ”=: 0.3

7m

0.1 0.05 0 100000 low00 loo00

0.15 0.1 0.05 0

1000

100

10

1oooO0 100oO0 lo000

1

CEMII-FA+= 0.35 I

1000

100

1

10

0

0

I

I

at three ages. I

I

I

1

1000

100

10

1

0.4

I

,

CEMIl-FA+SF at three ages

,

0.3

0.25 0.2

0.15

0.1

0.05 0

100000 1oO000 10000 n

100000 lo0000 lo000

1000

100

10

1

0

Fig 4 Pore refinement effect of pozzolanic powders used in HP SCM with CEM 11

RHA and SF seem to reduce porosity, more than FA alone, at the age of 7 days and beyond. It is obvious that the maximum pore size decreases with age and also with inclusion of RHA and SF in FA. The pore size distribution shifts towards smaller end of the pore range confirming their pore refinement effect. Table 8 gives the pore sizes and pore ranges allowing significant mercury intrusion in HP SCM formulations. At 7 days age, FA+RHA formulation shows a bi-modal distribution of pore sizes. Table 8 MIP pore size characteristicsof HF’SCM formulations with CEM 11-51

CONCLUSIONS

The following conclusions are drawn ffom this investigation on HP SCM systems. 1. The flow of SCM systems especially through continuously changing cross-sections depends very heavily on the shape of powder particles with spherical particles being more suitable. 2. LSP gives the lowest strength, highest early shrinkage and maximum pore size. 3. The increased strengths of SCM formulations with FA and its partial substitutions can be simply quantified by using LSP as a base line and relatively inert material. 4. Formulations with ashes expand if moisture movement is prevented with environment.

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5. Incorporating 20% medium reactivity amorphous RHA and SF in FA improves the microstructure with SF giving slightly improved microstructure than RHA. 6. FA +RHA seem to be a nice combination for the early shrinkage control in covered SCCS.

ACKNOWLEDGEMENTS The authors are thankful to Mr. Karl Kiser, Plant Manager, Agrilectric International Technologies, Lake Charles, LA USA for providing the RHA used in this investigation. We are grateful to Mr. Javed Bashir Malik, Associate/Structural group leader, Carter &Burgess, Houston Texas, USA for bearing the expenses of the ash transportation.

REFERENCES 1. 2.

3.

4. 5.

6. 7 8.

9. 10. 11.

Rizwan, S.A and Bier, T.A.,” Inclusion of mineral admixtures in cement pastes for high performance concrete”, pr0c.2”~ International conference on Concrete & Development, Tehran, Iran, May 2005.CD7-004,pp 1-12. Rizwan, S.A and Bier, T.A., “Role of mineral admixtures in high performance cementitious systems”, 2nd all Russian RILEM, ACI, CEB-FIP conference on “Concrete & Reinforced Concrete-Development Trends”, 5-9 September 2005 .Mosco, Vo13 “Concrete Technology”, pp 727-732. Wiens, U., Briet, W and Schiessl, P.,” Influence of high silica fume and high fly ash contents on alkalinity of pore solution and protection of steel against corrosion”, Proc. Fifth International Conference on Fly Ash, Silica Fume, Slag and Natural Pozzolans in Concrete, ACI SP 153-39 Vol 2(Ed . V. M. Malhotra) Milwaukee, Wisconsin, USA 1 9 9 5 . 741-761. ~~ Bouzoubaa, N et al.,” Development of ternary blends of high performance concrete”, ACI Material Journal, V. 101, No 1, January-February 2004. Sanchez de Rojas, M.I. and Frias, M.,” The influence of silica fume on the heat of hydration of Portland cement”, Proc. Fifth International Conference on Fly Ash, Silica Fume, Slag and Natural Pozzolans in Concrete, ACI SP 153-44 Vol 2(Ed . V.M.Malhotra) Milwaukee, Wisconsin, USA 1995.p~829-843. The European Guidelines for Self-compacting Concrete, May 2005. 63 pp. Cyr, M; Lawrence, P: and Ringot, E.,” Efficiency of mineral admixtures in mortars: Quatification of the Physical and Chemical Effects of Fine Admixtures in Relation with Compressive Strength”, Cement and Concrete Research 36 2006 264-277. Magarotto, R., Moratti, F., and Zeminian, N., Characterization of lime stone and fly ash for a rational use in concrete”, Proc. Int. Conf. 5-7 July 2005, Dundee, UK. Cement combinations for durable concrete, Editors (R K Dhir, T. A Harrison and M.D. Newlands), Thomas Telford Publisher, pp 71-80. Tangpagasit, J. et al.,” Packing Effect and Pozzolanic Reaction of Fly-Ash in Mortar”, Cement and Concrete Research, 35 2005 1145-1151. Detwiler, R.J: and Mehta, P.K.,”Chemical and Physical Effects of Silica Fume on the Mechanical Behavior of Concrete”, ACI Materials Journal, V, 86, No 6, NovemberDecember 1989. pp 609-614. Yu, Qijun; sawayama, K; Sugita, S; Shoya, M: and Isojima, Y.,” The Reaction Between Rice Husk Ash and Ca (OH)2 Solution and the Nature of its Product”, Cement and Concrete Research, 29 1999 37-43.

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12.

Maeyama, A; Maruyama, K, Midorikawa, T: and Sakata, N.,” Characterization of Powder for Self-compacting Concrete”, Proc. International Workshop on SCC, 23-26 August 1 9 9 8 . 19 ~ ~1-200. Okamura, H; and Ouchi, M.,”Self-Compacting Concrete”, Invited Paper, Journal of Advanced Concrete Technology, Vol 1, No 1, 5-15 April 2003.Japan Concrete Institute. Safawi, M, I; Iwaki, I: and Miura, T.,”The Influence of Flowability and Viscosity in Vibration of High Fluidity Mortar”, Journal of Cement Science and Concrete Technology, Japan, No 56,2002,582-589. Tsivilis et al.,” Properties and Behavior of Limestone Cement Concrete and Mortar”, Cement and Concrete Research 30 2000 1679-1683. Chang, P. K: and Peng, Y. N.,”Influence of Mixing Techniques on Properties of High Performance Concrete”, Cement and Concrete Research, 3 1 2001 87-95. Chopin, D; de Larrard, F: and Cazacliu, B.,” Why do HPC and SCC Require Longer Mixing Time?”, Cement and Concrete Research, 34 2004 2237-2243. Wada, I et al.,” The Strength Properties of Concrete Incorporating Highly Reactive Rice-Husk Ash”’ Transactions of Japanese Concrete Institute, Vol21, 1999, pp 57-62. Brandt, A.M.,”Cement Based Composites”, Materials, Mechanical Properties and Performance, E& FN Spon Publishers, 1995, pp 299-303. Baroghel-Bouny et al.,” Autogenous deformation of Cement Pastes. Part 11-W/C Effects, Micro-Macro Correlations, and Threshold Values”, Cement and Concrete Research, 36 2006 123-136. Massazza, F.,” Pozzolana and Pozzolanic Cements” in “Chemistry of Cements and Concrete “by Lea., 4‘hEd. (Editor: P.C.hewlett), pp 553. Mehta, P.K: and Pirtz, D.,” Use of Rice Hull Ash to Reduce Temperature in High Strength Mass Concrete”, ACI Journal title No 75-7, February 1978, pp 60-63.

13. 14. 15. 16. 17. 18. 19. 20. 21. 22.

Proc. Int. Symp. “Brittle Matrix Composites 8” A.M. Brandt, YC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

DISCUSSION ON SECONDARY FLEXURE IN UNIAXIAL TENSION TEST OF CONCRETE Hiroshi AKITA”, Hideo KOIDE”, Yoshio OZAKA” ’) Tohoku Institute of Technology Taihaku-ku Yagiyama Kasumicho 35-1,982-8577 Sendai, Japan e-mail: [email protected], [email protected] )’ Tohoku-Gakuin University Chuo 1-13-1,985-8537 Tagajou, Japan e-mail: [email protected]

ABSTRACT The fracture process of concrete is analyzed by using a finite element model combined with the tension softening curve of the concrete. The curve is obtained exactly by only a uniaxial tension test.

However, there are still

four misunderstandings about secondary flexure that inevitably occurs in the test. They are the followings. 1.

secondary flexure is avoided by avoiding load eccentricity

2.

secondary flexure appears after peak load

3. uniform stress distribution is realized by a simple load-control loading system 4.

secondary flexure is avoided by adopting fixed boundaries

Those misunderstandings are obstacles to establish a standard test method for the uniaxial tension of concrete. In addition, there are still papers based on those misunderstandings even recently. In this paper, those misunderstandings are discussed, cleared and confirmed to be erroneous by experimental results and theoretical considerations.

Keywords uniaxial tension, tension softening, secondary flexure, test method, fracture mechanics

INTRODUCTION The information of the tension softening process of concrete is essential to analyze fracture behavior and to estimate concrete properties. One of the best ways to investigate the tension softening process is testing under uniaxial tensile loading because of the simultaneous investigation of both tensile strength and softening curves from an identical specimen.

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Hiroshi AKITA, Hideo KOIDE, Yoshio OZAKA

Additional tests or calculations, such as inverse analysis, are not required for the uniaxial tension test. The authors have developed and reported a unique test procedure for the uniaxial tension test of concrete [ 11. This test procedure provides solutions for four common problems of the test, namely unstable fracture, secondary flexure, multiple cracks and overlapping cracks. First, the problem of unstable fracture can be avoided by employing a deformation-controlled loading process with an appropriate measuring length for the deformations of the specimen. Second, a secondary flexure caused by heterogeneous nature of concrete as well as unexpected flexures due to load eccentricity are eliminated by adopting a specifically designed adjusting gear system. The gear system was operated manually or automatically in order to equalize deformations of each two opposite faces separately during the test. Next, multiple cracks are prevented by the application of primary notches on the middle of two identical faces of a specimen. There is a misunderstanding that strong stress concentration due to notches prevents us from obtaining the exact tensile strength, but the stress concentration occurs only in the elastic range of concrete and diminishes after the concrete in the notched section is softened [2]. Thus, notches do not prevent us from obtaining exact tensile strength. At last, overlapping cracks are avoided by adopting additional notches, called a guide notch, on the middle of the other sides (cast and bottom faces). The prevention of the secondary flexure is the most significant among these four issues, because the flexure reduces the measured peak load up to 20% [2]. Nevertheless, many researchers have paid little attention [3-41 or have sometimes failed in the prevention of it [5-61, because of misunderstandings on the effect of the secondary flexure. The main issue of this study is to discuss four misunderstandings on the flexure and to clarify them experimentally and theoretically.

CONFIRMATION BY EXPERIMENT

Akita et al. have developed the test method for the uniaxial tension of concrete to eliminate secondary flexure by an adjusting gear system as illustrated in Fig. 1 and 2. In these figures, there can be seen a notched prismatic specimen, Q type extensometers, load cells connected to steel rods, the boxes with DC motor accommodating the adjusting gear and tensile loading attachments of the loading machine. In combination with universal joints, the gear system can eliminate both secondary flexure and unexpected flexure caused by load eccentricity. When a certain side of a testing specimen is elongated more than the opposite side, the more elongated side should be contracted by turns of the adjusting gear on that side until its elongation is balanced properly to the opposite one. In this test procedure, it is possible to leave the flexure to develop freely without operating the gear system and to compare the results with those obtained by eliminating the flexure.

Discussion on secondaryflexure in uniaxial tension test of concrete

189

4

Grip end Universaljoint

\

Fig. 1 Experimental set-up

Grip end Unit : mm

4

+

100 ch3

J

U

Fig. 2 Detail of experimental set-up

Unit : m m

CastPce notch

1

6 1

T/

chl

Bottom i c e

Guile notch'

Fig. 3 Notched section of specimen

0

10

20

30

40

P/Pmax

50

60

70

80

90

66)

Fig. 4 Load of secondary flexure initiation

Hiroshi AKITA, Hideo KOIDE, Yoshio OZAKA

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In order to detect and eliminate the secondary flexure, the deformations on four side faces of a prismatic specimen were observed during the test. Their measuring points called chl to 4 are illustrated in the notched section of a specimen as shown in Fig. 3. The secondary flexure is recognized as the deformation difference between two opposite side faces, i.e. chl and ch3 or ch2 and ch4.

MISUNDERSTANDING 1 There is a misunderstanding that secondary flexure can be prevented by removing load eccentricity or that no flexure occurs unless load eccentricity exists. In order to clarify this point, the uniaxial tension test without elimination of secondary flexure was performed using 20 ordinary concrete specimens. Figure 4 shows the applied load when secondary flexure initiated in the specimen during the test. The applied load is described by the ratio to maximum load and 0 means less than lo%, 10 means less than 20% and so on. Ordinate expresses the number of specimen among the twenty within the respective range. The secondary flexure occurred in all the twenty specimens. This means that the effort to eliminate load eccentricity by using universal joints which consist of two pins in mutually perpendicular directions is not enough to prevent the secondary flexure. The secondary flexure inevitably initiates by damage at the weakest zone of the concrete specimen due to heterogeneity. The damaged zone is softened and elongated and then more softened and more elongated. Thus, the secondary flexure continues to develop during the increase of the applied load unless some counteractions are executed.

4.2

0

0.2

0.4

0.6

0.8

4.04

0

0.04

0.08

0.12

0.16

0.2

6 (mm)

6 (mm)

Fig. 5 P-6 curves without elimination (1)

Fig. 6 P-6 curves without elimination (2)

MISUNDERSTANDING 2 Another misunderstanding is that secondary flexure occurs after peak load and therefore produces insignificant error in the measured tensile strength. This is easily confirmed to be

Discussion on secondaiy flexure in uniaxial tension test of concrete

191

erroneous by the analysis of the experimental results. Figure 4 also shows an experimental evidence to indicate that the secondary flexure initiates before peak load, because the flexure initiated during the increase of the applied load before the peak in all the twenty specimens. The secondary flexure initiates after 50% of peak load in most of specimens, but sometimes it initiates at less than 40%. The initiation of secondary flexure was determined as the time when the difference of the two opposite side deformations exceeds O.OOlmm, because we can recognize the deviation of the two deformations in the load-deformation curves. Figures 5 and 6 show examples of load-deformation curves without elimination of secondary flexure concerning to two opposite face deformations. In Fig. 5, it seems that secondary flexure occurred just before the peak load. However, when we zoom in 4 times on the abscissa (Fig. 6), it shows that secondary flexure appeared when the applied load became about 13 kN. At the threshold of 0.001 mm, secondary flexure occurred at the load of 13.2 kN and 72% of the maximum load for this specimen. In addition, secondary flexure cannot be detected by only the observation of average deformation but is detected by that of individual deformation of four side faces. Figure 7 shows an example of two load-deformation curves by the elimination of the flexure concerning to individual deformation of ch2 and ch4. There can be seen no deviation in the two curves showing that secondary flexure was eliminated completely. As Figure 8 also shows a specious load-deformation curve, it comes from the same specimen of Fig. 5 and 6 concerning to the average of four side deformations. It means that secondary flexure cannot be detected without the observation of the individual deformation of four side faces.

15 h

5

v

10

a 5

0'

0.05

0.1

0.15

0.2

0.25

6 (mm)

Fig. 7 P-6 curves with elimination

0.3

0

0

0.05

0.1

0.15

02

0.25

0.3

0.35

6 6nm) Fig. 8 P-6 curves without elimination (3)

MISUNDERSTANDING 3

It is also a misunderstanding to consider that it is unnecessary to use a closed-loop loading machine or strain control loading process only if tensile strength is examined without obtaining a tension softening curve [7]. As secondary flexure initiates before peak load and

Huoshi AIUTA, Hideo KOIDE, Yoshio OZAKA

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develops gradually, stress distribution in the cross section at peak load is far from a uniform one. Then, tensile strength obtained by dividing the peak load by the cross sectional area includes a significant error such as 10% usually and sometimes 20% as shown in Fig. 9. This figure shows the actual reduction of tensile strength comparing both the experiment in which a secondary flexure was left to develop freely and was eliminated completely. In Fig. 9, the experimental and analytical results are shown with respect to tensile strength ft by the ratio to true tensile strength fme The true tensile strength in the experiment was assumed as the average from 4 or 5 specimens being tested by complete elimination of the secondary flexure. This result shows that exact tensile strength cannot be obtained without elimination of secondary flexure. In a simple load increasing process, fracture occurs suddenly being called unstable fracture. In such fracture, a crack initiates at the weakest zone in the specimen and propagates in the cross section with very high speed. The fracture cannot be controlled and stress distribution in the cross section is far from uniform one as is obtained when secondary flexure is eliminated completely. Fig. 10 shows the stress distributions in the notched section when secondary flexure is left obtained by numerical analysis combining a finite element model and the tension softening curve [ 8 ] . In the figure, the curves show stress distributions during softening zone propagates and spreads across the cross section. The curve with number 19 correlates to the maximum load which reduces 11% from the true maximum load. The notch in the model produces the heterogeneity of concrete; the deeper notch is applied the larger is the secondary flexure. 1

alft

3

P/Ptrue=0.65

-1.0-

O=

Fig. 9 Tensile strength reduction

'! 3

i

I :I %

P/Ptrue=O.89 ; 31 P/Ptrue=0.78 '' 3 4 ~ / ~ t r u e = 0 . 4 2 ,, 3 8 P/Ptrue=0.13 19

1'

o;

40 $0 X (mm)

s'o

,I,,

Discussion on secondaryflexure in uniaxial tension test of concrete

193

MISUNDERSTANDING 4 In order to prevent the secondary flexure, fixed boundaries were proposed and specimen ends were glued to loading platens in some studies. However, it is also a misunderstanding that the fixed boundary can be realized by gluing specimen ends to the loading platens. As the rotation of loading platens along some horizontal axis allow a deformation gradient, even a small rotation of about lo4 radian will invalidate the fixed boundary and easily produce the secondary flexure (see Fig. 9). In fact, the deformation difference of mutually opposite faces of the specimen can be observed at the peak load in such experiments [3, 51. The deformation difference was observed even when the rotation of both platens was restricted by using a guiding system [6]. Thus, the fixed boundaries cannot be realized by gluing the specimen ends to the loading platens, and the secondary flexure cannot be eliminated by this procedure. When a specimen is stiffened by steel bars fixed at both upper and lower parts [9-lo], the fixed boundary is realized. However, relatively small rigidity of specimen cannot prevent the secondary flexure when the specimen is fairly long, as reported by Hordijk et al. [6]. Whereas, a short specimen cannot guarantee a uniform stress field in the cross section near the middle height. Consequently, only by maintaining the opposite side deformations mutually equal within relatively short measuring length, the fixed boundary can be produced and the secondary flexure can be eliminated. In addition, this method can adopt fairly long specimen which realize a uniform stress field in the cross section of the middle part of the specimen.

CONCLUSIONS There are still four misunderstandings in relation to secondary flexure in the uniaxial tension test of concrete, in spite that more than 40 years has passed since the test was investigated. These misunderstandings were discussed and clarified through experimental evidences and theoretical considerations.

REFERENCES Akita, H., Koide, H., Tomon, M. and Sohn, D., A practical method for uniaxial tension test of concrete, Materials and Structures, 36,2003, pp 365-371 Akita, H., Koide, H., Tomon, M. and Han, S. M., Three misunderstandings in uniaxial tension test of concrete, In: Proc. ACI 5th Int. Conf. “Innovations in Design and Materials”, V. M. Malhotra ed. ACI international SP-209,2002, pp 405-414 Li, Z., Kulkami, S. M. and Shah, S. P., New test method for obtaining softening response of unnotched concrete specimen under uniaxial tension, Experimental Mechanics, 33, 3, 1993, pp 181-188 Li, Q. and Ansari, F., High-strength concrete in uniaxial tension, ACI Materials J., 97, 1, 2000, pp 49-57

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Hiroshi AKITA, Hideo KOIDE, Yoshio OZAKA

5 . van Mier, J. G M., Schlangen, E. and Vervuurt, A., Tensile cracking in concrete and

6.

7. 8.

9. 10.

sandstone: Part 2- Effect of boundary rotations, Materials and Structures, 29, 1996, pp 87-96 Hordijk, D. A., Reinhardt, H. W. and Cornelissen, H. A. W., Fracture mechanics parameters of concrete from uniaxial tensile test as influenced by specimen length, In: Fracture of Concrete and Rock, SOC. Exp. Mechanics, S. P. Shah and S. E. Swartz eds, 1987, pp 138-149 Zheng, W., Kwan, A. K. H. and Lee, P. K. K., Direct tension test of concrete, ACI Materials J., 98, 1,2001, pp 63-71 Akita, H., Sohn, D. and Ojima, M., Simulation study of secondary flexure versus fracture behavior of concrete under uniaxial tension loading, In: Proc. 6th Int. Symposium “Brittle Matrix Composites 6”, A. M. Brandt, V. C. Li and I. H. Marshall, eds, WOODHEAD Publishing LTD, Warsaw, Poland, 2000, pp 371-378 Phillips, D. V. and Binsheng, Z., Direct tension tests on notched and un-notched plain concrete specimens, Magazine of Concrete Research, 45, 1993, pp 25-35 Lenke, L. R. and Gerstle, W. H., Tension test of stress versus crack opening displacement using cylindrical concrete specimens, ACI International SP-201,2001, pp 189-206

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

IDENTIFICATION OF UNIAXIAL TENSION TESTS OF CONCRETE BASED ON MACHINE LEARNING TECHNIQUE Dariusz ALTERMAN'), Janusz KASPERKIEWICZ'), Hiroshi AKITA'), Mitsuo OJIMA') ')

Tohoku Institute of Technology,

35-1 Yagiyama Kasumicho, Taihaku-ku, Sendai 982-8577, Japan

e-mail: [email protected], [email protected], [email protected] ')Institute of Fundamental Technological Research, Polish Academy of Sciences Swiqtokrzyska 2 1,OO-049 Warsaw, Poland e-mail: [email protected] ABSTRACT The paper is dedicated to presentation of possibilities of Machine Learning techniques, (ML), in uniaxial tension tests on brittle matrix composites. This method enables extracting in an automatic way the knowledge hidden in examples. Concrete is weak in tension but proper evaluation of its tensile strength allows better understanding of its possibilities. It is important to identify whether its tensile properties were measured without or with elimination of uncontrolled flexure which may occur in uniaxial tension tests. In the last time, many different tests on uniaxial tension with elimination of such secondary flexure have been performed in Tohoku Institute of Technology. The results have been collected in a database. The ML experiments bring an answer to a question how can be correctly identified the two types of uniaxial tension tests - those obtained without and those obtained with a detrimental effect of the secondary flexure.

Keywords Secondary flexure, uniaxial tension, tensile strength, brittle matrix composites, Machine Learning, concrete databases processing, automatic rules generation

INTRODUCTION The best way to understand the behaviour of concrete under tensile loading is by applying tension directly to a concrete specimen, (direct tension test), because it is possible then to obtain immediately both the tensile strength and the tension softening curve from specimens loaded in a possibly uniform way. The situation does not occur in other experiments dedicated to tension, such as three- or four-points bending, or in splitting tests. To obtain reliable results it is necessary, however, to prevent or to minimize the unacceptable phenomenon of unwanted flexure - known as secondary flexure, which may impair the resulting data, especially the load-deformation curve, but which affects also various fracture parameters that are to be evaluated. Many researchers ignore the secondary flexure, [l-21. Problem was noticed by Akita et al. in 2000, when a specially designed, adjusting gear systems has been developed to eliminate the secondary flexure, [3-51. Series of the experiments were done - with and without elimination of the secondary

196

Dariusz A L T E W , Janusz KASPERKIEWICZ, Hiroshi AKITA, Mitsuo OJIMA

flexure, and the results have been evaluated by Machine Learning techniques. This allowed obtaining rules valuable for differentiating the two phenomena. Artificial Intelligence methods, (AI), like Machine Learning techniques, (ML) are relatively little used in the field of Materials Science in general, and in the cement based materials in particular. A1 opens new possibilities to combine knowledge stored in the experimental tests databases with the traditional knowledge, including also the qualitative data; the approach was recently successfully applied in concrete technology, [6-91. Interesting possibility opened here is that ML can generate concepts automatically, by learning fiom observation of objects and situations. UNIAXIAL TENSION TESTS

The uniaxial tension tests were performed on prismatic specimens of dimensions 1 0 0 ~ 1 0 0 ~ 4 0mm. 0 The notches introduced to prevent multiple cracks and called primary notches were cut on two side faces, perpendicular to the cast and bottom planes. Besides these, other notches called guide notches were made on the cast and bottom faces in order to prevent overlapping of cracks. The different depth of notch was used, and the attribute of the notch depth was introduced into the database description, (Table 1). The extensometers, aligned at the centre of the prisms, were attached on the all four side faces. Experiments were performed in the close-loop loading machine equipped with the adjusting gear system, designed by Akita et al, [3-41, and also in the machine without such a gear system. The adjusting gear system combined with the universal joints as shown in Fig. 1 can eliminate any unexpected flexure which can produce a significant error during uniaxial tension testing of concrete.

Figure 1. Adjusting gear system and experimental set up.

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The problems encountered when investigating tension-softening behaviour under the uniaxial tensile load are unstable fracture, secondary flexure, multiple cracks, and overlapping cracks. The undesirable flexure may appear for two reasons: due to the load eccentricity andor a one side cracking, because of the heterogeneity of concrete; this problem was discussed by Akita, [5]. The elimination of the unwanted flexure is done as follows. If one side of the specimen is more stretched than the opposite side, this side should be constricted, and this is done by action of a special adjusting gear, fixed on this side, and activated until a proper balance in elongation is reached. For the operation, it is necessary to observe the deformations (elongation) on all four sides of the prism under loading. When a certain side is compressed, the opposite side should be completely loosened as not to introduce unnecessary forces into the specimen. The effect of such unwanted flexure can be seen in Fig. 2, where an example of a certain actual load-deformation curve is shown, (diagram P - 4 . It is assumed that the material outside the softening zone behaves elastically after passing the maximum, i.e. when the specimen loading is on the softening branch of the curve. The process is analysed by crack opening displacement, ( Wc) and the deformation is evaluated by the following equation: PL Wc = 6--- 6 r EA

In this equation: 6- the observed deformation, P - the applied load, L - measuring length, E Young's modulus, A - area of the initial cross section (of the ligament), and &- residual deformation, when the load from the maximum value decreases to zero.

pt

Pmax

triFngle

I

L

~

EA Figure 2. Schematic explanation of the crack opening displacement, (Wc),as observed in the load-deformation curve, ( P - 4 . The Young's modulus is calculated from the slope of the P-6 curve in region between 10% and 65% of the maximum load. The feature triangle is obtained from extrapolation of the curve after the specimen was broken. Without elimination of the flexure the obtained tensile strength has wide scatter and the reduction of the tensile strength reaches sometimes even more than 20%. In addition, the load - deformation curves are not as regular as in Fig. 2, which can be seen in Fig. 3. The conclusion is that the tensile strength and the load deformation curves obtained from tensile tests without elimination of such unexpected flexure are misleading. However,

Dariusz ALTERMAN, Janusz KASPERKIEWICZ, Hiroshi AKITA, Mitsuo OJIMA

198

estimation of the actual curve of the tensile test is possible by taking an average of analytical curves, obtained for deformations measured on the four sides of the specimen, [4].

20

-

15

10

5 0 -0.01

0.04

0.09

0.14

0.19

Wnm)

Figure 3. An example of the load-deformation curves for the primary notch sides, when there is no elimination of the secondary flexure. Symbols ch2 and chl concern the gauges attached on two opposite sides of the prismatic specimen. UNIAXIAL TENSION TESTS DATABASE The data for the present investigation generally consist of the results of various uniaxial tension tests, the details on the composition and certain other results from mechanical tests on concrete. The attributes in the data can be qualitative and quantitative. The database for the experiments has been collected from the results from tests on concrete mixes prepared at Tohoku Institute of Technology during last 6 years. All the results of the experiments represent together a certain knowledge base, (KB). Its subset - the experimental results selected and extracted from the KE3 represent a database, (DB). DB is a dataset of properly formatted records composed of attributes, (components or variables). Databases in which records are only numbers can relatively easily be processed with various mathematical algorithms, first of all with popular statistical programs, but a database with quantitative attributes must be treated by more sophisticated soft computing tools. One possibility is the technique of Machine Learning, (ML), applied in this paper. The finally collected database consists of more than 200 records concerning uniaxial tension tests of ordinary concrete, recycled, early age, or high performance concrete, and of reinforced mortar. Taken into account were 24 attributes which are described in Table 1. The layout of the database can be seen in Table 2, (the structure of the database is as explained in the Table 1).

Identification of uniaxial tension tests of concrete based on machine learning technique

199

Table 1. An example of structure of database.

by extrapolahngofthe

In the above database example the observed attributes concern components of concrete composition (C, W, FA, CA, Add), fresh mixture parameters (air, slump), compressive and splitting strength v c , fs), characteristics of the uniaxial tension tests (days, notch, surface, E, &, P,,, ft, triangle, Wc, shape, k, GA control) - generally described in previous chapter (uniaxial tension tests), also date of casting (casting) and type of concrete (type). Two attributes are of derived attribute types; they result from the algebraic operations on the original components. These are the water-cement ratio (WC) and the sand equivalent (S-E) of

200

Dariusz A L T E W , Janusz KASPERKIEWICZ, Hiroshi AkYTA, Mitsuo OJZMA

the concrete mix. In the database the nominal attribute control is a Boolean variable: yes or no, and the label type attribute date of casting is actually a data type - a qualification valid only in See5; in aq19 this parameter must be ignored. Table 2. The layout of the uniaxial tension tests database of the structure as defined in Table 1 (the row at the top specifies symbols of the attributes) No r_

I

source ]days1

notch

, ,P

1

R

I... I

fc

I fs I ... I S-E I air Islump1

shape

I.. 1

Gf lcontrol

In the table the question marks, (?), mean lack of information, and the source column contains names of original data files. MACHINE LEARNING TECHNIQUES In Machine Learning techniques, (ML), the data are treated in terms of classes and relations among them, also in terms of qualitative descriptors. The system can automatically construct classifications- rules or decision trees for groups defied by mutually exclusive phenomena, for example non-control and control of the secondary flexure. ML techniques identify rules belonging to different groups of records - by examples gathered in respective classes. In the present studies, two ML programs were applied aq19 and See5, [lo-121. Processing of data by aq19 and See5 involves categorizing all the test results under consideration into groups of examples selected according to a certain criterion. The aq19 family can refer to syntactic simplicity of the rules - measured by number of rules, number of conditions in the rules - selectors, by simplicity of the conditions, or by a combination of these factors, and also to certain estimate of the cost of data procurement. The See5 technique is able also to construct decision trees, which can be an easy presentation of the rules. The processing of the data goes like in the symbolic example in Figure 4.For both classes of examples ML system generates separate rules.

20 1

Identijication of uniaxial tension tests of concrete based on machine learning technique

non-control # notch E 6 r ft shape Gf 1 2

3

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

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

3 -

A

I

inference machine

control # notch E 6r

1

2

...

(rU119,W

ft shape Gf

......

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

A

A

I .1

1

1

IUIAES gification trees)

III 1-

3 .................. ...... ...

Figure 4. Generalization of the experimental results by ML techniques. Database is analyzed for two classes of examples, described by various combinations of the attributes, which were defined previously, (Tables 1 and 2). In the example above the system was to identify rules for discrimination of records submitted as belonging to two different classes. The property of correct organization of the experiment in the analyzed uniaxial tension tests - and the results were qualified as two different kinds of test - depends mainly on three characteristics: observed residual deformation (&), recorded tensile strength cft) and fracture energy (Gf).

OBTAINED RESULTS The data concerning only the ordinary concrete has been separated from the whole collected database. The experiments were performed using a919 and See5 programs for 123 records (100 obtained with the elimination of secondary flexure - named control, and 23 records without such elimination - named non-control). This corresponds to setting the attribute control in the Tables 1 and 2 as - respectively -yes and no. An excerpt from the output file is shown in Fig.5. In the example the program generated two rules (so called: complexes) for identification of non-control records and only one rule for control records. The rules are presented as output hypotheses (outhypo): non-control-outhypo # rule 1 Idrc8.481 [ft>2.711 [Gf<81.70] (t:7, u:7, n:O, q:O.55) 2 [E<30.51] [dr>9.25] [Pmax<19.95] (t:4, u:2, n:l, q:O.36) control-outhypo # rule 1 [dr<3.53] (t:58, u:58, n:O, q:O.76)

Figure 5. Rules concerning analyzed classes in an output file from aq19. In parentheses there is information about the total number of positive examples covered by the rule - t, number examples covered only by this rule - u, number of negative cases (opposite class) covered by the rule - n, and prediction accuracy - q. Parameter ‘ & was coded as ‘dr’ in input and output files. The rules in Fig. 5 translated into the spoken language are:

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Janusz KASPERKIEWICZ, Hiroshi AHTA. Mitsuo OJIUA

Dariusz ALTE-,

The experiment was performed wiihout elimination of the secondaryj7mure if the following relations are observed: r & < 8 . 4 ~ 1 AND r j b 2 . 7 1 1 AND r ~ f < 8 1 . 7 i OR rE< 30.511

r s.> 9.251

AND

The experiment was performed with elimination of secondalyflexure when: &< 3.531

r

The respective units of the parameters were described in Table 2. The predictive accuracy of the rules was found high, respectively 55%, and 76% for first rules (indicated in Fig. 5 by the factor q=0.55 and q=0.76). The See5 programs generated two decision trees, which allow evaluating uniaxial tension experiments. An example of the results of using See5 is presented in Fig. 6. Decision tree: Gf c = 74.5: non-control ( 7 ) Gf > 74.5: :...dr c = 6.1: control (62) dr > 6.1: :...dr c = 7.8: non-control ( 8 ) dr > 7.8: :...E C = 22.2: non-control (2) E > 22.2: :...Pmax c = 24.45: control (29/1) Pmax > 24.45: non-control (2)

Figure 6. An example of the output file - decision tree, obtained fiom See5. The decision tree corresponding to this result is shown in Fig.7

\ Young’s modulus

control E92.2

3

E222.2

Maximum load

LIzl control (29/1)

Figure 7. The decision tree. The values in brackets specify number of records in a given class. As previously it is possible to formulate rules corresponding to the above decision tree in a spoken language. The rules concern the two classes of records, as below:

Identification of uniaxial tension tests of concrete based on machine learning technique

203

The experiment was performed without eliminationof secondaryflemrre if: G f s 74.451 OR G p - 74.51 AND rdi.€(6.1;7.8>1 OR Gf> 74.51 AND rdi.> 7.81 AND rE5 22.21 OR

r

r

r

r Gp74.51

AND r&=-7.81 AND rE>22.21 AND rpm,> 22.21

The experiment was performed with elimination of secondaryj7exure if:

r Gf>74.51

r cf>74.51

AND r & r 6 . 1 1 OR AND rdi.>7.81 AND rE>22.21 AND rpm,s22.2i

The accuracy of the prediction was very high, about 95%. Only one record from non-control class was classified to control class. Thus the obtained rules exactly described uniaxial tension test with and without elimination of the secondary flexure. The program See5 produced somewhat different decision tree, but also with a very high accuracy. The decisions are described by three parameters: Young's modulus, residual deformation and maximum load, (Fig. 8).

I

Young's modulus

I

Figure 8. The See5 decision tree. The obtained results were confirmed by direct verification of the rules in Excel, by applying its data filter function. CONCLUSIONS By applying the ML technique it was possible to evaluate quality of the uniaxial tension tests. Certain valuable rules including parameters of fracture mechanics were obtained in what concerns ordinary concrete. Such rules allow understanding differences between two kinds of experiment: with and without elimination of secondary flexure, (control and non-control tests). Previously it was noted only, that the tensile strength is sometimes less than 20% in tests without elimination of the unacceptable flexure. The examples of obtained rules show how the techniques of AI, especially ML can be applied successfully to analyze databases on the composition, or on the properties, or on the applied

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Dariusz ALTERMAN, Janusz USPERKlEWCZ, Hiroshi AKITA, Mitsuo OJIMA

experimental techniques, or on the combination of all these in analysis of engineering materials. It can be noted that even if the programs may generate correct rules imperceptible to human observers they will not extrapolate the knowledge beyond the original source database domain. It should be added that further experiments using various A1 tools should make possible an improvement of the quality of data. This will be obtained by elimination of outliers, by avoiding problems of missing values, by applying clustering and re-ordering of the data, etc. ACKNOWLEDGMENTS The first author is gratehl for the financial support awarded by the Japan Society for the Promotion of Sciences (JSPS), in form of the grant for Postdoctoral Fellows. REFERENCES 1. Li, Q. & Ansari, F., High-Strength concrete in uniaxial tension, ACI Materials Journal 97( l), 49-57. 2. van Mier, J.GM. et al., Tensile cracking in concrete and sandstone, Part 2 - Effect of boundary rotations, Materials and Structures 29, 1996, 87-96. 3. Akita, H. et al., Simulation study of secondary flexure versus fracture behavior of concrete under uniaxial tension loading, 6'h International Symposium on Brittle Matrix Composites BMC6, Woodhead Publ. Ltd. (Cambridge) and ZTUREK Research - Scientific Institute, Warsaw, October 2000,371-378. 4. Akita, H. et al., A testing procedure for assessing the uni-axial tension of concrete, Fracture Mechanics for Concrete Materials: Testing and Applications, In C. Vipulanandan & W.H. Gerstle (eds), Farmington Hills: ACI, 2001, 75-91. 5. Akita H., Koide H., Tomon M., Sohn D., A practical methods for uniaxial tension test of concrete, Materials and structures, Vol. 36, July 2003, 365-371. 6. Kasperkiewicz J., Alterman D. Artificial intelligence in predicting properties of brittle matrix composites, 6*h International Symposium on Brittle Matrix Composites BMC6, Woodhead Publ. Ltd. (Cambridge) and ZTUREK Research - Scientific Institute, Warsaw, October 2000,485-496. 7. Alterman D., Evaluation of concrete materials by automatic reasoning (in Polish), doctoral dissertation, manuscript, IFTR PAS, Warsaw 2005, 180 pp. 8. Kasperkiewicz J., Alterman D. On effectiveness of pre-processing by clustering in prediction of C.E. technological data with ANNs, Intelligent Information Systems 2003, Int. Conf.: New Trends in Intelligent Information Processing and Web Mining, Springer - Verlag Company, Zakopane, June 2-5,2003,261-266. 9. Brandt A.M., Kasperkiewicz J. - Eds., Diagnosis of concretes and high performance concrete by structural analysis, (in Polish), IFTR PAS - NATO Sci. Aff. Div., Warsaw, 2003, 218 pp. 10. Michalski R.S., Kaufman K.A., The AQ19 system for machine leaming and pattern discovery: a general description and user's guide, George Mason University 2001, MLI 01-2, 39 PP. 11. AQ19 - Machine Learning programs, at http://www.mli.gmu.edu/mdirections.html. 12. See5/C5.0 - demonstration, at http://www.rulequest.com/download.html.

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and WoodheadPubl., Warsaw 2006

VIRTUAL 3D NONLINEAR SIMULATION OF UNIAXIAL TENSION TEST OF CONCRETE Drahomir N O V k '), Jan PODROUiEK 'I,Hiroshi AKITA ') Institute of Structural Mechanics Faculty of Civil Engineering, Bmo University of Technology Veveri 95, 662 37 Bmo, Czech Republic, e-mail: [email protected] )' Department of Civil Engineering Tohoku Institute of Technology, Sendai, 982-8577 Japan, e-mail: [email protected] ABSTRACT The paper shows possibilities of nonlinear fracture mechanics simulation to capture results of uniaxial tension experiments and discuss problematic aspects of modelling. The occurrence of secondary flexure is a fundamental problem studied experimentally by third author [l]. The virtual 3D numerical simulation of those experiments eliminatinglleaving secondary flexure was performed. The aim of modelling is to give a better insight into uniaxial tension behavior which cannot be obtained directly from experiments including the formation of fracture process zone, the influence of load eccentricity and heterogeneity of concrete.

Keywords Concrete, uniaxial tension, nonlinear fracture mechanics, secondary flexure, tension softening, stress concentration factor. INTRODUCTION Fracture-mechanical parameters, tensile strength and fracture energy, play the decisive role when using nonlinear fracture mechanics FEM simulation. If these parameters are determined well the numerical simulation of a particular task is realistic. It mainly means capturing ultimate load, post-peak of load-deflection curve and size effect phenomenon. It is widely accepted nowadays that the best way to investigate the tension softening process is to apply a uniaxial tension force directly on a concrete specimen. Several problems appear in such experiment, one is the secondary flexure. The problem was studied intensively by the third author, e.g. in [ 1-31. The test procedure eliminating secondary flexure has been suggested [ 11. The aim of this paper is to show possibilities of nonlinear fracture mechanics simulation to capture uniaxial tension experiments and to discuss problematic aspects of FEM modeling. Software ATENA [4] is used for this purpose. The software tool is based on special constitutive models for the finite element analysis of concrete structures and is developed in both 2D and 3D versions. Tensile behaviour of concrete is modelled by nonlinear fracture mechanics combined with the crack band method and smeared crack concept. Randomness of material parameters in ATENA computational model for heterogeneity modelling can be considered too [S]. The following topics are tackled in the paper: The numerical simulation of selected experimental results from Tohoku Inst. Technology - load-deflection curves by eliminatinglleaving secondary flexure; formation of fracture process zone in concrete specimen; the influence of load eccentricity

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Drahomir NOVAK, Jan PODROUjEK, Hiroshi AKITA

and heterogeneity of concrete; stress concentration factor. The aim of numerical virtual simulation is to give a better insight into uniaxial tension behavior which cannot be obtained directly from experiments due to many obstacles, like expensive experiments, the limitation to “reasonable” specimen sizes and numbers of tests. EXPERIMENT The prismatic specimen of 100x100x400 mm with notches (the depth 7 mm and thickness 3 mm) on all four side faces is subjected to uniaxial tensile force, Fig. 1 [2]. Fig. 1 shows the experimental set-up with gear system to eliminate both secondary flexure and the flexure caused by load eccentricity. This elimination was executed in such a way that the more elongated side was given sufficient contraction to reach a proper balance by tightening the gear system during the test. Extensometers of length 70 mm were attached on all four side faces aligning at the center. Fig. 2b shows load-deformation curves (1-d) of two opposite deflections of specimen, curves coincides, which means that there is no difference between the both opposite sides deformations and the secondary flexure was effectively eliminated. When the secondary flexure was left to develop freely, the deformation monitored from one side increases monotonically (Fig. 2a, channel ch-2), whereas opposite side first increases, then decreases and finally becomes compressed (Fig. 2b, ch-4). Fig. 1 Experimental set-up

25 -

-0.02

0

0,02 0,04 Wmm)

0,06

0,08

04 0

0.05 0.1

0.15

0.2

0.25

0.3

WW

Fig. 2 1-d curve a) by leaving secondary flexure b) by eliminating secondary flexure

Ertual3D nonlinear simulation of uniaxial tension test of concrete

207

COMPUTATIONAL MODEL Several 3D computational models were developed and tested using ATENA 3D. One of the suitable topology is presented in Fig. 3. Note, that the node incompatibility in notched specimen is effectively solved by master-slave approach. Steel plates on the boundaries of the model of thickness 50mm are considered in the model (elastic isotropic material) to transfer the point loading in order to prevent unrealistic cracking of concrete near boundaries. The concrete specimen consists of two macroelements of 255 finite elements each and a macroelement which represent the notched part. The number of elements with respect to stress concentration and fracture propagation was selected to number 784 in one row. Note, that higher number of rows was practically impossible due to the computational demands (two rows only would result in 6498 finite elements!). Average estimates of material parameters from experimental testing [2] were directly used for 3D nonlinear cementitious material model: Young modulus of elasticity 22 GPa, tensile strength of concrete 2.87 MPa and fracture energy 112.5 N/m. Brick finite elements and nonlinear solution based on Newton-Raphson and/or Arc-Length methods were used. Several finite elements meshes were tested and the sensitivity of results to different meshes was rather small. Note that as localization limitor the crack band method is used in ATENA software to prevent the spurious mesh sensitivity. Secondary flexure was easily eliminated in computational model - applying prescribed deformation in four points symmetrically around the axes of uniaxial tension and using the fixed-end support at the bottom of the specimen. In the second case of leaving secondary flexure just one point was used for deformation and hinged support. Deformations are monitored in four points exactly according to the experimental set-up, compare Fig. 1 and 5, their difference has to be considered in order to be comparable with experiment. The secondary flexure case was modeled by one tensile force and hinged support, crack patterns are shown in Fig. 4. In order to model a heterogeneity of real material and thus to speed up the secondary flexure phenomena occurrence, the initiation and propagation of fracture process Fig. 3 Scheme zone, a weaker material was considered along one side of the of computational notch, Fig. 5. The material was simply weakened by smaller model tensile strength and fracture energy of concrete. The influence of small eccentricity of prescribed deformation was also investigated.

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Drahomir NOVAK, Jan P O D R O m K , Hiroshi AKITA

Fig. 4 Crack patterns during the secondarv flexure

Fig. 5 Detail of notch with weakened part and 8 monitors

NUMERICAL RESULTS Eliminating secondary flexure modelling

This alternative is easier fiom computation point of view and resulted in a good agreement of the trend, Fig. 6. The experiment is plotted by solid curve and dashed curves represent virtual simulations (1746 and 7460 elements). Some disagreements can be observed in both pre-peak and post-peak regions. Let us emphasize that the average material parameters from more 25 1 experiments were used here in a -exp ch-1 -__ exp ch-3 straightforward way without any -----.camp 1 7 4 6 modification. Better agreement can - - -comp 7 4 6 0 certainly be achieved by heuristic parameters updating or by sophisticated inverse analysis material model parameters identification [ 6 ] . Statistical I simulation of Monte Car10 type considering random scatter of 01 material parameters could provide a 0 0.05 Wmm) random scatter of 1-d diaaams which describes real behaviour, as ,,ideal Fig. 6 Experimental and simulated l-d experiment" is contaminated by curves: eliminating secondary flexure many uncertainties, e.g. [ 5 ] .

-

-

209

Virhral3D nonlinear simulation of uniaxial tension test of concrete

r=0.22

~0.79

Fig. 7 shows normal stresses in direction of uniaxial tension, the stress redistribution due to cracking and formation of fracture process zone is visualized for increasing load level till the peak. The level of load is expressed by ratio of of actual level of uniaxial force P normalized by maximal force at peak Pmm, r=P/Pmm The stress concentration factor near the notch tip was calculated for these cases, the 3D calculation resulted in the starting value of the factor 4, Fig. 8. Strong stress concentration occurs only in the elastic range and diminishes when notched section becomes softened.

0

0,2

0.4

0,6

0,8

1

r

Fig. 8 Stress concentration factor vs. uniaxial force level ratio

Fig. 7 Normal stresses (in uniaxial tension direction) at the cross-section in the notch; isoareas represent the stress concentration, the maximum is moving from the periphery cornem ( d . 2 2 ) towards the core (for r=lthe maximum slightly surrounds the core square); colorful visualization available on the enclosed CD

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Drahomir NOVM, Jan PODROU&K, Hiroshi AKITA

Leaving secondary flexure modelling Virtual simulation of this case is, in the contrary to reality, more difficult. The simple usage of ,,ideal computational model only " with boundary condition enabling flexure is not enough fracture process zone initiation with 25 consequent secondary flexure develops very slowly just as a result of rounding-off 2o errors during the numerical simulation. Imperfections (geometrical and material) 15 should be considered in order to model random fracture process zone propagation 'O and the speed-up of its initiation. Two main reasons for secondary flexure occurrence are usually mentioned in literature: the eccentricity of uniaxial O tensile loading and the heterogeneity of -0.02 0 0.02 0.04 0.08 0.08 material (spatial random variabiliv along 61mml -\..---., the volume of specimen). Fig. 9 Experimental and simulated 1-d c&es: leaving secondary flexure

F

Influence of loading eccentricity The influence of loading eccentricity can be easily modelled by changmg the pomt where prescribed deformation is applied by some small value e. For e=O mm the pre-peak part of diagram is the same for all four monitors. Increasing e value will influence 1-d diagram in such a way that the pre-peak part changes first, path of two main opposite channels starts to deviate very soon while one operates in tension and the other one in compression. This behaviour is shown in Fig. 10 for three considered eccentricities, e = 2, 1 and 0.2 mm. In spite of the fact that real experiments are always contaminated by some errors in the sense of eccentricity, it is obvious that considering eccentricity in virtual numerical simulation cannot capture I-d diagram of real experiment, Fig. 2b.

25

Fig. 10 1-d diagrams from models using different eccentricities of loading

20

4

&I5

C

.Q

ix3

10

5 0

-0.23

-0.03

0.17 Displacement [m]

0.37

Virtual 3 0 nonlinear simulation of uniaxial tension test of concrete

21 1

Material heterogeneity influence The best way how to simulate material heterogeneity would be to use a statistical approach, the theory of random fields and an efficient statistical simulation of Monte Car10 type [ 5 ] . In a simplified way the heterogeneity of real material can be modeled by a weaker part of material along one side of the notch in order to speed up the secondary flexure phenomena occurrence, the -initiation and propagation of fracture process zone, Fig. 5. The material was simply weakened by 20.5 smaller tensile strength and fracture energy of concrete, e.g. to 70, 90, 99.5 % of their initial value. This simple 20 procedure really resulted in satisfactory E simulation of secondary flexure, weakening to 99.5 % provided the best 1 19.5 results, Fig. 9. If this value of weakening was used, one branch of the 19 curve deviates after reaching the peak of 1-d diagram and continues in 18.5 compression direction (snapback) 0.0069 0.0089 0.0109 0.0129 creating “a loop” in similar way as in Displacement [mm] the experiment. The observed Fig. 11 The detail around the peak of 1-d phenomenon is highlighted in Fig. 11 diagram for different virtual weakening of by scale of the axes. In spite of the fact material that the size of this virtual loop is very small, we can assume that this numerical result can be a kev to understand “a loop” observed in experiment. Very small weakening of material in numerical model leads to slower propagation of fracture process zone and the second monitor path follows the first monitor path further beyond. In case of drastic decrease of material properties (e.g. to 70 % of original values), the fracture process zone propagation is very fast, it means that the paths of monitors almost immediately deviate before reaching the peak.

C

I

CONCLUSION

Nonlinear 3D modeling of uniaxial tension test of concrete could capture the experimental 1-d diagram including post-peak branch rather well. The agreement was good for the case of eliminating secondary flexure. Modeling of leaving secondary flexure was more difficult, but the trend has been captured. For this case the modeling of material imperfection is necessary, the paper discuss the role of eccentricity of loading and the material heterogeneity influence in a simplified way.

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Drahomir NOVAK, Jan PODROUaK, Hiroshi AKITA

ACKNOWLEDGEMENTS This outcome has been achieved with the financial support of the Ministry of Education, Youth and Sports, project No. lM680470001 (1M0579), within activities of the CIDEAS research centre. The solution utilized partially theoretical results obtained within the framework of project VITESPO, No. lET409870411, supported by Czech Academy of Science.

REFERENCES 1. Akita, H. Koide, M., Tomon, M., Han, S. M., Three misunderstandings in uniaxial tension test of concrete. Proc. Of ACI 5th Int. Conf. Innovations in Design with Emphasis on Seismic, Wind, and Environmental Loading; Quality Control and Innovations in Materialsmot-Weather Concreting, 2002, pp 405-414 2. Akita, H., Koide, H., Mihashi, H., Experimental validation in the effect of secondary flexure in uniaxial tension of concrete. CD-ROM Proc. of 1lth Int. Conf. on Fracture, Turin, Italy, 2005 3. Akita, H, Sohn, D., Ojima, M., Simulation study of secondary flexure versus fracture behavior of concrete under uniaxial tension loading. Proc. of 6th Int. Symp. Brittle Matrix Composites, 2000, pp 371-378 4. Cervenka,V., Pukl, R., ATENA Program documentation. Cervenka Consulting, Prague, http:ffwww.cervenka.cz ,2005 5. Novhk, D., Vofechovsky, M., Lehky, D., Rusina, R., Pukl, R. and Cervenka, V., Stochastic Nonlinear Fracture Mechanics Finite Element Analysis of Concrete Structures. In: G. Augusti and G.I. Schueller and M. Ciampoli (Eds.) Proc. of ICoSSaR '05 the 9* Int. Conference on Structural Safety and Reliability, Millpress Rotterdam, Netherlands, Rome, Italy, 2005, pp 781-788 6. Novbk, D., Lehky, D., Neural network based identification of material model parameters to capture experimental load-deflection curve. Acta Polytechnica, Vol. 44, No. 5-6/2004, Prague, Czech Republic, 2004, pp 110-116

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, V.C. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

COMPARISON OF BASIC MECHANICAL-PHYSICALPROPERTIES AND FROST RESISTANCE OF COMMON FINE-GRAINED CONCRETE AND BRICKCONCRETE WITH FIBRES AND WITHOUT FIBRES Hana HANZLOVA, Jaroslav d B O R N Y , Jan VODIeKA Faculty of Civil Engineering, Czech Technical University in Prague ThAkurova 7,166 29 Praha 6, Czech Republic e-mail: [email protected]

ABSTRACT Results of experiments of standard fine-grained concrete and brickconcrete specimens with and without polypropylene fibres and different water dosage are presented in this paper. Bulk densities, flexural strengths, compressive strengths, tensile-splitting strengths and frost resistance are compared. From the results of analyses is obvious positive effect of fibres on tensile strength and frost resistance. The experimental analyses results show that recycling of rubble brings interesting possibilities for sustainable building.

Keywords Brickconcrete, fibres, compressive strength, flexural strength, splitting strength, frost resistance INTRODUCTION Concrete with aggregate from recycled materials, which enables saving of a natural aggregate, is considered to be an advanced structural concrete. Dispersed synthetic fibres added to concrete matrix strengthen texture of concrete and brittle behaviour of material changes to tough one. The new material has enhanced tensile strength and ductility [ 13. The ductility of brickconcrete must be examined with different types of fibres and different amount of fibres in a mixture and prove durability of the concrete, as present experience in capillarity and absorbability of brickconcrete indicates that brickconcrete without fibres are not frostresisting enough and they cannot be exposed by negative temperatures if they are saturated with water.

EXPERIMENTAL PART A determination of mixture proportion of brickconcretes with higher strengths and durability was one of the aims of experiments. Basic characteristics of brickconcretes and normal concretes were compared. A coincidence of a grading of crushed bricks and mined coarse of small size (0/4mm) and crushed coarse (aggregate size 4/8mm) was proved (fig. 1). Therefore properties of normal fine-grained concrete with natural gravel and properties of brickconcrete (aggregate were crushed bricks only) were examined and compared [2].

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Hana HAI'VZLOVA, Jaroslav V f B O W f Jan VODICKA

Characteristicsof particular parts: a) pure crushed bricks (clay brick with vertical holes P10) - sieve test - see grain-size distribution curve (fig. 1) - weight of coarse aggregates in a compacted condition Pt,k = 1275 kg/m3 - bulk density Pv,k = 1770 kg/m3 - absorbability 35% b) sand coarse size 0/4 mm (Ostroiska Nova Ves) and natural crushed gravel 4/8 (Zbraslav) - grain-size distribution curve (fig.1) - weight of coarse aggregates in a compacted condition ptk = 1885 kg/m3 - bulk density pvk = 2568 kg/m3 c) portland fly-ash cement CEM II/B - V 32,5R d) Fibres FORTA FERRO - 1% volume content

0,oo

0,25

0.50

1,oo

2,oo

4.00

8,OO

sleve wlth sqare openings [mm] +sand

-Cpure crushed bricks

Fig. 1 : Grain-size distribution curves of fine-grained concrete and brickconcrete In a mixture proportion (tab. 1) water content was changed and,the amount of cement was reduced from the basic content to minimum value set in a Code CSN EN 206- 1. In order to minimise cost an optimal dosage of polypropylene fibres was determined as 1% of volume content eg. 9,l kg/m3. A set of mechanical- physical experiments of standard specimens 100/100/400mm was performed after 28 days (fig. 2-9). A fiost resistance experiment was performed 3 months after manufacturing of specimens. The compressive strength and tensile splitting strength were examined on halves of specimens remaining after a four-point flexural test. An annex A of a code CSN EN 12390-5:2001 says that a flexural test with three point bending shows strengths by 13% higher than four-point bending test. That is why a threepoint flexural test was performed to compare results.

Comparison of basic mechanical-physical properties and frost resistance of common jne-grained ...

fibres B 2H concrete with fibres C OT brickconcrete without fibres C 1 T brickconcrete with fibres C 2H brickconcrete with fibres

2 15

223

424

1225

617

991

21 1

260

1225

617

991

369

424

1196

369

424

1196

971

300

260

1196

991

Table 1: Mixture proportions of fine-grained concrete B 01, B lT, B 2H and brickconcrete C OT, C 1T and C 2H

Fig.. 2 Specimens from concrete andbrickconcrete

Fig. 3

Compressive strength of brickconcrete

Fig. 4

Fig. 5

Splitting strength of brickconcrete

Tensile strength of brickconcrete with fibers

Hana HANZLOVA,Jaroslav VfBORNf Jan VODI&KA

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Fig. 6

Details of destroyed specimens (concrete without fibers)

Fig.. 7 Details of destroyed specimens (brickconcrete without fibers)

Fig. 8

Details of destroyed specimens (concrete with fibers)

Fig. 9

Details of destroyed specimens (brickconcrete with fibers)

RESULTS OF EXPERIMENTS Results of mechanical physical testing of standard specimens 100/100/400mmfrom normal concrete and brickconcrete with fibres and without fibres are presented in a table 2.

specimen

Volume weight [kg/m31

I

Property Tensilestrength I Compres. strength on 1 load 2load fragments

I

1

Splitting strength fragments on

Table 2: Particular values of mechanical physical properties of investigated concretes (mean value of three observations)

Comparison of basic mechanical-physical properties and frost resistance of common fine-grained...

2 17

Frost resistance test was performed after 50, 75 and 90 freezing cycles according to the code CSN 73 1322: 1969 in laboratories of Technical University in Klokner institute. Three months old specimens with fibre reinforcement hlly saturated by water were frozen in the above mentioned cycles. They were measured and weighed before mechanical testing and their bulk density was determined. The strengths after particular cycles are listed in a table 3, coefficients of frost resistance are in a table 4.

with fibres C 2H brickconcrete with fibres

1996

2004

1998

1997

3,95

Compressive strength [ma1

1 B 1T concrete with fibres B 2H concrete with fibres C IT brickconcrete with fibres C 2H brickconcrete with fibres

I

freezing

2,96

3,19

3,31

Splitting strength [MPa]

I cycles I cycles I cycles I freezing I cycles I cycles I cycles

54,7

52,8

51,3

50,8

4,95

5,50

4,50

2,70

23,4

22,8

22,3

23,4

2,70

2,60

2,50

2,50

71,7

64,7

593

63,2

4,40

3,70

3,50

3,40

25,4

21,7

20,9

23,l

2,75

2,20

2,30

2,70

Table 3: The strengths after 3months of fine-grained concrete and brickconcrete with fibres determined on standard specimens 100/100/400 mm and their fractions

Hana HANZLOVA,Jaroslav * B O W , Jan VODItKA

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C 2H brickconcrete with fibres

0,75

0,81

0,84

0,85

0,82

0,91

0,80

0,84

0,98

THE COMMENT ON RESULTS The analysis proved positive effect of fibres on a tensile and flexural strength of both finegrained concrete and brickconcrete. The strength of brickconcrete was lower than the strength of fine-grained concrete. It is due to the character of inert part e.g. crushed brick. Nevertheless the strength of brickconcrete is sufficient to be used in low-exploited structures. A properly designed mixture will possess the desired homogeneity of brickconcrete both with fibres and without fibres. The fibre reinforcement in brickconcrete changes a failure mode of specimens. Fibres influence values of the flexural strength in three and four point bending test. According to the Code CSN 73 1222 Determination of concrete frost resistance the criterion of frost resistance is frost resistance coefficient. Concrete is frost resisting if the coefficient is smaller than 0,75. The brickconcrete with fibres with higher dosage of cement has higher coefficient of frost resistance compared to the normal fine-grained concrete. That means higher resistance to cyclic freezing and thawing due to positive effect of fibres on tensile strength. A reduction of cement dosage decreases frost resistance. Frost resistance of normal fine-grained concrete determined in compression test and tensile splitting test is higher compared to brickconcrete. The results in should be assumed as approximate values because of the small number of experiments. The table 4 lists them as an introductory set for prospective experiments.

CONCLUSIONS Present results of concretes with crushed brick coarse experiments testify that utilisation of brickconcretes with fibres in every-day life is possible and more than without plasticizer and other admixtures. Exploitation of the brickconcrete, namely the brickconcrete with coarse composed of crushed brick and gravel enables regulation of concrete characteristics by a careful proportioning and mixing of these ingredients. Thus will be attractiveness of this composite material achieved as cement is the most energy demanding component in concrete mixture manufacturing and changes of the brickconcrete material properties will not be dependent on cement dosage increase. A sprayed or pumpable brickconcrete with fibres are suitable in highway construction, namely layers of pavement, slope stabilization, in hydraulic engineering for the strengthening of dam crests and in structural engineering for layers of floors in commercial halls. In general for structures where restriction of cracking is required.

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ACKNOWLEDGEMENT The contribution was elaborated with support of research project VZ 04 Sustainable building MSM 684 077 0005.

REFERENCES 1. Vyborny, J., VodiEka J., Hanzlova H., Kolaf K., PorovnPni z&ladnich mechanickofyzikalnich vlastnosti oby6ejnbho betonu a cihlobetonu bez vliken a s v l h y . In: Sbomfk pHsp5vM 3. konference “Specialni betony”, Malenovice, zAfi 2005, C W T v Praze a Sekurkon Ostrava, ISBN 80-86604-22-5, str. 98-105 2. Vfborny, J., VodiEka J., Hanzlova H., Vyuiiti cihelnbho recyklitu k vfiob8 vlaknobetonu. In: Sbornik pHsp8vM workshopu VZ 04 “Udriitelna vystavba 29.1 1.2005, FSV C W T v Praze, EdiEni stfedisko C W T v Praze, ISBN 80-01-03395-3, str. 39-47.

Proc. Int. Symp. "Brittle Matrix Composites 8" A.M. Brandt, VC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

PLENARY INWTED PAPER

RECENT DEVELOPMENTS IN THE MODELING OF MATRIX CRACK PROPAGATION IN BRITTLE MATRIX COMPOSITES Henrik Stang, Lars Dick-Nielsen, Peter Noe Poulsen and John Forbes Olesen Department of Structural Engineering and Materials Building 118, Technical University of Denmark DK-2800 Lyngby, Denmark [email protected], [email protected], [email protected], [email protected]

ABSTRACT The mechanical performance of brittle matrix composites in general and cement based composite materials in particular is heavily depending on the interaction between matrix crack propagation and fiber bridging action. In order to understand this interaction a significant amount of modeling has been carried out over the years resulting in micro-mechanical models, cohesive laws for fiber reinforced concretes on the meso-scale and various models in the meso-scale for different aspects of high performance materials behavior such as strain hardening capability, the strain hardening process and overall strain capacity. The modeling of the fiberhatrix interface obviously plays a significant role and much work has been done in order to characterize this interface from an experimental and modeling point of view. However, the modeling of the matrix crack itself also plays a central role. While it is generally agreed on that cohesive laws (or fictitious crack models) are applicable when the matrix is regular concrete, modeling of paste and fine mortar is often carried out using Linear Elastic Fracture Mechanics. Recently, however, is has been showed that cohesive crack modeling is appropriate even for pure cement paste. The present paper discusses these results in particular in respect to modeling of initial defects, crack opening profile, process zone, matrix cracWfiber interaction during crack propagation, and possible construction of the resulting composite cohesive crack model. Reference is made to investigations based on a combination of analytical models and non-linear E M .

Keywords Fracture mechanics, Cementitious Composite Materials, Fictitious Crack Model, Bridged Crack Model, Cohesive Crack Model, Initial defects, Crack initiation, Crack propagation.

INTRODUCTION Coarse mortar and concrete are typically characterized as quasi-brittle materials. Cement and fine mortars are sometimes considered brittle and sometimes quasi-brittle materials. The proper modeling of this important family of materials have been under debate for more than 30 years. At this point in time it is generally agreed on that the cohesive crack model proposed by Hillerborg, [l], the so-called Fictitious Crack Model, FCM, provides a reasonably consistent framework for the modeling of Mode I crack propagation in concrete and mortar. Furthermore it

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Fig. 1.

Hen& STANG, Lars DICK-NIELSEN Peter NOE POULSEN, John FORBES OLESEN

Thin section of unidirectional, polypropylene fiber reinforced cement paste loaded .- -. vertically in the direction of the fibers and epoxy impregnated while under stress (corresponding to 2% axial strain). The bright horizontal lines are the epoxy filed cracks and the dark particles are unhydrated cement particles. The cracks show similar characteristics to cracks found in concrete in terms of branching and bridging, using similar techniques on a larger scale.

has been shown in recent years that the FCM is also applicable in analysis of crack propagation problems in j b e r reinforced concretes, e.g. steel fiber reinforced concrete, e.g. [2]. The FCM model has been used in analytical and semi-analytical work as well as in finite element models of problems where the crack path was known in advance (problems for which the model was initially intended). Presently significant work is carried out in order to introduce the FCM in the concept of Extended Finite Elements (XFEM) [3], [4] in order to deal with problems where the crack path is not known a priori. The FCM can be characterized as cohesive crack modeling with a crack closing or cohesive stress which depends on crack opening and which eliminates stress singularities at the crack tip. The crack propagation criterion associated with FCM is that the tensile strength at the crack tip should never be exceeded. The reason for the success of the FCM in modeling crack propagation in fiber reinforced concrete and plain concrete and mortar is the significant contribution to the overall energy dissipation in these materials associated with crack formation from micro-cracking and frictional effects on a relatively large scale. These frictional effects are traditionally ascribed to aggregate interlock and fiber bridging during crack opening and the conceptional equivalence of these effects to the closing or cohesive stress of the FCM is obvious. Recent developments in fiber reinforced cementitious composites have focused on materials with matrices consisting of cementitious paste or very fine mortars with sand particles less than a few millimeters, see Fig. 1, and commercial versions of such materials are available today (Densit@ and Ductal@ - to mention a few). It has been demonstrated that such materials - when properly engineered - are capable of exhibiting multiple cracking giving rise to strain

Recent developments in the modeling of matrix crack propagation in Brittle Matrix Composites

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Applied stress

Fig. 2.

The geometrical concept for stress free initial defects in cement based materials: a pore with irregular walls and crack like structures is modeled as a slit-like crack with length 2ao corresponding to the diameter of the pore plus irregularities.

hardening in uniaxial tension even with relatively low fiber contents, see e.g. [5], [6] for an introduction to so-called ECC materials. Studies of these materials are conducted on several levels including crack initiation, single crack propagation and multiple crack formation. In these studies fracture mechanics of cementitious materials are pushed to the limits, since little knowledge is available on the applicability of fracture mechanics on the paste and fine mortar level. Typically, Linear Elastic Fracture Mechanics (LEFM) is applied in modeling crack initiation on paste and mortar level, see e.g. [7] and [8], however a realistic description of the dependency of initial defect size on overall strength of cementitious materials is not obtained through the application of LEFM, which is particular evident if artificial defects are introduced in the material. Typically, unrealistically large initial defects are predicted from LEFM combining measured fracture toughness and first crack strength. Further, though it is usually assumed that the tensile strength of cementitious materials is determined by the largest crack-like flaws, pores, weak boundaries and shrinkage induced cracks, there is no consensus on what microstructural features in the hardened cement paste or mortar should be identified as the initial defects. On the other hand, however, it is found through simulations and experimental work that introduction of initial defects in the matrix of fiber reinforced composites and the spacial and mechanical statistical distributions of these defects are quite significant for the mechanical properties of the composite material, [9] [lo]. Also on the composite material level, i.e. in the study of single crack propagation and multiple crack formation, fracture mechanics are being pushed to the limits, in particular in single crack propagation as it is evident that crack propagation involves fracture processes on somewhat different scales, ranging from the fracture of the matrix at the crack tip to the subsequent fiber bridging action, see Fig. 1. In the same picture it is also shown that the matrix cracks show similar characteristics to cracks found in concrete in terms of branching and bridging, using

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Henrik STANG, Lars DICK-NELSEN, Peter NOE POULSEN, John FORBES OLESEN

similar techniques on a larger scale. The present paper summarizes recent results on the applicability of cohesive crack models in description of crack initiation and single crack propagation in cement paste, fine mortar and fiber reinforced fine mortar. Both a so-called bridged crack model (where cohesive stresses exist together with singular stresses at the crack tip) and a pure cohesive crack model (FCM) have been applied and differences in their behavior are pointed out. The results are based an a semi-analpcal model for a single crack in an infinite sheet and FEM simulations. In the studies the concept of initial stress-free defects from LEFM is maintained even though such defects are not necessary when the FCM is applied. However, when bridged crack models are applied initial defects are necessary. Further, cementitious material are always porous and initial, stress free defects with a length of 2ao are in a fracture mechanical sense a good approximation to a pore with radius slightly smaller than a0 with tiny cracks or irregularities radiating from the surface, [111, see Fig. 2.

COHESIVE CRACK MODELS FOR CEMENT PASTE AND FINE MORTAR Fictitious crack model In the FCM non-constant closing stresses or cohesive stresses, a,(w),are applied to the crack surface. These cohesive stresses depend on the crack opening, w, and vary from the tensile strength of the material, ft,at zero crack opening at the tip of the crack to zero at a characteristic crack opening, w,. It is assumed that the cohesive stresses close the crack smoothly, thus even when the un-cracked material is considered linear elastic - which is often the case in the modeling of cementitious materials - stresses are finite in the un-cracked material at the crack tip. In other words, during Mode I crack propagation, the stress intensity factor KI is zero and the condition for crack propagation is that the stress at the crack tip has reached the tensile strength, ft.Thus during crack propagation the following conditions are fulfilled, see Fig. 4(a):

and at

a,=a,(w)

x s a

with ~ ~ ( 0=)

ft

a,(w)

0

=

at IC = a for w > w,

(3)

The zone in which the cohesive stresses are present was originally called the micro-cracked zone, [11, and later thepmess zone orfracture zone, [121, when it was realized that the cohesive stresses were due not only to stress transfer in micro-ligaments between micro-cracks, but also to frictional stresses in various bridging configurationsconsisting of aggregates, fibers and other inhomogeneities, see Fig. 1. Thus, the cohesive law implicitly contains information about the micro-structure of the material. In the FCM there is no a priori assumption regarding the length of the process zone, lp, there is no direct connection between the length of the process zone and the opening, w. of the process zone and in particular it is not assumed, that the fracture process zone is small compared to the pre-existing macro-crack, total crack length or a characteristic dimension of the structure in question. In order to maintain smooth crack closure the distribution of closing stresses along the crack faces is essential in the FCM, as pointed out by Karihaloo, [13].

Recent developments in the modeling of matrix crack propagation in Brittle Matrix Composites

“w

h

225

4

Smallest length scale

@

Fig. 3.

Largest length scale

Schematic illustration of a multi-scale cohesive law. Note that the x-axis is logarithmic. The resulting cohesive law (not shown) is obtained by superposition of the individual cohesive laws. From [ 191.

The features described above distinguish the FCM from other well-known cohesive models. The Barenblatt model, [ 141, assumed characteristic cohesive stresses distributed over a short process zone, resulting in a model essentially equivalent to Linear Elastic Fracture Mechanics (LEFM) models, while the Dugdale model, [151 assumed constant closing stresses. All three models reduce to LEFM only in the limiting case of small process zones (compared to the total crack length and/or characteristic dimension of the structure in question). In the original formulation of the FCM it is implicitly assumed that for any given cohesive law, a,(w) - whether it is experimentally determined or theoretically assessed - the process zone length and opening can be adjusted in such a way that smooth crack closure, i.e. finite stresses at the crack tip, can be achieved. Otherwise the model formulation would break down and the crack propagation criterion would be inadequate. To the authors’ knowledge it has never been rigorously proven that for any cohesive law, the process zone can always be adjusted to meet the smooth crack closure condition, however, implementations in numerical analysis support this assumption.

Bridged crack models Bridged crack models, [16], [17], [18], are formulated in order to deal with the fact that many crack propagation problems are in fact multi-scale problems. Real materials are inhomogeneous with many size scales which make crack propagation problems multi-scale problems with the atomic bond being the smallest scale and phenomena such as aggregate and fiber bridging being the largest scale in typical cementitious materials. Assuming that FCM type cohesive laws can be applied on all length scales, the length scale of the various phenomena is introduced through the various characteristic crack openings, w:. The concept of a multi-scale cohesive law is illustrated in Fig. 3. On each scale the cohesive law for a cementitious material represents a characteristic mechanism reflecting the average nature of the bond including the presence of defects: atomic bond separation on the smallest scale, separation of grain interfaces, micro-crack ligament stress-transfer, and aggregate or fiber bridging at the largest scale. The resulting cohesive law (not shown) is obtained by superposition of the individual cohesive laws. It follows that the magnitude of the cohesive stresses at the various length scales in general

226

Henrik STANG, Lars DICK-NIELSEN. Peter NOE POULSEN, John FORBES OLESEN

sass

I\

Fig. 4. Sketch showing the process zone in (a) a cohesive law with smooth crack closure and (b) a bridged crack model. can be of different orders of magnitude. The significance of the cohesive law at the various length scales can be characterized by the energy that it represents, i.e. the area under the curve at the various length scales. Thus, in general it is not possible a priori to disregard any part of the cohesive law. In practice, however, it is not possible to solve a crack propagation problem by taking detailed information about the cohesive law into account on all length scales because of the finite resolution of the solution for the displacement fields in the solution methods applied. This problem can be solved by lumping all energy corresponding to length scales smaller than a certain scale, wa, corresponding to the resolution of the solution for the displacement field, into a single point. This corresponds to applying LEFM to those small scales and applying the cohesive law only to the larger length scales, i.e. applying the bridged crack model. It follows that the FCM or any other cohesive crack model with smooth crack closure can be considered as a special case of the more general bridged crack model where cohesive stresses are assumed to exist together with a stress singularity at the crack tip, i.e. smooth crack closure is not required. In bridged crack models the crack propagation criterion is KI = K I , and in crack propagation problems the length and the opening (and thus the cohesive stress) of the process zone are adjusted so that this criterion is fulfilled in the un-cracked material at the crack tip, see also Fig. 4 (b). Thus, with this interpretation of the bridged crack model, and assuming that LEFM is adopted for all length scales smaller than w:, the crack propagation criterion is: KI = KI, with

(4)

--

and

CF= J

0

ow(ut)dw

(6)

while the cohesive law is enforced for w > w:. Here, E' is Young's modulus, E, for the un-cracked material in plane stress and E / (1 - v2) in plane strain, with v denoting Poisson's ratio.

In practice - since UJ;N 0 - bridged crack propagation problems are solved by requiring equation (4)to be fulfilled in the uncracked material at the crack tip together with a cohesive

Recent developments in the modeling of matrix crack propagation in Brittle Matrrk Composites

227

law governing the cohesive stresses on the crack faces, see Fig. 4(b):

x -+a+

KI = K I , at

(7)

and aw=o,(w)

at

x
with

aw(0) = ft at x + a(9) aw(w) = 0 for w > w, where w, corresponds to the largest cohesive size scale and where ftdenotes the tensile strength. In such a calculation the stress intensity factor is a function of the specimen geometry, stiffness, external load and the cohesive law. The multi-scale concept of cracking in materials has been discussed extensively in the literature, see e.g. [20] and [21] and previous applications of bridged crack models deal typically with crack propagation in various types of reinforced or fiber reinforced materials, e.g. [22], [231,[241, P51, [261 and P71. Bridged crack model versus FCM Clearly, the FCM is a special case of the bridged crack model and applicable when small scale energy dissipation is insignificant compared to the larger scale dissipation. If small scale energy dissipation is insignificant it is expected that the small scale energy dissipation can be suitably described trough the tensile strength and the initial slope of the cohesive law. If, on the other hand, small scale energy dissipation is significant the bridged crack concept should be introduced. It is evident, however, from the above arguments that the values for KI,, and ft are somewhat arbitrarily chosen, and propably cannot be regarded as independent material parameters. Further, when small scale energy dissipation is significant, defect size becomes important, which - as will be shown in the following - is not the case when FCM can be applied. A SEMIANALYTICAL MODEL FOR COHESIVE CRACK PROPAGATION As a generic example the problem of a centrally cracked infinitely large sheet is considered, see Fig. 5. The un-cracked material is assumed to be linear elastic with Young's modulus E', (see

Fig. 5.

Figure showing the geometry of the generic problem under consideration together with the superposition scheme applied in the semi-analytical approach.

228

Henrik STANG, Lars DICK-NIELSEN Peter NOE POULSEN, John FORBES OLESEN

above) while a bridged crack model is governing fracture. Thus, the crack propagation problem is governed by the equations (7) to (9). In [28] a semi-analytical approach was suggested to solve the presented crack propagation problem in order to investigate the effect of initial crack length on first crack strength of cement mortar and paste assuming smooth crack closure, while a study with emphasis on the effect on crack length and opening profile of the stress intensity factor in the crack propagation criterion in the bridged crack model was presented in [19]. The model will be summarized here. In the infinite sheet an initial stress free defect of length 2ao is present. The total length of the crack is 2a, see Fig. 5 . Note also the coordinate system with the x-axis in the crack plane and with origo at the center of the crack. The cohesive crack is assumed to propagate when the stress intensity factor KI is equal to KI,. The complete solution for a given crack length a and a given far field uniaxial tensile stress, a,can be found by superposition of two fundamental solutions as shown in Fig. 5. The first fundamental solution is trivial with the traction a1 = a on the crack surfaces and boundary of the specimen, KI = 0 and u1 --= 0 where u is the deformation of the upper crack face, equal to half the crack opening and where the subscript refer to the solution. The second fundamental solution is a crack in an infinite sheet loaded only on the crack surfaces with the traction u2(x).The stress intensity factor and the crack opening displacement is obtained by integration of the fundamental solution of two opposite forces on a crack surface, see [ 111: 1

-w(x) 2

= u2(x,0) = -a

KI = K I z ( f a ) = -

a TE'

a

L

cosh-'

a2 - Ex 4 alx - 61

~

a2(x)dz7dx a

x

aFx

where 6 is an integration variable along the x-axis, ~ ( xis )the total crack opening and E' = E for plane stress. The stress free condition of the initial defect and the cohesive law of the propagating crack require that: ~ ( x )a1 = 0 for 1x1 < a0 (12)

+

and

+

a,(w(x))= a2(x) a1 for

a0

< 1x1 < a

(13)

Introduction of Equation (10) in (13) determines a2together with (12) and thus the crack opening displacement through Equation (10). The stress intensity factor can then be determined from Equation (1 1). Finally, Kr = KI, can be achieved by adjusting the far field, uniaxial tensile stress, al for fixed crack length, a. As a result corresponding values for crack length and far field stress are obtained given the cohesive law and the crack propagation criterion K I = KI,. The cohesive crack model is obtained from the special case KI, = 0.

SIGNIFICANCE OF SMALL SCALE DISSIPATION Investigations were carried out in [19] to investigate the significance of small scale energy dissipation by introducing the critical stress intensity factor as a fraction of the total energy

229

Recent developments in the modeling of matrix crack propagation in Brittle Matrix Composites

2 1.5

1

.

*-

c-

D

0

-

1

0.5

0

2

1.5

"0

0.5

1.5

2

(b) Small initial defects.

(a) Large initial defects.

Fig. 6.

1 a I ]dl

a 1,

Figure showing dimensionless far field stress as function of dimensionless crack length for different values of a and for (a) large initial defects, p = 55, and (b) small initial defects, p = 222. From [19].

dissipation associated with the cohesive law:

KI,= adE'GF with

W C

G F = ~gW(w)dw The investigations were carried out with a simple linear descending cohesive law:

a,(w) = ft(l- alw) for 0 5 w 5 wc= l/al

(16)

Results were presented in terms of far field stress o versus crack length a curves (loadcrack propagation curves) and in terms of crack opening at the center of the initial defect w(0) versus crack length a (crack opening versus crack propagation curves). It was found that results can be represented on dimensionless form if the following dimensionless parameters are used representing dimensionless far field stress, crack length, small scale energy dissipation, crack opening and initial defect size: 0

where the characteristic length, h

a -1

ft

lch

, is

w(0) a,wc

1P=

-1

lch -

a0

given by:

E'Gp

lch =

f,"

Typical results for load-crack propagation curves are shown in Fig. 6 where bridged crack model results are shown together with a LEFM prediction with a critical stress intensity factor corresponding to the fracture energy of the cohesive law (corresponding to a = 1 in equation (14)). Results are shown for different values of a and for large and small defects ( p = 55 and 222, respectively). For small values of the small scale energy dissipation, crack propagation is stable until a has reached about half the characteristic length after which the load gradually drops and has a significant drop when a has exceeded the characteristic length by approximately

230

Henrik STmG, Lars DICK-NIELSEN, Peter NOE POULSEN, John FORBES OLESEN

.

I

.

.

m:

. . . . . . -?,-lo

18.

8-111

16.

14.

. h

I'

06,

o p r

'

0

02

.

'

.

04

06

08

a 'ch

.

.

.

.

I

l l

14

16

I8

'

2

a 'ch

(a) Dependency on a, small defects.

Fig. 7.

.

(b) Dependency on p, a = 0.

Figure showing dimensionless center crack opening as function of dimensionless crack length for (a) p = 222 and various values of a , and (b) a = 0 and various values of beta, from [193.

30%. After this drop the LEFh4 solution is gradually approached. Interestingly, the significant drop in the load-crack propagation curves is seen to correspond to full development of the process zone, see Fig. 7(a) where corresponding crack opening versus crack propagation curves are shown for small defects. For small values of the small scale energy dissipation, the crack opening is very small for quite long cracks and increases abruptly when the characteristic length is exceeded and sensitivity to initial defect size is weak. The load-crack propagation behavior is relatively sensitive to small scale energy dissipation and the length at which cracks propagate stably is rapidly decreased as a increases. Further, as expected the a sensitivity is increased as the initial defect size is decreased. As seen in Fig. 7(a) and (b) cracks propagate with very small crack opening in particular when initial defects are small and small scale energy dissipation is small. In such cases the crack can propagate to he characteristic length with maximum crack openings in the order of 0.05~~.

CRACK PROPAGATION IN MORTARS In [28] studies of crack propagation and opening where conducted using material data obtained from inverse analysis of wedge splitting tests on an ECC mortar material previously investigated in [ 101 (mix 3). The FCM was applied together with a bi-linear cohesive law, which proved to be a close approximation to the measured cohesive law: 1 - alw

1 - b2 for 0 5 w 5 w12 = - a1 > 0 a1 - a2

ft b2

- a2w for w12 < w

5 w,

b2

=-

(18) a2

>0

a2

Parametric studies of the influence of the tensile strength where carried out where the total fracture energy (14 Nm/m2) was kept constant while at the same time adjusting the slope of the first linear branch, al. The characteristic lengths of the material with tensile strengths of 3 MPa and 5 MPa are calculated to 54 mm and 17 mm respectively, while the first linear, descending

Recent developments in the modeling of matrix crack propagation in Brittle Matrix Composites

-

p -a0 = 0.5 mm, -aO=l.Omm -a0 = 1.5 mm I -a0 = 2 0 mm -LEFM

E. 5 0

t

40

3:o

f9 :: 0.0

0

20

40

60

a crack length [mm]

(a) Tensile strength, ft = 3 MPa.

Fig. 8.

23 1

60 50

3

2.0

‘i 1 0 0.0

5

0

20

40

60

a crack length [mm]

(b) Tensile strength, ft = 5 MPa.

The influence of initial defect size on the load-crack propagation curves in a fine mortar for two different tensile strengths, 3 MPa and 5 MPa as predicted by a bi-linear FCM. From [28].

branch of the cohesive law extends to crack openings, w12 (see equation (18)), of approximately 5 pm and 2.5 pm, respectively. In Fig. 8 (a) and (b) the influence of initial defect size on the load-crack propagation curves are shown for two different tensile strengths, 3 and 5 MPa. The overall curve shapes and stress levels are as expected and it is evident that the influence of initial defect size on first crack strength (the maximum stress of the load-crack propagation curves) is larger the smaller the characteristic length. Overall the influence of defect size is small - in the expected order of magnitude - and evidently much smaller than predicted by LEFM. When comparing the results in Fig. 8 (a) and (b) with the dimensionless results from the previous section, for a = 0, it is evident that the extent of stable crack propagation is shorter than predicted by the dimensionless curves in Fig. 6 (a) and (b), where the significant drop in load takes place after crack length, a, has exceeded the characteristic length. This phenomenon can easily be explained, however, by examining the crack opening associated with crack propagation. In Fig.9 the crack opening profiles are shown for a crack extension of 7, 12.5 and 25 mm in the case of ft = 5 MPa. It is seen that for this tensile strength and for crack lengths smaller than about 10 mm, the crack propagation is governed only by the first branch of the bi-linear cohesive law, i.e. the characteristic length should be calculated based on the fracture energy associated with this branch and not the total fracture energy. This will be the case also for smaller tensile strengths. Re-calculating the characteristic lengths so that only the first branch is taken into consideration, the values 30 mm and 9 mm are obtained for the tensile strengths 3 MPa and 5 MPa, respectively. Clearly, with these characteristic values, the dimensionless curves and the curves obtained based on real mortar characteristics match nicely in the initiation phase and for prediction of first crack strength. It is noteworthy that the FCM predicts that for fine mortars, crack initiation and first crack strength is independent of the total fracture energy and only depends on the initial part of the cohesive crack, op to crack opening of 2.5 pm to 5 pm.

MATRIX CRACK PROPAGATION IN CEMENTITIOUS COMPOSITES When studying matrix crack initiation and propagation in fiber reinforced composites, the relevant physical system to model is illustrated in Fig. 1. Clearly, the bridging action of the fibers must now play a major role. Assuming, as substantiated in the previous section, that the FCM

232

Henrik STANG, Lars DICK-NIELSEN, Peter NOE POULSEN, John FORBES OLESEN

5.0E-08

-!? f

1 8

4.OE-06 3.OE-06

1.OE-06

%

O.OE+OO 0

Fig. 9.

10 20 a crack length [mn]

30

Crack opening profiles in a fine mortar as predicted by a bi-linear FCM for a crack extension of 7 mm, 12.5 mm and 25 mm in the case of ft = 5 MPa, 2 ~ 1 2x 2.5 pm. From [28].

can adequately model crack initiation and propagation in fine mortars, the FCM should also be used for crack initiation and propagation in fiber reinforced materials, as long as the largest crack openings associated with crack propagation in the fiber reinforced material do not prevent an accurate solution at the crack tip as well. As it will be shown in the following, this is typically not the case for cementitious composites. For regular fiber reinforced concrete it has been suggested to superpose the cohesive laws for the matrix and the smeared, average fiber bridging law due to debonding and pull-out, [29]. In that study it was suggested that such superposition cannot the done without taking into consideration initial fiber debonding taking place during matrix crack propagation, the so-called Cook-Gordon effect, [30]. In recent, detailed FEM simulations of crack propagation in fiber reinforced fine mortars (ECC-materials), [31], it has been shown that direct superposition of the cohesive law for the matrix and the smeared, average fiber bridging law is possible due mainly to the very small crack openings associated with matrix crack propagation next to a fiber, which allows the matrix crack to propagate past and around a fiber before initiation of the debonding and subsequent pull-out process. In [32] a resulting cohesive law for a single crack in a ECC type material was calculated using direct superposition of the matrix cohesive law (the same as used for the pure mortar investigations in the previous section, now with a tensile strength of 2.8 m a ) and a calculated, smeared, average fiber bridging law for a typical PVA fiber, based on the fiber pull-out model by Lin et al. [33]. The matrix cohesive law, the fiber bridging law and the resulting cohesive law is shown in Fig. 10. Implementing this cohesive law in the semi-analytical model outlined above with the usual FCM assumptions (a = 0), load-crack propagation and crack opening versus crack propagation curves where calculated in [32]. The curves are shown in Fig. 11. Interestingly, the charcteristic shape of the dimensionaless curves, Fig. 6 (a) and (b), reappear. It should also be noted that if the characteristic length of the composite material is estimated based on the first initial descending branch of the resulting cohesive law, a value of approximately 45 mm is obtained, which matches with the crack length at which a significant drop in the load is observed (about 50 mm). Again, the background for this phenomenon is that the crack opening associated with the first stable crack propagation is so small that crack propagation is governed by the first descending branch of the cohesive law. Further, it is interesting to note that crack does not reach a steady

233

Recent developments in the modeling of matrix crack propagation in Brittle Matrix Composites

..................;........................ 71

j.

........

iECc)

v&Total

2 ....................... .......................... Mortar

“0

50 w -t w l

j

.........

100

Fig. 10. Cohesive law for single crack initiation and propagation in an ECC-type material obtained from direct superposition of the the cohesive law for the matrix and the smeared, average fiber bridging law. From [32].

state propagation mode even at half crack lengths of 200 mm, and that at those crack lengths the maximum crack opening is predicted to be less than 15 pm. Finally it should be noted, that the predicted influence of initial defect size of the first crack strength is close to what has been observed in the literature, [lo].

15

0.5fflm

IOmm h:!?“== 1’5 mm I

O;

50

100

a - [mm]

150

2;o

(a) Far field stress as function of crack length.

‘0

50 a

100

- [mrn]

150

200

(b) Crack opening at the center of the initial defect versus crack length

Fig. 11. The influence of initial defect size on the load-crack propagation curves (a) and crack opening versus crack length curves (b) for a single crack in an ECC type material. From [321. It follows from the above that the crack length for which stable crack propagation takes place is heavily influenced by the initial slope of the resulting cohesive law, which again depends on the initial slope of the cohesive laws of the matrix and the fiber bridging law respectively. This dependency was also demonstrated in [32], where a variation of the initial slope of the matrix was performed (a1 = 78 mm-l, 156 mrr-l, 311 111111-l and 622 mm-’), keeping the tensile strength constant, resulting in the cohesive laws for the composite material shown in Fig. 12. As expected, the smaller the slope of the initial resulting cohesive law, the larger the char-

Henrik STANG, Lars DICK-NIELSEEN,Peter NOE POULSEN, John FORBES OLESEN

234

10

0'

20

30

40

50

Fig. 12. Cohesive law for single crack initiation and propagation in an ECC-type material obtained from direct superpositionof the the cohesive law for the matrix and the smeared, average fiber bridging law. A variation of the initial slope of the matrix has been performed. From [32].

acteristic length and the larger the crack length during stable crack propagation, see Fig. 13 (a). Interestingly, a drop in the first crack strength is observed for very large slopes. This phenomenon is associated with the fact that the hardening slope of the composite cohesive law is initiated at crack openings of 1.3 pm, 2.6 pm, 5.2 pm and 10.4 pm, respectively for the various cohesive laws, thus by comparing with the crack openings shown in Fig. 13 (b), we can see that in particular for the two cohesive laws with the largest slope, almost all the crack propagation is governed by a combination of the descending and the hardening part of the cohesive law invalidating the underlying assumptions for the dimensionless curves.

I 0 ;

So

100

a - [mm]

150

do

(a) Far field stress as function of crack length.

:0

50

a

-100 [mm]

150

do

(b) Crack opening at the center of the initial defect versus crack length.

Fig. 13. The influence of initial slope of the composite cohesive law on the load-crack propagation curves (a) and crack opening versus crack length curves (b) for a single crack in an ECC type material. From [32].

Recent developments in the modeling of matrix crack propagation in Brittle Matrix Composites

235

DISCUSSION AND CONCLUSIONS The applicability of cohesive crack models in the modeling of crack initiation and propagation in fine cementitious pastes, fine mortars and composites has been discussed and the consequences of applying such modeling have been pointed out. The most general form of cohesive crack models is the co-called bridged crack model which contains the Fictitious Crack Model (smooths crack closure, no energy dissipation at the crack tip) and LEFM (all energy dissipation is small scale dissipation, which consequently takes place at the crack tip) as special cases. The problem of initiation and propagation of a stress free slit in an infinite sheet under uniaxial tension was studied as a generic example, [28], [ 191. First, a bridged crack model with a simple linear cohesive law was studied. It was found that the stability of crack propagation and the crack opening associated with crack propagation is quite sensitive to small scale energy dissipation. While crack propagation in LEFM is always unstable, the FCM predicts that cracks grow to a total length of about the characteristic length, lch, before they become unstable. At this length the crack opening is only a fraction of the critical crack opening, w,,for typical flaw sizes. Thus, at critical crack length the FCM predicts the crack to be a weak plane in the matrix, rather than a stress free crack. The first crack strength is quite insensitive to initial flaw size in contrast to the prediction by LEFM. The initial flaw sensitivity is rapidly increased as small scale energy dissipation is introduced in the bridged crack model and the flaw sensitivity predicted by the FCM seems to be in line with experimental findings even though more date is needed to make firm conclusions about the applicability of bridged crack models and FCM. Single crack propagation in fiber reinforced cementitious materials such as ECC materials has been investigated with the FCM using a composite cohesive law obtained by superposition of the matrix cohesive law and the smeared, average fiber bridging law, [31], [32]. For typical micro-mechanical parameters the nature of crack propagation is surprisingly similar to pure matrix propagation. The stability of crack propagation is dominated by the initial part of the cohesive law and at loading corresponding to first crack strength, the crack opening is typically a few micron, depending on the initial flaw size. Even at total crack lengths of up to 400 mm, the maximum crack opening is of the order of 10 pm and steady state crack propagation is not achieved in the studies conducted here. As matrix toughness is reduced the first crack strength and the length of stable cracks are reduced. The crack opening during crack propagation, however, remains of the same order of magnitude. Influence of the shape of the load-crack propagation curves on the multiple cracking ability of the composite material has not been investigated in the current studies and is not yet clear. It is possible, though, that the early instability and initial flaw size sensitivity which follows from a small characteristic length of the matrix promotes multiple cracking in line with experimental observations, [lo].

REFERENCES 1. A. Hillerborg, M. ModCer, and P.E. Petersson. Analysis of crack formation and crack growth in concrete by means of fracture mechanics and finite elements. C e m . Concl: Rex, 6(6):773-782, 1976. 2. RILEM-Committee-TDF-162.Test and design methods for steel fiber reinforced concrete. design of steel fibre reinforced concrete using the 0 - w method: Principles and applica-

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Henrik STANG, Lars DICK-NELSEN, Peter NOE POULSEN, John FORBES OLESEN

tions. Matel: Struc., 35(249):262-278, 2002. Prepared by RILEM-Committee-TDF-162, Chairlady L. Vandewalle.

3. G.N. Wells and L.J. Sluys. A new method for modelling cohesive cracks using finite elements. International Journal for Numerical Methods in Engineering, 50( 12):2667-2682, 2001. 4. N. Moes and T. Belytschko. Extended finite element method for cohesive crack growth. Engineering Fracture Mechanics, 69(7):813-833,2002. 5. V.C. Li. On engineered cementitious composites (ECC). A review of the material and its applications. Journal of Advanced Concrete Technology, 1(3):215-230, November 2003.

6. V.C. Li and T. Kanda. Engineered Cementitious Composites for structural applications. Journal of Materials in Civil Engineering, 10(2):66-69, 1998. 7. B.F. Dela and H. Stang. Wo-dimensional analysis of crack formation around aggregates in high-shrinkage cement paste. Engineering Fracture Mechanics, 65(2-3):149-164,2OoO. 8. V.C. Li and C.K.Y. Leung. Steady-state and multiple cracking of short random fiber composites. Journal of Engineering Mechanics, 1 18(11):2246-2264, 1992.

9. H.-C. Wu and V.C. Li. Stochastic process of multiple cracking in discontinuous random fiber reinforced brittle matrix composites. International Journal of Damage Mechanics, 4(1):83-102, 1995. 10. S. Wang and V.C. Li. Tailoring of pre-existing flaws in ECC matrix for sturated strain hardening. In Kaspar J. Willam Sarah L. Billington Victor C.Li, Christopher K. Leung, editor, Fracture Mechanics of Concrete Structures, volume 2,pages 1005-1012.The Fifth International Conference on Fracture Mechanics of Concrete and Concrete Structures, IaFraMCos, 2004. 11. H. Tada, P. Paris, and G. Irwin. The Stress Analysis of Cracks Handbook. Paris Productions Incorporated, 226 Woodbourne Dr., St Louis, Missouri, USA, 1985.

12. A. Hillerborg. Analysis of fracture by means of the fictitious crack model, particularly for fibre reinforced concrete. The Int. J. Cem. Comp., 2(4): 177-184, 1980. 13. B. L. Karihaloo. Fracture Mechanics & Structural Concrete. Concrete Design and Construction Series. Longman Scientific & Technical, Harlow, Essex, England, 1995.

14. G. I. Barenblatt. The mathematical theory of equilibrium cracks in brittle fracture. Adv. Appl. Mech., 756-129, 1962.

15. D. S.Dugdale. Yielding of steel sheets containing slits. J. Mech. Phys. Solids, 8:lOO-104, 1960. 16. B.N. Cox and D.B. Marshall. Stable and unstable solutions for bridged cracks in various specimens. Acta Metallurgica et Materialia, 39(4):579-589, 1991. 17. B.N. Cox. Scaling for bridged cracks. Mechanics of Materials, 15(2):87-98, 1993.

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18. B.N. Cox and D.B. Marshall. Concepts in the fracture and fatigue of bridged cracks. Acta Metallurgica et Meteriala, 42:341-363, 1994. 19. H. Stang, J.F. Olesen, P.N. Poulsen, and L. Dick-Nielsen. On the application of cohesive crack modeling in cementitious materials. Materials and Structures, 2006. Accepted for publication. 20. V. C. Li and M. Maalej. Toughening in cement based composites. part i: Cement, mortar and concrete. Cement & Concrete Composites, 18:223-237, 1996. 21. G. Bao and Z. Suo. Remarks on crack-bridging concepts. Appl. Mech. Rev., 45(8):355-366, 1992. 22. R. Ballarini and S. Muju. Stability analysis of bridged cracks in brittle matrix composites. Journal of Engineering for Gas Turbines and Powel; Transactions of the ASME, 115(1):127-138, 1993. 23. B. L. Karihaloo, J. Wang, and M. Grzybowski. Doubly periodic arrays of bridged cracks and short fibre-reinforced cementitious composites. Journal of the Mechanics and Physics of Solids, 44(10): 1565-1586, 1996. 24. A. Carpinteri and R. Massabo. Continuous vs discontinuous bridged-crack model for fiber-reinforced materials in flexure. International Journal of Solids and Structures, 34(18):2321-2338, 1997. 25. A. Carpinteri, G. Ferro, and G. Ventura. The bridged crack model for the analysis of fiber-reinforced composite materials. Advances in Composite Materials and Structures VII. Seventh International Conference. CADCOMP VII, pages 301-10,2000. 26. E. K. Gamstedt and S. Ostlund. Fatigue propagation of fibre-bridged cracks in unidirectional polymer-matrix composites. Applied Composite Materials, 8(6):385-410,2001. 27. G. Ferro. Multilevel bridged crack model for high-performance concretes. Theoretical and Applied Fracture Mechanics, 38(2):177-190, 2002. 28. L. Dick-Nielsen, P.N. Poulsen, H. Stang, and J.F. Olesen. Semi-analytical cohesive crack model for the analysis of first crack strength of mortar. In Proceedings of the 17th Nordic Seminar on Computational Mechanics, pages 183-1 86. KTH Mechanichs. Stockholm, Sweden, 2004. 29. V.C. Li, H. Stang, and H. Krenchel. Micromechanics of crack bridging in fiber reinforced concrete. Mat. and Struc., 26( 162):486-494, 1993. 30. J. Cook and J. E. Gordon. A mechanism for the control of crack propagation in all brittle systems. Proc. Roy. SOC.,282A:508-520, 1964.

31. L. Dick-Nielsen, H. Stang, and P.N. Poulsen. Micro-mechanical analysis of fiber reinforced cementitious composites using cohesive crack modeling. In Proceedings of the Knud Hjgaard conference. Department of Civil Engineering, Technical University of Denmark, 2005.

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32. L. Dick-Nielsen, H. Stang, and P.N. Poulsen. Condition for strain-hardening in ecc uniaxial test specimen. In Proceedings of EFC16, PMMMA Special Symposium, 2006. 33. Z. Lin, T. Kanda, and V.C. Li. On interface property characterization and performance of fiber-reinforced cementitious composites. Concrete Science and Engineering, 1: 173-1 84, 1999.

Proc. Int. Symp. 'Brittle Matrix Composites 8" A.M. Brandt, re. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ.. Warsaw 2006

COMPRESSIVE TOUGHNESS OF FIBRE REINFORCED CONCRETE UNDER IMPACT LOADING Lihe ZHANG and Sidney MINDESS Department of Civil Engineering University of British Columbia 6250 Applied Science Lane Vancouver, British Columbia V6T 124, Canada e-mail: [email protected]

ABSTRACT Test standards for plain concrete and fibre reinforced concrete (FRC) under static compressive loading are well defined. However, there is no standard test method for either plain concrete or FRC under compressive impact loading. In the present research, a method for determining both the compressive strength and the compressive toughness of FRC under impact loading was developed, using an instrumented drop weight impact machine, supplemented by a high speed video camera. Using this equipment, it was possible to determine the performance in compression of FRC subjected to impact loading. It was also possible to estimate the internal damage to the specimens due to impact loads that were not high enough to cause complete failure of the specimens. It was found that the total deformation of FRC specimens under impact was greater than that of plain concrete specimens. However, deformation to failure under impact loading was less than that under static loading.

Keywords Fibre reinforced concrete, compression, impact loading, compressive toughness INTRODUCTION It is easy to measure the compressive strength of concrete under quasi-static loading. However, because of the brittle nature of concrete, it is very difficult to measure the post-peak load vs. deflection curve in compression, even with a stiff, closed-loop testing machine; the behaviour obtained seems to depend in large part on the characteristics of the particular test apparatus. While there exists a standard test method for determining the compressive toughness of fibre reinforced concrete [I], which is much less brittle than plain concrete, such tests are rarely carried out (in large part because no one knows what to do with the results). There has been even less work on determining either the compressive strength or the compressive toughness of concrete under impact loading because of the experimental difficulties, since the entire loading event is typically less than five milliseconds (ms). There have been a few studies reported in the literature on both plain concrete and fibre reinforced concrete (FRC) [2-81. For instance, Bischoff and Perry [3] found that under compressive impact loading, the compressive strength of plain concrete increased by 85loo%, but the critical axial strain values ranged from a decrease of 30% to an increase of 40%, compared to static loading. For FRC tested using a split Hopkinson pressure bar [2-3,

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Lihe ZHANG, Sidney MINDESS

81, it was found that post-peak ductility was absent at strain rates greater than ~ O S - ' , apparently because the fibres could no longer bond to the concrete fragments [8]. Unfortunately, the reported results have been inconsistent, since the various investigators used quite different testing techniques and instrumentation. Some investigators [2-3,5, 81 used strain gauges to measure surface strains. While this method provides an accurate measure of the stress vs. strain response up to the peak load, the brittle nature of the fracture makes it impossible to measure the strain (and hence the toughness) beyond the peak load. Other investigators [4, 6-71 used a direct measurement of the displacement between the loading platens to estimate the concrete strains. However, this provides misleading results, as the platen-to-platen deformation is not the true material deformation; it reflects the total displacement of the testing machine (see below). In the work reported here, a method is described that permits the true deformation of the specimen to be measured, in both the pre-peak and post-peak regions, using a high speed video camera. EXPERIMENTAL PROCEDURES Specimens The test specimens were standard lOOmm x 20Omm cylinders, with a maximum aggregate size of 19mm. Three different mixes were tested, with static 28-day compressive strengths ranging from about 60MPa to about 120 MPa. Some specimens were also reinforced with fibres (either steel or polypropylene), at volume fractions of 0.5% and 1.O%. Impact machine An instrumented drop weight impact machine (Fig. l), designed and constructed at the University of British Columbia, was used for the impact tests. The machine, which has been described in detail in [9], is capable of dropping a 578kg mass from heights of up to 2.3m. A cylindrical load cell, with a diameter of l O O m m , was mounted on the bottom of the falling hammer to obtain the applied loads.

Two independent systems were used to record the impact event. The load vs. time data were recorded using a high speed data acquisition system, which recorded the load at a frequetcy of lo5Hz.The deformation vs. time history was obtained using a high speed video camera at 20,000 kames per second. Commercial softwaret was used for the image analysis. The complete load vs. deformation curve was then obtained by combining these two records.

Phantom V 4.2, Model No: VR408 1888V42M, produced by Vision Research, Inc., USA TEMA Photosonics Company, Sweden

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Compressive toughness offibre reinforced concrete under impact loading

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Fig. 1 Schematic view of instrumented impact machine.

Deformation vs. time To obtain a good estimate of the deformation vs. time history, ten points on the specimen surface were tracked simultaneously, located as shown in Fig. 2. The two lines of points were l O O m m apart, so that the data could be compared to that obtained from static tests carried out as described in [I]. The true deformation of the l O O m m high central portion of the cylinder is then the difference between the displacement of the upper line of points and that of the lower line, as shown in Fig. 3. The appearance of a plain concrete specimen while undergoing impact failure is shown in Fig. 4.

Lihe ZHANG, Sidney MINDESS

242

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243

Compressive toughness offibre reinforced concrete under impact loading

Fig. 4 Impact failure of high strength plain concrete It should be emphasized here that deformation measurements based on the total displacement between the drop hammer and the machine base significantly overestimate the true specimen deformation, as they are much greater than those determined from the high speed video record. This is shown in Fig. 5 for a steel FRC specimen under a drop height of lm. The platen-to platen deformation may include not only the specimen deformation, but also local crushing and settlement of the specimen on the machine base, and machine deformations. It is thus essential that direct measurement of the specimen deformation be obtained for a proper analysis of compressive impact behaviour.

- - Def-Hammer De f-Cylinder

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Fig. 5 Displacement comparison of drop hammer and FRC cylinders under impact

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Lihe ZHANG, Sidney MINDESS

Load vs. time A typical load vs. time curve for a FRC specimen is shown in Fig. 6. It may be seen that the total impact event had a duration of only about 2.7 ms, and that the duration of the post peak portion of the event is very short, of the order of 0.2 ms. This reflects the very brittle failure of the FRC under impact loading. 800 r

700 600

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300 200 100

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600 500

400 .c1

8

300 200 100 0 0.0

0.1

0.2

0.3

0.4

Deformation (mm)

Fig. 7 Load vs. deformation for FRC with a 500 mm drop height DISCUSSION OF RESULTS Deformation The deformation results from these tests are shown in Table 1. It may be seen that the total deformation of the FRC specimens under impact loading is significantly greater than that of plain concrete; the differences between the steel fibre concrete and the polypropylene fibre

245

Compressive toughness offibre reinforced concrete under impact loading

concrete are not significant. These deformations are about 25% lower than the deformations at peak load under static loading, again showing the increased brittleness of the material under impact loading. It may also be seen from Fig. 7 that the pre-peak behaviour of the steel FRC is somewhat irregular and non-linear. This reflects, in part, the anisotropic nature of the concrete itself. Previous research [lo] has also shown this type of non-linear behaviour. It may be that, under impact, considerable microcracking develops, distributed randomly within the specimen; these cracks then localize in several planes, as shown in Fig. 4, leading to failure. However, this is an area that requires fiuther investigation. Table 1. Deformation of plain concrete and FRC

Steel FRC, 0.5%

Steel FRC, 1.0%

PP

PP

FRC, o.5%

FRC,

0.10

0.2182

0.2025

0.2124

0.2486

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0.2703

0.2655

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Plain concrete Total deformation, under impact compression,

1.O%

(mm) Deformation at peak load, under static compression, (mm)

Internal damage Some of the specimens remained intact, though damaged, after impact loading. A typical load vs. deformation curve of such a steel FRC specimen is shown in Fig. 9. Again, it may be seen that deformation is non-linear up to the maximum load reached, and then decreases also in a non-linear fashion; there is a considerable amount of irreversible deformation, in this case equal to about 0.2mm. This indicates that there has been significant internal damage to the specimen, primarily in the form of microcracks, as mentioned earlier. Upon static reloading such specimens, a considerable reduction in elastic modulus was found, consistent with internal microcracking.

246

Lihe ZHANG, Sidney MINDESS

Before contact,

2ms,

5ms,

Oms, load increases

3ms, Peak load is reached

8ms,

lms,

4ms, load drops to zero

20ms,

Fig. 8 Progressive failure process under 500mm drop height at 20,000 frameslsec

247

Compressive toughness offibre reinforced concrete under impact loading

600 500

I-1

200

100 0 0

0.1

0.2 0.3 Deformation (mm)

0.4

0.5

Fig. 9 Normal strength steel fiber reinforced concrete under a 250 mm drop height Scabbing For those specimens that did fail under impact loading, about two-thirds of the specimen mass was lost through scabbing, as shown schematically in Fig. 10. The appearance of the cylinder just sfter impact is shown in Fig. 11. Using the high speed video camera and the associated software, it was possible to track the average velocity of the scabbed particles. Typical velocities were in the range of 2 - 6 d s . It was found, however, that the kinetic energy consumed in scabbing represented less than 5% of the fracture energy of the FRC cylinders (i.e., the energy lost by the falling hammer).

FRC cylinders at peak load FRC cylinders after peak load

8% c*l

0 0

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248

Lihe ZHANG, Sidney MINDESS

Fig. 11 Tracking the scabbing velocity of FRC cylinders. C0NCLUSI 0N S 1. Determining the specimen deformation under compressive impact using platen-to-platen measurements yields deformations that overestimate the true specimen deformations. Direct measurement using a high speed video camera provides a much more accurate deformation vs. time history. 2. The total axial deformation under impact loading is substantially less than that under static loading. 3. FRC specimens underwent more deformation than plain concrete specimens before failure. 4. Only a small amount of the energy lost by the drop hammer was consumed by scabbing of the specimens.

REFERENCES 1. Japan Society of Civil Engineers, Method of tests for compressive strength and compressive toughness of steel fiber reinforced concrete. Standard SF-5, JSCE Concrete Library, No. 3,1984, pp. 63-66. 2. Bischoff, P.H., Perry, S.H., Compressive strain rate effects of concrete. In: Symposium Proceedings Vol. 64, Cement-Based Composites: Strain rate Effects on Fracture, S. Mindess and S.P. Shah eds. Materials Research Society, Pittsburgh, PA 1986, pp. 151-165. 3. Bischoff, P.H., Perry, S.H., Compressive strength of concrete at high strain rates. Materials and Structures, 24, 1991, pp. 425-450. 4. Campione, G., Mindess, S., Compressive toughness characterization of normal and highstrength fiber concrete reinforced with steel spirals. In: ACI SP-82, Structural Applications of Fiber Reinforced Concrete, N. Banthia and C. MacDonald eds. American Concrete Institute, Farmington Hills, MI 1999, pp. 141-161. 5. Bischoff, P.H., Perry, S.H., Impact behavior of plain concrete loaded in uniaxial compression. J. Engineering Mechanics, ASCE, 121, 1995, pp. 685-693. 6. Celik, T., Marar, K., Eren, O., Relationship between impact energy and compression toughness energy of high-strength fiber-reinforced concrete. Materials Letters, 47, 2001, pp. 297-304. 7. Fujikake, K., Mindess, S., Xu, H.F., Analytical formulation for concrete confined with steel spirals subjected to impact loading. In: Design and Analysis of Protective Structures against

Compressive toughness offibre reinforced concrete under impact loading

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ImpactlShock Loads, T. Ohno, T. Krauthammer and T.C. Pan, eds. Daioh Co., Lte., Tokyo, 2003, pp. 500-507. 8. Lok, S.T., Zhao, P.J., Impact response of steel fiber-reinforced concrete using a split Hopkinson pressure bar. J. Materials in Civil Engineering, 16,2004, pp. 54-59. 9. Banthia, N., Impact Resistance of Concrete. Ph.D. Dissertation, University of British Columbia, Vancouver, Canada, 1987. 10. Hamelin, P., Razani, M., Impact behavior of metallic fiber reinforced concrete and mortar. In: Symposium proceedings, Vol. 2 11, Fiber-Reinforced Cementitious Materials, S . Mindess and J. Skalny eds. Materials Research Society, Pittsburgh, PA, 1990, pp. 133-137.

Proc. Int. Symp. “Brittle Matrix Composites 8” A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

BEHAVIOR UNDER COMPRESSION AND BENDING LOADS OF MULTI-SCALE HIGH PERFORMANCE STEEL FIBER REINFORCED CEMENT BASED COMPOSITES Flavio de A. SILVA, Sidiclei FORMAGINI, Romildo D. TOLEDO FILHO and Eduardo de M. R. FAIRBAIRN Department of Civil Engineering Albert0 Luiz Coimbra Institute - Graduate School and Research in Engineering (COPPE) Federal University of Rio de Janeiro (UFRJ) P.O. Box 68506, ZIP-CODE: 21945 - 970, Rio de Janeiro, RJ, Brazil e-mail: [email protected]

ABSTRACT In this work high performance cementitious composites reinforced with randomly dispersed steel fibers with single and multi-scale geometry were developed and their mechanical behavior was characterized. The steel fiber volume fraction ranged &om 2% to 3.5% presenting a meso and macro scale with an aspect ratio of 65 and lengths of 12 mm and 35 mm, respectively. In the micro scale level the wollastonite fiber was used as reinforcement. To design the ultra-compact cementitious matrix (compressive strength of 160 MPa, elastic modulus of 47 GPa and equivalent elastic post cracking bending stress of 35 MPa) the concept of maximum granular packing of grains was used. The matrix was of river sand with particle size ranging from 150 to 600 pm, silica flour, slag cement, silica fume and a polycarboxilate superplasticizer. The waterhinder ratio of the self-leveling composite was 0.17. The behavior under tension loads was determined from four point bending tests. Compression tests were performed to determine the composites’ modulus of elasticity and its ultimate strength.

Keywords Cementitious composites, steel fiber, multi-scale reinforcement INTRODUCTION Historical background The first works on fiber reinforced concrete (FRC) were realized in the fifty’s and sixty’s decades of the last century with the aim of understanding the mechanical behavior of steel fiber reinforced concrete [1,2]. Since that period, other fibers have been evaluated as reinforcement in concrete elements, but the steel is still the most used fiber. Its popularity is associated with the fact that steel presents a good affinity with concrete, which was proven in rebar and pre-stressed concrete, the ease of use, the high toughness and resistance to static and dynamic loads. In the last twenty-five years significant improvements in the development of cement based materials have been observed originating the high performance concrete that can present uniaxial compressive strength ranging from 150 to 400 MPa [3,4,5]. This improvement was

252

F. de A. SILVA, S . FORMAGINI, R. D. TOLEDO FILHO, E. de M. R. FAIRBAIRN

only possible due to developing techniques of cement paste microstructure densification using efficient superplasticizing chemical additives and ultra-fine particles [5]. As Buitelaar stated, in the beginning of the sixty’s, the concretes presented high porosity which resulted in a low mechanical resistance and hence, in a low durability, once the chemical agents penetration was easy. The usual uniaxial compressive ultimate strength at that time was 15 MPa and the waterkement ratio ranged from 0.50 to 0.75. According to Buitelaar, in 1964 Bache from The Concrete Research Laboratory of Denmark obtained a concrete with compressive strength ranging from 60 to 80 MPa and waterkement ratio of 0.30 [5]. Between 1967 and 1972, several researches were realized combining pressure and vibration to compact the concrete and, with this technique, compressive strengths of 100-130 MPa were obtained by Bache and his collaborators, for concretes produced with low cement and high aggregate ratios. In the beginning of the seventy’s the superplasticizerwas invented and, around 1975, it was possible to produce in Japan, concretes with waterkement ratio of 0.25 and compressive strength up to 120 MPa. In the end of the sixty’s, with the development of the superplasticizers, it became possible to disperse ultra-fie particles as well as to enhance the dispersion of cement particles in water solution allowing a dense packing of the grains among the cement particles [5]. Bache and his collaborators began their studies aiming the densely compacted mixes using blends of normal cement and ultra-fine cement particles. In a more advanced stage, blends of cement and silica fume were used and, in 1987, Bache et. a1 obtained concretes with compressive strength of 128 MPa after one day thermal cure. Later on, using high strength aggregates, like calcined bauxite, it was possible to mix concretes with compressive strength up to 280 MPa. The dificulty of improving the concrete compressive strength was then overcame. Nevertheless, another problem was beginning to gain importance as the compressive strength was raised: the fragility of the material. To overcome this problem, the solution was to add multi-scale fiber reinforcement [4,5,6]. According to Rossi [4] the fiber addition to the densely packed matrices, known as DSP (Densified Small Particles), resulted in the Ultrahigh Performance Fiber Reinforced Concrete (UHPFRC) and Ultra High Performance Fiber Reinforced Cement Composites (UHPFRCC). Fiber Reinforced Concrete: taxonomy and characteristics

Several categories of fiber reinforced cementitious composites (FRCC) have been developed over the past three decades presenting different mechanical properties. Conventional fiber reinforced concretes (FRC) presents an increase in the ductility when compared with the plain matrix showing a strain softening behavior after the appearance of the first crack and in some cases a decrease in the ultimate strength. On the other hand the high performance fiber reinforced cementitious composites (HPFRCC) exhibit a strain-hardening type of response accompanied by multiple cracking in tension which leads to an improvement in strength and toughness compared to the non-reinforced matrix. New terms have been suggested to classify FRC that presents a multiple cracking behavior. Ductile fiber reinforced cement composites (DFRCC) was proposed by several researchers to describe a class of FRC that presents a multiple cracking in bending [7,8,9]. A class of ultra ductile FRCC was developed by Li [ 10, 11,121 and is called engineered cementitious composites (ECC). The ECC is reinforced by synthetic fibers with a fiber volume fraction as high as 2% and presents a multiple cracking behavior in bending and direct tension achieving up to 5 MPa and 5% of strains in direct tension [121. The slurry infiltrated fiber concrete (SIFCON) and slurry infiltrated mat concrete (SIMCON) are a special class of concrete produced by infiltrating slurry in pre-placed steel fibers

Behavior under compression and bending loads of multi-scale high perjomance steel fiber reinforced ...

253

formwork. Its fiber volume fraction can reach up to 20% producing a concrete with high compressive strength that can reach 2 10 MPa [ 131. Ductal is a type of reactive powder concrete produced using the concept of optimal packing theory. Its mechanical properties are characterized by high compressive strength (210 MPa), high bending strength (45 MPa) and high ductility [14]. A multi-scale reinforced cement composite was developed by Rossi at LCPC-France [151. Two types of materials were developed by using the multi-scale concept. The MSCC (Multi Scale Cementitious Composites) which was reinforced by 7% of two metal fibers of different geometries, and the CEMTEC that was reinforced by 11% of three classes of steel fibers [ 16,171. Both materials present a tension hardening behavior but the MSCC can achieve up to 15 MPa under direct tension while the CEMTEC up to 20 MPa . A multi-scale reinforcement was also proposed by Kawatama [ 181 using polyethylene mixed with steel cords and PVA mixed with steel cords, using 2% of fiber volume fraction. Ultimate tensile strength of 4 MPa and strains up to 4% were achieved. In the present research the multi-scale reinforcement approach was used to produce the Multi Scale High Performance Cement Composite (MSHPCC). Two geometries of steel fibers were used: hooked end 35 mm and plain 12 mm fibers, both presenting an aspect ratio of 65. The wollastonite fiber was used to counteract the micro-cracks. Three types of materials were developed with steel fiber volume fraction ranging from 2% to 3.5% and constant wollastonite fiber volume fraction of 2.6%. The compressible packing model developed by de Larrard at LCPC [19] was used to design the self leveling matrix used for all the produced composites. The rheology in the fresh state of the material was determined from the L-box and its spreading trough the inverted Abram’s cone. Compression and bending tests were performed to determine the first crack strength, ultimate strength and toughness of the composites. THE COMPRESSIBLE PACKING MODEL

The compressible packing model (CPM) was developed by de Larrard and his collaborators and used in this research to design the matrix of the MSHPCC [19,20]. Composite materials like concrete are made up of grains embedded in matrix. The aim of the design is to use the least possible amount of binder by combining these grains in order to minimize the concrete porosity [19]. The equation representing the virtual packing density of a granular mix containing n classes of grains, ordered in such a way that its diameters are d , 2 d, 2 ..... 2 d , 2 d,+,2 ..... 2 dn, when the class i is dominant, is expressed by the following Equation:

where:

f ) is the virtual packing density when the ifh class is dominant; yi is the volumetric fraction of the h‘i class; fl is the virtual packing density of the ithclass; it represents the volume of grains contained in an unitary volume, compacted with an ideal compaction energy that would correspond to a maximum virtual packing; aid and bij represent the loosening effect and the wall effect exerted by the grains, respectively; they can be determined either experimentally or by the following formulas:

254

E.de A. SILVA, S. FORMAGINI, R.D. TOLEDO FILHO, E. de M. R. FAIRBAIRN

-/,

a,,,=

b,,j = 1- (1 - di/ d j ),.so

The virtual compactness of the mix can be found by using the formula:

where inf indicates the least value. The actual compactness depends on three main parameters: the size of the grains, the shape of the grains, and the method of processing the packing. The compressible packing model allows making the transition ffom virtual compactness, which can not be obtained in practice, to the actual compactness of the mix, which depends on the energy being applied at the time of placing. A scalar K called compaction index enables connecting the virtual compactness (9 with the actual compactness (@. This scalar is strictly dependent on the protocol implemented for the particular mix. As K tends to infinity, the compactness 4 tends to the virtual compactness The general shape of the compaction index equation, for n classes of grains, is as follows:

where 4 is the actual compactness of the granular mix. The values of index K are calculated from the binary mixes for each placing processes. K assumes a value of 4.5 when the compaction process is the simple pouring, 6.7 for water demand and 9 when the placing process is vibration plus lOkPa compression [ 191. If the actual compactness for a single granular class i (6)is experimentally determined, by means of a compaction process having compaction index K,it is possible to use equation (5), derived from equation (4), to determine the virtual compactness of the granular class i.

pi=-(1+K) 4; K

Equation (4) is an implicit equation in 4 and allows the determination of the actual compactness since the other variables are all known. To use the model it was determined the virtual compactness, size grading distributions and specific gravity of the constituents as well as the cement contribution to compressive strength and the saturation dosage of the chemical additive.

Behavior under compression and bending loads of multi-scale high performance steel fiber reinforced. ..

255

EXPERIMENTAL PROCEDURE Materials

The matrix was designed using a Brazilian slag cement type CPIII 40, river sand with two classes of particle size: one ranging from 150 to 300 pm and the other from 425 to 600 pm, silica flour (ground quartz), silica fume, and a polycarboxilate superplasticizer with solid contents of 32.5%. The waterhinder ratio of the self-leveling composite was 0.17. The Figure 1 shows the grain size distribution of the powder materials.

I

10

1

I00

1000

Diameter (pm)

O.l

Figure 1 - Grain size distribution of the powder materials. The steel fibers were produced by Dramix. Two lengths of steel fibers were used: 12 mm and 35 mm (with hooked ends) both presenting an aspect ration of 65. The mineral micro-fiber of wollastonite JG was obtained from Energyarc and used as reinforcement in all materials. The MSHPPC mix compositions are presented in Table 1. Table 1 - Composition of the MSHPPC. Composites

SC

S.Flour SJume

SP

SI

@g/m3)

@g/m3) @g/m3)

@g/m3)

@g/m’)

S2

SP

Water

Steel Fiber

Steel Fiber

W

@g/m3) @g/m3) fight3) (12mm) (35 mm)

2% M1 1011 - 2.6% 80 58 50 60 823 50 175.5 M2 1011 80 58 50 60 823 50 175.5 2% 1% 2.6% M3 1011 80 58 50 60 823 50 175.5 2.5% 1% 2.6% SC = Slag cement, %Flour = Silica flour,S.fume = Silica fume, S1 = sand (150 - 300) S2 = sand (425 - 600), W = wollastonite, SP = Superplasticizer

Processing

The MSHPCC was produced using a 100 liters planetary mixer. The exact time of pouring the materials into the mixer was monitored and controlled by the required power demand. This procedure was determined in a previous work [21], with the objective of standardizing the exact time of adding the powder materials, water, superplasticizer and fibers. For the present work the dry cementing materials and the wollastonite fibers were first poured inside the

256

F. de A. SILVA, S . FORMAGINI, R.D. TOLEDO FILHO, E.de M.R. FAIRBAIRN

mixer. The mixer was turned on for the homogenization of the dry mix for one minute. Following this, half of the water plus half of the superplisticizer were slowly added to the running mixer. After nine minutes the other half of the water was added. The remaining superplisticizer was added at 17 minutes, and at 20 minutes the steel fibers were carefully poured into the mixer.

Test Set-Up The self-leveling capacity of the MSHPCC was determined using the L-box test (cf. Figure 2 (a)). In this test, 12 liters of MSHPCC was mixed and poured inside the L-box. The time elapsed for the concrete to flow to the other section of the box as well as the different hights were measure. The spread of the material was determined trough the inverted Abram’s cone test (cf. Figure 2 (b)).

Figure 2 - Rheology tests set-up: (a) L-Box test and (b) Inverted Abram’s Cone. A Shimadzu UH-F 1000 kN was used to perform the compression and bending tests. The compression tests were carried out at a crosshead rate of 0.0025 W m i n whereas for the bending test a crosshead rate of 0.5 d m i n was selected. Three cylindrical samples with 50 mm of diameter and 100 mm height were tested under compression load. The deflections were measured using two electrical transducers (LVDT) positioned as shown in Figure 3 (b) and the loads and corresponding deflections were continuously recorded using a 32-bit data acquisition system taking four readings per second. Three specimens with nominal dimensions of 50 mm x 50 mm x 230 mm (width x thickness x length) were tested under bending (180 mm span) as shown in Figure 3 (a). Deflections at mid-span were measured using a LVDT and the data acquisition system was the same as the one used for the compression tests.

Behavior under compression and bending loads of multi-scale high pe$ormance steel fiber reinforced...

257

Figure 3 - Mechanical tests set-up: (a) four-point bending and (b) uniaxial compression. RESULTS AND ANALYSIS The results of the rheology tests are presented in Table 2. It can be seen that the composites M1 and M2 can be classified as self compacting concrete (SCC) since their spreads were over 55 cm and the ratio H2/H1 obtained from the L-box test was above 0.8. The inclusion of 1 % of the 35 mm fiber did not impair the spreadability and flow capacity of the composite. The M3 composite did not leveled in the L-box test and presented a spread below 55 cm, therefore it can not be classified as SCC. Table 2 - Rheology results: L-box and inverted Abram cone tests. Composites L-box (timefor leveling in sec.) L-box (HJH,) 30 1 MI M2 40 0.83 Did not leveled 0 M3

Spread (cm) 13 64 41

The results of the four-point bending tests for the different composites studied in this work: M1, M2 and M3 can be seen in Figure 4. For all the composites it was noticed a multiple of the composite MI cracking behavior with strain-hardening. The first crack strength (sr) was 152 % and 184 % higher than the M2 and M3, respectively. There were no benefits in the ultimate (s;)strength when the fiber volume fraction was increased as can be seen in Table 3 (determined of the maximum load carried out by the composites after the first crack event using the bending formula $I = 6M/bd2, where b is the width and d is the thickness). The average ultimate strength ranged from 31.58 to 35.99 MPa for the studied composites. If the standard deviation is taken into account, it can be concluded that the three composites presented the same magnitude of the ultimate bending strength.

F. de A. SILVA, S. FORMAGINI, R. D. TOLEDO FILHO, E. de M.R.FMRBMRN

258

4%

5

'

'

'

8

'

'

8

'

'

'

'

'

L

40 -

O

~ 0

., 1,

. ,

2

. 3, . 4,

. ,

5

. 6, . 7,

,

r

0

8

1

2

Displacement (mm)

3

4

5

6

7

6

Displacement (mm)

(a)

(b)

45 40

Displacement (mm)

(c) Figure 4 - Four-point bending test results: (a) composite MI, (b) composite M2 and (c) composite M3. The displacement which corresponds to the ultimate strength (d,) was approximately 2.5 times higher for the M2 and M3 composites when compared to the M1. This behavior indicates a higher capacity of energy absorption when the fiber volume fraction was increased from2 % to 3 %and to 3.5 %. Table 3 - First crack and post crack average strength and their respective displacements obtained in the four-point bending tests. ComPosites

M1 M2 M3

First Crack SD &,. (I&&) ( m a ) (mi) 26.30 3.13 0.044 1.37 0.068 17.24 1.17 0.051 14.25

o,,

Post-crack SD (MPa) ( m a ) 35.99 2.65 36.51 4.16 2.96 31.58 0 ,

6,, doer (m) 0.252 1.37 0.605 2.11 0.639 2.21

The toughness of the composites was calculated using the RILEM recommendations [22] and the ASTM (21018 [23]. Following the RILEM the toughness was calculated from the area under the load-deflection curves obtained under bending up to a post-peak deflection corresponding to 40 YOof the peak load and the results are presented in Table 4.The inclusion

Behavior under compression and bending loads of multi-scale high pe~ormancesteel fiber reinforced...

259

of the 35 mm fiber to produce the multi-scale reinforcement in the composites M2 and M3 resulted in a toughness increase of approximately two times when compared to the composite M1 which was only reinforced by the 12 mm fiber. The 35 mm fibers were able to bridge the cracks formed after the ultimate strength. The additional 0.5 % included in the M3 composite did not improve the toughness calculated following the RILEM recommendations.

M2

23.13

3.73

The toughness indexes were calculated following the ASTM C 1018 and the results are presented in Table 5. The indexes 15, 110, I20 and I 30 were calculated as the ratio of the area under the load-deflection curve up to 3, 5.5, 10.5 and 15.5 times the displacement corresponding to the first crack strength (dcr) by the area calculated under the same curve up to the first crack event, respectively. It can be seen that for small displacements, which occurs before the ultimate bending strength (I5 and IlO), the toughness indexes were approximately in the same range for the three composites. For displacements after the ultimate strength (I20 and 130) the toughness indexes presented a considerable increase. For the I20 index M2 and M3 showed an increase of 13.6 % and 11.6 %, respectively, when compared to M1. In a similar behavior the I30 toughness index showed an increase for M2 and M3 of 13.7 % and 12.8 %, respectively. This behavior confirms the assumption that for displacements beyond the ultimate strength the 35 mm fibers increase the capacity of energy absorption by bridging the macro cracks and an increase beyond 2% in the 12 mm fiber does not enhance the toughness performance. Table 5 - Toughness indexes calculated using ASTM C 10 18. Composites

M1 M2 M3

1 ' '

(kN) 14.00 11.97 9.89

'cr Is (mm) Average

0.044 0.068 0.051

4.87 5.95 4.77

110

SD Average 0.69 11.60 0.76 14.71 0.002 11.83

120

SD Average 1.20 25.00 1.53 34.08 0.22 29.18

I30

SD Average 2.70 37.00 2.89 50.93 2.57 47.42

SD 4.10 5.77 5.78

The curves obtained in the uniaxial compression tests for all the composites studied in this work are presented in Figure 5. For the M1 composite it was impossible to continue the test after the ultimate strength and for that reason it was not possible to calculate the toughness. The M2 and M3 composites presented a strain softening behavior after the ultimate strength, indicating that the 35 mm fiber bridged the macro cracks promoting a higher capacity of energy absorption. As can be seen in Table 5 the modulus of elasticity was in the same range for all the classes of composites, which shows that the fiber content and geometry did not influence the results. The compressive ultimate strength decreased with the increment in the steel fiber volume fraction. The composite M1 showed a compressive ultimate strength 11.25 % and 11.57 % higher than the one presented by M2 and M3, respectively.

E de A. SILVA, S . F O M G I N I , R.D. TOLEDO FILHO, E. de M. R. FAIRBAIW

260

180 160

l180

8

O

1

140 120

100 80 80 40 20 0

0

2

8

4

10

12

Strain (%)

Strain (%)

(a)

(b)

Strain (%)

(c) Figure 5 - Uniaxial compression test results: (a) M1 composite, (b) M 2 composite and (c) M3 composite. The results obtained in the compression tests indicated that the inclusion of the 35 mm fibers and the addition of 0.5 % of the 12 mm fibers reduced the ultimate strength of the composites. This reduction in strength may have been caused by an increase in the porosity caused by the higher fiber volume fraction presented by M2 and M3 or by the inclusion of flaws as a result of the loss of workability in the fresh concrete that was also caused by the higher fiber volume fraction. Table 5 - Ultimate strength and modulus of elasticity obtained in the uniaxial compression Composites Modulus of Elasticity Compressive Strength @a) (ma) M1 47.70 f 1.40 162.10 f 3.10 M2 44.20 f 0.67 144.17 f 8.34 M3 49.78 f 9.38 140.00 f 8.30

Behavior under compression and bending loads of multi-scale high peflormance steel fiber reinforced...

26 1

CONCLUSIONS In this work single and multi-scale steel fiber reinforced cementitious composites proportioned using the compressible packing model were developed and mechanically characterized. Three different fiber volume fractions were studied, denoted in this study as M1, M2 and M3. The four-point bending test results indicated that the composites presented the same range of ultimate strength, around 35 MPa. The toughness was calculated using the RILEM recommendations and the ASTM C1018. It was noticed that the addition of the 35 mm fiber in M2 and M3 increased the capacity of energy absorption after the ultimate strength when comparing to the M1 composite that was reinforced only by the 12 mm fiber. The highest compressive strength (162 MPa) was presented in case of the M1 composite. Increasing the fiber content decreased the compressive strength to 144 MPa and to 140 MPa for the M2 and M3 composites, respectively. ACKNOWLEDGEMENTS The authors would like to acknowledge the CNPq, CAPES, FAPERJ and FINEP for their financial support. REFERENCES [ 11 Bentur A., Mindess, S. Fibre Reinforced Cementitious Composites. Elsevier Applied Science, England, 1990,449 pp. [2] Balaguru, P. N., Shah, S.P. Fiber-Reinforced Cement Composites, McGraw-Hill, New York, USA, 1992,530 pp. [3] Naaman, A. E., Reinhardt, H. W. Characterization of High Performance Fiber Feinforced Cement Composites - HPFRCC. Proceedings of the 2"d International RILEM Workshop on High Performance Fiber Reinforced Cement Composites (HPFRCC2). Edited by A. E. Naaman and H. W. Reinhardt, E & FN Spon, London, 1996, pp. 1-24. [4] Rossi, P. Ultra-High Performance Fibre Reinforced Concretes (UHPFRC): An overview. 5'h International RILEM Symposium on Fibre-Reinforced Concretes (FRC) - BEFIB 2000, Edited by P. Rossi and G. Chanvillard, RILEM Publications, 2000, pp 87-100. [5] Buitelaar, P. Ultra High Performance Concrete: Developments and Applications during 25 years. International Symposium on UHPC, Germany, 2004. [6] Van Mier, J.G.M., Stang, H. Ramakrishnan, V. Practical structural applications of FRC and HPFRCC. Proceedings of the 2"d International Workshop on High Performance Fiber Reinforced Cement Composites (HPFRCC4). Edited by A. E. Naaman and H. W. Reinhardt, 1996, pp 443-459. [7] Naaman, A.E. Strain hardening and deflection hardening fiber reinforced cement composites. In: Fourth International Workshop on High Performance Fiber Reinforced Cement Composites (HPFRCC4), Ann Arbor, USA, 2003, pp. 95-1 13.

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[8] Matsumoto, T., Mihashi, H. JCI-DFRCC Summary report on DFRCC terminologies and application concepts. Proceedings of the JCI International Workshop on Ductile Fiber Reinforced Cementitious Composites (DFRCC), Takayama, Japan, 2002, pp. 59-66. [9] Naaman, A.E. Toughness, ductility, surface energy and deflection-hardening FRC composites. Proceedings of the JCI International Workshop on Ductile Fiber Reinforced Cementitious Composites (DFRCC), Takayama, Japan, 2002, pp. 23-57. [lo] Li, V.C. Reflections on the research and development of engineered cementitious composites (ECC). Proceedings of the JCI International Workshop on Ductile Fiber Reinforced Cementitious Composites (DFRCC), Takayama, Japan, 2002, pp. 1-22. [l 11 Li V.C. High performance fiber reinforced cementitious composites as durable material for concrete structure repair. International Journal for Restoration 2004; 10; 163-180. [12] Li V., Wang S. and Wu C. Tensile Strain-Hardening Behavior of Polyvinyl Alcohol Engineered Cementitious Composites (PVA-ECC). ACI Materials Journal 2001,98,483-492. [13] Reinhardt, H.W. and Fritz, C. Optimization of SIFCON Mix, Fibre Reinforced Cements and Concretes, Recent Developments, 1989, pp. 11-20. [14] Orange,G., Acker, P., Vernet, C. A new generation of UHP Concrete: Ductal damage resistance and micromechanical analysis. Fifth RILEM Symposium on Fiber-Reinforced Concretes (FRC), Lyon, France, September, 2000,78 1-790. [15] Rossi, P., Acker, P., Malier, Y. Effect of steel fibers at two stages : the material and the structure, Materials and Structures, vol. 20, 1987, pp. 436-439. [16] Boulay, C., Rossi, P. Tailhan, J.L. Uniaxial tensile test on a new cement composite having a hardening behaviorh: 6* Rilem Symposium on Fibre-Reinforced Concretes (FRC)BEFIB 2004, ,Varenna, Italy, pp. 61-68. [171 Parant E. Mkcanismes d’endommagement et comportements mkcaniques d’un composite cimentaire fibrk multi-Cchelles sous sollicitations sdveres : fatigue, choc, corrosion. DSc. thesis, Ecole Nationale de Ponts et ChaussCes, France, 2003. [ 181 Kawatama, A., Mihashi, H., Fukuyama, H. Material design of hybrid fiber reinforced composites manufactured by extrusion molding, Proceedings of the JCI International Workshop on Ductile Fiber Reinforced Cementitious Composites (DFRCC), Takayama, Japan, 2002, pp.75-84. [ 191 De Larrard, F., Concrete mixture proportioning: a scientific approach, Modem Concrete Technology Series, E&FN SPON, London, 1999. [20] Sedran, T., Rhkologie et rhCom&ie des bktons. application aux bktons autonivelants, Doctoral Thesis of Ecole Nationale des Ponts et Chausskes, 1 9 9 9 , 4 8 4 ~ ~ . [21] Formagini, S., Scientific mix design and mechanical characterization of ultra high performance fiber reinforced concrete, Doctoral Thesis, CCOPPE/Federal University of Rio de Janeiro, Rio de Janeiro, Brazil, 2 0 0 5 , 2 8 4 ~In ~ .Portuguese. [22] Rilem Technical committee 19-FCR. Testing methods for fiber reinforced cement-based composites. Materiaux et Constructions, vol. 17, 1984, pp. 441-456. [23] ASTM C 1018 - 92. Standard Test Method for Flexural Toughness and First-Crack Strength of Fiber-Reinforced Concrete (Using Beam with Third Point-Loading), ASTM 1992 Annual Book, Vol. 04.02, ASTM, Philadelphia.

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

FLEXURAL RESPONSE OF REINFORCED BEAM WITH HIGH DUCTILITY CONCRETE MATERIAL Maria M. SZERSZEN*, Aleksander SZWED**, and Victor C. LI* *Department of Civil and Environmental Engineering, University of Michigan, USA e-mail: [email protected], [email protected] **Civil Engineering Faculty, Warsaw University of Technology, Poland e-mail: [email protected]

ABSTRACT This paper reports on a study of flexural behavior of steel reinforced ductile engineered cementitious composite (ECC) members. ECC materials show extraordinary levels of strain ductility in tension (2-4%). Multiple microcracking in ECC delays fracture localization typically observed in normal concrete. Based on experimental stress-strain curves for ECC and reinforcing steel, a typical elastic-plastic model is assumed to derive the moment-curvature relation for reinforced beams in flexure. The resulting closed-form formulas are used in prediction of ultimate flexural capacity and ductility of beams made of ECC. Substantial difference in beams performance shows beneficial features of ductile ECC material. Direct design examples for beams and slabs using ductile or brittle materials present quantitative comparison of flexural behavior of structural members for typical design cases.

Keywords Composites, strain-hardening, fibers, ductility, ECC, RC, flexure

INTRODUCTION Ductile engineered cementitious composite is characterized by an ability to sustain equal or higher levels of loading after first cracking, while straining is significantly higher than the elastic limit. The high strain ductility can be achieved by controlling the composite ingredients of fiber, matrix and interface so that cracks initiated from defect zones do not result in fracture localization [ 11. Instead, the bridging fibers transfer the tensile load back into the matrix to create additional microcracks. One such ECC material extensively studied contains 2% of Polyvinyl Alcohol (PVA) fibers of 4 0 p n in diameter and 12mm long have been demonstrated to achieve strain ductility exceeding 3% under uniaxial tension [1,2]. During tensile straining of a ECC specimen, steady state crack propagation occurs with multiple dense cracking developing, and preserving stable crack opening at about 6 0 p n . Typical experimental stress-strain curves in uniaxial tensile test are shown in Fig.1. The experimental curves show a non-softening trend with strain capacity over 300 times that of a non-reinforced concrete matrix. Application of a new type of high performance material in engineering practice requires extensive analysis of the structural members response in order to develop design guidelines. The high ductility and tight crack width features of ECC is expected to provide

264

Maria M.SZERSZEN, Aleksander SZWED, Victor C. Ll

significant advantages in ultimate and serviceability limit states of structural members under tensile or flexural loading. Full benefits of such material, such as increased moment or shear capacity of structural members, can be addressed in design procedures using a material model different from that established for ordinary concrete.

j

8

-ExponmenmlTsstNo 3 ExpcnmarmlTsst No. 4 +EIW~-plasU~ldealmhoo

.-

Fig. 1. Typical experimental stress-strain curves for ECC in uniaxial direct tension test. I

smin

I

0000

0005

0010

0015

0020

0025

0030

0035

In this paper, flexural analysis of reinforced ECC slender beams is presented. Utilizing available uniaxial tension and compression tests results for ECC, as well as properties of reinforcing steel, a simple idealization of stress-strain curves is assumed. Using strain compatibility and constitutive relationships of constituent materials, closed-form formulas for moment-curvature relations are derived. Closed-form formulas are used in prediction of ultimate flexural capacity and curvature ductility of reinforced beams made of ECC, and they are very useful in parametric study. Comparison of designs performed for ductile ECC and ordinary brittle concrete highlights beneficial properties of ECC when used in structural members under flexural loading. BASIC ASSUMPTIONS

Based on experimental stress-strain curves for uniaxial tension (Fig.1) and uniaxial compression tests for ECC, some idealizations and simplifications in material descriptions were carried out. Linear elastic and then perfectly plastic one-dimensional material model was assumed because of its simplicity and easy application. The behavior of ECC is defined as,

- a,, for - E,, Ec& for

a,, for

I

E

<-E~,

-cot I E I E ~ ~ E,, < E I E ~ ,

where E ~ ,and E,, are elastic limit strain in compression and tension, a,, and a,, are the elastic limit stress in compression and tension. Elastic (Young's) modulus for linear elastic range is defined by formulas, E, = a,, /cot = ooT / eOT.Ultimate strains in tension and compression, E,, E,, are defined to control stable straining of material. The behavior of reinforcing steel is assumed to be linear elastic and then perfectly plastic, and is defined by the following relationship, - f,

for

E,E, for fY for

<-E,

-E,

-E, E,

I

E,

I E,

< Es I E",

Flexural response of reinforced beam with high ductiliw concrete material

265

where E, is elastic limit strain, f , is elastic limit (yield) stress, E, = J,/E, is elastic modulus for steel in linear elastic range. The ultimate strain in reinforcing steel E , is typically several times larger than the ultimate tensile strain in ECC, and this parameter will not be critical in the performed analysis. Assumption of plane cross-section of a beam is adopted in the following derivation of basic equation for bending of slender beams. The assumption states that beam cross-section remains plane in any stage of beam deformation. Neutral axis (or generally a curve) is defined as the position of points of a beam’s cross-section where no axial strains and stresses occur, and its vertical displacement is given by the h c t i o n ~ ( x )According . to the assumption, the horizontal displacement, u , of a point in distance of z from the neutral axis is expressed as, u(x, z) = -z- d 4 x ) = -z tg@z -z@(x)

dx

(3)

Then, the axial strain is defined as,

4.)

is the curvature of a bent beam, and @(x)is the slope. The relation between axial where strain and curvature given by (4)is often called the strain compatibility. MOMENT-CURVATURE RELATION FOR SECTION

The moment-curvature relationship for cross-section is determined by considering compatibility conditions of the section, (4), and the stress-strain relations of the constituent materials (1) and (2). For simplicity, a rectangular cross-section with single layer of reinforcement will be assumed in the following analysis, but presented approach can be applied to any symmetrical cross-sections with multiple layers of reinforcement. Perfect bond between reinforcing steel and ECC is assumed, what is experimentally observed [4]. Assuming pure bending, derivation of the moment-curvature relationship involves the assumption of various levels of strains in the cross-section, corresponding to the different stages of loading. Stress distribution defined according to the constitutive low is integrated over the cross-section area. Zero axial force condition defines the compression zone depth, which subsequently is used to obtain the moment capacity of the cross-section. The complete moment-curvature relationship is subdivided onto five phases depending on the load level and reinforcement ratio. Phase I is determined by the linear elastic behavior of constituent materials in the cross-section. Phase I1 is defined with the assumption of elastic-plastic behavior of ECC in the tension zone, and elastic behavior in the compression zone, while steel is in elastic range. Phase 111, with elastic-plastic behavior of ECC in the tension zone, and elastic behavior in the compression zone; steel reaches yield stress. Phase IV,valid only for high reinforcement ratios, covers elastic-plastic behavior of ECC material in both, tension and compression zones, while steel is in elastic range. Phase V occurs with elastic-plastic behavior of ECC material in both, tension and compression zones, and steel is plastic. The basic idea of the proposed phase subdivision is explained in Fig.2, and Fig.3, what summarizes the analysis performed in this section. In the following derivations, each phase is considered separately and closed-form solutions for moment-curvature relations are obtained. Limits of application for each phase are defined, and some characteristic values of moment and curvature are given. Formulas for other parameters involved, such as: the depth of compression zone, and cracking and crushing zone depths are also given.

vlaria M.SZERSZEN, Aleksander SZWED, Victor C. LZ

266

)I.\;-

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

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

............. ..- ...._............I

T t & h......l....-..........l....d... - c

...

.............. .9 ...................... ......................................

ES <EY

b)

~

................... ................ T' T0'>

0 s
...............

a:..:.::.:::.:: h-c

4 'EY

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

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

pc

c)

' 7

'

T O '

UaT

Fig. 2. a) Phase I: All materials are elastic; b) Phase II: ECC elastic-plastic in tension; c] Phase 111: ECC elastic-plastic in tension and steel yields; d) Phase IV:ECC elastic-plastic --+ in tension and compression and elastic steel; fY e) Phase V: ECC and steel elastic-plastic.

Phase I. In the case of strains and stresses linearly distributed in beam cross-section (Fig.2a), the moment-curvature relationship and compression zone depth, c , can be expressed as, 1+ 256 2(1+ 5)

c, = -

where scaled moment m = M I McE and curvature k = K /K, are used throughout this ~ moment, , McE,for an un-reinforced rectangular analysis. The elastic limit curvature, K ~ and ECC cross-section are given by formulas,

. The elastic bending stiffness of an un-reinforced beam is then: EcZc = McE/ K ~ The dimensionless parameters p , 5 and S in (5) stand for reinforcement ratio, modularreinforcement ratio and effective depth ratio, respectively:

Further notation is explained in Fig. 2. The actual moment, M , and actual curvature, K , in (5) are lower than the elastic limits in phase I. The end of elastic behavior of the cross-section is reached when in exterior fibers of beam tensile stresses are equal to the elastic limit stress ooT. Elastic limit curvature and moment of reinforced cross-section are, 1+ 45[1- 3S(1- S)] k, = 1+5 mE = 1+ 25(1-S) ' 1+25(1- 6) . "1

267

Flexural response of reinforced beam with high ductility concrete material

Phase 11. Phase I1 of cross-sectional behavior begins when strains in the ECC matrix exceed the elastic limit, E,, , in the tension zone (Fig.2b). The condition of zero axial force yields to the evaluation of depth of the compression zone, c ,and then cracking zone depth r :

The moment-curvature relationship can be expressed as follows, m, = 3 + 6{(1+ { + 6)+45[2{’ +36(6+ 2{$-

2[1+ {+45(26+ {)k

1J

5(26+ {)+% k

.

(10)

When the elastic limit stress in steel, f , , is reached, the end of elastic behavior of ,the corresponding curvature is, steel defines the end of the phase. Using the ratio, y = E ,

kyT=

1+ g+6(y- 1)+ J(l+ y5)[1+ g + 26(y-l)] 26’

(1 1)

In the case of highly reinforced cross-section, the end of elastic behavior of ECC in compression region can be reached, before steel starts to yield. At this stage of loading, compressive stress in external fibers is o,,, and using parameter p = ooc / o,,, the curvature is,

Balance point (see Fig.3, point B), reflecting simultaneously reached limits, f , for steel and o,, , for concrete, can be evaluated comparing (11) and (12). The balance point reinforcement ratio, curvature and moment for this point are given by the formulas,

c,,

= ripe ,

’+’

k, = 26 ’

m B =3(1-26)+(2P+3y-l

)“~~~I. ~

(13)

Phase 111. Phase I11 governs when tensile strains in concrete composite are greater then elastic limit, E,, , in a part of beam cross-section, and reinforcing steel yields in tension, whereas concrete remains elastic in compression, see Fig.2~.This case is typical for tensile controlled failure path of a cross-section. The compression zone depth, c , and moment-curvature relationship is then as follows:

The end of elastic behavior in compression is reached when compressive stress in exterior fibers is equal too,, . Then, the curvature and moment for the end of the phase I11 are,

Maria M.SZERSZEN, Aleksander SZWED, Victor C. LI

268

Phase IV. In phase IV, strains are greater then elastic limit in tension, and in compression, strains are greater then elastic limit in compression, E,, , while reinforcing steel remains in the elastic range. This case is appropriate for compression controlled failure path of a crosssection. Axial force and bending moment can be calculated using stress diagram shown in Fig.2d. The compression zone depth, c, and crushing zone depth, s, can be obtained as, CIV

=

P’ -1+4k(1+2g/Oh 4k(l+,8+2ck)

’

sW =cIV--h, P 2k

while moment-curvature formula can be expressed as

When yield stress in steel, f, ,is reached, the end of phase IV occurs, and curvature is,

Another possible end of this phase occurs, when the end of concrete composite ductility in compression, E,, ,is reached before steel starts yielding. Then, introducing R = E,, I E,, ,the value for curvature is,

Simultaneous reaching of E,, in concrete composite and f, in steel can be derived by comparing (18) and (19), (see Fig.3, point C). Then the reinforcement ratio for this state is,

6, =

6- 2y- 2p41- 6)+(2n - 1)6#’ 2YtY+fl4

= npc,

and the characteristic values of curvature and moment are given by the formulas,

Phase V. In the case of strains in compression greater then E,,, and stress in steel equal to f, , phase V of cross-sectional behavior occurs, Fig.2e. The depth of Compression zone and moment-curvature are,

When ultimate strains, E,, ,are reached, the ultimate curvature and moment are,

269

Flexural response of reinforced beam with high ductiliw concrete material

For a very lightly reinforced section, compression failure mode controls the ultimate strain , fulfill the following inequality, capacity when the strain ratios, f ,and a = E~ I 1 2

2a-1 2p2

a<-+-=AB4,

,

where AB4/ is the balanced strain capacity failure factor, reflecting simultaneous reaching of strain ductility of ECC in tension and compression. In cases other than (24), tension failure mode controls sectional failure. Then, the ultimate curvature and moment are as follows,

Tension point (see Fig.3, point T), for simultaneous reaching ultimate strains in compression and tension in ECC material, can be evaluated by comparing (23) and (25). Then, T point reinforcement ratio (for A < f B , , , ) , curvature and moment can be expressed as,

-

Asymptotic ultimate capacity of reinforced ECC cross-section is reached when K 4 . Then, the estimation of ultimate bending moment and depth of compression zone give values,

In Fig.3, the relationship between moment and curvature with explanation of notation used in foregoing derivations is shown. Curvature and moment are scaled by elastic limits (6), defined for an un-reinforced ECC cross-section. All characteristic parameters for curvature and moment are plotted for the five phases of behavior previously described and discussed. In phase I, moment-curvature relationship is linear, while in other phases 11, 111, IV and V is nonlinear. Four moment curvature relationships for characteristic reinforcement ratios p = 0, pr, pB,pc are plotted. Boundaries between phases are also plotted in Fig.3 with given characteristic strain in concrete or steel, related to the appropriate limit.

Fig. 3. Subdivision of momentcurvature relationship onto five phases in function of reinforcement ratio with identification of characteristic boundaries and points, where r n = M f M c Eand k = K f K c E . 0

10

20

30

40

270

Maria M.SZERSZEN, Aleksander SZWED, Victor C. LI

ANALYSIS OF RESULTS AND PARAMETRIC STUDY

In this section, basic analysis of moment-curvature relations and graphical interpretation of results are presented. Typical data for constituent materials are assumed and characteristic parameters involved in the model are obtained. Parametric study is mainly focused on the influence of reinforcement ratio, as well as on moment and curvature capacity for beams cross-section. Based on experimental uniaxial compression and tension tests for ECC composite, characteristic strain and stress values have been estimated and used in the following calculations. Compressive strength and modulus of elasticity are estimated as: f', = 6 0 M a ,. Tensile strength of concrete is: ooT = 5.3MPa, and structural compressive strength is assumed to be: o,,= 0.88f', = 53MPa. Characteristic values for elastic limit strains are calculated as: E , ~= ooT I E, = 0.0003 and E ~ = , oo, I E, = 0.003. Ultimate compressive and tensile strains are estimated as: E,, = 0.0045 and E, = 0.033. Based on these values, basic dimensionless material parameters can be obtained #=Ia,, I ooT = 10 A = E,, I cOc= 1.5 , and a = E, /EoT= 110. Properties of reinforcing steel are: E, = 200GPa f, = 420 MPa and E, = 0.00021. Then, n = E, I E, = 11.4 and y = E, IEoT = 7.0 can be calculated. For this study, concrete cover ratio is assumed to be: 8 = 0.85. Balanced failure factor is ABo,= 1.6, which defines compression controlled failure for lightly reinforced cross-sections. Reinforcement ratios for characteristic points are: pB= 2.54%, pc = 4.34%. The relationship between moment and curvature for several reinforcement ratios is shown in Fig. 4, where characteristic boundaries and points are explained. Characteristic values of curvatures and moments are: K~ = 10.00 K , ~ ,MB = lO.OOM,, K, = 12.94 K , ~ , M , = 14.73McE.For the given material data, A < ABo,,the compression controlled failure occurs for un-reinforced cross-section, without T point in Fig.4. Ultimate capacities of curvature and moment for un-reinforced cross-section are: kuc = 57.75, mu, = 2.70. )

)

16 14 12

10

Fig. 4. Moment-curvature curves for several reinforcement ratios.

8

6 4 2

0

10

20

30

40

50

From the figure, it is easy to read an increase in moment capacity due to amount of reinforcement applied, as well as decrease in curvature ductility as a function of reinforcement ratio. Curves given in Fig. 4 clearly identify that there is no sudden drop in moment capacity after the first cracking. A sudden drop in moment capacity is typical for ordinary concrete, especially for lightly reinforced sections [3]. In case of ECC matrix applied in reinforced beam, moment-curvature curves indicate hardening behavior in the full range of curvatures, up to ultimate capacity. This occurs because of the continued tensile load carrying capacity of the ECC material, beyond its elastic limit.

27 1

Flexural response of reinforced beam with high ductility concrete material

The compression zone depth as a function of curvature and reinforcement ratio is presented in FigSa. Rapid decrease of compression zone in the cross-section is observed in phases 11, I11 and V, for lightly reinforced sections. For highly reinforced sections, compression zone remains on relatively constant level with low curvature ductility, which is characteristic of sudden compression failure mode. In FigSb, cracking zone in tension and crushing zone in compression, are plotted as a function of curvature. In phase 11, rapid development of cracking zone is observed. In phases 111, IV and V, development of cracking zone is not as rapid as in phase 11, and stabilizes for highly reinforced sections. Crushing zone develops rapidly for highly reinforced sections, especially in phase IV,what reflects brittle failure mode. trlh

k Fig. 5. a) Compression zone depth; b) Cracking and crushing zones development. 16

60

tm

14

50

I2

40

10

30

8

6

20

4

10

2

0

001

002

003

004

005

006-

b)

P 0

001

002

003

004

005

006

Fig. 6. Characteristic curvatures (a) and moments (b) as functions of reinforcement ratio. Relationships between reinforcement ratio and characteristic curvatures, and corresponding moments are shown in Fig. 6. Fig.6a shows a decrease in curvature capacity (ductility) due to the amount of reinforcement applied. Different trends for characteristic curvatures can be observed for different curves. Elastic limit and ultimate ductility of ECC in compression decreases in hyperbolic manner with an increase of reinforcement ratio. These limits typically govern design in structural applications, when tension controlled failure (ductile) is a desired behavior for a structural member at ultimate load. Fig.6b shows an increase in moment capacity due to the amount of reinforcement in the cross-section. It can be observed, that the moment increase is close to linear for all phases of cross-section behavior. Two characteristic points, B and C, initiate deviations in the trend of moment increase. However, there is very little increase in elastic capacity due to reinforcement increase.

272

Maria M.SZERSZEN, Aleksander S Z M D , Victor C. LI

DESIGN EXAMPLES FOR BEAM WITH ECC AND ORDINARY CONCRETE Moment capacity is calculated for two beams, made of ordinary concrete and ECC, with identical cross-section and reinforcement equal to maximum reinforcement permitted by ACI 318 Code, to enforce ductile failure mode. This maximum reinforcement should be no more that0.75pb, where pb is the balance failure percentage of reinforcement equal to 4.74% for ordinary concrete beam, and 5% for ECC beam. Compressive strength, equal to f, ’ = 8700psi (60M a ) is the same for OC and ECC material; reinforcing steel with yielding strength of f , = 60,OOOpsi (Grade60) is used in both beams. For the design calculation below, the percentage of reinforcement used in both beams, p ,is assumed equal to, p =0.75~0.0474 = 0.035 (3.5%)

H~,=~=fJE,=0.00207

H w oar ~=~,=fJE.TO.00207

Fig. 7. Distribution of strains and stresses in Fig. 8. Distribution of strains and stresses in the cross-section of single reinforced ordinary the cross-section of single reinforced ECC beam. concrete beam. Moment capacity of ordinary concrete beam (Fig. 7), calculated according to ACI 3 18 Code standard procedure is as follows,

C=0.85f,’xaxb

T, = Asfy C=T,

3 a=

4fY 0.85f, ‘xb

A, = 0.035~ b x d = 0.035~15~18.5 = 9.71in2

a=

9.71~60000 =525in 0.85~8700~15 =9,248,775 Ib-in

Calculation of moment capacity of ECC beam presented below is modified comparing to standard code procedure, to account for strain and stress capacity of the composite in tension zone (Fig. 8). Stress block parameter,p,, for ordinary concrete equal to 0.65 (for f,’ = 8700psi ), is recalculated for ECC material, and becomes yl = 0.66 .

Flexural response of reinforced beam with high ductiliiy concrete material

273

A

A, = ~ , x E , , x ~ . ~ + ( E , - E , , ) ~ ,

E,

=0"=0.00303 E C

4 =o,xx 4=4 XEocx0'5+(Euc

X=

-Eoc)ooc

=0.00303x0.5+(0.0045-0.00303) =0.00299

( J ,

x

y1 ---

E,

-0.00299 - -= 0.66 0.0045

a = y,c = 0 . 6 6 ~

A, = 9.71in2 C = 0.85f, 'ab

T, =oOT (h-c)b T, =A& C=T,+T,

0.85f, ' ~ 0 . 6 6 = 6 oOr ~ ~( h - C )b + A,& C=

'sfy

c=

0.85S,'x0.66b+ ooTb

770 ~ 2 0 x 1 + 59.71~60000 = 9.6 in 0 . 8 5 ~ 8 7 0 0 ~ 0 . 6 6+~770x1 15 5

a = 0.66 x 9.6 = 6.3in

Mcop= T,

Mcop

(F-4)+ 4)

=OoT

+c

(h-c)b[

T, ( d -

h-c+2c-a) + A f 2

;)

d--

= -1[ ( J o , ( h - ~ ) ( h + ~ - a ) b + A s fy(2d-a)] 2

1 McopEK = -[ 770 (20 - 9.6)(20 +9.6 - 6.3)15+ 9.7~60000 (2x18.5 - 63)] = 10,342,308Zb 2

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Maria M.SZERSZEN, Aleksander SZWED, Victor C. LI

There is 12% increase in moment capacity for a beam made with ECC material if compared to the same beam, but made of ordinary concrete. The sensitivity analysis presented earlier show that the significance of ECC contribution in tensile zone increases for lower percentages of reinforcement. If similar comparison of moment capacities is performed for concrete bridge slabs with small amount of reinforcement ( p 3 0.003 (0.3%)for empirical design), the difference in moment capacities is even more significant. Standard thickness (7.5 in) ECC bridge deck indicates 230% increase in moment capacity if compared to the same, but ordinary concrete deck. SUMMARY AND CONCLUSIONS

High ductility and stable multiple crack development in ECC gives a beneficial behavior in flexure of structural members made of this material. Simplified material models for the constituent materials in terms of stress-strain curves, allow for derivation of the analytical moment-curvature relations for beam. The resulting complete, closed-form formulas can be easily used for predicting the ultimate flexural capacity and curvature ductility of beams made of reinforced ECC. The resulting formulas are very useful in developing reliable design procedures for ECC, to aid structural engineers in applying such materials in field applications. The proposed rational derivation of moment-curvature relations gives qualitative and quantitative description of flexural behavior of reinforced ECC cross-section. The results can be applied to any beam or one-way slab, and extended to two-way slabs. Precisely defined phases in sectional behavior, and limits of application, can serve as a design guide for flexural elements. From the present analyzes, a distinct difference between reinforced ECC (RECC) and common reinforced concrete can be observed. The sudden drop in moment capacity typical in ordinary concrete, especially for lightly reinforced sections, is absent when ECC is applied in reinforced beam. Moment-curvature curves indicate hardening behavior in whole range of curvatures, up to ultimate capacity. Although the modulus of elasticity for ECC is about 2025% lower than for ordinary concrete with the same compressive strength, the flexural stiffness (in post cracked phase II) of reinforced cross-section is of the same order as for RC, and even higher for lightly reinforced sections. The ultimate flexural capacity is always higher for RECC than for RC, especially when low reinforcement ratio is applied. For instance, forp = 1% , the capacity of RECC is higher by more than 50%. The curvature ductility of RECC is always higher than that of RC, even for highly reinforced sections, for example, if p = 3% ductility of RECC is higher by 40%. REFERENCES

1 Li V.C., Wang S., Wu C.: Tensile Strain-Hardening Behavior of PVA-ECC. ACI Materials Journal, V. 98, No. 6, pp. 483-492,2001. 2 Wang S., Li V.C.: Polyvinyl Alcohol Fiber Reinforced Engineered Cementitious Composites: Material Design and Performances. Proc. of International Workshop on HPFRCC Structural Applications, Hawaii, May 2005. 3 Bosco C., Carpinteri A., Debernardi P.G.: Minimum Reinforcement in High-Strength Concrete. Journal of Structural Engineering, V. 116, No. 2, pp. 427-437, 1990. 4 Fischer, G., and V.C. Li, ‘‘Influence of Matrix Ductility on the Tension-Stiffening Behavior of Steel Reinforced Engineered Cementitious Composites (ECC),” ACI structural J., Vol. 99, No. 1, pp. 104-1 11,2002.

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

EFFICIENCY OF MULTI FILAMENT REINFORCEMENT IN CEMENTITIOUS COMPOSITES Frank JESSE Institute of Concrete Structures Technische Universitat Dresden Helmholtzstr. 10,O 1062 Dresden, Germany e-mail: [email protected]

ABSTRACT The use of continuous fiber reinforcement in a cementitious composite is very efficient concerning strength and ductility. Using technical textiles made of AR-Glass or Carbon in a cementitious matrix leads to textile reinforced concrete and is a way of doing this cost-effectively. In an extensive study, mechanical testing was combined with microscopic observations to identify basic mechanisms of load transfer between multi- filament yarns and matrix. Uniaxial tension tests on the composite were carried out to evaluate mechanical performance of textile reinforced concrete reinforced with different types of fabrics. Thin section petrography was used to find specific bond properties. The paper discusses failure mechanisms of textile reinforced concrete and how they are influenced by different micromechanical bond parameters. It has been found that bond characteristics between single filaments in the roving structure and between filaments and matrix and how they are distributed across the filament bundles and along the fiber-matrix interface are of major importance for composite performance. Bond properties are affected by several parameters during textile processing, for instance yarn material, yam titter, size, weaving pattern, and others. Understanding these dependencies leads to recommendations for optimizing textile reinforcement.

Keywords Textile reinforcement, fabric, AR-glass, filament, yam, stress-strain-curve, cracking, ultimate load, bond strength

INTRODUCTION Textile reinforced concrete (TRC) [ 11 has become an extremely effective and promising cementitious material that can be used for structural application for new buildings and infrastructure as well as for retrofitting measures in existing structures [2]. The material has been proved to allow efficient strengthening in different applications, e.g. flexural strengthening of slabs and beams [3], shear strengthening of beams and t-beams [3], confining of columns [4]and others. Stitch bonded fabrics or other non woven fabrics are preferably used for textile reinforced concrete. A wide variety of fiber materials has been used in many studies. Among them, ARglass fibers and carbon fibers are the most interesting ones because of their high strength, high modulus (especially carbon), non corrosive properties and economic attractiveness. Glass and carbon fibers are supplied in form of bundles of fibers, called rovings or yams, consisting of several hundred or thousands of elementary fibers, called filaments. Numerous studies have

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Frank JESSE

shown that the interaction of the filaments of a yam and the interaction between yams is very complex. Thus there is a lack of a detailed understanding of the mechanical mechanisms in the load bearing behavior of TRC. The main impact factors are bond properties between filaments and filaments and between matrix and filaments. This is mainly because multifilament yams are supplied by the manufacturer in a condition with only minor bond between filaments. As a consequence, bond properties are affected in each subsequent technological step of fabric and composite production. Peled et al. have studied the impact of composite production technology [5]. During fabric production the yarn’s characteristic is changed in a multiple manner. Filaments may be damaged and can fail during textile processing. This may lead to strength reduction. The geometry of yam cross section will change heavily, which affects the bond between yam and matrix and thus the load bearing behavior. The stitching yarn leads to a confinement of the filament bundles and thus will affect inner bond properties. At the crossings of warp and weft yarns there is not matrix contact. The same applies to the contact areas between yam and stitching yam. This paper aims to study the load bearing behavior of TRC made of fabrics produced using different weaving patterns. In order to achieve a good comparability, all parameters except the weaving pattern were kept constant during this study.

EXPERIMENTAL PROGRAM

Materials and sample production Fine grained concrete with a maximum grain size of 1 mm, type I11 cement, fly ash, micro silica, water and super plasticizer were used as matrix. Compressive strength and flexural strength were measured with mortar prisms according to DIN EN 196-1. Mean values were approximately 76.3 N/mmz and 7.1 N/mmz respectively. An AR-glass multi filament yarn from Nippon Electric Glass Ltd., Japan (NEG) was used as fiber material (Table 1). The yarn was processed on a Malimo, type 14024 from KarlMayer Chemnitz, Germany, to produce biaxial warp knitted fabrics with different bindings, namely tricot, double-tricot and pillart-tricot (Fig. 1). During fabric production all parameters except the type of weaving were identical. Warp and weft yams had a spacing of 7.2 mm. For comparison purposes, specimens with unidirectional yarn reinforcement were also manufactured and tested using identical procedures. Table 1: Properties of AR-glass multifilamentyarn VET-ARG1200-02 according to [7] Identifier NEG-ARG310-01 NEG-ARG620-0 1 AR620S-800/TM Manufacture name AR3 lOS-8OO/DB 309 tex 614 tex Yarn count Number of filaments 800 1600 12.94 pm 12.76 pm Filament diameter 2271 N/mm2 2076 N/mmz Filament strength (mean) 1357 N/mmz 1341 N/mmz Yam strength (mean)

277

Eficiency of multifilament reinforcement in cementitious composites

(a)

mcot, front

(d) tricot, back

(

(e) double-tricot, back

(0 pillart-tricot, back

Fig. 1: View on the three different weaving patterns used for this study: (a)+(d) tricot, (b)+(e) double-tricot, (c)+(f) pillart-tricot, (a)-(c) top side, (d-e) bottom side, warp yarn in vertical direction The test specimens were produced using a simple hand lay-up process. Thin layers of tine grained concrete and fabric were incorporated alternately using a formwork with stainless steel with dimensions of 1200 mm x 600 mm x 8 mm (length x width x thickness). After three days, the specimens were removed from the formwork and stored in water. From the seventh day of curing until prior testing, they were stored in a climate chamber at 22 "C and 65 % r. h. Small specimens with dimensions of 500 mm x 100 mm x 8 mm (length x width x thickness) were cut from the large slab using a water-cooled diamond saw. In a previously published paper [6] it has been shown that the strength is inversely correlated with the fiber content. Thus the fiber volume content was kept constant for all tests at approximately 1.9 %, which is achieved by using three layers of yams spaced 4.5 mm and five layers of fabric reinforcement, respectively.

Tensile tests A hydraulic standard testing machine with a capacity of 20/100 kN from WPM Leipzig (Germany) was used for the examination. The load introduction was realized via a wedge clamping anchorage. To avoid stress peaks in the load introduction between the steel plates and the coflcrete samples, sheets of 0.5 mm thick rubber were placed in between. The load was applied deformation-controlled using an external linear variable differential transducer (LVDT) with a rate of 1 mm/min. The strains were measured with a clip-on extensometer using two strain transducers DD1 from HBM (Darmstadt, Germany). During the test, load, crosshead displacement and specimen's deformation at front and back were measured. All data, including the controlling LVDT, were recorded with a MGC data acquisition system from HBM, Darmstadt, Germany. From these data, stress-strain curves were generated for evaluation.

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Thin section petrography For the microscopic observations, standard thin section petrography was used. Photographs were taken from these samples, using an optical microscope with transmitting light, equipped with a digital camera. Subsequently, a digital image analysis was performed with a procedure described in detail in [8, 91 in order to obtain cross sectional area, contact perimeter between matrix and fibers and packing density and other geometrical parameters.

RESULTS AND DISCUSSION Perimeter and aspect ratio of the reinforcement Perimeter and packing density are crucial parameters for analyzing the bond behavior of textile reinforcement. They have been estimated by means of digital image analysis from thin section petrography (Table 2). Comparing the measured cross sectional area At,, does not give a realistic view, because the 3 10 tex yarn was used as double yarn in the fabrics in order to allow comparison with the 620 tex yam. For that reason, packing density has been calculated, i.e. the theoretical fiber cross section Af,, over the measured fiber cross section In most cases, packing density of fabric reinforcement is higher when compared with packing density of the pure yarn. The increase is not dramatic for the 3 10 tex yam. For pillarttricot we find a decrease of the packing density. With the 620 tex yam the packing density is considerably increased for all fabrics when compared to the pure yam. For both fibers the highest increase can be asserted for the double-tricot fabric, followed by pillart-tricot and tricot binding. This means that during the manufacturing of the fabric the yam’s cross section was changed and that the packing density increased in principal. This can be attributed to the confining action of the needle thread. Obviously it is most pronounced for double-tricot, followed by pillart-tricot and tricot binding. Another important parameter for the bond behavior is the contact perimeter Pf,,between fibers and matrix (Table 2). The largest perimeters have been measured for the pure yarn. For both fiber types, the perimeter is (accidentally)identical with Pf,,,, = 11.51 mm. The measured perimeters for the fabrics are considerably smaller. Compared to the 310 tex yarn (620 tex yarn), the perimeter of the tricot fabric is 62 % (88 YO), for the double-tricot fabric 32 % (30%) and for the pillart-tricot fabric 46 % (54 YO)of the values measured for the unprocessed Yarn. Table 2: Geometrical properties from thin section petrography Yam and fabric type

A,,

in -2

AS,,, COV A,,/Af,, in mmz i n % in %

P,, COV in mm in %

NEGARG310-01

Yam

109.7

184.7

3.3

59.9

11.51

Tricot

219.4

361.9

5.5

60.8

7.18

18.2 13.1

Double-tricot Pillart-tricot

219.4 219.4

344.3 391.2

4.6 10.7

63.8 56.6

3.71 5.29

13.7 27.2

218.0 218.0 218.0 218.0

482.8

13.7

45.9

11.51 21.6

421.1 336.4 406.5

5.0

9.0 9.8

52.3 65.3

10.11 12.2 3.42 22.0 6.24 13.2

NEGARG620-01

Yam Tricot Double-tricot Pillart-tricot

54.1

Efficiency of multifilament reinforcement in cementitious composites

279

This means that two effects appear during the manufacturing of the textile reinforcements: The filament bundles become more dense. A more dense fiber bundle is supposed to improve the bond performance between the filaments and hence is assumed to result in an increase of the composite strength. The shape of the yam bundles is heavily changed. In the delivered condition the yams have a shape like a very flat ellipse with an aspect ratio of 1 : 10 (NEGARG310-01) and 1 : 7 (NEG-ARG620-OI), respectively. For the double-tricot fabric, the aspect ratio is approximately 1 : 2 (NEG-ARG310-01) and 1 : 1.6 (NEG-ARG62O-0 1). If every filament could be bonded directly to the matrix and if bond quality is assumed to be constant, the same strain would apply to all filaments of the yam. In that case, the highest possible strain could be achieved before the yarn fails. However, this is clearly a very theoretical case since usual yam does not allow direct contact between filaments and matrix for all filaments. If the shape of the yams is more compact, the load has to pass more filament-filament-interfaces from the fiber-matrix-interface to the core filaments. Because of the deformations of the filament-filament-interface, the filament becomes smaller the more interfaces are passed. If we assume a very large number of layers of filaments, the strain in the filaments will become smaller the more we move to the center of the yam bundle. Failure of the fiber bundle will always start if the outer (highest stressed) filaments exceed their ultimate strain. This leads to a decrease of the ultimate capacity of the fiber bundles with increasing compactness of the bundle shape. Thus, a more compact shape is associated with smaller strength and ultimate load, respectively, of textile reinforced concrete.

Stress-strain behavior In principle, all specimens behave almost identical concerning the general run of the curve. The well known multi-linear behavior of reinforced concrete with a high stiffness in the uncracked state I, a very low stiffness during the multiple cracking in state IIb, and an increased stiffness with an almost linear section in state IIb is visible for all specimens. There are many differences in detail between straight yams on the one hand and all types of fabrics on the other hand as well as between the different fabric types. The first cracking stress for the yam is higher than for all types of fabric. Pillart-tricot and tricot have the first cracking stress between 4 and 5 N/mmz. Double-tricot shows the lowest first cracking stress. This reduction in first cracking stress is generally attributable to the presence of the transverse yams. Because the load bearing capacity in transverse direction is very limited for all yam types, including the PP-yams, the transverse yams act as a pore and reduce the concrete section. Furthermore, they might cause stress concentrations in the matrix. This results in a reduced tension capacity of the concrete section and thus in a reduced first cracking stress compared to those specimens with only unidirectional yams [9]. The run of the curve during multiple cracking is almost identical for pure yams and tricot binding, although the stress level is considerably lower for the latter. All curves show an almost identical incline, but whereas the curve of the pure yam is relatively smooth, the curves of the double-tricot and pillart-tricot fabric reinforced specimens clearly show a pronounced up and down. This is due to larger crack spacing and accordingly larger crack opening. Cracking has finished for yam and for tricot fabric when the composite strain is between 3 and 5 %o. For pillart-tricot and double-tricot fabrics, the strain at which cracking is completed is clearly higher. Especially in Fig. 2 (b) it looks as if (primary) cracking stops at a comparable strain at approximately 3 %O and further (secondary) cracks appear afterwards until the cracking stress level of the specimen with pure yam is reached. On the contrary, this behavior is less obvious in Fig. 2 (a).

Frank JESSE

280

P

If

B

i

‘O

I

O

f

&&h%

10

(a) NEG-ARG62O-0 1

I5

0

5

&&h%

I0

If

(b) . , NEG-ARG310-01

Fig. 2: Typical stress-strain curves ffom uniaxial tension tests (5 specimens tested for each fabric)

After cracking has finished, the stiffness of the curve rises again and in most cases the fabric reinforced specimens show almost the same incline as the yam reinforced specimen. In all cases, absolute strain values of fabric reinforced specimens are higher compared to the yam reinforced specimen at same stress level. In some cases this difference is very pronounced. Cracking pattern The specimens with unidirectional yam reinforcement show a very dense cracking pattern with a narrow crack spacing of 6.7 mm for the 620 tex yam and 3.9 nun for the 310 tex yam respectively (Fig. 3 (a)). For the 3 10 tex yam it should be stated that the crack spacing is not directly comparable because doubled yams were used in the fabrics whereas the results for the pure yam have been obtained using single rovings. It is assumed that a “doubled” 310 tex yam would result in a crack spacing between the values for the corresponding tricot fabric and the single 3 10 tex yain. Actually, crack spacing is underestimated with the single 3 10 tex yam when compared with the “double” 3 10 tex yam used in the fabrics. Unfortunately, a “double” 3 10 tex roving is not available. For the fabric reinforced specimens, the cracking pattern changes considerably. Mean crack spacing as well as the cracking pattern give similar results for the tricot-binding as with pure yam reinforcement. Crack spacing is slightly increased compared to the specimens with pure yam reinforcement, with respect to the underestimated crack spacing for the single 3 10 tex yam. For the pillart-tricot, the crack spacing is more than doubled and for the doubletricot binding crack spacing is more than trebled. A detailed analysis of the specimens after cracking shows that when a fabric was used as reinforcement, the cracks appear almost exclusively at the position of transverse yams. During the production of the specimens, the mutual position of the individual fabric layers had not been considered. However, the cracking pattern seems to be completely controlled by the transverse yams. (Fig. 4) In some cases it has been observed that the cracking pattern is fully identical with the position of the transverse yams [9]. That cracking is controlled by transverse yams is also indicated by the considerably reduced cracking stress during state IIa (Fig. 2). The reason for this behavior can be found in the yam structure. Yams are not able to carry appreciable loads in transverse direction. With respect to the low aspect ratio for yams in the fabric it becomes obvious that transverse yams directly reduce the tension capacity of the material.

28 1

Efficiency of multifilament reinforcement in cementitious composites

1 x NEGARC62001

2 x NEG-ARG310-01

(a) mean crack spacing

1 xNEG-ARG620-01

2xNEG-ARG310-01

(b) mean bond force

Fig. 3: Mean crack spacing for different fabric types 10 mm

(a) 620 tex vam. tricot bindine

Fig. 4: Cracking patterns for different fabric types made of NEG-ARG620-01

In terms of bond we can calculate a bond force related to length. According to [ 101 this is

where z is the bond stress, Uf is the yam perimeter, V, is the matrix volume fraction, V , is the fiber volume fraction, om,is the matrix tensile stress, X is the mean crack spacing and Af is the cross section per yam. The related bond force z Uf is a bond indicator that describes the load a yam (of a fabric) can transfer per unit length. As expected from the cracking pattern,

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Frank JESSE

the yarn reinforcement achieved the highest related bond forces (Fig. 3 (b)). Again the tricot binding shows better bond performance than pillart-tricot and double-tricot whereas doubletricot reaches only 50 % of the related bond force of pillart-tricot, 25 % and 20 % of the related bond force of tricot and pure yarn respectively. Comparing these results with the surface available for bond between matrix and filaments (Fig. 1, Table 2), a strong correlation between available surface and bond performance in terms of related bond force can be postulated. That area of the fiber surface that is masked by the needle threat is not available for bond between matrix and fibers. Reduced bond performance can be the reason for increased strains after multiple cracking compared to the yam reinforced specimens, because reduced bond performance will also reduce the tension stiffening effect. Some studies have also noticed crimped fiber geometry [9],which can also lead to increased strain after multiple cracking. Composite strength The test results are very different concerning strength. For both yarn types, unidirectional yams and the fabric with tricot binding obtain almost identical values. Furthermore, these values are much higher than those obtained from double-tricot and pillart-tricot fabrics. As is typical of double-tricot, the lowest strength has been obtained from this fabric, which is less than 50 % of the reference with straight yams. The strength of pillart-tricot is in between tricot and double-tricot for both yams. Bond distortion hypothesis When comparing Fig. 1 and the stress-strain-curve it becomes obvious that there is a strong correlation between the geometry of the AR-glass yarns and the binding. The yarn geometry of fabrics with tricot weaving is very similar to that of unprocessed yarn. Their cross sections show a flat ellipse. Specimens reinforced with these two types of reinforcement show a very similar stress-strain-curve. Compared to tricot, using double-tricot and pillart-tricot leads to a much more compact cross section of the yam. But this change in cross section does not explain the reduced strength of the fabric reinforced specimens when compared to the one reinforced with yarn. As is well known from different studies, a more compact yarn cross section improves the inner bond performance, which leads to an increasing strength. Consequently, there must be at least one other mechanism that contradicts this effect. One of the major differences between the yarn and the different fabric is the amount and the quality of the fiber-matrix-interface (again see Fig. 1). For the yarn, the filament-matrix interface is assumed to be macroscopic homogeneous. Using tricot binding, the needle thread crosses the reinforcing yarn in an angle of approximately 45'. This means that there are parts of the yarn surface that have no direct contact to the matrix. Using pillart-tricot or doubletricot binding, much larger areas of the filament surface are covered by the needle threat and thus these areas have no direct contact to the matrix. In order to explain what happens to the strain distribution in the filaments, a simple model, originally developed by Ohno & Hannant [ 111, is being used. For simplification it is assumed that the bond stress is constant in undisturbed areas of the filament-matrix interface and zero in disturbed areas, i.e. no load is transferred at the filament-matrix-interfacein disturbed areas. The situation for a bond distortion caused by transverse yarns is illustrated in Fig. 5. The strain is constant in the contact area between longitudinal and transverse yarn because there is no stress transfer over the bond interface. Outside the disturbed area the curves for disturbed and undisturbed filaments have the same incline because the bond stress is assumed to be constant. Equilibrium requires an identical area under both strain curves. This results in a higher strain of undisturbed filaments in the area of bond distortions. But higher stress filaments fail first and thus the ultimate capacity of the material is reduced by bond

283

Efficiency of multifilament reinforcement in cementitious composites

distortions. The same situation applies in principle in the case of bond distortions caused by needle threads crossing the yam.

transverse yam longitudinalyam

1

-crackspacmg _ _ _ _ ~

I

X

-4

Fig. 5: Strain distribution of marginal filaments with and without bond distortion caused by transverse yams

CONCLUSIONS By means of an extensive experimental study different types of bindings for textile reinforcements have been studied. It has been found that the stress-strain-behavior of the composite undergoing uniaxial tension and the cracking behavior can change considerably. The stress level during the multiple cracking is reduced for all types of fabric reinforcement. A detailed analysis of important geometrical parameters of the yam shows that the yam section geometry can change heavily depending on the binding technology used during textile manufacturing. In fact the perimeter is reduced, packing density is increased and the aspect ratio is changed towards a more compact yam shape. This means that textile reinforced concrete is very sensitive to bond performance of the multifilament yams used in the fabric reinforcement. On the one hand, amore compact fiber geometry decreases the area bonded to the matrix and thus increases crack spacing and crack widths. Simultaneously, the strength is reduced because load transfer to the inner filaments causes larger differences of outer and inner filaments. On the other hand, a higher packing density improves the inner bond and thus reduces strain differences between inner and outer filaments of a fiber bundle. This should result in a strength increase. But the test results clearly show that the former aspect is dominating. For a hrther optimization of the material, the bond performance has to be increased considerably in order to achieve a better overall performance concerning cracking and strength. Furthermore, an additional effect has been found to be responsible for the dramatic drop in material strength. Based on a simple theoretical model, the hypothesis is introduced that

284

Frank JESSE

bond distortions caused by transverse yams and crossing needle threads also lead to a strength decrease.

ACKNOWLEDGEMENTS The authors gratefully acknowledge the support of the German Research Foundation (DFG) in the framework of the Collaborative Research Center 528 entitled “Textile Reinforcements for Structural Strengtheningand Repair”.

REFERENCES 1. Curbach, M., Jesse, F., High-Performance Textile-Reinforced Concrete. Structural Engineering International: Journal of the International Association for Bridge and Structural Engineering (IABSE) 9, 1999, (4), pp. 289-291 2. Jesse, F.; Curbach, M., The present and the fiture of textile reinforced concrete. Burgoyne, C. (Edt.), FRPRCS-5 Fibre-reinforced plastics for reinforced concrete structures. London : Thomas Telford, S. 593-4305 3. Briickner, A., Ortlepp, R., Curbach, M., Textile reinforced concrete for strengthening in bending and shear. Materials and Structures, accepted for publication, doi: 10.1617/14283 4. T. C. Triantafillou, T. C., Papanicolaou, C. G., Zissimopoulos, P. and Laourdekis, T., Concrete Confinement with Textile-Reinforced Mortar Jackets. ACI Structural Journal, 103,2006, pp. 28-37 5. Peled, A., Mobasher, B., Sueki, S., A comparison of processing technologies for the manufacture of textile cement-base composites. Kovler, K., Marchand, J., Mindess S., Weiss, J., International RILEM Symposium on Concrete Science and Engineering: A Tribute to Arnon Bentur. RILEM Publications SARL, 2004, 187-202, doi: 10.1617/29 12143586.017 6. Jesse, F., Curbach, M.: Strength of Continuous AR-Glass Fibre Reinforcement for Cementitious Composites. Naaman, A. E. (edt.); Reinhardt, H.-W. (edt.): High Performance Fibre Reinforced Cementitious Composites HPFRCC-4. Proceedings of the Fourth International RILEM Workshop. Bagneux, Francs, RILEM Publications s.a.r.l., 2003, p. 337-348 7. Abdkader, A., Characterization and modeling of AR-glass filament yarn used as concrete reinforcement. (in German) - Technische Universitilt Dresden, PhD-Thesis, 2004 8. Jesse, F., Curbach, M., A new approach for determining geometrical properties of glass fibre reinforcement in grc composites. In: di Prisco, M., Felicetti, R., Plizzari, G. A. (edt.), Fitwe-Reinforced Concretes :Proceedings of the “Sixth International REEM-Symposium - BEFIB 2004“, Varenna, 20.-22.9.2004. Bagneux : RILEM, 2004, S. 267-278 9. Jesse, F., Load Bearing Behaviour of Filament Yarns in a Cementitious Matrix (in German). PhD-Thesis, Dresden, Faculty of Civiel Engineering, Technische Universitilt Dresden, 2004 - urn:nbn:de:swb:14-1122970324369-39398 10. Purnell, P., Buchanan, A. J., Short, N. R., Page, C. L., Majumdar, A. J., Determination of bond strength in glass fibre reinforced cement using petrography and image analysis, J. Mater. Sci. 35,2000, pp. 4653-4659 11. Ohno, S., Hannant, D. J., Modeling the Stress-Strain Responce of Continuous Fibre Reinforced Cement Composites. ACI Materials Journal 91, 1994, NO. 3, p. 306-312

Proc. Int. Symp. ‘Y3rittleMatrix Composites 8” A.M. Brandt, KC. Li and I. H.Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

VOLUME CHANGES OF FIBRE CONCRETE WITHSTEEL AND SYNTHETIC FIBRES DuSan S P h ” , Jan VODICKA”, Marian ABRAMOWICZ” Department of Concrete Structures and Bridges Faculty of Civil Engineering, Czech Technical University in Prague T h h o v a 7,166 29 Praha 6, Czech Republic e-mail: [email protected],[email protected],cz 2, Warsaw Technical University Armii Ludowej 16,OO-661 Warsaw, Poland

ABSTRACT The verification and precise formulation of volume changes of structural concretes (creep, shrinkage) are very difficult (see models like: B3 - Bazant [8] or EC2 - European Standard 151, etc.). However, many structural analyses cannot be performed without considering the effects of those properties. Similar situation is in the analysis of concrete structures with steel, synthetic and others fibres and the accurate formulation of volume changes of fibre concrete is much more challenging because of diversity of FRC types. Structural fibre concretes FRC are used since many years and the fresh mix is blended and mixed particularly for a specific application, according to required properties. In the paper steel and synthetic fibres used for FRC, their volume changes measured and calculated are discussed. The results of measurements carried on during last twenty years are summarized. Basic differences in results of synthetic and steel fibre concretes are obtained from laboratory testing and from analytical model, which should replace experimental tests. The volume changes from tests are confirmed by measurements of Young’s modulus and compressive strength for FRC with different types of fibres (synthetic and steel).

Keywords Creep, Shrinkage, Compressive strength, Structural Synthetic Fibres INTRODUCTION

In the first stage of practical application of fibres in concretes, rapid development was observed only for steel and other metallic fibres. Because of that, steel fibres were used at the beginning of this experimental investigation. Results and other information (type of specimens, used fibres and measurement methods) related to this part of research were published during past conferences of BMC and the model for calculation was presented at BMC6 and BMC7 [l], [2]. Synthetic fibres (SF) were used at a greater scale in this investigation since 2000 and at present they are most often known as “Structural Synthetic Fibres”. Therefore, the paper is focused on verification of mechanical properties of fibre concrete with SF and the volume changes measured during last five years. The applicability of the previously proposed model for creep (per analogy to shrinkage) is shown for fibres called Structural Synthetic Fibres. This is presented using papers published in the volume of BMC7 [2].

286

Jan VODI&A, DuSan SPORA, Marian ABRAMOWICZ

The strength and volume changes of specimens with steel fibres and with Structural Synthetic Fibres are compared, using selected series of tests carried on in the years 1996 2003 that are presented in Table 1.

EXPERIMENTAL PART The results of the measurements were partly published in BMC 7 [2] and in RILEM Symposium BEFIB 2004 [3]. New original measurements of FRC with SF are presented below. All used specimens were prisms 150x150~600mm loaded by constant compressive force within the elastic strain zone. Mix composition with fibres FERRO FORTA: Cement: CEM JJ 32,s R ..... Sand: fraction 0-4 ..... Coarse aggregate: fraction 8-16 ..... Water: ..... Plasticizer: ..... Fibres: FerroForta(l%)* ..... Fibres: Ferro Forta (0,5%)* .....

340 kg 980 kg 890 kg 170 kg 2.4 kg 9.1 kg 4.55 kg

Mix composition with steel fibres of Bohumin producer with parameters length/thickness 60 mm 10.60 mm : Cement: CEM II32,5 R ..... 491 kg Fine aggregate/Sand: fraction 0-8 ..... 791 kg Coarse aggregate: fraction 8- 16 ..... 198 kg Coarse aggregate: fraction 16-22 ..... 745 kg Water: ..... 167 kg Plasticizer: Umaform SM ..... 4.9 kg Fibres: Steel fibres (1%y ..... 78.5 kg Fibres: Steel fibres (0,5%)* ..... 39.25 kg *.

in % of volume.

The following series of test results were used for comparison between measured and calculated values. Selected series of specimens are: Series with plain concrete 1/96,4/03 (12.6) no fibres .... 0.5% steel fibres, 2/96 Series with steel fibres 60/0.60 3/96 .... 1% steel fibres, 1 /03 (3.6) . . .. 0.5% fibres FERRO FORTA, Series with structural synthetic fibres 2/03 (10.6) .... 1% fibres FERRO FORTA The following graphs show records of creep strain (Fig. 1, Fig. 3) and conditions during measurements (temperature and humidity) (Fig. 2, Fig. 4)

Volume Changes of Fibre Concrete with Steel and Synthetic Fibres

0,000

-7-

History of strain (creep) of serie 96

---

I so

200

-_~___ -I

7 7

'F

S?

287

270

300

400

35

i

Ynr (days)

4,050 I I

' unloading

[part

i

4,100

I

I I

i

8 4.150 -0.200

K

...- ....

-0.250

--.

+II y

1%-3/%

'

!

Plsineooeme

m

4.300

steel fibres 6010 6 (0.5 %) ~...i rteelfibre360/0.6(1%)

~

~

~

\ 0%. 1/96

4.350

J

Fig. 1 Strain due to creep during 1 year measurements of series 1/96 (plain concrete), 2/96 (0.5% steel fibres) and 3/96 (1% steel fibres).

Hstory of humidity and temperature Series 96 23.0

1

i 22.0

I

/i

21.0

60

Fig. 2 Curves of humidity and temperature recorded during test of series 1/96, 2/96 and 3/96. The negligible influence of fluctuations of humidity and temperature on the curve of creep strain is evident from Fig.1 and Fig. 2. Because variations of humidity had low influence on creep, it was possible to calculate the creep strain taking into account weighted average of humidity during test period.

288

Jan VODI&,

+

DuSan SPORA, Marian ABRAMOWZCZ

+

History of creep measurement 0.W 1

rnS(D.y.1

t

Jw

4.m

4.1W

z

0,150

E

4,WO

4.250

Fig. 3 Strain due to creep during 1.5 year measurement of selected series 1/03 (plain concrete), 2/03 (0.5% fibres Ferro Forta) and 3/03 (1% fibres Ferro Forta) Hktofy of humidityand temperatun of Serbs 03

Amount 0.5 year 80.0

80.0 I

l.;,O; 21.0

I 1 I II

I

19.0

- .

-i

E

18.0

50.0

Fig. 4 Curves of humidity and temperature recorded during test of series 1/03,2/03 and 3/03 Even a strong rise of humidity had only limited influence on the fluctuations of curve of creep strain as it is evident from Figs.3 and 4.Like in Fig. 2 humidity had a small influence on creep strain and it was possible to calculate it with weighted average of humidity during period of the test.

Volume Changes of Fibre Concrete with Steel and Synthetic Fibres

289

MODEL AND HYPOTHESIS ON DETERMINATION OF VOLUME CHANGES

Hypothesis The model for calculation of volume changes of FRCs consists of two parts. The first part reflects the fact that the main share on the volume changes is related to the composition of cement matrix. For this composition the calculation using the mentioned model is certainly correct as it exactly reflects all the influences and with acceptable accuracy determines the volume changes of plain concrete. The influence of fibre reinforcement forms the other part of the calculation model. This part should characterize the remaining facts derived from the set of results. The ratio of basic strengths of FRCs vs. plain concrete with identical cement matrix composition (compressive strength for creep calculation and splitting tensile strength for shrinkage) seems to be an acceptable expression. The advantage of application of these basic strengths is that they are measured on the same shape of specimens like the plain concrete, for example cubes of 150 nun. The main advantage of that model for FRCs is that the determined strengths reflect the complete FRC technology, which in comparison with plain concrete is much more complex (mix design, maintaining of the homogeneity, etc.), [3] Model Basic step is based on the database of experimental results for determining K

J ~ , (rcJ,h) . ~ ~ ~

Formulation of hypothesis for creep strain of FRCs:

E/.ereep,exp.

Ec.c'eep,exp..

fce,exp.K f .creep,exp.

(1)

ffi,exp.

Hypothesis analogy for shrinkage strain determination for FRCs:

where: rcjCreep. .. coefficient of agreement of the hypothesis for creep . .. coefficient of agreement of the hypothesis for shrinkage KJsh (Coefficient rcfdescribes a variance of laboratory measured results and theory of hypothesis.) Here index exp. indicates data measured in experiment and symbols without index exp. indicate theoretically calculated values. Final step is based on determined coefficients in database of results K theoretical calculation of volume changes.

J (KJsh) ~ and ~ full ~ ~

Creep strain determination of FRCs Ef

fcc.

.creep

= Ee,creep..-.Kj,creep f/c.

and by analogy, the same for shrinkage of FRCs f .sh.

= 'c,sh

fc,,l '-

' Kf,sh.

(2 '>

f/,SPl.

From the detailed analysis of measured strengths and strains of specimens with structural synthetic fibres collected in the database of the test results it is possible to predict that the

~

290

Jan V O D I m , DuSan SPURA, Marian ABRAMOWICZ

basic hypothesis of strain (ECreep, &sh) is applicable also for structural synthetic fibres. During these analysis it was discovered also the reduction of Young’s modulus and an increase of strain due to the added fibres reinforcement, in this case structural synthetic fibres, r41.

Simplified formulae for determining volume changes are shown below, without agreement (correction) coefficients K J for~ measured ~ ~ and ~ calculated creep strain and KJsh for shrinkage strain published in BMC7 [2] and BEFIB 2004 [3]. The model is presented in the simple formulae because the value of coefficient KJ is approximately equal to one (h 159’0) and hence the coefficient KJ significance for the practical structure design is small. f

Analogy for shrinkage (in this paper is not presented in detail).

- ‘c,sh

‘f,sh

fct,spl ‘

f.,SPl

Where:

~ __.._.. .

E E

J

....... ~ .......

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

strain ~ due~to creep ~ of ~plain concrete ~ strain of FRC ~ ~ due~to creep ~ strain due to shrinkage of plain concrete strain due to shrinkage of FRC compressive strength of plain concrete measured on cubes (shape 150 mm x 150 mm x 150 mm) compressive strength FRC concrete measured on cubes (shape 150 mm x 150 mm x 150 mm) splitting tensile strength of plain concrete measured on cubes (shape 150 mm x 150 mm x 150 mm) splitting tensile strength of FRC measured on cubes (shape 150 mm x 150 mm x 150 mm)

Table 1 Selected results of series 96 and 03 with very good agreement between measured and calculated creen values Serie (original labeling)

Reinforced (type and dosage in volume %)

Compressive Rntlo of Compressive at pge

elastic

Volume Changes of Fibre Concrete with Steel and Synthetic Fibres

29 1

Values of strain due to creep are presented in Table 1. These values were measured and calculated according to the expression (3). This calculation is very simple in the comparison with the model for calculation of the strain, according to B3 or EC2. The calculation may be used in the case when fibre concrete is applied, because it expresses quite well relation between an increase of strain from creep and an influence of volume of fibres in fibre concrete. The strength of fibre concrete is the main factor influencing the strain. The formula (3) does not include all conditions of environment, but mainly the effect of humidity during test. The values of strain presented in Table 1 were measured on series shown in Figure 1 when the humidity was mostly constant and in Figure 2 when the humidity was variable due to environmental changes. The values of strains of series shown in the Figure 3 were measured when the humidity was significantly variable during one and a half year. A difference between the measured and the calculated values of strain is therefore partly significant and the difference of results between the measured and calculated values of strain increased. For FRCs with SF calculated values of strain were greater than experimentally measured values. The smaller measured values may be explained by 0.5 year increase of humidity. (See highlighted part in Fig. 4.) Strain values calculated according to the model may be assumed as acceptable because of small differences between experimental and calculated values (smaller than 8 %). SUMMARY AND CONCLUSIONS The aim of this paper is to present the original model for calculation of shrinkage and creep of fibre concrete which respects the influence of various types of fibres, their shapes and bonding to the matrix. The model was made on the basis of the test results of strain from creep and analogically for shrinkage. The results were verified on different types of fibres during 15 years and their are considered acceptable until a more perfect model will be proposed. Small influence of variations of humidity and temperature on the curves of creep strain were identified and used to simplify the calculation of humidity influence in the model. There is always a possibility to calculate the influence of humidity with weighted average values during considered period of time. The test results presented in Table 1 show that fibres can change strain in the case when the steel fibre concrete is produced properly (produced properly mean standard or high standard properties: good homogeneity without agglomerating of steel fibres, good compaction and porosity, minimum dispersion of mass density) and with synthetic fibres and for these fibres that condition is not valid (lower problems with homogeneity, compaction or porosity than with steel fibres). A decrease of creep strain with increase of the volume of steel fibers have to be expected (See Tab. 1 - series 2/96 (0,5%) and 3/96 (l%).) The obtained results are confirmed by other records and calculations of many years’ standing database. Small difference of decrease experimentally measured strain of SF in comparison with calculated strain (in table 1) is caused by long time (amount half of year see Fig. 4)high humidity. (See highlighted part in Fig. 4.) On the other hand, the strain is increasing when the synthetic fibres are used in a large volume. Both of these effects could be see on the compressive strength of fibre concrete. Therefore, the compressive strength which this reality predestines was accepted as a basic quantity in the model which expressed strain of fibre concrete.

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Jan VODICkA, DuSan SPORA, Marian ABRAMO WZCZ

Proposed model for calculation of creep strain for steel fibre reinforced concretes may be applied for synthetic fibres as well. The calculation model can be so used for design the FRC structures. There were small differences between calculated and measured creep strain recorded during 15 year tests and the proposed model may be considered as confirmed. ACKNOWLEDGEMENT

This investigation was supported by Grant Agency of the Czech Republic No. 103/06/0685 and by the Research Programme VZ 01 MSM 6840770001. REFERENCES

1. Kratkq, J., Trtik, K., VodiCka, J., Spiira D., Abramowicz, M. ,,Determination of Creep and Shrinkage of Steel Fibre Reinforced Concrete". Proceedings of the Sixth International Symposium on BMC6, Warsaw, Poland, October. 2000, pp 352-356 2. Sp6ra D., VodiEka, J., Krhtkjr, J., Abramowicz, M. ,,Method of Determination of Creep and Shrinkage of SFRC ". Proceedings of the Seventh International Symposium on Brittle Matrix Composites (BMC7), Warsaw, Poland, October. 2003, pp 515-521 3. Spiira D., Kratlj, J., VodiEka, J. ,,Model for Calculation of Creep and Shrinkage of Fibre Reinforced Concrete". Proceedings of the Sixth International RILEM Symposium on FibreReinforced Concretes, BEFIB 2004, Varenna, Italy, October. 2004, pp 895-902 4. VokaC M., BouSka P., VodiEka J. "Ductility of Fibre Reinforced Concrete (Pfetvarnk vlastnosti vlhobetonu)" in Proceedings 3d International Conference FC 2005 ,Malenovice, Czech Republic, June 2005, pp 153 -156 (in Czech) 5. European standard - EN 1992-1 (2nd draft), ,,Eurocode 2: Design of concrete structures Part 1 : General rules and rules for buildings", (September 2001) 6. ACI 209, ,,Prediction of Creep, Shrinkage, and Temperature Effects in Concrete Structures", (Manual of Concrete Practice, ACI, 1994), pp. 209R-1 to 209R-47 7. Spura D., Kratky J., Vodicka J.: ,,Volume Changes in Fiber Concrete - Experiment, Calculation", in Proceedings of the Second International Symposium of ,,Fiber Concrete and HPC".(Sekurkon, Prague 2,2003), pp. 144-150 8. Baiant, Z.P., Baweja, S.,: ,,Creep and Shrinkage Prediction Model for Analysis ad Design of Concrete Structures": Model B3. Presented at the Adam Neville Symposium, Atlanta, Georgia, November 1997

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, V.C. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

IMPACT BEHAVIOR OF FABRIC-CEMENT BASED COMPOSITES Efrat BUTNARIUa,Alva PELEDb,and Barzin MOBASHER' Material Engineering Department, Ben Gurion University, Beer Sheva, Israel, [email protected] Structural Engineering Department, Ben Gurion University, Beer Sheva Israel, [email protected] Civil and Environmental Engineering Department, Arizona State University, Tempe, AZ, USA, [email protected] a

ABSTRACT A drop weight three point bending test machine was used to study the dynamic behavior of fabric-cement based composites. Hybrid sandwich specimens made from combinations of short fibers and fabrics were prepared by manual hatch process. In addition, laminated composites were produced by the pultrusion process. The acceleration, deflection and loads were recorded to compute the strength and energy absorption. Composites made from short polypropylene and polyvinylalcohol fibers are compared with polypropylene and AR glass fabric composites. The specimens were tested in three different directions: parallel and perpendicular to the fabric length as well as through the specimen thickness (x and z directions of the specimens). The microstructure of failure surfaces were studied using optical and scanning electron microscopy (SEM) and observations were correlated with the mechanical properties of the different systems. It is demonstrated that fabrics are very promising reinforcements for cement-based elements exposed to dynamic loading. Composites reinforced with fabrics showed significantly greater impact behavior compared to those reinforced with fibers. The best behavior was reported for the pultruded composites made from PE knitted fabrics.

Keywords Impact, Fabric, Cement composite, Textile, Fiber INTRODUCTION Cementitious materials may be subjected to dynamic loading for a variety of reasons including: fast moving traffic, wind gusts, blast explosions, missiles, earthquakes, projectiles, and machine vibrations. Due to the inherent brittleness and low tensile strength of most cement-based elements, such impact loadings can cause severe damage, resulting in loss of functionality. Fiber reinforcement is one of the most effective means of enhancing the impact and blast resistance, in both strength and energy [l-21. The research to date on fabric-cement composites clearly demonstrates a significant improvement in the energy absorption capacity of fabric reinforced cement composites under static loading as compared to plain concrete materials and other fiber cement composites [4-71. Therefore, fabric reinforced cement products are expected to exhibit high resistance under impact loads as well. The aim of this work was to study the dynamic behavior of fabric-cement based composites. Hybrid sandwich specimens made from combinations of short fibers and fabrics

294

Efrat BUZVARIU, Alva PELED, Banin MOBASHER

in addition to laminated composites produced by the pultrusion process were prepared and examined. Comparison was made with composites made from short polypropylene and polyvinylalcohol fibers (PVA) were compared with polypropylene and AR glass fabric composites. Specimens were tested in three different directions: parallel and perpendicular to the fabric length as well as through the specimen thickness (x and z directions of the specimens). The microstructure of the different elements was studied using a scanning electron microscope (SEM) and an optical microscope, these observations were correlated with the mechanical properties of the different systems. EXPERIMENTAL PROGRAM Fabric and fibers This work examined cement based composites reinforced with different combinations of short chopped fibers and two dimensional fabrics (2D). The short fibers used were polypropylene (PP) and Polyvinyl alcohol (PVA) having 6 and 2 mm lengths, respectively. Two types of 2D fabrics were used in this study: short weft knitted (Fig. la) and bonded (Fig. 1b). In the short weft knitted fabric the warp yams are knitted into stitches and bind together a set of yams, which are laid-in intermittently in both the weft and the warp directions (in a zigzag form). In the bonded fabric two sets of straight perpendicular yams are glued together at the junction points. The knit fabric was made from polyethylene yarns (PE) and the bonded was made of AR (Alkali Resistance) Glass. Table 1 summarizes the mechanical properties of the different materials.

1 1 1 1 1 1 1 Fig 1: Fabric geometry: (a) warp knitted short weft, (b) bonded T reinforcement

Preparation of specimens In all cases, cement paste with a waterkement ratio of 0.4 was used as the matrix. Dimensions of all tested specimens were 160~40x20mm. Four different series of specimens

Impact behavior of fabric-cement based composites

295

were prepared using both a conventional casting method and the pultrusion method as described in the following: Short Fibers - short fibers, PP and PVA were added to the cement paste during mixing and cast into moulds manually ; 2) Sandwich of fabrics - In this case one layer of fabric was placed at the bottom of the specimen over a thin layer of paste (2 mm). The mold was then cast with cement paste and a second layer of fabric followed by a 2 mm thick layer of cement paste placed on the top layer. These sandwich specimens were prepared from AR glass fabric only; 3 ) Sandwich of fabrics with shortfibers- This type of sandwich element was prepared using the cast method as described above, however, in this case the cement paste in between the fabric layers contained 1% by volume of short PP fibers or PVA fibers. Knitted PE fabrics were used to prepare these specimens; and 4) Pultruded - The fourth specimen series was made by the pultrusion process. In the pultrusion process fabrics were passed through a slurry infiltration chamber, and then pulled through a set of rollers to squeeze the paste in between the fabric openings while removing excessive paste. The fabric-cement composite laminate sheets were then formed on a plate shaped mandrel. Each cement board contained 20 layers. Knitted PE fabrics were used to produce the pultruded cement composites. This technology allows good penetration of the cement in between the fabric spaces, homogeneous laminated composite, and high fabric contents. For more details on the pultrusion process see Refs. 6-7. All the specimens were cut 24 hours after casting or pultrusion process. All specimens cast and pultruded were cured in moist (100% relative humidity) for 21 days and tested in impact after 28 days from casting. Table 2 summarized all the tested specimens as well as their volume fraction of reinforcement.

Specimen type PP fibers PVA fibers PE sandwich +PP fibers PE sandwich

Test direction

Vf %

-------

1.o 1.o 1.6 1.6 1.6

Vertical Horizontal Horizontal

I

5.7

I

I AR glass sandwich I Horizontal I

1.5

I

Pultrusion

I

Vertical

Impact strength MPa 9.8 8.7 15.1 14.8 12.4

Efficiency factor

Toughness N*mm

1.4 1.o 1.8 1.7 1.2

1069 77 2839 1448 517

31.8

2.1

I

2492

TESTING Drop weight impact tests A three point bending drop weight impact test using a free-fall drop of an instrumented hammer on a specimen was used to characterize the impact behavior is shown in Fig. 2. The drop height was 5 cm. The experimental set-up was include: two load cell of 89 KN capacity one placed on the hammer and the order one under the supports; a linear variable displacement transducer with a range of 23.17mm; one accelerometers placed on the hammer with a range of +log; one accelerometers placed on the specimen with a range of +lOOg; and

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296

a total hammer weight of 155 N. A rubber pad was placed between the hammer tip and the specimen. Two loading arrangements were carried out: the load was parallel to and perpendicular to the fabric layers as presented in Fig.2. The loading schedule for each system is presented in Table 2. Stress deflection curves were evaluated for all specimens and for each series, a typical stress deflection curve was chosen for comparison purposes. In all cases the impact strength at peak and toughness were calculated. In order to compare between the different composites composed from the different materials, fabrics and fibers, efficiency factors were calculated, by dividing the impact strength of the composite with the volume fraction and strength of its reinforcement.

Accelerometer

Hammer Rubber pad

Lever arm

Specimen Accelerometer

LVDT

Fig. 2: Impact set up: (a) specimen loaded horizontally to fabric layers, and (b) specimen loaded vertically to fabric layers Microstructure characteristics Microstructure characteristics of the different composites were observed using Scanning Electron Microscopy (SEM). For these observations, fiagments of specimens obtained after impact tests were dried at 60°C and gold-coated. The failure mechanisms of the composite as well as the matrix penetration in between the opening of the fabrics were evaluated.

RESULTS Table 2 summarizes the impact strength and toughness for all tested specimens. Effects of fabric types, short fibers, and differences between the cast and pultruded specimens are discussed in the following sections based on the data provided. Short fibers composites Fig. 3 compares the impact behavior of composites made with PP and PVA fibers. The performances of the PP composite are much better than that of the PVA composites in terms of slight improvement in strength and pronounced improvements in toughness. It should be noted that the bond development between the cement matrix and the PP is much lower than that developed with the PVA fibers. Bond strength values of 0.81 MPa and 3.59 MPa were obtained for the PP and PVA fibers, respectively [8]. The differences in the bond strengths, may lead to differences in failure mechanisms of the two composite systems; fiber pullout in the case of PP and fiber fracture in the case of PVA. The fracture and pullout mechanism of the PVA and PP fibers are observed in Fig.4, which shows the edge of the fiber at the failure

297

Impact behavior of fabric-cement based composites

surface of the composite. The surface of the PP fiber is relatively flat and with no damage while the surface of the PVA fiber is damaged. Fracture of the PVA fiber can lead to the brittle behavior of this composite, and the pull out of the PP fibers can result in a tougher and more ductile behavior of the composite. Also note that the length difference of the PP fibers (20 mm) as compared to the PVA (6 mm) may also contribute to some of the differences in the impact behavior of the two composites.

10 $

8

0 0

1

2

3

Deflection, m m Fig. 3: Comparison of impact behavior of composites with short fibers, PVA and PP

Fig. 4: SEM images of the different fibers at the fiacture surface of the composite: (a) PP and (b) PVA Effect of fabric - sandwich composites Fig. 5 presents the effect of the knitted PE fabrics on the impact behavior of the sandwich composites. Sandwich composites containing PVA and PP short fibers are compared. In both composites the load was applied horizontally to the fabric layers. The sandwich composite with the PP short fibers performed much better in strength and toughness than that with the short PVA fibers (Table 2). This correlates with the results presented in Fig. 3 i.e., the improved behavior of the short PP fibers composite as compared with the short PVA fibers composite. The improvement of the sandwich with the PP fibers is most significant in toughness, exhibiting toughness value of 1448 MPa for the PE fabric-PP sandwich compared to toughness value of only 517 MPa for the PE fabric-PVA sandwich. In the case of the PE-

Efrat BUTNARIU Alva PELED, Barzin MOBASHER

298

PVA sandwich, the peak stress is observed at very low deflection (first crack), and is only slightly smaller than that of the PE-PP sandwich, 12.4 and 14.8 MPa, respectively. However, after reaching the peak, the stress is significantly reduced in the case of the PE-PVA sandwich. On the other hand, the PE fabric-PP sandwich exhibits relatively high impact stresses at the multiple cracking region. The fabric mainly with the PP fibers also exhibits relatively ductile behavior as compared with short fiber composites.

16 14 E 12 m" 8 10

,

-

-

-

_ _ _ _ _ _ __ __

_-

_-

-

--

-- - - -

-

-

h

I s 8 c

Z 6 2

CI

4 2 0 0

1

2

3

4

5

6

Deflection, mm Fig. 5: Impact behavior of sandwich composites made from PE fabric sandwich and PP or PVA short fibers The benefit of the PE-PP sandwich composites is also observed when comparing the efficiency factors of the different systems. This value takes into account the volume content of the reinforcement and its tensile strength and allows comparison of the different systems on the same basis. Efficiency values of 1.7 for the PE-PP sandwich compared to a value of only 1.2 for the PE-PVA sandwich composite was calculated (Table 2). When comparing the efficiency factors of the PE-PVA sandwich to its fiber composite, i.e., short PVA fiber composites, the values are greater for the sandwich (fabric) composite, 1.0 and 1.2 for the short fiber composite and sandwich composite, respectively. This indicates that the fabric layers at the top and bottom of the specimen improve the composite performance. Similar trend is observed with the PE-PP sandwich, the efficiency factor of the sandwich (fabric) composite is greater than that of the PP short fiber composite, 1.7 and 1.4, respectively, indicating that the fabric layers contribute to the composite performance. Laminated pultruded composites Fig. 6 compares the stress - deflection curve of pultruded elements loaded in the two directions of horizontal and vertical to the fabric layers. The excellent impact behavior of the pultruded composite is obvious in this figure for both loading directions. Note that the content of reinforcement in the pultruded composite is relatively high, 5.7% by volume, as compared to all the other tested systems, PE fabric-PP sandwich (Vt=1.6%) and short PP fibers ( v ~ l % )However, . when the reinforcing efficiency factors of the different systems are compared (Fig.7) the pultruded composite is the best followed by the PE fabric-PP sandwich, while the short fiber composite is the lowest. This is observed for both loading arrangement systems, vertical and horizontal, indicating the advantage of the pultrusion process to produce

299

Impact behavior of fabric-cement based composites

cement laminated composites as discussed in earlier publications [6-71. The pultrusion process improves the bond between the fabric and the cement matrix as it allows better matrix penetration in between the fabric openings. This improved impact behavior suggests that pultruded composites are very attractive for elements exposed to dynamic loading. 40

9 3530 d

25

m 9)

h 20

m

15

g 10 d

- 5

0 0

1

2

3

4 5 Deflection, mm

6

7

Fig. 6: Impact behavior of pultmded composites made from PE fabrics, tested horizontally and vertically to the fabric layers

3.00 L

2S0

Horizontal

0

u

g

2.00

Lr

5, 1.50 E

w

g

. r (

w

1.00

0.50 0.00 -

PE Sandwich Pultruded +PP fibers Fig. 7: Efficiency factors of composites with PP short fibers, PE fabric sandwich + PP composites and pultruded composites, tested at different loading directions PP fibers

When comparing the impact behavior of the pultruded composites loaded in the two directions, vertically and horizontally, the strength of the composite loaded horizontally is slightly greater than that loaded vertically (Table 2 and Fig. 6). The stiffness of the composite is highly influenced by the loading direction, when the composite is loaded vertically (perpendicular) to the fabric layers its behavior is much stiffer than that loaded horizontally (parallel) to the fabric layers (Fig. 6). Also, the toughness of the vertical composite is much

Efrat BUlXXIUU, Aha PELED, Banin MOBASHER

300

lower than that of the horizontal composite, giving much more brittle behavior. At horizontal loading the fabric layers are stretched and the crack develops through the fabric layers, which can lead to the ductile behavior of the composite. Also, tensioning of the fabrics affects the stiffness of the laminated composite, i.e., reduces the stiffness. At vertical loading the crack is developing along the fabric layer mainly through the cement matrix leading to a more brittle behavior. When the composite is loaded perpendicular to the fabric layers the cement matrix is dominant, since cracks cause a variable state of stress throughout each fabric layer. Since the un-cracked matrix has a relatively high modulus, it limits the composite ability to flex, resulting in increased stiffness. When the sample is loaded in a horizontal mode (plate mode) each fabric layer is loaded homogeneously as the matrix cracks resulting in a more ductility and a uniform straining of the layers.

Fabric types Fig. 8 compares the impact behavior of sandwich composites reinforced with PE fabric- PP sandwich composite and AR glass fabrics composite. In both cases the volume content of reinforcement is similar at about 1.5% (Table 2). The impact behavior of the sandwich PE fabric-PP composite is relatively high mainly at low deflections up to 1 mm. This is despite the much higher modulus of the AR glass fabric as compared with the PE-PP reinforcement (Table 1). The AR glass composite exhibits better performance at deflections above 1 mm. The reinforcement efficiency factor of the PE-PP sandwich composite is much greater than that of the AR glass composite, giving a value of 1.7 for the PE-PP composite and only 0.8 for the AR glass composite (Table 2). The improved performance of the PE sandwich composite can be explained by observing SEM microscope images seen in Fig 9. Fig.9a shows low penetration of the cement in between the filament of the glass bundle making up the fabric. While in Fig 9b good penetration of the cement in between the stitches of the knitted PE fabric is observed, providing improved bonding of the PE fabric by mechanical anchoring. Strong bond can lead to the efficiency of the PE sandwich during impact loading.

16 14 m" 12 10 z 8 *I, 0 a 6

AR glass fabric

d

E

a a

U

PE Sandwich +PP fibers

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Deflection, mm Fig. 8: Impact behavior of PE - PP sandwich composite and AR glass composite

Impact behavior offabric-cement based composites

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Fig. 9: SEM images of: (a) glass bundle (part of a fabric) and (b) PE knitted fabric, in the cement matrix

CONCLUSIONS Fabrics show high potential as reinforcement of cement based composite exposed to dynamic loading, as compared to short fibers. Pultruded elements exhibit excellent behavior under dynamic loading, due to high content of reinforcement and improved bonding. Low modulus fabrics are very attractive as reinforcement for cement based composite under impact loading, as observed in this work with the PE fabrics, for the different systems, pultruded and sandwich composites. The PE-PP fabric sandwich exhibits even better impact behavior than that of the glass fabric composite, when comparing their efficiency factors. The loading direction has a significant influence on the behavior of the composite under dynamic loading. When the composite is loaded parallel to the fabric layers, a large numbers of fabric layers are uniformly involved and the composite is more ductile, tougher and exhibiting relatively low stiffness, due to stretching of the fabric during loading. A vertically loaded composite exhibits relatively brittle behavior and stiffer composite.

ACKNOWLEDGMENT The authors would like to thank Polysack LTD, Israel , SAINT-GOBAIN Vetrotex Cem-Fil, and Karmray INC for their cooperation for providing the fabrics and fibers used in this study. The National Science Foundation, program 0324669-03, and the BSF (United States Israel Binational Science Foundation) are acknowledged for the financial support in this research.

REFERENCES 1. Banthia, N., Crack growth resistance of hybrid fiber reinforced cement composites. Cement and Concrete Composites, 25(1), 2003, pp 3-9. 2. Gupta, P., Banthia, N., Fiber reinforced wet-mix shotcrete under impact. J. of Materials in Civil Engineering (ASCE), February 2000, pp 81-90 3. Suaris, W., Shah, S. P., Properties of Concrete Subjected to Impact. J. of Structural Engineering, 109(7), 1983, pp 1727-1741

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4. Peled, A., Bentur, A., Geometrical characteristics and efficiency of textile fabrics for reinforcing composites", Cement and Concrete Research, 30,2000, pp 78 1-790. 5. Peled, A., Bentur, A., Yankelevsky, D., Flexural performance of cementitious composites reinforced by woven fabrics. Materials in Civil Engineering (ASCE), 1999, 11(4), pp 325330 6. Peled, A., Mobasher, B., Pultruded fabric-cement composites. ACI Materials J., 102(l), 2005, pp 15-23 7. Peled, A., Mobasher, B., Sueki, S., Technology methods in textile cement-based composites. In: "Concrete Science and Engineering, A Tribute to Arnon Bentur", K. Kovler, J. Marchand, S. Mindess, and J. Weiss, eds., March 2004, RILEM, Chicago 2004, pp 187202 8. Zagori, E. M.Sc., Thesis, Hebrew University, Israel, 2006

Proc. Int. Symp. 'BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

MODELING OF STRESS WAVE PROPAGATION IN REPAIR SYSTEMS TESTED WITH IMPACT-ECHO METHOD Andrzej GARBACZ' and Leslaw KWASNIEWSKI' Faculty of Civil Engineering, Warsaw University of Technology Armii Ludowej 16,OO-637 Warszawa, Poland 'Department of Building Materials Engineering, e-mail: [email protected] 'Institute of Structural Mechanics, e-mail: [email protected]

ABSTRACT Impact-echo method - a stress wave based method - is one of the most promising methods of a repair efficiency assessment. Propagation of stress waves in a repair system is complex phenomenon and depends on acoustic properties of repair material and its thickness as well as a quality of interface. In this work finite element (FE) method was used to simulate stress wave propagation for various models of repair system. Ls-Dyna, an explicit finite element program dedicated for transient dynamics was applied. Several cases of impact-echo test with different parameters of layer system were analyzed. Among considered model parameters were acoustic properties, layer thickness of repair material, concrete roughness, and presence of delamination at the interface. The Fourier transforms and wavelet analysis were applied to characterize signals obtained for particular repair systems. The effect of repair system parameters on wave propagation was discussed.

Keywords: repair system, impact-echo, stress wave propagation, FEM modeling, wavelet

INTRODUCTION A growing interest in application of non-destructive techniques (NDT) for evaluation of concrete structures is noted recently. This is a result of quality requirements increase for concrete in new structures. On the other hand, the percentage of repair and rehabilitation in the total building market has significantly increased [l]. The European Standard EN 1504-10 [2] and ACI Concrete Repair Manual [3] give guidelines for repair efficiency evaluation. According to those documents, the bond strength and quality of bond are the main features of the repair system necessary to be assessed. The pull-off test is commonly used to test bond strength. However, the non-destructive methods (NDT) are preferred for this purpose. A majority of NDT methods mentioned in EN 1504-10 and ACI Concrete Repair Manual for repair efficiency assessment are based on propagation of stress waves. Impact-Echo (I-E) method is treated as one of the most promising methods for this purpose. In this method, an analysis of frequency spectrum (Fast Fourier Transformation, FFT, of a signal) is used. Recently, a new tool for signal analysis - wavelet analysis - is being implemented in NDT assessment, also in the case of concrete structures. It is a multiresolution time-scale methods. Polymer-cement composites are one of the most common repair materials [4]. However, their properties, including acoustic, are different from those of concrete substrate. This can affect the wave propagation in repair system and influence results interpretation. The aim of this work was analysis of effects of repair system parameters on the stress wave propagation.

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Andrzej GARBACZ, Leslaw KWASNIEWSKI

EXPERIMENTAL Impact echo principle Impact-echo is a method for non-destructive evaluation of concrete [ 5 ] , based on the use of an elastic, low energy impact on the surface generating low frequency stress waves (mainly below 60MIz). These waves consist of compression (P), shear (S) and Rayleigh (surface) (R) components which propagate through the structure and are reflected by interfaces within the material (internal flaws such as voids, honeycomb, cracks, delaminations) or external boundaries. I-E method is very often used for quality control of various types of repair, eg. injection of cable ducts, overlays and patches, etc. (eg. [6,7]). As the stress waves generated in the I-E method (Fig. la) have low frequencies (in comparison to eg. ultrasonic) this method is less sensitivity to natural heterogeneity of concrete. Additional feature of I-E method is an application of frequency analysis besides a time-domain analysis (Fig lb). The selection of the impact source is a critical factor for defect detection in multilayer system with impactecho [8]. The impact duration determines the frequency content of the stress wave and determines the minimum flaw depth that can be detected (Fig. lc). 900,

I

;LeM Fast Fourier Transform

A!

I6 2

4

6

8

10

12

14

16

ball diameter [mm]

Figure 1. Scheme of impact-echo method (a), examples of waveform and corresponding frequency spectrum (b) and (c) effect of ball size on depth and size of defect possible to detect; fT - a thickness frequency, fD - a defect frequency Nondestructive testing of repair system is more complex than a concrete slab. Beside various defects in concrete substrate or repair material microstructure an interface quality can significantly affect the wave propagation. In the case of a multilayer system the propagation of stress waves depends also on differences in acoustic impedances (Z = pulse velocity x volume density), of both repair material and concrete substrate [9]. For two dissimilar materials, part of the energy of the vibration is refracted into the new one, Am while the other part is reflected back, A,. The wave reflection is characterized by a reflection coefficient, R

Experimental and numerical investigations with I-E method have shown that usually an interface is “visible” if absolute value of R coefficient is higher than +0.24 [ 101.

Modeling of stress wave propagation in repair systems tested with impact-echo method

305

Assumption of the FEM models Comprehensive numerical time-domain studies were carried out by using Finite Element (FE) explicit code LS-Dyna [ 1 11, dedicated for transient dynamic problems. All developed models refer to an infinite plate with total thickness of 200 mm. The FE models represent segments of the cylinder with radius 260 mm, cut off about the vertical axis positioned along the impact direction. Depending on the problem, a quarter (with two symmetry planes indicated) or a half (with one symmetry plane) of the cylinder was considered (Fig.2). In the both models two nodes are identified: node 1 - where the load was applied and node 7 - located 26.5 mm along x-axis corresponding to the vertical displacement detected by receiver. Non reflecting boundaries were applied to the cylindrical surface to represent the connection to the infinite media. Symmetry boundary conditions were assumed on the vertical cross-sectional planes, where normal displacements were constrained. To defined well reflection conditions vertical displacements were constrained on the bottom surface to simulate a connection of the concrete layer with a relatively rigid (such as steel) medium.

Figure 2. FE models (a) and contours of pressure and deformation (b) for model of repair system without (on the left) and with delamination (on the right) All FE meshes are built of regular wedge six-node elements, each with the same vertical dimension of 2 mm. The quarter model consists of 276 050 elements, the half-cylinder model has twice more elements. It is assumed that considered plate can be built of maximum 6 layers. Five upper layers, each with thickness of 20 mm, can have different material properties. The last, bottom layer of thickness 100 mm is supposed to represent concrete. As only elastic material properties are taken into account the following three parameters should be determined; E-elastic modulus, V-Poisson ratio, and p density. The material parameters assumed for the simulations are given in Table 1. It presents also calculated acoustic properties: P-wave speed C, [9]:

C, = 0.96

E ( l - V) p(1+ v)(l - 2v)

and reflection coefficient R for normal incidence:

where: Z I , Z -~acoustic impedance of concrete and repair material respectively; Z=Cp*p -acoustic impedance

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Andrzej GARBACZ, Lesiaw KWASNIEWSU

Table 1. Assumed material properties Property elastic modulus E [GPa] Poisson ratio y [-3 Density p [kg/m3] P-wave speed [ d s ] acoustic impedance Z [kg/(m2s)] reflection coefficient R [- 3

Material #1 10 0.2 2190 2162 4.7.106 0.35

Material #2 20

0.2 2190 3058 6.7.106 0.18

Material #3 30 0.2 2190 3745 8.2.106 0.08

Concrete 38 0.2 2400 4027 9.7.106 0.00

The implemented impulse loading corresponds to the impact of a dropped steel ball with a diameter of 5 mm. The time interval was determined using the experimental formula [ 121:

where: tc - impulse time interval in ,us, D - diameter of the ball in mm. Based on [ 5 ] , variation of the loading in time was approximatedby half-cycle sine curve, with the assumed maximum value of 200 N.

RESULTS AND DISCUSSION Variation of repair material properties Finite element analysis enables for comprehensive parametric study, where the influence of some parameters on final results is tested through repeated calculations. One of such parameters is a difference between material properties of the repair layer and the concrete, expressed by acoustic impedances Zi and reflection coefficient R. The cases considered in this section refer to plates with one repair layer (20 mm) made of three different materials (see Tab.1). Three calculated curves are compared with the result for the plate made of concrete only. Figure 3a shows initial disturbances with the maximum displacements caused by Rwave. Wave fronts reach node 7 at different times depending on the stress wave velocities. Figure 3b presents the FFT transformation results for all four cases, in an amplitude spectrum. It shows the relative amplitudes of the various frequencies contained in the waveform, calculated as a ratio of the absolute value to the sum of all magnitudes. The solid curve, representing amplitude spectrum for the half-sine impact impulse, was added to show the relationship between the impact impulse and the resultant signal. It was assumed that the bottom of the plate is connected to the rigid media so the wave changes its phase only when reflecting from the upper surface. For the two layer plate with one repair layer of thickness Ti and made of a material “i” there can be at least two dominant modes. One corresponds to the P-wave traveling within the first layer. If the reflection coefficient R is positive, reflecting in the interface does not change the wave’s phase and the corresponding one layer thickness frequency is [ 5 ] : ‘pi

f c =-

4q

(5)

where: I;. is a thickness of the repair layer made of the material “i ” with the velocity Cpi. There can also be recognized the mode corresponding to the waves traveling through both materials (repair material i and concrete). Expected frequency is equal to the reciprocal of the

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Modeling of stress wave propagation in repair systems tested with impact-echo method

time needed for the wave to cross both layers four times as the wave has to cross the entire thickness four times to return to the origin mode. The corresponding frequency is: 1 fMT,+C

=

4TMTi CPi

; 4Tc CPCi

Table 2 shows good correlation of calculated peak frequencies with the expected values for full frequency thickness considering the fact that presented frequency analysis is based on the discrete representation of the waveform. The waveforms were calculated with the sampling which is equal to the time integration step used in the explicit finite period At= 0.301pUs, element calculations [ l l ] . The Fast Fourier Transform was applied here for the first 8192 samples giving accuracy of frequency estimation [ 6 ] :Af=l/ N.At = 406 Hz. Figure 3b shows that for smaller values of the reflection coefficient the peak corresponding to the one thickness frequency disappears. The curve has many local maximums but no one can be recognized as a representing a dominant mode.

Relative displacement

Relative magnitude

Frequency [kHz]

Figure 3. Time histories of the relative vertical displacement at node 7 (a) and (b) corresponding amplitude spectra

Repair

reflection coefficient

Expected

Calculated

Expected

full thickness

full thickness

one thickness

Calculated one thickness

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Variation of repair layer thickness

The peak notifymg one thickness (FigAa) appears moves toward lower frequencies as the repair thickness increases and can appear close to the total thickness frequency as it occurred for T1=100mm (Fig.4b). It means that for thicker layers the thickness frequency can be not recognizable even for the reflection coefficient R>24%. This is in a good agreement with experimental observation that the interface is invisible if the reflection coefficient R< 0.24 [lo]. Additionally, the differences between expected and calculated value are less than 10 %, what is close to the results obtained by Abrams [13]. Local delamination at the interface surface

To simulate delimination (see Fig.2), the adjacent region elements are not connected and there is no contact between them, allowing that elements to penetrate. This dilatation represents a relatively large flaw according the classification presented in [S]. Two cases are taken into account, in the first, all layers are made of concrete and in the second the top four layers are made of the repair material #1 (Table 1) and the rest of concrete. The reflecting stress wave changes its sign, in the same way as reflecting from the free (concrete - air) surface. To show the deformation, displacements were enlarged 30000 times.

0'015

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Figure 4. Time histories of the relative vertical displacement at node 7 (a) and (b) expected and calculated frequencies for different repair thickness TI On the Figure 5 amplitude spectra for concrete plate and repair system (T=80mm) with and without delamination are shown. Theoretical frequencies of the wave bouncing between the top surface and the delamination are: CPC material#I: fm,, =-- -3513Hz andconcrete: fc =--5169Hz. 2T2 2T,

The maximum magnitudes appear for lower frequencies corresponding to modes where deformation extends on the entire volume between the delamination and the free face, as it is

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Modeling of stress wave propagation in repair systems tested with impact-echo method

shown in Fig.5. As the ratio of the thickness to the horizontal dimensions is smaller these modes are more similar to the modes of vibrating plate. Due to the delamination the material above it is much more flexible than the entire monolithic plate connected to the steel base. WAVELET ANALYSIS Wavelet principle

Recently, a new tool for signal analysis - wavelet analysis - is being implemented in NDT assessment, also in the case of concrete structures, eg. to analyze results of acoustic emission, detection of crack generation, etc. [7,14,15]. Wavelets are composed of a family of basis functions that are capable of describing a signal in a localized time and frequency (or scale) domain. The main advantage gained by using wavelets is the ability to perform local analysis of a signal, i.e. to zoom-in on any interval of time giving possibility to detect some hidden features of the signal. 0.06 7

Relative magnitude :: 7.68

delamination between repair material and concrete

no delamination

0

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50

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Figure 5. Comparison of amplitude spectra for concrete plate and repair system (T=80mm) with and without delamination In wavelet analysis a family of functions which are derived from a single generating function called the 'mother wavelet'. It is dilated (stretched or compressed) by a and translated in space by b to generate a set of basis functions as follows:

The function is centered at b with a spread proportional to a. Wavelet transform can be categorized into continuous (CWT) and discrete (DWT). The CWT is the sum over all time of the signal multiplied by a scaled and shifted version of a mother wavelet (Fig.6a). The discrete parameter wavelet transform uses scale and position values based on dyadic scales

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and positions. In this case, the scale and time parameters are described as follows: b=nb,,hn, m,n - integers. In practice this approach is realized be using low and highpass filters. Filtering process decomposes signal (Fig.6b) into the approximation (low frequency content) and details (high frequency components).

T

c

B

8

T A3

time

b

A4

Figure 6 Scheme of: (a) continuous wavelet transform and (b) discrete wavelet transform Wavelet analysis of selected repair systems

In this work the MATLAB environment for wavelet analysis of I-E signal obtained for FEM model of repair systems was used. CWT and DWT were performed using Daubechies wavelet (dbl0). It is clearly visible that character of time-scale diagrams depends on acoustic properties of repair material (Fig.7). The CWT and DWT for repair material of 30GPa are very closed to those for concrete plate, while for E=lOGPa is different. This confirmed that for low value of reflection coefficient the repair system can be treated as a "solid" concrete plate. This implies that NDT procedures elaborated for concrete structure can be applied for such repair system.

Modeling of stress wave propagation in repair systems tested with impact-echo method

311

Cnncrete substrate /T=ZOOmm: E=38GPa)

Concrete substrate lT=180mm: E=38GPa) + RM fl=20mm. E=30GPa)

Fig.7. CWT (left) and DWT (right) for repair system created by concrete substrate and repair material (RM) of different elasticity modulus Results of the DWT transform - details Dl-D3 and reconstructed signal were statistically analysed using standard procedure of the MATLAB. A mean absolute deviation (MAD) was selected for characterization of wavelet coefficients distribution (Fig.8).

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I /

I

l l

I

Fig.8. Mean absolute deviation of the reconstructed signal (a) and ie wavelet coefficients distribution for (b) detail 1, (c) detail 2, (d) detail 3 for repair systems with materials of different elasticity modulus Figures 7 and 8 show that as the elasticity modulus of repair material increases the CWT and DWT are close to those characteristic for concrete. Repair system with material of low elasticity modulus is characterized by higher MAD value for reconstructed signal as well as details 1 and 2. Similar analysis of CWT and DWT was performed for the systems with different thickness of repair material (Fig.9). CWT and DWT for various thickness of repair layer are close to each others and different fkom the obtained for the concrete plate.

2.0607 1.5507

1

1.OCO7 5.0608

o.oE+oo

Fig.9. Mean absolute deviation of the reconstructed signal (a) and the wavelet coefficients distribution for (b) detail 1, (c) detail 2, (d) detail 3 for systems with different thickness of repair layer The MAD value for the concrete plate with delamination was four order of magnitude higher than for the plate without delamination (Fig.10). The same trend was observed in the case of the delamination between repair material and concrete substrate even for the high value of R. In this case, the MAD value was 1.5 lower than for concrete plate.

Fig. 10. Mean absolute deviation of the reconstructed signal (a) and the wavelet coefficients distribution for (b) detail 1, (c) detail 2, (d) detail 3 for systems with and without delamination

Modeling of stress wave propagafion in repair systems tested with impact-echo method

313

CONCLUSIONS The simulation results show that repair material thickness and its acoustic properties are important factors influencing propagation of stress wave through repair system. The results obtained are in a good agreement with practical observation towards application of impact echo for assessment of concrete structures. This confirms that developed model is usefil for simulation of stress wave propagation trough repair systems evaluated with impact-echo method. Due to the variety of parameters characterizing possible tested objects, experience is required to interpret impact-echo test results. The results of wavelet analysis of I-E signal indicate that this approach is promising for assessment of bond quality in repair systems. Mean absolute deviation of reconstructed signal and wavelet coefficients properly characterized the stress wave propagation trough various repair systems simulated in this work. The usefulness of this approach for real systems should be verified on the experimental data. However, the developed FEM models of repair systems can be treated as a “ideal” reference repair systems. ACKNOWLEDGEMENTS The authors are thankful to Professor L.Czarnecki, Head of the Building Materials Engineering Department at the Warsaw University of Technology for his valuable discussion and remarks and Tomasz Piotrowski, MSc., for his assistance in the wavelet analysis with MATLAB. The research project was granted by Polish Scientific Research Committee - grant number KBN 4 T07E 027 27 (2004-2006). REFERENCES 1. Czarnecki L, Emmons P.H. 2002. Repair and protection of concrete structures (in Polish). (Polski Cement ed., Krakow, 2002), pp. 2. European Standard EN 1504-10:2003. Products and systems for the protection and repair of concrete structures - Definitions - Requirements - Quality control and evaluation of conformity - Part 10: Site application of products and systems and quality control of the works 3. Concrete Repair Manual. 2003. ACI International, Farmington Hills, MI, USA, 2003 4. Czamecki, L., Polymers in Concrete; Personal reflections on the edge of the new century, Concrete International, August, 8,2005, 1-7 5 . Sansalone M. J., Street W. B.: Impact-echo. Nondestructive evaluation of concrete and masonry. Ithaca NY, Bulbrier Press 1997 6. Sansalone M., Carino N.J. 1989. Detecting delamination in concrete slabs with and without overlays using the impact-echo method, ACI Materials Journal 86 7. Garbacz A.: Non-destructive assessment of repair efficiency with impact-echo and ultrasonic methods - an overview, ICCRRR 2005 International Conference on Concrete Repair, Rehabilitation and Retrofitting (Eds. H. Beushausen, F. Dehn and M.G. Alexander), Cape Town, South Africa, 1027-1031 (2005) 8. NOTES for DOCter Impact-Echo Course, GERMAN INSTRUMENTS NS, Denmark, January, 2000 9. Krautkramer, J., Krautkramer, H., Ultrasonic Testing of Materials, 4th Ed., SpringerVerlag, New York, 1990

3 14

Andrzej GARBACZ, Leslaw KWASNIEWSk7

10. Lin, J. M., Sansalone, M. "The Impact-Echo Response of Hollow Cylindrical Concrete Structures Surrounded by Soil or Rock, Part 1-Numerical Studies," ASTM Geotechnical Testing Journal, V. 17, No. 2, (1994), 207-219 1 1. LS-DYNA Theoretical Manual. Livermore Software Technology Corporation: Livermore, California, 1998 12. Goldsmith W., Impact: The Theory and Physical Behavior of Colliding Solids, Edward Arnold Press, Ltd., pp. 24-50 13. Abraham O., Impact-echo basse frequence. Pour la detection de vide dans les gaines de precontrainte, Controles-Essais-Measures,2002,56-6 1 14. Kurz J.H., Finck F., Grosse Ch.U., Reinhardt H-W. 2003. Automatic analysis of acoustic emission measure-ments on concrete. In Proc. BB 85-CD (ISBN 3-93 138149-8) International Symposium (NDT-CE 2003) Non-Destructive Testing in Civil Engineering, Berlin 15. Ovanesova A.V., Suarez L.E., Applications of wavelet transforms to damage detection in came structures, Engineering Structures, 26 (2004), 39-49

Proc. Int. Symp. “BrittleMatrix Composites 8” A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

STRENGTH DEVELOPMENT AND EPOXY RESIN-CEMENT INTERACTION IN HARDENER-FREE EPOXY-MODIFIED MORTARS

Yoshihiko OHAMA” ,Mitsukazu OCHI” , Shinsuke KUMAGA13),Masahiro OTA4’ ‘)Professor of Architecture, College of Engineering, Nihon University Koriyama-shi, Fukushima-ken, 963-8642 Japan e-mail: [email protected] 2, Professor of Applied Chemistry, Faculty of Engineering, Kansai University 3-3-35, Yamate-cho, Suita-shi, Osaka-fu, 564-0073 Japan, e-mail: [email protected] 3)Researcher,Urawa Research Laboratory, Konishi Co., Ltd. 5-3-35, Nishibori, Sakura-ku, Saitama-shi, Saitama-ken, 338-0832 Japan e-mail: [email protected] 4, Graduate Student, Graduate School of Engineering, Nihon University Koriyama-shi, Fukushima-ken, 963-8642 Japan e-mail: [email protected] ABSTRACT In conventional epoxy-modified mortars and concretes, proper hardeners have been used for the hardening of the epoxy resin. However, the authors found out that even without any hardener the epoxy resin can harden in the presence of cement hydrates in the cement mortars at ambient temperature. The purpose of the present study is to examine the strength development and epoxy resin-cement hydrate interaction in epoxy-modified mortars without any hardener. Hardener-free epoxy-modified mortars using an epoxy resin without any hardener and three types of cement are prepared with various polymer-cement ratios, and tested for flexural and compressive strengths, hardening degree of epoxy resin, infrared spectrum analysis and microstructures. As a result, the flexural strength of the hardener-free epoxy-modified mortars reaches a maximum at polymer-cement ratios of 5 to 15% regardless of the type of cement, and the hardening degree of epoxy resin in the mortars is decreased with an increase in the polymer-cement ratio. According to the infrared spectrum analysis results, the epoxy group does not exist in the mortars at the higher hardening degree of epoxy resin, and the hardening of the epoxy resin is confirmed. The microstructures of the hardener-free epoxy-modified mortars are formed by the combination of cement hydrates, hardened epoxy resin films and sand.

Keywords hardener-free epoxy-modified mortar, strength development, degree of hardening of epoxy resin, microstructure, infrared spectrum.

Yoshihiko O H M . Mitsukazu OCHZ, Shinsuke KUMAGM, Masahiro OTA

316

INTRODUCTION

Conventional epoxy-modified mortars and concretes have an inferior applicability due to the two-component mixing of the epoxy resin and hardener, the toxicity of hardeners such as polyamine and polyamide and the obstruction of cement hydration by the hardeners. The author’s group found out that even without any hardener the epoxy resin can harden in the presence of the alkalis or hydroxide ions produced by the hydration of cement in the epoxy-modifiedmortars as expressed by the following fomula(1) (2):

O-CH2-CH-CH2

I

OH

EDOXY resb

hydration

Hardened eDoxy resin

This means the development of the epoxy-hydraulic cement systems of new concept. However, the detailed behavior of the epoxy resin without any hardener in the epoxy-modified mortars has been made clear till now. The objective of this study is to clarify the mechanism of the strength development and epoxy resin-cement interaction in the epoxy modified mortars without any hardener. In this paper, hardener-free epoxy-modified mortars using an epoxy resin without any hardener and three types of cements are prepared with various polymer-cement ratios, and tested for strength development and epoxy resin-cement interaction. MATERIALS Cement Ordinary portland cement (OPC) and low-heat Portland cement (LHPC) as specified in JIS (Japanese Industrial Standard) R 5210 (Portland cement) and aluminous cement (AC) as specified in ex-JIS R 25 11 (Aluminous cement for refiactories) were used as cements. The physical properties and chemical compositions of the cements are given in Tables 1 to 3.

Table 1 Physical properties and chemical compositions of OPC. Blaine specific surface Setting time Compressive strength of mortar Density (h-min) (Mpa) area (dcm3) (cm2/g) Initial set Final set 3d 7d 28d 3.16 3270 2-22 3-20 28.4 42.7 59.4 Chemical compositions (%) Si02 A1203 Fez03 CaO MgO SO3 ig. loss Total 21.52 5.16 2.97 63.62 1.42 2.01 1.89 98.59

3 17

Strength development and epoxy resin-cement interaction in hardener-free epoxy-modified mortars

Table 2 Physical properties and chemical compositions of LHPC.

Final set Chemical f%). . . . ~ comDositions ~ .. ~ ~ . CaO MgO 0.90 64.00 ~

Si02 26.40

A1203 2.80

Fez03 2.60

~~~

~

~

~

3d

7d

28d

SO3 2.30

ig. loss 0.80

Total 99.80

~~~

Table 3 Physical properties and chemical compositions of AC. Blaine specific surface Setting time Compressive strength of mortar Density (MPa) area (h-min) (g/cm3> Initial set Final set Id 3.13 4740 5-50 6-35 36.8 Chemical compositions (%) SiOz A1203 Fez03 Ti02 CaO MgO ig. loss Total 0.49 73.09 1.20 0.03 25.16 0.13 0.21 100.31

Fine Aggregate Toyoura standard sand as specified in ex-JIS R 5201 (Physical testing methods for cement) was used as a fine aggregate. The properties of the fine aggregate are listed in Table 4.

Size

(mm) 0.300-0.106

Table 4 Properties of fine aggregate. Fineness Bulk densitv Density modulus (kg/l) (g/cm3, 2 0 ~ ) 1.52-1.36 1.52 2.63

Water absomtion (YO) 0.11

Polymeric admixture A diglycidyl ether of bisphenol A (epoxy resin or epoxy) was used as a polymeric admixture. The constitutional formula of the epoxy resin is expressed in Figure 1. The properties of the epoxy resin are given in Table 5.

Figure 1 Constitutional formula of diglycidyl ether of bisphenol A (epoxy resin). Note,* : Average degree of polymerization (n=O. 1-0.2).

318

Yoshihiko OHAMA, Mitsukazu OCHI, Shinsuke KUMAGAI, Masahiro OTA

-

..

Epoxide equivalent 190

-

Table 5 Properties of epoxy resin. Density Viscosity i v i ~ i e ~ u i aweight weignt r Molecular m ~ a s ,2 0 ' 99 (g/cm3, 2 0 ~ ) ((m~as, 380 1.17 12000 TESTING PROCEDURES

Preparation of specimens According to JIS A 1171 (Test methods for polymer-modified mortar), hardener-free epoxy-modified mortars were mixed with a mass ratio of cement to fine aggregate 1 : 3, polymer-cement ratios (P/C) of 0, 5, 10, 15 and 20%, and their flows were adjusted to be constant at 170*5 in the water-cement ratio range of 70.0 to 82.0%. Beam specimens 4 0 x 4 0 ~160mm were molded with the hardener-free epoxy-modified mortars, and subjected to a 2d-2O0C-90%(RH)-moistplus 5d-20°C-water plus 2 1d-2OoC-60%(RH)-drycuring. Flexural and compressive strength tests Beam specimens were tested for flexural and compressive strengths in accordance with JIS A 1171. SEM observation of microstructures The samples 5 ~ 5 ~ 7 r n which rn had been taken from the beam specimens were etched and unetched with 5% hydrochloric acid and 47% hydrofluoric acid, washed with distilled water, and then dried by the D-dry method. The microstructures of the etched and unetched samples after drying were observed by scanning electron microscopy(SEM). Determination of degree of hardening of epoxy resin The beam specimens were crushed and ground into powders, and the powders were passed through a sieve of 1.2mm for samples. The unhardened epoxy resins in the samples were extracted by tetrahydrofuran[(CH&O], and their mass was determined. The degree of hardening of the epoxy resin was calculated by the following equation: Degree of hardening of epoxy resin(%)=(Eui-Eue)/ Euix 100 where Eui :mass(g) of the unhardened epoxy resin contained in each sample, and Eue : mass(g) of the unhardened epoxy resin extracted from each sample Infrared spectroscopy for epoxy group The hardened epoxy resins were extracted by treating the same samples (as those for the degree of hardening of epoxy resin) with inorganic acids. The infrared spectra of the extracted hardened epoxy resins were characterized in the wavenumber range of 4000 to 450cm-' by infrared spectroscopy, and the existence (presence or absence) of epoxy groups was confirmed at a characteristic wavenumber of 916cm-'. TEST RESULTS AND DISCUSSION

Figure 2 shows the polymer-cement ratio vs. flexural and compressive strengths of hardener-free epoxy-modified mortars using different cements. Regardless the type of cement, the flexural strength of the hardener-free epoxy-modified mortars increases with increasing polymer-cement ratio, and reaches a maximum at polymer-cement ratios of 5 to 15%, and the maximum flexural strength is about 1.2 to 2.1 times that of unmodified mortars

3 19

Strength development and epoxy resin-cement interaction in hardener-free epoxy-modifed mortars

(with a polymer-cement ratio of 0%). Irrespective of the type of cement, the compressive strength of the hardener-free epoxy-modified mortars increases with increasing polymer-cement ratio, and reaches a maximum at polymer-cement ratios of 5 to 10%. However, the incorporation of the hardener-free epoxy resin does not cause a marked improvement in the compressive strength compared to the flexural strength. The compressive strength of the hardener-free epoxy-modified mortars using AC is much higher than that of the hardener-free epoxy-modified mortars using other cements at any polymer-cement ratio. In spite of the polymer-cement ratio, the flexural and compressive strengths of the hardener-free epoxy-modified mortars are (high) AC>OPC> LHPC(1ow) in the order of the cements used with a few exceptions. The use of the cements with the higher setting or hardening rate and faster early-strength development tends to cause a higher strength development as the polymer-modified mortars. The reasons for this are that the timing of the interaction between the cement hydration and the hardening of the hardener-free epoxy resin becomes good, and then the strong comatrixes composed of the cement hydrates and the hardened epoxy resin are formed.

- 10.Or 9.0 8. 7.

Type of Cement

6. 5. 4. 3.

AC

10.0 5.0 U

0

5 10152025303540

Polymer-Cement Ratio (YO)

OPC LHPC

I

l

l

l

l

l

,

,

5 10152025303540 Polymer-Cement Ratio (%)

0

Figure 2 Polymer-cement ratio vs. flexural and compressive strengths of hardener-fkee epoxy-modified mortars using different cements. Photo 1 shows the microstructures of unetched and etched hardener-free epoxy-modified mortars using different cements at a polymer-cement ratio of 10%. In the microstructures, irrespective of the type of cement, the hardener-free epoxy resin in the hardener-free epoxy-modified mortars with a polymer-cement ratio of 10% can harden, and tough epoxy resin films formed can be observed. The formation of the tough epoxy resin films is a sufficient evidence of the hardening of the hardener-free epoxy resin by the catalytic action of the alkalis or hydroxide ions[3]. From the microstructure observation, it is found that the epoxy resin films bind the cement hydrates and fine aggregate together to form a monolithic composite structure in which the epoxy resin film phase interpenetratesthroughout the cement hydrate and aggregate phases in the mortars. In particular, the network structures of the epoxy resin films improve the interfacial transition zone. It appears that such a composite structure is closely related to the strength development, waterproohess, durability, etc. of the hardener-free epoxy-modifiedmortars.

320

Yoshihiko OHAMA. Mitsukazu OCHI, Shinsuke KUMAGAI, Masahiro OTA

OPC

LHPC

AC

LHPC AC Etched Photo 1 Microstructures of unetched and etched hardener-free epoxy-modified mortars using different cements at a polymer-cement ratio of 10%. OPC

Figure 3 represents the polymer-cement ratio vs. degree of hardening of hardener-free epoxy resin in hardener-free epoxy-modified mortars using different cements. Irrespective of the type of cement, the degree of hardening of the epoxy resin in the hardener-free epoxy-modified mortars decreases with increasing polymer-cement ratio. The degree of hardening of the epoxy resin in the hardener-free epoxy-modified mortars using AC is lower than that of the mortars using OPC or LHPC. This is attributed to the differences in the chemical compositions or cement hydrates between OPC or LHPC and AC. Type of Cement

100.0 M

.9

80.0 60.0

! 0

40.0

n 0

10 20 30 40 Polymer-Cement Ratio (YO)

Figure 3 Polymer-cement ratio vs. degree of hardening of hardener-free epoxy resin in hardener-free epoxy-modified mortars using different cements.

Strength development and epoxy resin-cement interaction in hardener-free epoxy-modified mortars

321

Figure 4 represents the infrared spectrum of the unhardened hardener-free epoxy resin used as a polymeric admixture. Figure 5 shows the infrared spectra of the hardened epoxy resins extracted from the hardened hardener-free epoxy-modified mortars with a polymer-cement ratio of lo%, using different cements. The epoxy group seen in the spectaum of the unhardened hardener-free epoxy resin in Figure 4 disappears completely in the spectra of the hardened epoxy resins extracted from the hardened hardener-free epoxy-modified mortars

C&-CH0 -'

Figure 4

Wavenumber (cm-') Infrared spectrum of unhardened hardener-free epoxy resin.

1'I I

3 1 1,

I

W I

?+i /

---'

.-_-- 4

'.__ ^--I_-

4000

3000 2000 1500 1000 Wavenumber (cm-l)

450

4000

3000 2000 1500'

->d

--

1000- 450

Wavenumber (ern-') LHPC ~I

opc

L..

3 6 6 0 - ~ f l 0 0 ~ ~ 5 0 0 1000 450 Wavenumber (cm-I) AC Figure 5 Infrared spectra of hardened epoxyresins extracted from hardened hardener-free epoxy-modified mortars with polymer-cement ratio of lo%, using different cements. 4dOO

'

322

Yoshihiko O H M , Mitsukazu OCHI, Shinsuke XUMAGAI, Masahiro OTA

using different cements, OPC, LHPC and AC in Figure 5. Consequently, it is proved by this fact that the hardener-free epoxy resin in the hardener-free epoxy-modified mortars can harden with the hydroxide ions (OH) produced from cement hydration irrespective of the type of cement as mentioned above. CONCLUSIONS The conclusions obtained from the above test results are summarized as follows: (1) Regardless of the type of cement, the flexural strength of hardener-free epoxy-modified mortars reaches a maximum at polymer-cement ratios of 5 to 15%, and the maximum flexural strength is about 1.2 to 2.1 times that of unmodified mortars (with a polymer-cement ratio of 0%). However, even if the polymer-cement ratio increases, the compressive strength of the hardener-free epoxy-modified mortars is hardly improved irrespective of the type of cement. (2) In hardener-free epoxy-modified mortars, the epoxy resin films bind the cement hydrates and fine aggregate together to form a monolithic composite structure in which the epoxy resin film phase interpenetrates throughout the cement hydrate and fine aggregate phases. It appears that the composite structure is closely related to the strength development, waterproohess, durability, etc. of the hardener-free epoxy-modified mortars. (3) In spite of the type of cement, the degree of hardening of hardener-free epoxy resin in hardener-free epoxy-modified mortars decreases with increasing polymer-cement ratio. (4) It is proved by the complete disappearance of epoxy group in the extracted hardened epoxy resins that the hardener-free epoxy resin in hardener-free epoxy-modified mortars can harden with the hydroxide ions (OH-) produced from cement hydration irrespective of the type of cement. REFERENCES

1. Ohama, Y., Demura, K., Endo, T., Strength properties of epoxy-modified mortars without hardener. In : Proceedings of the 9th International Congress on the Chemistry of Cement, Vol.V, Performance and Durability of Concrete and Cement Systems, National Council for Cement and Building Materials, New Delhi 1992, pp5 12-516 2. Ohama, Y., Demura, K., Uchikawa, H., Interaction between epoxy resin and cement hydrates in epoxy-modified cementitious compositions (in Japanese). In : JCA Proceedings of Cement & Concrete, No.49, Japan Cement Association, Tokyo 1995, pp252-257 3. Lee, H., Neville, K., Handbook of Epoxy Resins. McGraw-Hill, New York 1967, PP(5-8)-(5-9)

Proc. Int. Symp. 'Brittle Matrix Composites 8" A.M. Brandt, rC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

NON-DESTRUCTIVE TESTING OF POLYMER MODIFIED CONCRETE

')

Jure GALIC '), Ivana BANJAD PECUR ') ')Engineering Design Bureau, Zagreb Prilaz B. Filipovica 21, 10000 Zagreb, Croatia, e-mail: [email protected] University of Zagreb, Faculty of Civil Engineering, Department of materials Kaciceva 2 1, 10000 Zagreb, Croatia, e-mail: [email protected] ABSTRACT

Addition of polymer to modify concrete (PMC) mixtures significantly changes its properties in fresh and hardened states. This is why the evaluation of results of non-destructive tests is not directly applicable to PMC. As PMCs are currently used in various applications, it is necessary to investigate the possibility of applying the existing non-destructive test methods developed for the evaluation of mechanical properties of ordinary concretes. These methods are quite simple and inexpensive, but data analysis and result interpretation require adequate experience. Direct determination of mechanical properties of concrete with non-destructive testing methods is one of the biggest tasks in modern civil engineering, and despite limitations significant improvements have been made in this area. In this paper the results of the study are presented in which twenty concrete mixtures with different cement amounts, w/c ranging from 0.35 to 0.65, different aggregate size and polymer amounts were tested. The relation between mechanical properties obtained by destructive and non-destructive testing methods of the laboratory cast PMC specimens is determined.

Keywords Polymer modified concrete, non-destructive testing, ultrasound velocity, Schmidt hammer rebound INTRODUCTION In the recent years the use of non-destructive methods for the diagnostics and defectoscopy of the state of construction has been on the increase. The advantage of non-destructive methods is that they are simple and do not damage the construction, except for the possibility of some minor surface damage. However, the interpretation of results is one of the most challenging tasks in modem building construction. The advantage of applying combined non-destructive methods is clear when the variation of a certain properties of concrete directly affects the results of testing by a non-destructive method, but not to the same degree. This can be seen on the example of the increase in the moisture content of concrete, where testing gives a lower value of Schmidt hammer rebound, which at the same time increases the ultrasound velocity and in this way reduces or increases the converted or derived concrete strength if only one method is used. [ 1-31 The known relations between quantities measured by non-destructive methods and, for example, strength are valid for ordinary concretes, however, it has been observed that such relations do not apply to polymer-added concretes [4-51. As polymer modified concretes are

324

Jure GALId, Ivana BANJAD PECUR

increasingly applied in practice, reliable models should be developed which enable effective determination of requested properties. Adding polymer into fresh concrete results in better workability of fresh concrete (cohesion, soaking angle to the surface, smoothability), lower value of slump for the same w/c ratio, reduced need for water for the sake of achieving necessary consistence, etc. With hardened concrete, it improves adhesion to surface, increases fluid impermeability, resistance to penetration of aggressive substances, freeze-thaw and de-icing salt resistance, as well as elongation and resistance to impact. Modulus of elasticity has been reduced and creep coefficient increased. [3] It can therefore be concluded that the addition of polymer changes the values of the results of testing PMC with non-destructive methods as there is a certain influence on the concrete modulus of elasticity and thus also on testing results. [6] THEORETICALBACKGROUND

Concrete as a composite material has markedly heterogeneous features. Variations in strength, elasticity module and all other features should be observed through the share of individual concrete components in the total volume. A prediction of concrete elasticity module can be made by means of a two-phase model consisting of cement stone and aggregate, or mortar matrix and coarse aggregate, respectively. For this it is necessary to know their modulus of elasticity and the amount of aggregate in the volume of concrete. Generally, by the application of non-destructive methods it is not possible to obtain the concrete strength, but the correct assessment of strength requires the knowledge of the relationship between the non-destructive testing results and compressive strength determined by destructive testing. Although many non-destructive methods for concrete testing have been developed, the most frequently used method is the combination of Schmidt hammer and ultrasound, particularly if it is necessary to optimize available means and testing technology. The basic purpose is to develop a simple, reliable and applicable diagram for compressive strength assessment. The combination of the above mentioned methods reduces errors which occur when concrete strength is assessed by use of one method only, which is in no way sufficient for determining a requested parameter. Y = axX+ b + E

where is the compressive strength of concrete is the result of non-destructive measurement are direction parameters E is error It is important to emphasize that direction parameter a is less influenced by model than parameter b so it is necessary to determine parameter b for every element of the construction that is tested. Due to that it is necessary to use destructive methods for testing small number of specimens (for instance drilled cores), and increase the number of non-destructive testing places, which will lead to increased efficiency. As the usage of rebound hammer or determination of ultrasound velocity is simple it is applicable, without greater increase in expenses.

X Y a, b

325

Non-deshuctive testing of polymer modified concrete

Testing results show that increased statistical sample depends on results dispersion of the applied non-destructive testing method, and that is possible to achieve the same error level if the number of results done by non-destructive testing method is progressively increased. On every tested sample before determination of the compressive strength, hammer rebound number and ultrasound velocity were determined. Regression analysis of the gained data was done with “Data analysis” program in “Microsoft Excell”. EXPERIMENTAL PROGRAM In the experimental part, concrete mixtures were produced which cover a wide spectrum of strengths and thus enable the calibration of sclerometer and ultrasound. The polymer quantity is also varied and pre-requisites are created for the development of a model based on a whole series of concrete mixture data. Selected concrete mix designs are shown in table 1. The quantities of the following components are varied: - cement: 320 kg/m3 and 400 kg/m3 - amount of polymer (styrene butadiene latex): 5 and 10 % of cement mass - wlc ratio: 0.35; 0.45; 0.55; 0.65 - maximum size of aggregate: 8 mm and 16 mm Concrete mixtures that were not produced are marked with “-“as it is shown in table 1. Also, for every concrete mix the following fresh concrete properties were determined: SLUMP test, density, and air content. Table1 Overview of the selected compositions of concrete mixtures wlc 0,35 0,45 0,55 0,65

c=320 kg/m3 P-10% P-5% P-10% P-5% P-5%* P-lO% P-5%* P-10%

c=400 kg/m3

P-5%

P-5% P-5%* P-5%*

c=320 kg/m3

P-lO% P-IO% P-lO% P-IO%

c=400 kgjm’

~

P-0% P-0% P-O%* P-O%*

them (due to the time needed for sifting, the mixtures are added a setting retarder 0,25 % mJ. From each concrete mixture following specimens were made, and testing methods were done: - cube 15~15x15cm - compressive strength, ultrasound velocity (for dynamic modulus of elasticity Ed), hammer rebound - cylinder h=30, 015 cm - modulus of elasticity E - prism 10~10x50cm - bending strength, compressive strength (on halves), ultrasound velocity (for dynamic modulus of elasticity E d ) Modulus of elasticity E was determined according to the prEN 13412. Ultrasound velocity and dynamic modulus of elasticity E d were determined according to the HRN EN 125042:2001.

RESULTS AND DISCUSION Mechanical properties Results of testing the mechanical characteristics (compressive strengthf,, bending strengthfb, modulus of elasticity E and dynamic modulus of elasticity Ed) of concrete mixtures and

326

Jure GALId, Ivana BANJAD PEeUR

WIC r

amount of polymer

0.35 0.45 0.55

b 0

0.55 0.65 0.45 0.65

I I

5%

fc

fb

(MW 40.43 34.74 27.21 23.36 36.61 31.61

6.21 5.50 4.41

E

Ed

(GW

(GPa)

30.41 31.03 29.50 28.50

44.54 39.54 36.85

4.90 28.48 32.98 29.81

I I

4.29 3.19

MORTAR MATRIX

27.50 27.60

1

43.54 45.05

I

Non-destructive testing of polymer modified concrete

327

Mixtures with maximum aggregate size of 8 mm have higher compressive strength than a mixture of the same composition but with maximum grain of 16 mm. As the smaller grain size enables better “packing” of the structure and the likelihood of the occurrence of defective grains is lower, such an increase in compressive strength is to be expected. It can be seen that with the increase of the quantity of polymer the bending strength of concrete is increased, too. With the reduction of the maximum aggregate grain the flexural strength is reduced, as larger grains bridge over the cracks (as a result of bending) better. With the smaller maximum aggregate size the tensile flexural strength is smaller, too. We may conclude that the modulus of elasticity of concrete decreases with the increase in the quantity of polymer. The modulus of elasticity of mortar matrix is lower than the modulus of elasticity of the concrete of the same composition, which can be explained by the influence of aggregate on the increaseldecrease of modulus of elasticity. As the aggregate has a higher modulus of elasticity than cement paste, it is obvious that a smaller amount of aggregate also results in a lower elasticity module. Dynamic modulus of elasticity is calculated from experimental results of ultrasound velocity measurement of different concrete. Increasing of ultrasound velocity give higher results of dynamic modulus of elasticity of concrete. With the equal wlc ratio the dynamic modulus of elasticity decreases with the increase of the quantity of polymer. Dynamic modulus of elasticity of concretes with 400 kg/m3 cement is higher than in concretes with 320 kg/m3 cement. With the decrease in the maximum aggregate size for the same composition of concrete, the dynamic modulus of elasticity decreases. We may conclude that for the same value of compressive strength the dynamic modulus of elasticity decreases with the increase in the quantity of polymer. In regard to the results presented (see the compressive strength), it can be seen that the dynamic modulus of elasticity is lower in mortar matrix samples although their compressive strength is higher than in concrete samples. The total volume amount of aggregate in mortar matrix samples decreases, which results in the lower dynamic modulus of elasticity. Smaller aggregate size enables better grain packing, which increases compressive strength. As with the static elasticity module, with the increase in the quantity of added polymer the dynamic elasticity module decreases. The same influence is shown by the decrease in the maximum aggregate size, as larger aggregate size increase the rigidity of matrix.

Schmidt hammer rebound Test results of Schmidt hammer rebound and compressive strength are shown on figure 1 an figure 2 for concrete and mortar matrix made with 320kg and 400 kg of cement and 5 % of polymer. With the increase in the quantity of cement the Schmidt hammer rebound decreases and compressive strength increase, because the total volume of aggregate in concrete with more cement are smaller. The reduction of the maximum grain size from 16 mm to 8 mm does not affect the value of the Schmidt hammer rebound to the same degree as it affects the increase in compressive strength. This results in an upward translation of the curves for the determination of compressive strength. As the aggregates with the maximum size of 8 mm are the ones that are most frequently used for the production of high-strength concretes, this confirms the results of the experiment. It is highly probable that the use of the grain with maximum size of 4 mm or less would bring about a decrease in the Schmidt hammer rebound.

328

Jure GALk Ivana BANJAD PECUR

Mortarmatrix Concrete

-CONCRETE

10,o

+ 25

20

I

I

I

30

35

40

Schmidt hammer rebound, N

Figure 1 Dependence of compressive strength and Schmidt hammer rebound of concrete and mortar matrix for concretes with 320 kg of cement and 5 % of polymer

45.0

0

3 40.0

-

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329

Non-destructive testing of polymer modified concrete

Regression function Compressive strength of concrete calculated according to regression models for the measured Schmidt hammer rebound and ultrasound velocity shows in the best way what the influence of polymer addition on output results is. Calculating regression function of compressive strength for different concrete mix design (amount of cement and polymer) based on testing results of compressive strength by non-destructive methods (Schmidt hammer rebound and ultrasound velocity) are present in table 3. Table 3 Calculating results of regression function

Amount cement I polymer (kg I Yo) 32010 32015 320110 40010 40015 400110

Regression function - compressive strength fc = -41,77+0,015XU+0,357XS fc = - 101,1+0,024XU+l XS fc = -56,91+0,014XU+0,99XS fc = -142,09+0,034XU+l,lXS fc = -58,739+0,02xU+O,612XS fc = -47,014+0,016~U+0,519xS

Correlation coefficient 0,899 0,808 0,861 0,825 0,899 0,529

where fc U S

is the compressive strength of concrete is the result of ultrasound velocity is the results of Schmidt hammer rebound CONCLUSION

The ultrasound velocity in polymer modified concretes is lower in comparison to the velocity in ordinary concretes. With each increment in the quantity of polymer in concrete the velocity of ultrasound is additionally reduced. Likewise, with polymer addition the dynamic modulus of elasticity decreases with the quantity of added polymer. As the compressive strength and modulus of elasticity decrease with polymer addition, it is obvious that all the above said affects output results of non-destructive testing. The reduction of maximum aggregate size results in the increase of compressive strength as it enables better “packing” of the structure, but decreases the modulus of elasticity as more cement paste with a lower modulus of elasticity is needed in order to wrap in all aggregate grains. Likewise, a reduction in the maximum aggregate size reduces the speed of ultrasound and the dynamic modulus of elasticity. The decrease in the ultrasound velocity may be explained with the increase in the amount of cement paste in which the speed of ultrasound is lower than in the aggregate. A comparison of concretes of various compositions (with and without polymer) clearly shows the influence of the grain size on the above mentioned values. Beside the quantity of added polymer, the ultrasound velocity and the Schmidt hammer rebound are also significantly influenced by the quantity of cement in concrete. Polymer film that is formed around cement stone has an irrelevant thickness in relation to other measurable thicknesses of the layer, but has a big influence on the results of non-destructive tests. As the polymer film that is formed around cement stone alters its features in macroscopic sense, it obviously also affects the measured values of Schmidt hammer rebound. It is generally true that the addition of polymer decreases the value of Schmidt hammer rebound, but the ratio between the percentage of the sclerometer index decrease and the quantity of added polymer

330

Jure GALIC, Ivana BANJAD PECUR

is not constant. It is obvious that even small quantities of polymer significantly alter the properties of concrete. Likewise, it can be concluded kom the above said that the test results of measurements do not decrease proportionallyto the increase in the quantity of polymer. REFERENCES 1. Ohama Y.: Handbook of polymer-modified concrete and mortars - Properties and process technology; Noyes publications 1995 2. Hisham Y.Qasrawi: Concrete strength by combined nondestructive methods Simply and reliably predicted, Cement and Concrete Research 30 (2000) 3.ACI Recommendation 548.3R-95Polymer modified concrete 4.Scientific project “Bridge reconstruction with high performance concrete overlay” for Croatian road, project manager Prof. V. Ukrainczyk, University of Zagreb, Faculty of Civil Engineering, Department of materials 2003,Zagreb 5. Execution design of reconstruction of bridge over river Mreznica in Ostarije, University of Zagreb, Faculty of Civil Engineering, Department of materials, 2001,Zagreb 6. Galic J.: Determination of polymer modified concrete properties by non-destructive methods, Master thesis, University of Zagreb, Faculty of Civil Engineering, 2005,Zagreb

Proc. Int. Symp. "Brittle Matrix Composites 8" A.M. Brandt, rC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

QUANTIFYING VARIABILITY IN ASSESSING THE RISK OF EARLY-AGE CRACKING IN RESTRAINED CONCRETE ELEMENTS Aleksandra RADLINSKA, Jason WEISS Purdue University 550 Stadium Mall Drive, West Lafayette, IN 47907-1284, USA e-mail: [email protected], [email protected]

ABSTRACT Civil Engineers frequently perform calculations based on material properties that represent either the 'worst case' or the 'average' properties. While these approaches are extremely helpful in simplifying computations, they can fail to quantify how variability from different sources combines to influence the overall performance of the system. This paper describes two approaches to quantify the variation in residual stress development and the risk of cracking in concrete. First, the paper illustrates the use of a Monte Carlo simulation procedure to quantify the residual stresses and age of cracking for a restrained concrete element due to variability in material properties, environmental conditions, and construction procedures. Second, the paper illustrates how statistical analysis can be directly applied to the problem using an approach that is similar to that used in the design of steel structures (a Load and Resistance Factor Design approach). This paper begins with a short review of a model that computes residual stress development in a restrained concrete element. The paper then describes a Monte Carlo modeling approach that was added to the model to account for variability in materials, construction practices, and environmental conditions. The results show that even low variability in material inputs results in substantial scatter in the predicted time of cracking. An alternative approach to predict the time of cracking is then presented that is similar to the one used for the design of steel structures (Load and Resistance Factor Design - LRFD). The LRFD approach enables the computation of a reliability index p which can quantify a material's potential for cracking. The results indicate that, as one may expect, as the magnitude of shrinkage coefficient increases, the potential for failure increases. Results also show that when the free shrinkage is reduced to a sufficient level, cracking will no longer be observed. The LRFD method enables the level of shrinkage to be quantified and this can then be used in the development of mixture proportioning procedures for mixtures containing shrinkage reducing admixtures (SRA).

Keywords Early-age cracking, Monte Carlo, LRFD, variability, probability REVIEW OF A MODEL TO PREDICT EARLY-AGE CRACKING Concrete is susceptible to cracking, especially at early ages. The likelihood for cracks to form depends on several time-dependent factors including material stiffness, shrinkage rate, shrinkage magnitude, stress relaxation, and material toughness. To account for the interaction between these properties, a model has been developed to estimate the potential for restrained shrinkage cracking in concrete elements [ 1-41. The following section describes this modeling approach in greater detail.

332

Aleksandra RADLZNSU, Jason WEZSS

Modeling Procedures Previous work [l] has shown that equation 1 may be used to estimate the stress development in a restrained concrete element:

where, &pennit(t) is the total strain that is permitted to develop in the actual restrained concrete (i.e., for complete restraint E(t) = 0 which will be assumed throughout this paper), E d { ) is the time dependent elastic moduli, EC is a reference elastic modulus (i.e., a 28 day value), Nt,4 is the creep coefficient [5], and a({) is the differential shrinkage with respect to time (t) (i.e.,

,

Erot-Shr

(t) =

ja(odc )' '

0

While the specific details on the forms of the material inputs (i.e., elastic modulus, shrinkage, creep) can be found in [3], the total shrinkage (which will be further discussed, later in the paper) is assumed to be the sum of the autogenous and d y n g shrinkage components, as illustrated in Fig. 1. In order to relate the ultimate shrinkage of paste to the shrinkage of concrete, equation 2 has been applied [6,7]: 'Shr-

= f l N (l-

(2)

F'

where f l is~ the shrinkage coefficient of the paste, VFis the volume fraction of the aggregate, and n is a coefficient that describes the stiffness of the aggregate and paste. The value of coefficient n is typically between 1.2 and 1.7 for normal strength and normal weight concretes (in this paper n was assumed to be 1.43 [S]). 600 )

0

10

20

30 40 50 60 70 80 Age of the Specimen (Days)

90

100

Fig. 1 : Components of the total shrinkage for concrete subjected to drylng initiated after 7 days (detailed description of equations and factors used can be found in [3]). The likelihood of early-age cracking is predicted by comparing the residual stresses that develop in the restrained concrete with a materials' resistance to cracking. While equation 1 can be used to compute the residual tensile stress, the tensile strength can be approximated with a time dependant h c t i o n as described in [3]. To predict the likelihood of cracking, a parameter called cracking potential, &R(t), has been defined as the ratio of residual stress (oft))and tensile strength (f',(t)), as shown in equation 3:

Quantzfiing variability in assessing the risk of early-age cracking in restrained concrete elements

333

Based on a deterministic strength of materials approach, cracking can be expected to occur when the cracking potential (equation 3) is equal to one. However, experimental procedures have shown that cracking can be observed for cracking potential values even as low as 0.7 [9111 and this discrepancy is thought to be related to the statistical variation in material properties [3,4, 101. An example of the deterministic computations for the time-dependant development of strength and stress in concrete is presented in Fig. 2. In the analyzed case, the aforementioned numerical model [3] was used to evaluate a concrete with water to cement ratio wlc = 0.5, subjected to drymg after 7 days, with a long-term elastic modulus of 27.5 GPa, a long term splitting tensile strength of 5 MPa, and a compressive strength of 33 MPa, and the volume of aggregate: Vagg= 68%. The rate of development for the elastic modulus was 2.0lday and for tensile strength it was 0.31day. A shrinkage coefficient CO,) of 3500 pf was assumed which corresponds to the long term shrinkage value of 600 pf (for the given Vaggand n). The rate constants for the autogenous and drymg shrinkage were 15lday and 3O/day, respectively. Other details about parameters can be found in [3]. '

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Fig. 2: Time dependent stress and strength development for concrete with shrinkage coefficient PN= 3500 pf (& = 600 p&)and subjected to drylng after 7 days. The use of a deterministic approach for predicting the time of cracking does not include the variability present in material properties and production procedures which can alter the time of cracking significantly. The following section describes how material property variability influences the results.

THE INFLUENCE OF MATERIAL VARIABILITY IN CRACKING PREDICTION Monte Carlo Method To include the effects of variability in the model, a Monte Carlo simulation method has been applied to the aforementioned model [3]. The material properties were randomly selected from predefined normal distributions and used to calculate the time-dependant strength and stress development. A series of simulations were performed varying the values of shrinkage, elastic modulus, and splitting tensile strength, assuming that coefficient of variation for every parameter equals to 5%. Using the same material properties as those used in the development

334

Aleksandra RADLIfiSKA. Jason WEISS

of Fig. 2, Fig. 3 was obtained which presents how variability can change the cracking potential as a function of time. Fig. 3a and Fig. 3b present typical output from the Monte Carlo simulation in terms of aprobability and cumulative density functions obtained for the predicted time of cracking. Fig. 3a presents probability density h c t i o n and a histogram of the results of the simulation. Each bar of the histogram has a width of 4 days and represents the number of specimens out of the 10,000 simulations that cracked over that period. It can be noticed, that the distribution h c t i o n contains the most frequent ages of cracking before the time (26.5 days) that was predicted by the deterministic approach (Fig. 2). Integrating the probability density function over time yields the cumulative distribution h c t i o n (CDF), shown in the Fig. 3b. It can be seen that after a certain period of time, cumulative function begins to plateau and reaches the final value of 90.4 %, meaning that out of the whole population, 90.4 % of all the samples will crack.

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b)

Fig. 3: a) probability density function and frequency distribution of predicted ages of cracking b) a cumulative distribution of the time of cracking. Load and Resistance Factor Design (LRFD) The previous section has described how Monte Carlo simulations can be used to provide information about probability of cracking in concrete. It should be noted, however, that alimitation of the Monte Carlo approach is that computations may be time consuming to perform, especially when more complicated geometries are evaluated. To overcome this limitation, an alternative approach has been proposed based on the Load and Resistance Factor Design (LRFD). This approach has been used in steel design [12, 161 and allows the reliability of a system to be estimated by treating the load (Q) and resistance (R) as random variables with assigned distribution, as presented in the Fig.4 [ 13-15]. The area of failure, i.e., the area when the load (Q) exceeds the resistance (R) may be represented by (however is not exactly equal to [12]) the amount of overlap of two probability density functions: fQ(x) and fR(x).

Quantifiing variability in assessing the risk of early-age cracking in restrained concrete elements

335

X

Fig. 4: Load (Q) and resistance (R) as the normally distributed variables. Instead of integrating the area under the two overlapping curves, it is more convenient to deal with a single curve that consist of the logarithm of R divided by Q (or the inverse of the cracking potential), as shown in Fig. 5. In this case, the area of failure is the region under the curve on the negative side of x-axis, marked with the cross-hatched region. Throughout this paper, the expression ‘failure’ is understood as the synonymous of cracking. The probability of cracking can be expressed as either the probability of resistance being lower or equal to load, as the difference between resistance and load being less than or equal to zero, or as the probability of the resistance to load ratio being less than or equal to one: p f = P(R S Q )

= P(R-

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Fig.5: The reliability index p. The LRFD approach enables the total variability of the system to be estimated through the calculation of the reliability index w(see Fig. 5):

where ,U~,,(W@ is the mean value of the natural logarithm of R divided by Q and SD,,,(WQ corresponds to the standard deviation of the natural logarithm of the R divided by Q. For the example presented earlier (Fig. 2), the maximum stress reaches a long term value of 5.50 MPa is equal to -0.11. and maximum tensile strength is 4.95 m a , therefore N,,(R/P) Rather than describing the variability in each material property, a simplified approach can be used to approximate the total coefficient of variation (COV) using equation 6 [ 161:

cov = Jcov: +cov;+cov,2,r

(6)

336

AleksandraRpDLIfiSKA, Jason WEISS

where the single coefficients of variation under the square root are those of the shrinkage (E), elastic modulus and splitting tensile strength (f',). It can be assumed that the S D I , , is~ ~ equal to the total coefficient of variation [16]. In the previous example (Fig. 3), all three parameters had a coefficient of variation equal to 5%, resulting in a total coefficient of variation that is equal to 0.0866. In that case, the reliability index can be calculated as:

(a,

p=- -0.1 1 = -1.22 0.0866

The probability of failure can be written as:

where Z is the standard normal variable and Cg denotes the cumulative density function, i.e., probability that a variable has a value less than or equal to In(WQ) [12]. The use of this approach to calculate the reliability index allows a rapid estimation of the cracking probability using commonly available standard normal distribution tables. For the reliability index calculated in equation 5b, the probability of cracking can be found as: pf = @(-P, = Cg( 1.22) = 0.8887

(7b) The probability of cracking in concrete depends on the coefficient of variation in the system. Fig. 6a presents how the probability of cracking changes depending on the magnitude of the shrinkage coefficient while the corresponding figure (Fig. 6b) shows the reliability index values for the same example. The horizontal dashed lines in the Fig. 6b denotes the maximum and minimum values usually reported in the mathematical tables for standard normal distribution, where for reliability index equal to 3.4, probability of failure is minimal:

pf = Cg(-P, = Cg(-3.4) = 0.0003

(8)

while for reliability index equal to -3.4, probability of failure is very high: pf = @(-P, = Cg(3.4) = 0.9998

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337

Quantifiing variability in assessing the risk of early-age cracking in restrained concrete elements

cracking risk at each age, but requires time-consuming computations. The LRFD approach permits the designer to assess the reliability of the system and does not require complicated simulation procedures. The differences between the results obtained using the three mentioned methods have been presented in Table 1. For a mixture with a shrinkage coefficient of 3000 p~ and COV = 0.0866, the deterministic analysis would indicate that no cracking is expected to occur. However, if the Monte Carlo simulations were performed for the same concrete, the output would indicate that 5% of samples will crack by the age of 18.6 days and by an age of 300 days, 31.5% of samples will be cracked. For a shrinkage coefficient of 3500 p~ and a COV of 0.0866, the deterministic analysis would result in the predicted time of cracking after 26.5 days (Fig. 2). If the Monte Carlo simulations were performed for the same concrete, the output would indicate that 5% of samples will crack by 12.2 days, 50% will crack by 27.1 days but some of them will never crack. If the LRFD analysis is conducted, the reliability index (p) would indicate that probability of cracking for the concrete with A=3500 p~ is 88.8%. Table 1: Modeling inputs Ultimate Shrinkage coefficient P N [PI

E . ~rwi ,.,

Time of ~i~~ of cracking (MI-) [days] cracking fiom deterministic m a l ~ s i s [ h ~ s l 5% 50% 95%

Probability of cracking at 300 days (MC)

Reliability index

P

Overall probability of cracking (LRW

31.5 % 90.4% 99.8 %

0.56 -1.22 -2.76

28.6 % 88.8 % 99.7%

I.,

3000 3500 4000

515 600 400

nocracking 26.55 16.55

18.6 12.2 9.5

500 27.1 17.1

500 500 120

The approach presented here enables calculating the risk of cracking and as such, it can be used to determine the shrinkage that corresponds to a specified level of cracking probability. If the shrinkage of the system can be reduced, the probability of cracking can be minimized to the specified level. The following sections will describe steps required in the design procedure.

REDUCING THE PROBABILITY OF CRACKING USING A SHRINKAGE REDUCING ADMIXTURE The use of shrinkage reducing admixtures (SRA’s) decreases the magnitude of shrinkage in a concrete mixture [ 17-23]. This reduction in shrinkage is due to a reduction in the surface tension of the pore solution [24, 251. Fig. 7 shows result from the experiments where the shrinkage coefficient was measured for a series of pastes containing different concentrations of SRA [26]. It can be seen that when the shrinkage coefficient is plotted versus the surface tension of water-SRA solution, a linear trend is observed. A relationship between the surface tension of water-SRA solution and the percentage of SRA was previously measured and the results are shown in Fig. 8. The surface tension decreases with addition of SRA (up to 15% SRA) and can be described using equation 10 [24]: if C s u < 15

338

Aleksandra R A D L I k X 4 , Jason WEZSS

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0

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10 15 20 25 90 SRA Concentration (%)

95

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Fig. 8: Surface tension measurements for deionized water-SRA mixtures for Tetraguard, Eclipse Plus and Eclipse Floor [24,26]. Since the addition of SRA decreases the shrinkage of concrete, one may want to compute the value of long term shrinkage that is required to reduce the probability of cracking. The following section will describe how the concentration of SRA can be determined to correspond with a specific level of performance. AN EXAMPLE OF USING THE LRFD PROCDURE TO DETERMINE THE

REQUIRED CONCENTRATION OF SRA IN A CONCRETE MIXTURE

Consider an example of a plain concrete mixture with a wlc = 0.5, volume of aggregate 68%, a total coefficient of variation COV = 0.0086 and a plain paste with a shrinkage coefficient of

Quantrfving variability in assessing the risk of early-age cracking in restrained concrete elements

339

3900 p (Fig. 7). It is assumed that elastic modulus and splitting tensile strength are 27.5 GPa and 5 MPa, respectively, while the rate parameters are the same as those used in the earlier example. The reliability index for this example can be computed using equation 5 (p = -1.29) and probability of cracking is equal to 99%. This denotes that the concrete will almost certainly crack and therefore the design of the concrete mixture requires some modification. In order to reduce the probability of cracking in concrete, a shrinkage reducing admixture could be used, the volume of aggregate in the mixture could be increased, or different cementitious materials could be incorporated to reduce the shrinkage coefficient. Here, a procedure is presented to estimate the amount of SRA that would be needed to decrease the potential for cracking to an acceptable level. If variability in the system can be estimated, the LRFD approach allows for the development of a graph relating shrinkage coefficient with probability of cracking, as presented in the Fig.9. If the designer was willing to accept a 10% risk of cracking, the shrinkage coefficient would have to be reduced from 3900 to 2820 p,which is equivalent to the reduction of long term shrinkage value ( E S ~ )from 670 to 480 p. As shown in Fig.9, lowering shrinkage to that value corresponds to 3% SRA concentration (for Tetraguard). The procedure presented in Fig.9 can be described in the following steps: Step 1: determine the shrinkage and variability in the system together with corresponding probability of cracking Step 2: determine the accepted probability of cracking (i.e., 5%, 10%) Step 3: determine the amount of SRA ('YO)needed to lower the shrinkage coefficient and the probability of cracking to the desired level

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Fig. 1: Steps in the shrinkage based design approach. The present example shows that in order to reduce the probability of cracking from 99% to lo%, a concentration of 3% SRA should be used in the mixture. To verify this approach using a laboratory experiment, a large number of samples would have to be prepared and tested. Alternatively, the Monte Carlo method can be used to perform simulations for a concrete with shrinkage value of 670 and 480 p (Fig. 10). It can be seen that cumulative function for Esh= 670 p approaches value of one at the age of 50 days, which means that by the time of 50 days almost all the samples (i.e., 990/) will crack. The response

340

Aleksandra RADLINSK4, Jason WEISS

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SUMMARY The work presented in this paper leads to the following concluding remarks: The ratio of time dependant stress and strength (i.e., the cracking potential) can be used in statistical analysis (i.e., the Monte Carlo method or Load and Resistance Factor Design approach) to incorporate variability into computations; While the Monte Carlo method allows an accurate estimation of the time and probability of cracking in concrete, it is a time consuming process. The Load and Resistance Factor Design approach permits more rapid assessment of cracking; The addition of shrinkage reducing admixture decreases the shrinkage and cracking potential in concrete. A method is proposed to estimate the amount of SRA that is needed to reduce the cracking risk of concrete to the accepted level.

ACKNOWLEDGEMENTS

The authors gratefully acknowledge support for this research which has come from Master Builders and the National Science Foundation (NSF) through Grant No. 0134272: a CAREER AWARD. Any opinions, frndings and conclusions or recommendations expressed in this material are those of the authors and do not necessarily reflect the views of the National Science Foundation (NSF).

Quantzfiing variabili@ in assessing the risk of early-age cracking in restrained concrete elements

341

REFERENCES 1. Weiss, J., Yang, W., and Shah, S.P., Shrinkage cracking of restrained concrete slabs, Journal of Engineering Mechanics - ASCE, 124 (7), 1998, pp 765-774 2. Weiss, J., Yang, W., Shah, S.P., Influence of specimen size/geometry on shrinkage cracking of rings, Journal of Engineering Mechanics - ASCE 126 (2000), pp 93-101 3. Radlinska A., Pease B., Weiss J., A preliminary numerical investigation on the influence of material variability in the early-age cracking behavior of restrained concrete, Materials and Structures - RILEM (accepted for publication, in press) 4. Radlinska A. and Weiss J., Determining early-age cracking potential in restrained concrete elements using a load and resistance factor design (LRFD) approach, International Symposium ‘Advances in concrete through science and engineering’, Quebec City, 2006 5. Muller, H.S., New Prediction Models for Creep and Shrinkage of Concrete, Creep and Shrinkage of Concrete: Effect of Materials and Environment eds. M.A. Daye and C.C.Fu, American Concrete Institute, Detroit, Michigan 1992 6. Pickett, G., Effect of Aggregate on Shrinkage of Concrete and Hypothesis Concerning Shrinkage, Journal of ACI, Vol. 52, 1956, pp. 581-590 7. L’Hermite, R. G., Volume Changes of Concrete, Fourth International Symposium on the Chemistry of Cement, Washington, D.C., 1960, pp. 659-694 8. Moon, J.-H., Rajabipour, F., Pease, B., Weiss, J., Autogenous Shrinkage, Residual Stress, and Cracking in Cementitious Composites: The Influence of Internal and External Restraint, ‘Fourth International Seminar on Self-desiccation and Its Importance in Concrete Technology’, Maryland, USA, 2005 9. Hossain A., Weiss J., Assessing residual stress development and stress relaxation in restrained concrete ring specimens, Cement and Concrete Composites, 26,2004, pp 53 1-540 l0.van Breugel, K., Lokhorst, S.J., Stress-based crack criterion as a basis for prevention of through-cracks in concrete structures at early ages’ RILEM Proceedings 23, International RILEM Conference on Early Age Cracking in Cementitious Systems - EAC’OI, Edited by K. Kovler and A. Bentur, 2001 1l.Schiel31, P., Rucker P., New results on early-age cracking risk of special concretes, ACBM Concrete Cracking Workshop, Evanston, IL, 2005 12.Melchers R.E., Structural reliability. Analysis and prediction, Baffins Lane, Chichester 1987 13.Stewart, M.G., Melchers R.E. Probabilistic risk assessment of engineering system, Chapman & Hall, London 1997 14.Stewart, M.G., Reliability-based assessment of aging bridges using risk ranking and life cycle cost decision analyses, Reliability Engineering and System Safety, 74, 2001, pp 236273 15.Ellingwood, B.R., Mori, Y. Probabilistic methods for condition assessment and life prediction of concrete structures in nuclear power plants, Nuclear Engineering Design, 142, 1993, pp 155-166 16.Salmon G.C., and Johnson, J.E., Steel structures design and behavior, 3rd Edn, Harper and Row, New York 1996 17.Shah, S. P., Weiss, W.J., and Yang, W., Shrinkage Cracking-Can It Be Prevented?, Concrete International, Vol. 20, No. 4, 1998, pp 51-55 18.Weiss, W. J., and Shah, S. P., Restrained Shrinkage Cracking: The Role of Shrinkage Reducing Admixtures and Specimen Geometry, Materials and Structures, Vol. 35, no. 246, 2002 19.Weiss, W. J., Schiebl, A., Yang, W., and Shah, S. P, Shrinkage Cracking Potential, Permeability, and Strength for HPC: Influence of W/C, Silica Fume, Latex, and Shrinkage

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Aleksandra RADLIfiSKA, Jason WEISS

Reducing Admixtures, International Symposium on High Performance and Reactive Powder Concrete, Sherbrooke, Canada 1998, Vol. 1,349-364 20.Weiss, W. J., Borischevsky, B. B., and Shah, S. P., The Influence of a Shrinkage Reducing Admixture on the Early-Age Behavior of High Performance Concrete, Fifth International Symposium on the Utilization of High Strength/High Performance Concrete Vol. 2, Sandefjord, Norway 1999, Vol. 2, pp 1418-1428 21 .Bentz D.P., Geiker M.R. Hansen K.K., Shrinkage-reducing admixtures and early-age desiccation in cement pastes and mortars, Cement and Concrete Research 31(7), 2001, pp 1075-1085 22.Bentz D.P., Jensen OM., Mitigation strategies for autogenous shrinkage cracking, Cement and Concrete Composites 26,2004, pp 677-685 23 .Bentz D.P., Curing with Shrinkage-Reducing Admixtures, Concrete International 27( lo), 2005, pp 55-60 24.Pease, B. J., The Role of Shrinkage Reducing Admixtures on Shrinkage, Stress Development, and Cracking, Master's Thesis, Purdue University, West Lafayette, IN, 2005 25.Ai, H., and Young, J.F., Mechanism of shrinkage reduction using a chemical admixture, Proceedings of the 10" International Congress on the Chemistry of Cement, Vo1.3, ed. Justnes, H., Gothenburg, Sweden 1997 26.Pease, B. J., Shah, H. R., Weiss, W. J., Shrinkage behavior and residual stress development in mortar containing shrinkage reducing admixture (SRA's), Shrinkage and creep of concrete, ACI SP-227, Special Publication on Concrete Admixtures, Farmington Hills, Michigan 2005 27.Weast, R.C., Astle, M.J., and Beyer, W.H., CRC Handbook of Chemistry and Physics 64th ed., CRC Press, Inc., Boca Raton, FL, 1983, pp. F-33

Proc. Int. Symp. "Brittle Matrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-2.5, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

STRUCTURAL PERFORMANCE AND CRACK CONTROL OF FIBRE CONCRETE BEAMS WITH CONVENTIONAL REINFORCEMENT Alena KOHOUTKOVA, Iva BROUKALOVA Department of Concrete Structures and Bridges Faculty of Civil Engineering, Czech Technical University in Prague T h h r o v a 7, 166 29 Praha 6, Czech Republic e-mail: [email protected], [email protected]

ABSTRACT Experimental study includes investigation of performance of concrete beams reinforced with conventional steel bars and with synthetic fibres. Several sets of beams have been tested to determine the influence of polypropylene fibre reinforcement on the mechanical behaviour of conventionally reinforced concrete beams in bending. The present study indicates that polypropylene fibre reinforcement can reduce crack width and deflections and improve ductility of beams and that the combination of polypropylene fibres, longitudinal steel bars (without stirrups) may meet strength and ductility requirements. The basic material characteristics were obtained from laboratory tests on prisms in four or three point bending and by an inverse analysis procedure the load-deflection relation were transformed into constitutive relations used in simulation. The test was simulated by means of a non-linear program system using various constitutive relations, which enable fibre concrete modelling.

Keywords Fibre concrete, inverse analysis, simulation

INTRODUCTION Significance of fibres in fibre concrete is not only in improvement of the performance of FRC in comparison to the plain concrete but also in application in reinforced concrete structures, high strength and ultra high strength concretes. For application of FRC in structural members it is necessary to ensure not only appropriate technology for production but also adequate methodology for design. Material properties of fibre concretes have been investigated for since 70's and benefits of fibre concrete are beyond all doubt. Lately a utilisation of fibre concrete in structural elements has been focused. First structures have been executed; e.g. a flat slab with only fibre reinforcement [l], a frame which formwork is manufactured from SIMCON (Slurry Infiltrated Mat Concrete) [ 2 ] etc. These applications are rather rare and only for exceptional and special structures though it is assumed that fibres in a structure improve structural behaviour under service load, fatigue resistance, enhance service life and provide advantageous failure mechanism due to higher ductility. It is also assumed that thanks to fibres the amount of shear reinforcement may be reduced.

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Alena KOHOUTKOVA, Iva BROUKALOVA

For extension of fibre concrete utilisation more investigations of structural behaviour have to be provided. Research presented in this paper intends to be a contribution to the structural fibre concrete development.

FIBRE CONCRETE STRUCTURALANALYSIS

Generally, the FRC characteristicsfor tension largely influence the bending resistance and the shear resistance properties of the structure. By estimating tension characteristics adequately, it is encouraged to produce the structure with efficient safety and also with the economical rationality. Prior to proposing a suitable methodology for fibre concrete structures design, demands on FRC analysis should be stated. Designing of FRC elements must be compatible to concrete structures analysis. Benefits of fibre concrete must be taken into account. A general routine for fibre concrete member is not standardized yet. There are many types of fibre concretes with diverse behaviour. These are the reasons to find methodology as simple and low-cost as possible. Analysis of any structure must be based on realistic material properties and suitable material model. The basic material characteristics shall be determined in common laboratory tests: compression test, tension test, flexural test.

RESEARCH OF SSFC BEHAVIOUR

In an experimental programme at our department a structural synthetic fibre concrete with polypropylene fibres has been investigated. The effect of a dosage and a type of fibres on load bearing capacity and service load behaviour with different types and amount of fibres was monitored. The programme was focused on finding of proper material model for design and evaluating material model influence on the structural analysis. Structural Synthetic Fibre Concrete (SSFC) was found advantageous. These positives of SSFC were proved in an analysis or are commonly known: SSFC with 1% volume content of polypropylene fibres showed good workability of the mixture with common fibre content, proper fibre dispersion was observed, no balling occurred. High tensile strength of the fibre enables effective use. With a small diameter of fibres is needed aspect ratio obtained with relatively short fibres. Polypropylene fibres ensure both good control of shrinkage and induced cracking (<0.5%) and enhancement of the tensile and flexural strength of SSFC. And an important positive of polypropylene fibres is contribution to fire safety of SSFS elements. In initial stages of research project a dosage of fibres was examined. A one percent volume content of fibres showed convenient properties of resultant material (ductility, residual strengths) and satisfactory met demands on both shrinkage cracking and layout of cracks at ultimate loading. In a following stage of the project a structural element was inquired into. Two mixtures were prepared to compare benefit of fibres in the concrete matrix: a conventional concrete and corresponding fibre concrete with polypropylene fibres Forta Ferro. Fibres Forta Ferro have strength c. 700 m a , length 50mm and bulk density 910kg/m3.

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A set of beams 1800 x 150 x 100 mm reinforced with two steel bars of 6 mm diameter both fiom concrete and fibre concrete were manufactured. Properties of the steel used as alongitudinal reinforcement were determined in a tensile test. The yield strength was 603 MPa. The steel had relatively low increase of strength from the value of yield strength to ultimate strength. The beams were bended in a laboratory test and a deflection at midspan was measured and development of cracking was observed. The beams were loaded and controlled with deformation. The speed of loading was low at first stages of experiment after crack formation the speed was increased. The load was applied until a failure was reached.

Figure 1 : Setup of a laboratory loading of the fibre concrete beam with common steel reinforcement

Figure 2: Load arrangement of the beam in the bending test Record of load-deflection relation is depicted in the Figure 4. The lines in dark grey are load-deflection curves of fibre concrete beams (SF PPx), the light grey lines are for concrete beam (Cx) load-deflection relation. Until a first crack arises the L-D curves are practically identical, also behaviour of fibre concrete beam and concrete beam was similar. Several cracks were created then one main crack developed. Crack spacing was measured for concrete and fibre concrete beams. A fibre concrete beam had more cracks than a concrete beam. That implies more favourable layout of cracks of a fibre concrete beam. The mean value of crack spacing was 63mm for the fibre concrete beam and 80mm for the concrete beam. In the

Alena KOHOUTKOVA,h a BROUKALOVA

346

Figure 3 are depicted regions near midspan of fibre concrete beam (above) and concrete beam (below).

Figure 3: Comparison of crack spacing of the fibre concrete and common reinforced concrete beam

16,OO 14,OO

-3

12,oo 10,OO

Y

z

8,OO

0

6,OO 4,OO 2,oo 0,oo 0

5

10

15

20

25

Deflection (mm)

Figure 4: Load - deflection curves for concrete beams (C) and fibre concrete beams (SFPP) recorded in a laboratory test

Structural peflonnance and crack control ofjibre concrete beams with conventional reinforcement

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The failure of a concrete beam came on with a fracture of steel reinforcement. The beam was broken into two pieces. Compared to it fibre concrete beam remained whole even after the steel rebar reinforcement was broken thanks to the fibres bridging a crack A higher loadbearing capacity of comparable fibre concrete beams is evident from the picture (Fig. 4). Comparing the ultimate part of the experiment different failure was observed for reinforced concrete beam and fibre concrete beam with conventional reinforcement. The failure of common concrete beam was rather brittle while the failure of fibre concrete beam had a ductile character.

NUMERICAL SIMULATION An additional aim of the project was a numerical simulation of fibre concrete member

behaviour. Structural analysis of the fibre concrete element could be performed in a traditional way of design where inner forces are calculated assuming elastic behaviour of the structure and other methods are used to embody non-linear or plastic behaviour in the analysis. Fibre concrete members behave strongly non-linearly. For that reason the potential profit from the nonlinear analysis of fibre concrete structures is much higher. The laboratory loading of the beams was simulated in a finite element program. For simulation of the beam behaviour a finite element program ATENA was used. The ATENA program is a tool for computer simulation of damage and failure of concrete and reinforced concrete structures. In material models involved in the program a fibre concrete material model is included. The tensile behaviour of concrete is modelled by nonlinear fracture mechanics combined with crack band method and smeared crack concept. Main material parameters for modelling of fracturing behaviour are tensile strength, fracture energy and shape of the stress-crack opening curve. Both beams were modelled in 2-dimensional version of the programme. Solid plane quadrilateral elements created FE mesh for concrete and fibre concrete. Reinforcement was modelled in a discrete form by truss elements embedded in concrete finite elements with axial stiffhess only, which is a usual way in FE programs to prevent problems in mesh generation. As a first step of the computer analysis the material model was determined and adjusted by means of an inverse analysis. Basic input parameters to inverse analysis were compressive and tensile strengths. Specimens recommended by RILEM were cast - cubes lOOxlOOxl00 mm and prisms 100x100x400 mm. In a compressive test of cubes compressive strength was taken. Values of tensile strength were determined in a splitting test. List of measured values of strengths is in Table 1. Table 1. Material properties of fibre concrete from testing

348

Alena KOHOUTKOVA, Iva BROUKALOVA

Other required material parameters were determined and fitted in the inverse analysis. A load - deflection curve recorded in a four-point bending test of a prism was a second source of the material parameters. The flexural test of a prism specimen was simulated in a FE analysis. Material properties for the first run of computer analysis were estimated. Results of FE simulation were compared with experiment and material input was adjusted to approximate the load - deflection curve from FE simulation to that obtained in laboratory test. This procedure was repeated for several times until acceptable coincidence of the curves was reached. Obtained material model was than used for numerical simulation of the reinforced fibre concrete beams.

Figure 5 : Comparison of load - deflection curves measured in a laboratory loading and the best approximation with material model obtained in inverse analysis There are many uncertainties in a structural analysis of a fibre concrete member. In a laboratory conditions relatively good distribution of fibres in a mixture was achieved; in real structures the probability of non-heterogeneity of the material is higher. Therefore a stochastic analysis was performed where both spatial uncertainties and material indeterminateness are taken into account. The variability of material properties was introduced by SARA-FREETATENA system for probabilistic assessment of engineering structures. The material input parameters from a deterministic analysis are used as mean values for random distributions of selected variables. Further stochastic parameters for selected material parameters were defined by their probability density functions. Selected material parameters were described by mean value, variance and other statistical parameters. The randomness of the input values reflected randomness regarding material properties. Correlation between random input variables was introduced in a form of the correlation matrix. In structural design, the current reliability assessment concepts, based on deterministic analysis should be replaced or supplemented (at least for more complex design) by probabilistic analysis.

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349

CONCLUSIONS The findings of this study can be summarized as follows: From the test emerges that despite bigger values of compressive strength and elasticity modulus of common concrete, due to the content of the fibres the reinforced fibre concrete beam had bigger flexural load bearing capacity. This can be seen from the data record of the laboratory experiment (Fig.4). The shape of the ruptured surface of the concrete beam was similar to a certain type of a brittle failure, whereas the ruptured surface of fibre-reinforced beams may be presented as a more plastic failure. In this case brittle character of the failure of the concrete beam might be explained by unexpected brittleness of high strength steel used in the test beams. In a stress strain diagram of such steel only a small increment of strength after yield point to ultimate strength can be observed. As brittle character of failure is not safe for the structural behaviour and therefore unallowable; in a case that it occurs fibres used in structural elements contribute to the safety of structures. Crack patterns of FRC beams exhibit more cracks with smaller spacing and with smaller crack widths, it means that fibre reinforced structural elements have more favourable response from point of view of SLS considering cracking and deflections. And higher service life may be assumed thanks to fibres resulted in higher ductility of the beam. Laboratory flexural test of both concrete and fibre concrete beam was modelled in a nonlinear finite element program ATENA developed by Cervenka Consulting especially for analysis of concrete structures. Using material model and material characteristics found in an inverse analysis a satisfactory coincidence of experiment and its simulation was reached. Only when using appropriate material parameters in design of FRC both load bearing capacity and serviceability could be assessed with liability. Number of samples is selected according to complexity of the problem to be solved and required quality of expected results. Already a small number of samples could give a reasonable estimation of stochastic parameters of the structural response. It was confirmed that the variability in properties, dimensions and other quantities could be expressed by random variables in a progressive method of structural design with probabilistic attitude.

ACKNOWLEDGEMENT This paper was prepared with financial support of the grant projects No. 103/05/2226 and 103/06/1559. The support of Grant Agency of the Czech Rep. is gratefully acknowledged.

REFERENCES 1. Destree, X.: Structural application of Steel Fibres as Only Reinforcing in Free Suspended Elevated Slabs: Conditions - Design - Examples, 6" W E M Symposium on FRC, Varenna, Italy September 2004 2. Krstulovic-Opara, N..: Use of partially-precast HPFRC systems for LNG tanks Workshop Fiber Reinforced concrete From theory to practice, Bergamo 2004 3. R. Pukl, J. Cervenka, V. Cervenka, D. Novb, M.Voiechovsk9 and D. Lehky: Deterministic and statistical models for nonlinear FE analysis of FRC-based structures,

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Proceedings of the 1st Central European Congress on Concrete Engineering, Graz, Austria, September 8-9,2005 4. A. Kohoutkova, I. Broukalovh: Experience with Experimental and Numerical Research of Fibre Reinforced Concrete Beams, Proceedings of the 1st Central European Congress on Concrete Engineering, Graz, Austria, September 8-9,2005 5. Suwada, H., Fukuyama, H.: FEM Analysis on the Shear Behaviour of HPFRCC Member, International Workshop on High Performance Fiber Reinforced Cementitious Composites in StructuralApplications, Honolulu, Hawaii, USA, May 2005 6. Li V. C., Rokugo, K.: Task group D conclusions - HPFRCC design assumptions, International Workshop on High Performance Fiber Reinforced Cementitious Composites in Structural Applications, Honolulu, Hawaii, USA, May 200

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

THE ANALYSIS OF CRACK FORMATION IN CONCRETE AND SLIGHTLY REINFORCED CONCRETE MEMBER IN BENDING Marta SLOWIK Lublin University of Technology Nadbystrzycka 40,20-6 18 Lublin, Poland, e-mail: [email protected]

ABSTRACT In the paper the author's own experimental investigation and numerical calculations concerning cracking moment in flexural concrete members are presented. Two kinds of members were examined non-reinforced concrete beam and slightly reinforced concrete one. The analysis of crack formation in such members was made by Finite Element Method and the influence of reinforcement on the process of crack formation was discussed.

Keywords Concrete structures; crack formation; fracture mechanics; strain-softening. INTRODUCTION Plain concrete and slightly reinforced concrete members in bending are treated in the same way during the dimensioning. In such members cracking moment plays the most important part at estimating their load carrying capacity. The results of experimental investigation, both the author's own [l] as others reported in the literature (e.g. D4browski [2]) show that cracking moments in slightly reinforced concrete members are greater than in concrete ones without reinforcement. It is important to explain why the values of cracking moment in non-reinforced and reinforced concrete beams differ in order to estimate cracking moment in such members properly. In the first place it is necessary to recognize the influence of reinforcement on the fracture process of concrete. The fundamental rules for crack formation and crack growth are given by fracture mechanics and it is possible to apply these rules to complicated cases using finite element methods. Many nonlinear fracture models have been proposed to predict cracking of concrete. One of them is the model given by Hillerborg, Modeer, Peterson [3] which seems suitable to describe crack formation in flexural concrete members. The basic idea of this model is shown in Fig. 1.

Fig. 1. The fracture model for concrete in tension [3].

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Marta SLO WIK

The crack starts to propagate when the stress at the crack tip reaches the tensile strength

fcmt. The stress does not fall to zero at once but it decreases with increasing crack width and it

reaches zero when crack width is W I . For that part of the crack where W < W I , there is a fracture process zone with some remaining ligaments which allow to transfer stress. As these ligaments are to be overcome during opening the crack, energy is absorbed. The amount of energy absorbed per unit crack area is w,

GF=j ~ d w 0

The value of the energy GF corresponds to the area under the curve C-w, taken eom tensile test (Fig. 2.a). For the application of the proposed model to numerical calculations the curve GW may be simplified by, for example, linear relation (Fig. 2.b).

Fig. 2. Typical variation of o-w taken from tensile test a). Example of possible assumption of o-w in practical applications b). Another proposition for practical usage, as far as concrete in tension is concerned, is given in the 1990 edition of CEB-FP Model Code [4]. The fracture model given in [4] suggests bilinear relation of o-w and non-linearity of concrete before the stress reaches the tensile strength (see Fig. 3). a)

0

.

9

. b)

(r

2

z

k

0.15

f,h

CEB-FIP Model Code

fcmt

I

0.00015 E

WI

W

3. Diagrams of stress versus strain a) and stress versus crack opening b) for the tension zone of concrete recommended by CEB [4].

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353

In the paper some numerical analysis, based on nonlinear fracture mechanics, has been made in order to provide better insight into the phenomena associated with tensile fracture of concrete in flexural member depending whether the member is or is not reinforced. That analysis has been verified by the author's own experimental results. EXPERIMENTAL INVESTIGATION Test program The author's own experimental investigation presented hereafter is aimed at the determination of cracking moment of flexural concrete members with and without reinforcement. The test program was carried out at the Laboratory of the Department of Civil and Sanitary Engineering of Lublin University of Technology. Six concrete beams were tested, with the rectangular cross-section and the following dimensions: width - b=O.15m; height - h=0.30 m; total length - L=3.00 m; span - 1=2.70 m. Three beams which were marked S7, S8, S9, were not reinforced and three, which were marked S1, S2, S3, were slightly reinforced. The longitudinal reinforcement consisted of three steel bars, e.5 mm in diameter, and it gave the ratio of reinforcement p=O.12%. The bars were anchored in compression zone of the beam. Stirrups were used only in regions near the supports. Beam geometry and the location of reinforcement in reinforced beam is shown in Fig. 4.

Fig. 4.Beam geometry and the location of reinforcement. The experiments were performed using the four-point bent specimens and specially designed test stand. The stand was constructed in such a way as to enable the observation of the beams' work in post-critical range. Beams were loaded by two concentrated forces, which were applied from bottom towards top by hydraulic jacks. During the test, the displacementcontrolled experimental procedure was used in order to slow down the cracking process and to observe crack formation more precisely. The static scheme of the test specimen is shown in Fig. 5.

L I

Y

-

+-L __ 13-- + .I

Fig. 5. Static scheme of the test specimen.

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Marta SLO WIK

During the subsequent loading stages of the test, beam's deflections, applied external forces and concrete strains were measured. Concrete strains were recorded on two levels, in tension and in compression zone of the beam, by Huggenberger's gauge. The arrangement of the bases for strain measurement is presented in Fig. 6. level of tensile strain I

I 1

I

I measurement

I Fig. 6. The arrangement of the bases (numbered through 9) for strain measurement. Concrete and steel properties The mechanical properties of concrete were tested by standard methods. The compressive strength of concretef, was measured on twenty-one cubes, 150 mm in side. The tensile strength of concrete hcSp was measured on twenty-one cubes, 150 mm in side, through splitting tensile test. The modulus of elasticity E, was measured on twenty-one cylinders, 150 mm in diameter and 300 mm in height. The results obtained on specimens were evaluated statistically and they are listed in Table 1.

Property

Mean value [ma1

Standard deviation [MPaI

Coefficient of variation v

The compressive strengthf, The tensile splitting strensthf,l,sp The modulus of elasticity E,

27.67

2.77

10.0

1.64

0.18

10.7

22118

1927

8.7

f,= (0.77 - O.OOlf,)f,

= 20.54

[%I

MPa

(2)

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355

The value of the mean axial tensile strengthfc,, was estimated from the mean splitting tensile strength according to the formula given in the CEB-FIP Model Code [4]:

The fracture energy was not measured experimentally. In the absence of experimental data the fracture energy G,v may be estimated from the equation proposed in [4]:

GF = a, :’f

=82.95 Nm/m2

(4)

where: a F - coefficient which depends on the maximum aggregate size d, (dm,=8 mm - a F =4, d,,=16 ~ll~lla F =6, d,,=32 ~ll -~ ~ ~ ll 1 0 ) . To reinforce beams S7, S8 and S9, A1 steel grade bars were used, 4.5 mm in diameter. The mechanical properties of longitudinal reinforcement were tested on eight test elements. The determined mean value of yield strength was: f,=274.5 MPa (n=8, s=12.94 MPa, ~ 4 . 8 % ) , and the mean value of tensile strength was:f~398.9MPa (n=8, s=8.92 MPa, ~ 2 . 2 % ) . The author’s own experimental results In concrete beams only one crack and in reinforced concrete beams two or three cracks appeared. The destructive crack in all the test beams formed near the cross-section where one of the forces was applied. The cracks always ran perpendicular to the centre line of the beam. As already mentioned, the displacement-controlled loading procedure was used during the test. Therefore, in non-reinforced concrete beams, after the maximum force had been reached, it was possible to continue the test at the decreasing force and to determine the concrete strain precisely. In the loading stage, when the maximum external force was measured in the tested beams S7, S8 and S9, the following values of concrete strain in tension zone within the base O . I ~ ~ X ~ OThus - ~ . the in which the crack formed, were obtained: 0 . 1 5 1 ~ 0. 146~10-~, average value of the ultimate tensile strain of concrete, calculated from the experimental data, 0 0 When the declining stress-strain relation was observed, the maximum was ~ l , , ~ O . 0 14. elongation was recorded and it gave the maximum strain value in the tension zone &,=0.275~10-~, but this value of the strain was due to the opening of microcracks. During the experiment the following cracking forces in plain concrete beams were measured: 6.53kN, 6.91kN, 5.90kN. Based on these the average experimental cracking moment was determined: Mc,~5.08kNm.After the crack had formed, a sudden fracture of the tested beam into two pieces occurred. In the reinforced concrete beams S 1, S2, S3 after the cracks had appeared the applied forces did not decrease as in the plain concrete beams. The forces increased a little bit and stabilized on the constant level but the opening width after forming of the crack was relatively big. During the experiment the following cracking forces in slightly reinforced concrete beams were measured: 5.59kN, 5.44kN, 5.21kN. The average experimental cracking moment which was determined for slightly reinforced concrete members: Mc,~5.39kNmwas greater than in concrete members. Also the fracture process went in different way in those elements. The cracks formed slowly and the destruction process in slightly reinforced concrete beams went progressively, with cracks developing little by little. Based on experimental data it may be said, that the presence of reinforcement influences cracking process in the concrete member. The mean value of cracking moment was greater in reinforced elements than in concrete ones, although the percentage of reinforcement was low.

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Comparison with other experimental results The similar results as the author’s own were obtained by Dqbrowski [2]. His experiment consisted of a few series of concrete and slightly reinforced concrete beams. All specimens had the rectangular cross-section and the same dimensions: width - b=0.2 m; height - h=0.2 m; total length - L=l .O m; span - 1=0.9 m. The beams were tested at four-point bent test. Two external forces were applied symmetrically at one third of the span. Tested specimens were made with different kinds of concrete. Based on the obtained results (reported in [2]), Dqbrowski proposed the following relations between cracking moment in slightly reinforced concrete member M p and cracking moment in plain concrete one Mcr:

- when ~ 1 0 . 0 8 %

(5)

- when 0.08% S p S pmin

(6)

where: a: PI, P2 experimental coefficients which for examined beams were equal respectively 1.15; 250; 62.5. The experimental results show that the cracking process in flexural member is influenced by reinforcement. In order to explain the difference between the crack formation in flexural member with and without reinforcement, the analysis of the author’s own test results using Finite Element Method (FEM) was made. The theory of fracture mechanics was applied in the analysis. NUMERICAL CALCULATION Modelling of the beam The numerical calculations were performed using the APAKO module of the commercial finite element program ALGOR. The FEM-analysis was performed on non-reinforced and reinforced concrete beam corresponding to the test specimens. The bulk material of the beam was chosen as linear elastic one and only in the region of fracture process zone the concrete was modelled as nonlinear material. The FEM-beam model was composed using brick and truss elements. The brick elements were used in the region of the linear material behind the fracture process zone and truss ones were used in the fracture process zone. The fracture zone was modelled in the region where the load was applied and where the biggest bending moment value was obtained. The width of the fracture process zone wc was taken 10 mm. The chosen value of w, was successfully verified and the results of that verification were presented in [6]. Since the four-point bent test is symmetrical, only half of the beam was used in calculations. The finite element mesh for the analysed beams is shown in Fig. 7. When performing FEM calculations, the same material properties as those obtained during the experiment were taken. To describe the fracture region of concrete the Straight Line model (linear relations of 0-Eand OW) and model CEB-FP shown in Fig.3 were used, for which the following parameters were assumed: j&=1.48 MPa, 0.9f,,=1.33 MPa, Ecm=22118 MPa, ~,,=0.00015,~ ( ~ . 9 f , ~ ~ , F 0 . 0 0 0G0~68, 2 . 9 5Nm/m2, w,=0.25 mm, wl=O.O37 mm.

The analysis of crack formation in concrete and slightly reinforced concrete member in bending

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B-B

’

-4

150

concrete beam

reinforced concrete beam

Fig. 7. The finite element mesh. As the result of FEM-calculations, the dislocations of nodes and stress components along three axes of the global coordinate system were obtained.

Comparison of numerical results with tests In the first step of the analysis, results of FEM-calculations were compared with experimental data. For this reason the diagrams of beam’s deflection a measured in the middle of the beam’s span versus elongations measured on two chosen bases shown in Fig. 7 were prepared: for the base number 3 where the destructive crack appeared, and for the base number 5 which was located outside the fracture process zone. In Fig. 8 the results of calculation and experiment for the plain concrete beam are presented and in Fig. 9 for the beam with longitudinal reinforcement, 3 bars 4.5 mm in diameter. Based on the diagrams presented in Fig. 8 and in Fig. 9, it may be said that FEMcalculations are in good agreement with the experimental results both for the concrete beam without reinforcement and for the beam with reinforcing bars. Also the value of cracking 5 .kNm is nearly the same moment in concrete beam obtained in FEM-calculation M c , . ~ ~ p18 as the experimental cracking moment M , , ~ 5 . 0 8kNm. As the bending moment due to weight of the beam is significant enough, it was taken into account when calculating the cracking moment. It appears that the concrete model used in FEM-calculations which is based on nonlinear fracture mechanics, gives realistic results. The fracture tensile model of concrete described in [3] is suitable for the analysis of the cracking moment of concrete flexural members. Based on FEM-calculations the analysis of crack formation in flexural concrete member can be done.

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0.5

0.5

0.4

0.4

0.3

T

I

E

0.3

a

d

I

0.2

0.2

0.1

0.1

O.O0

O.O1

'.O2

elongation [m]

0.03

0.00

d

++ FEMcalculation,model CEB-FP 0

0.01

0.02

0.03

FEMcalculation, model SfmightLine Testresults-beamS7 Test results- beam S8 Test results- beam S9

0.04 0.05 0.06 elongation[mm]

0.07

0.08

0.09

0.10

Fig. 8. Diagrams of beam's deflection a versus elongation on base 5 , (a), and on base 3, @), for the plain concrete beam.

0.1 -

ao

,

Test results- b a r n S1 Test results- beam S2

I

Fig. 9:Diagrkns of beam's deflection a versus elongation on base 5, (a), and on base 3, (b), for slightly reinforced concrete beam. Analysis of FEM-results Figure 10 shows the diagrams of the calculated normal stress om along the fracture zone of the non-reinforced and reinforced concrete beam in different loading stages. The curves are linear up to the force F= 3.0 kN. At this load state the tensile concrete strengthfch, is exceeded in an upper side of the beams. The bending moment Mo, which is obtained at &=Arm,would be the failure moment if the material was elastic and perfectly brittle. As the concrete is not an elastic material, the bending moment is able to rise above A40 and then the crack starts to form in the fracture process zone. At the following load stages the stress in an upper site of the beams decreases and the crack tip runs through the fracture zone, so in deeper laying points of the fracture zone the stress reaches the tensile concrete strength. Based on the stress diagrams corresponding to forces higher than F=3.0 kN, the strain-softening of tension concrete in flexural member may be described.

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a)

-3500 -3000 -2500 -Zoo0 -1500 -loo0 -500

0

1000

500

1500

b) 0.30

n 74

-A . r

F=3.95 kN

A

bl8

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1-7

0.12 0.06

0.00 -4Ooo

-3000

-lm

-2000

0

1 m

ZOO0

a,Fpal

Fig. 10. Normal stress diagrams in the fracture process zone at different load stages for concrete beam a), and for reinforced concrete beam b). Differences in tension stress diagrams between non-reinforced and slightly reinforced concrete beam appear at load stages when strain soRening starts to develop. In the concrete beam stresses decrease quicker than in the beam with reinforcement. The process of crack formation goes slower in the reinforced beam. It is especially seen on the level of reinforcement where the decrease of stress is lower than on the levels next to it. Probably it is caused by the bond between concrete and reinforcing bars. The difference in crack formation in the beam with and without reinforcement is clearly seen when the stress diagrams are compared at the same load state which is presented in Fig. 11. 0.30 0.24

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Fig. 11. Tensile stress distribution for non-reinforced and reinforced concrete beam at the same load state in the fracture process zone.

Marta SLO WIK

3 60

CONCLUSIONS The initiation and propagation of a continuous crack in concrete occurs after the formation of the fracture process zone in concrete. In this zone the strain-softening of concrete takes place. Based on the presented numerical analysis, it may be said that the fracture process in flexural concrete member is related to strain-softening which takes place in tension zone of the cracking element. The cracking moment should be calculated with regard to this phenomenon and the influence of reinforcement should be taken into account. In flexural concrete beams the failure mechanism changes when the beam is reinforced. In order for the crack to form in concrete member, two conditions must be fulfilled. These conditions, given in literature [7], are: -strength condition: n 2 S,, or E 1 E ~ , ,

au

-energetic condition: -= G,b

ac

in which: c - the length of the crack (fracture zone), U - the total amount of potential strain energy. The cracking process is connected with the order in which the above conditions are fulfilled. When the strength condition is reached first, then the crack forms slowly in a static way. This kind of crack formation is observed in reinforced concrete members. When the energetic condition is met first, the brittle fracture takes a place. Such case of cracking is typical of non-reinforced concrete. In slightly reinforced concrete members the way of crack formation may go in a different way. The presence of reinforcement, even when reinforcement ratio is low, changes cracking process which influences cracking moment in flexural members. It is important to recognize more broadly the mechanism of crack formation in slightly reinforced concrete members according to the reinforcement ratio. This problem requires further analysis. REFERENCES 1. Slowik M., Analysis of Load Carrying Capacity and Serviceability of Slightly Reinforced Concrete Structures (in Polish). PhD Thesis, Lublin 2000, pp. 187 2. Dqbrowski K., Slightly Reinforced Concrete Members of Rectangular Cross Section (in Polish). Towarzystwo Naukowe Ekspertdw Budownictwa w Polsce, Warszawa 1962, pp. 80 3. Hillerborg A., Modeer M., Petersson P. E.: Analysis of Crack Formation and Crack Growth in Concrete by Means of Fracture Mechanics and Finite Elements. Cement and Concrete Research, Vol. 6, 1976, pp. 773-782 4. CEB-FP Model Code 1990 Bulletin d’information No. 196 5 . The Polish Code: PN-84B-03264, Concrete, Reinforced Concrete and Prestressed Concrete. Design Rules. 6. Siowik M., Blazik-Borowa E.: The Influence of the Width of the Fracture Process Zone on Numerical Calculations Concerning Concrete Beams. 33-rd Solid Mechanics Conference, Zakopane 2000, Volume of Abstracts, pp. 361-362 7. Wolinski Sz.: Tensile Behaviour of Concrete and Their Applications in Nonlinear Fracture Mechanics of Concrete (in Polish). Scientific Works of Technical University of Rzeszow, vol. 91, Rzesz6w 1991, pp. 210

Proc. Int. Symp. "Brittle Matrix Composites 8" A.M. Brandt, re. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

MULTIPLE CRACKS BRIDGED BY MULTIFILAMENT YARNS: IMPACT OF LOCAL SCATTER ON ULTIMATE LOAD Miroslav VOkECHOVSKy , Martin KONRAD**and Rostislav CHUDOBA** 'Institute of Structural Mechanics, Bmo University of Technology Vevefi 95,602 00 Bmo, Czech Republic, e-mail: [email protected] ** Chair of Structural Statics and Dynamics, Aachen University of Technology Mies-van-der-Rohe-Str. 1,52056 Aachen, Germany e-mail: [email protected], [email protected]

ABSTRACT The paper shows the correspondence between the short-range size effect occurring in a crack bridge and the long-range size effect observed on a textile reinforced concrete (TRC) tensile specimens. First an approach to modeling crack bridge behavior is introduced exploiting the evaluation of the mean Performance of a multifilament bundle. The influence of imperfections in the material structure is discussed using several parametric studies. The disorder in the filament bundle and the heterogeneity in the interface layer are described in form of statistical distributions. The crack bridge model provides the basis for studying the effect of the multiple cracking. In this analysis we exploit the fact that the specimen acts as a chain of crack bridges in the failure state so that it follows the weakest-link statistics. As a consequence, direct link between the local scatter in the material structure and the global size effect on ultimate strength can be easily established.

Keywords fiber bundle model, statistical size effect, chain of bundles INTRODUCTION The material structure of cementitious composites reinforced with multifilament yams such as textile reinforced concrete exhibits a high degree of heterogeneity and imperfection that requires special treatment in the development of numerical models. The micromechanical model with the fine resolution of filaments (bond-layer model) used for studying the crack bridge performance under monotonic loading has been thoroughly de-scribed in [ 11. The previous studies using the bond layer model have shown the significant influence of local imperfections on the performance of the crack bridge. These imperfections are exemplified by nonparallel orientation of filaments within the bundle or by varying bond quality between filaments and matrix across the bundle. As a result, the damage localization process of textile reinforced concrete exhibits interactions between elementary failure mechanism in the matrix, in the reinforcement, and in the bond. The presence of imperfections in the yam and in the bond structure makes it necessary to account for the effect of scatter of the material properties on the performance of the crack bridges. The crack bridges are the hot-spots of deformation and interaction of damage with localized damage at the material interfaces of matrix and reinforcement. In a way the crack bridge is equivalent to a tensile test on yam with an extremely short length equipped with a shear-lag-like clamping of filaments. Due to the varying penetration profile along the yam,

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Miroslav VOi?ECHOVSKf Martin K O N W , Rostislav CHUDOBA

the quality of the shear lag clamping exhibits high scatter. What is sought is a true statistical representation of the material structure able to reproduce the crack bridge behavior in different configurations and loading histories [2]. As documented in [3] on experimental and numerical studies, these sources of randomness can lead to a substantial reduction of the bundle strength especially for extremely short nominal lengths occurring in a case of a crack bridge. In this paper, the local statistical crack bridge model is presented in the context of the tensile specimen with multicracking with the goal to evaluate the effect on the ultimate load. In the final stages of the tensile loading with a finished crack pattern the specimen may be viewed as a chain of crack bridges with its strength governed by the "weakest-link" concept. The experimental results on tensile TRC specimens show a significant loss of bundle efficiency that are usually ascribed to damage or low penetration of the bundle by the matrix combined with insufficient anchorage in the boundary layers [4]. This paper contributes to the discussions about the reasons for the strength reduction by studying the weakest link effect in a chain of crack bridges with scatter of strength. The exploited fiber bundle model (FBM) representing a crack bridge is introduced first and illustrated on an example of scattered fiber lengths. Next, we introduce a bond model and plug it into the FBM. Parametric studies are utilized to illustrate the effect of scatter of bond quality together with varying bond free length. Effect of correlation of these two properties is briefly reviewed and the model is applied to the experiment. Finally, in order to demonstrate the correspondence between the statistics of the global response and the local scatter in the crack bridge we review the weakest link concept. The simple modeling framework allows us to study the change of the ultimate strength with an increasing number of cracks N for different levels of scatter of local filament properties in the crack bridges. APPROACHES TO MODELING CRACK BRIDGE BEHAVIOR

The filament bundle models provide the stepping stone for robust modeling of the failure process in the bond layer between the multifilament yam and the cementitious matrix. In the following the existing models for cracks bridged by fibers or fiber bundles will be classified based on the considered variability of material parameters. The effect of the irregular bond structure on the performance of the crack bridge has been represented by a distribution of the effective filament lengths in [5]. The work [6] additionally considers a finite bond and the abrasion of the filaments during the debonding. In [7] a micromechanical model has been developed to characterize the interfacial properties at single fiber pullout. The effect of fiber alignment on the pullout load was investigated in [8]. The influence of fiber alignment on in-situ strength has been characterized in [9]. In [7] Eq. (1) has been used to derive the crack bridging stress versus crack opening relation by averaging over the contributions of the filaments. With the strain based formulation of statistical fiber bundle model [lo] it is possible to evaluate the mean asymptotic response of the bundle under tensile loading analytically:

Where q ( e ; 0 ) represents the filament constitutive law with global strain e and the vector of material parameters 0 and Ge(0) is the cumulative probability distribution function of the parameter(s) 0 . We note that this class of models is limited to the local load sharing between filaments upon filament break. In other words, the load of the lost filament gets uniformly distributed

Multiple cracks bridged by multifilament yarns: impact of local scatter on ultimate load

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across the bundle. Based on the experimental evidence of low friction between filaments presented in [ 111 we neglect the interaction effect between the filaments with effective lengths the in range of several millimeters as they occur in the crack bridge. A thorough study of sources of randomness/disorder in the multi-filament yarn relevant for its performance in textile reinforced concrete has been presented in [3]. For the analysis of bundles with infinite number of fibers a continuous analytical model with refined kinematical relation has been developed. The considered distributions of material properties include both the variations of parameters from filament to filament and the variations of local strength and material stiffness across the bundle. In the companion paper [12] the variations of material properties over the length are considered.

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Figure 1. Mean force-strain diagrams of one crack bridge with uniform (left) and normal (right) distribution of additional fiber length 1E (0,1=$.1-= a) 0.0, b) 0.25, c) 0.5, d) 0.75, e) 1.0, f ) 1.25 and in case g) 1.5 plotted with a scatterband (mean f standard deviation). The right hand figure is computed with debonding (4 =3.38 MPa). The scatter of filament stiffness parameters leads to a reduced strength of a bundle with a very short length. In a crack bridge the strength reduction is in particular caused by the scatter of filament lengths and their delayed activation (slack). Besides that, the scatter of filament strength across the bundle leads to a further reduction of the bundle peak force [13]. In order to demonstrate the effect on an example we consider a crack bridge with a scatter of filament lengths. In particular, we introduce the relative difference of the filament length with respect with a uniform distribution, i.e. G, ( A )= A / & , to the minimum length as A = ( Z - l i ) / Z where 0 I A I &. The other parameters of the filament, i.e. Young's modulus E, area A and breaking strain ( are considered constant. For the chosen distribution, it is possible to derive analytical formulas for the bundle mean strength and its variance (see [3]) at a given control strain e as

Miroslav VO&CHOVSKf, Martin KONRAD, Rostislav CHUDOBA

3 64

The calculated mean load strain diagrams for a bundle with 1743 filaments and constant diameter D = 25.5pm, Young‘s modulus E = 70 GPa and breaking strain {= 1.79 % are exemplified in Fig. 1 left for several levels of scatter represented by &, The introducing example demonstrates the modeling concept used in the first part of this paper. In addition to the scatter of filament lengths, we shall include the scatter of the bond quality across the bundle. In order to justify this decision we first discuss the issue of appropriate choice of material parameters in general. The overall goal in the construction of the model of a crack bridge is to predict its performance in a broad application context, i.e. in an arbitrary configuration of cracks and textile fabrics. These configurations are represented by the crack spacing on the one hand and by the angle at which the roving crosses the crack on the other hand. In this paper we consider only tensile response and limit the discussion to the aspects of bond parameters that should be independent of crack spacing (Fig. 2).

I

I

Figure 2. Heterogeneity in the material structure with the sought distributions of micromechanical parameters. The minimum number of structural parameters is given by their capability to reflect the changes in the material structure during the evolution of local damage. As shown in the example in Fig. 1 left, it is possible to find such a distribution of filament lengths that can reasonably reproduce the experimental response. Indeed, the performance of a single crack can be reproduced by a proper choice of length profile resulting in the corresponding number of breaks at each level of control displacement. However, this distribution of smeared effect of length and bond variations cannot be considered representative in the context of multiple cracks. There are especially two reasons: The crack bridge with structural variations ascribed solely to the effective length would exhibit a shorter stress transfer length than under the assumption of finite bond. Therefore the fitted crack bridge stiffness would not be consistent with the bond stiffness and, thus, inconsistent with the crack spacing. 0 Even if the crack spacing would be defined a priori, e.g. using the experimentally measured crack distribution, the effective length profiles obtained from a single crack bridge

Multiple cracks bridged by multifilament yarns: impact of local scatter on ultimate load

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experiment would lead to an underestimated stiffness in the case of unloading of existing cracks at the formation of a new crack. This is due to the fact the stiffness reduction during the opening of a crack bridge can only be appointed to the failure of filaments. Therefore the crack will unload with the resulting secant stiffness. A consideration of plastic deformations in the interface leads to significantly higher stiffness. Separate consideration of the scatter of lengths and of the bond quality across the bundle in the vicinity of a crack bridge allows us to capture interactions between cracks. It would certainly be possible to resolve the material structure in even more detail. On the other hand, a pragmatic material resolution requires only capturing the components and interfaces that dominate in the disintegration of the material structure.

EFFECT OF FINITE BOND STRENGTH The procedure for evaluating the total strength described in the second section can be used with more complex idealizations of a crack bridge taking into account fiuther failure and damage mechanisms. In addition to filament rupture considered in the previous model we now include the influence of debonding between filament and matrix. The performance of a multifilament yam bridging a crack is evaluated by replacing the constitutive law q ( e ; 0 ) in Eq. (1) with a filament pull-out response pu( q 0 ) that could be derived analytically using a shear lag model [ 141 with a cohesive interface between the filament and matrix. With G, (A) defined as in the last example, we now include the debonding of a filament from the matrix as an additional effect characterizing the interface between a perfectly embedded filament and the cementitious matrix. In [6] filament matrix bond laws for perfectly embedded filaments have been derived as multi linear functions z, (s) (with s representing the slip) using the filament pull-out test and the cohesive interface model. In the particular case of the Vetrotex AR-glass roving with 2400 tex the derived function shows that no adhesion bond is developing and thus the interfacial properties are predominated by friction. Therefore z, (s) can be represented in the following bilinear form: z, ( s ) = z, .s / sCdt if 0 Is Isent and z, (s) = z, for s > sent

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Figure 3. Mean force-strain diagrams (kone standard deviation) of one crack bridge with debonding model ze and with uniform distribution of additional fiber length 1 E (0, ,Imx).

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Miroslav VO&CHOVSKf, Martin KONRAD, Rostislav CHUDOBA

Again we keep all other parameters including the interface characteristics, constant. Fig. 3 left shows the load displacement diagrams for the varied range of lengths with /2, E [0,0.25, 0.5, 0.75, 1, 1.25, 1.51 and a constant bond stress z,=3.38 m a . It exemplifies that the crack opening at peak load increases due to the debonding of the filaments. To demonstrate this effect we now keep &a = 1.5 constant and vary the bond stress in the range z, = c . 3.38 MPa, C E [lo, 1, 0.5, 0.25, 0.1251. The resulting load displacement diagram Fig. 3 right shows that the mean crack bridge strength increases significantly with decreasing bond stress z,. The reason for this increased strength is the homogenizing effect of the debonding causing a more uniform stress distribution across the bundle (more filaments can act simultaneously before they break). We also note that there is a significant reduction in the scatter of strength compared to Fig. 3 left: 184 N for perfect bond and 50 N with included debonding and the lowest bond stress. pv=a) 0.75, b) 0.5, c) 0 . 2 5 , =O. ~ ~ 125 ~ 1200

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Figure 5. Mean load-displacement diagrams of one crack bridge with constant bond free length ( l- = 0 ) and uniform (top) and normal (bottom) distribution of bond quality p [-I EFFECT OF SCATTER IN BOND STRENGTH Due to the irregular penetration of the matrix into the yarn the quality of the interface between matrix and filament varies. The reason for this scatter of bond quality are the varying contact area between filament and matrix and variations in the quality of the matrix itself due to the heterogeneous infiltration of aggregate particles. To evaluate the influence of the scatter in the interface we introduce the bond quality p of an individual filament as a dimensionless scaling factor for the bond law. The bond stress slip relation for this filament in the bundle is defined as z(s,p)=p.z,(s).

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The influence of a bond quality uniformly distributed over the bundle cross section with G, (p) and , I , =0 (all filaments have the same bond free length) is exemplified in Fig. 5 top

left. The bond quality p has significant influence on the stiffness. With decreasing mean bond quality p , the stiffness decreases due to the large amount of debonding (Fig. 5 top left). The leads to reduced peak peak load is reached at larger displacements. An increased scatter (0,) load and more ductile post peak behavior as a result of more inhomogeneous stress transfer in the interface (Fig. 5 top right). The same effect can be observed for normal distribution bond quality p, see Fig. 5 bottom left and right.

EFFECT OF CORRELATION SEM micrographs of a mutlifilament yarn embedded in the cementitious matrix suggest that there is a positional dependence for the two studied parameters 1and V, . Regarding the outside filaments of the cross section we see that they are very well embedded (v, = 1) while their free length is very short ( A = 0). For the inner filaments we can observe the opposite. This raises the question of correlation between these two random parameters. A strong negative correlation leads to a slightly reduced crack bridge performance. The negative correlation leads to more inhomogeneous stress transfer reducing the peak load an associate crack opening to higher stress in the softening branch. This corresponds to the combination of short filaments with high bond and long filaments with low bond. We remark if it would be technically possible to achieve the other extreme and combine low bond quality with large bond free lengths and vice versa, it would be possible to obtain an increased bond performance. In [ 151 we studied the effect of correlation p between the additional length and bond quality p and reported that the effect of both positive and negative correlation is surprisingly low.

I

I

I

Figure 6. Profiles describing the heterogeneity in the material structure.

APPLICATION TO THE EXPERIMENT In order to relate the discussed issues to experimental data we shall identify the parameters characterizing the material structure using a pull-out experiment. We shall simplify the model in such a way that the variability is considered only deterministically. This is justified by relatively low effect of correlation discussed above. If we assume p + -1 we can interpret the

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Miroslav VOkECHOVSKf Martin KONRAD, Rostislav CHUDOBA

inverse cumulative density functions of our distributions as profiles of heterogeneity varying across the bundle with the random nature replaced by the parametric coordinate 5 (Fig. 6 ) representing the position of the filament in the bundle. Based on numerical and experimental studies and for the sake of completeness we introduce the third parameter considered in the model: The delayed activation of filaments (slack) within the bond free length. We further assume negative correlation between bond quality and delayed activation. The interaction of the delayed activation with the bond free length and the bond quality has been described in [I]. For the purpose of numerical calibration using the experimental data we approximate the introduced profiles by the following polynomial functions:

The parametric coordinate 6 = 0 represents the idealized center of the bundle and { = 1 coincides with the outer layer of the bundle. With the assumption of uniform distribution of imperfections in the bond and yarn structure along the specimen the effect of imperfection on the performance of each crack bridge can be considered the same. Thus, the quantification of this effect can be performed with an isolated analysis of a pull-out experiment in connection with the bond layer model shown in Fig. 2. As can be seen in the figure the profiles of heterogeneity are directly used for discrete representation of the crack bridge. We emphasize that the fine resolution of the material structure further away from the crack bridge is not needed due to the increasingly homogeneous stress profile. In this way we construct the discrete bond layer model (Fig. 2) with the interface layer between the yarn and the matrix represented by a set of laminas. Each lamina interacts with the matrix trough a given bond law. It represents a subset of filaments with the same characteristics and is coupled with the matrix using zero thickness interface elements [ 161. The bond law distributed over the bond layer is then constructed as z( s,5) = z, (s (5))p(5) . The parameters describing the profiles are obtained using the calibration framework of automated calibration procedures based on evolutional strategies to avoid convergence to local minima. GENERAL DETERMINATION OF TRC STRENGTH STATISTICS As documented in Fig. 7, the tensile specimen exhibits very fine crack pattern. Obviously, the ultimate failure is governed by the weakest-link statistics. Therefore, the survival probability of the chain with N cracks may be obtained as a product of survival probabilities of the indi, where Pf = F; (6)represents the failure probability of a vidual cracks: 1- P,,N= (1 - P,,,)N single crack bridge (cumulative strength distribution). The load level D for a given number of crack bridges N and probability of failure Pf ,N can be computed with the inverse cumulative

,,

strength distribution function of a crack bridge: o = 4-'(1- .JI-p/,,). In case of normally distributed strength of a crack bridge (short filament bundle) the load level of a chain of bridges with failure probability Pf , N reads simply:

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This formula provides a general procedure for estimating the strength statistics of TRC with N crack bridges, each with the failure probability distribution P,,,= F, (a). This distribution is obtained using the statistical strain-based bundle model [ 101 as follows Given a single-filament response function (constitutive law) q ( e ; 0 ) as a function of strain e and a vector of random (or deterministic) quantities e with their corresponding distribution functions GB,i(4) compute the mean response of a filament within the bundle (normalized bundle force) as a k-fold internal over k-number of nondeterministic variables of the model q (e;0) uskg the Eq. (1): Find the local maximum of the mean response (bundle strain e = e’ at which the maximum force is attained). This can be done either by seeking the stationary point of & (e; 0) in case of analytical expression, see examples in [3] and [12], or numerically by seeking the peak force value. Evaluate the mean bundle response function at bundle strain e = e’ to get the mean bundle strength: pi = po(e’ ) and compute the bundle strength variance as:

-

Estimate the whole cumulative density functionP,,, . In most cases it suffices to consider Gaussian distribution for the “middle” part (core) (0.1 to 0.9 percentiles) so that the probability of failure at a given load level u reads:

where @ ( 0 )

= standard

Gaussian cumulative distribution function.

Note that the distribution (of bundle strength P,,, ) can be estimated numerically by means of Monte Carlo simulation. One can evaluate the bundle response N,,,times and save the peak forces (of all simulations). Then the CDF of bundle strength can be estimated by an empirical cumulative histogram of the peak sample. EXAMPLES FOR SELECTED CRACK MODELS IN A CHAIN In order to demonstrate the procedure and the significance of the size effect in a TRC specimen we now provide two examples of a crack bridge model with the response function represented both analytically and numerically. Example 1. Bundle model with perfect clamping: Such an example has been studied in the introduction and the asymptotic load-strain diagram was given in Eq. (2). Obviously, the crack bridge strengthlefficiency rapidly decreases with an increasingil- (see Fig. 1 left). In the derivation of the crack bridge strength distribution we shall exploit the fact that the maximum mean bundle strength is attained at the global strain e* = 5 for the uniform distributions with A,,,, I 1.71 and has the simple form:

The corresponding variance (see [lo]) is obtained as

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Miroslav VOkECHOVSKy:Martin KONRALI, Rostislav CHUDOBA

With reference to the central limit theorem we can expect the convergence of the crack bridge strength distribution to the normal distribution, see Eq. (4). The verification of the convergence to the Gaussian distribution has been done by Monte Car10 simulation in connection with the deterministic bundle model described in [3]. With the crack strength distribution at hand we can approach to the quantification of the chain statistics. Using Eq. (3) we quantify the interaction of the short-range size effect due to a local scatter (;k,) with the chaining of crack bridges in a tensile specimen. In Fig. 4 we plot the crack bridge efficiency (reduction of strength with respect to a perfect crack bridge) for the failure probability P,,N = 0.5 (median) and the levels of the scatter parameter A,,= studied previously in Fig. 1 left.

/

1

\

Figure Left: Crack pattern of a failed tensile specimen reductions by studying --e we;.sst link effect in a chain of crack bridges with scatter of strength. Right: median chain strength for varying number of cracks and scatter 1 ( PJ,N= 0.5)

-

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Example 2. Bundle model with debonging: The procedure for evaluating the total strength described above can be used with more complex idealizations of a crack bridge taking into

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account further failure and damage mechanisms. In addition to filament rupture considered in the previous model we now include the influence of debonding between filament and matrix. The response function q ( e ; 8 ) is represented by a finite element idealization of the shear lag with a cohesive interface between the filament and matrix. In order to evaluate the integrals (Eqs. 2 and 5) a general numerical integration tool has been implemented [15] to obtain the statistical characteristics of crack bridge strength. We remark that the crack bridge model provides the possibility to study the impact of the variability in any of the model parameter(s) on the statistics of the overall bundle response (load displacement diagram), not only on the length that is used in this paper. With GA(A)defined as in the example 1, we now include the debonding of a filament from the matrix as an additional effect. Again, we keep all other parameters including the interface characteristics, constant. The resulting load-displacement diagram displayed in Fig. 3 right for Am = 1.5 shows significantly higher mean crack bridge strength (1008 N, see Fig. 3 right) than in the case of a perfect bond (648 N, see Fig. 1 left g). The reason for such an increased strength is the homogenizing effect of the debonding causing a more uniform stress distribution across the bundle (more filaments can act simultaneously before they break). We also note that there is no significant reduction in the scatter of strength: 173.99 N for perfect bond and 145.08 N with included debonding. The performance of a chain of crack bridges with and without debonding is compared for A,,,== 1.5 in the semi-logarithmic plot in Fig. 8. Due to a similar amount of scatter, the slope of the two size-effect curves is almost the same. In other words, while the local debonding improves the mean strength by introducing stress redistribution during the failure process, the decay of strength with the increasing number of cracks remains almost the same. The size effect curve is simply shifted upwards. CONCLUSIONS In this paper a statistical fiber bundle model for the evaluation of the mean performance of a crack bridged by a multifilament yam has been introduced. The paper shows the correspondence between local scatter in a crack bridge of a textile-reinforced tensile specimen and the resulting reduction of the specimen tensile strength. The presented approach provides a rough estimation and explanation of the strength reduction of tensile specimens with dry yam reinforcement with a high amount of imperfections in the bundle structure and in its bond to the cementitious matrix. The effect of the disorder in the filament bundle has been described in form of distributions of relative extra filament length. We have shown that the scatter of filament lengths leads to a reduced performance of the crack bridge. The introduction of finite bond strength has a homogenizing effect leads to an increasing strength at larger displacements. The heterogeneity of the interface layer was represented in form of a statistical distribution of bond quality. Increased scatter of the bond quality leads to reduced crack bridge strength performance. If we consider a usual range of lengths of tensile structural elements of order of magnitude 1-6 m and average crack distance of 0.01-0.02 m, the realistic range of N is 50-600. As demonstrated by the two examples, the local scatter in a crack bridge significantly affects the load bearing capacity of textile reinforced specimens and, thus, the statistical size-effect resulting from the weakest link failure must be an inherent part of dimensioning rules for the discussed type of composite. While the size effect was demonstrated with a single source of randomness, in reality the crack bridge exhibits several mutually interacting sources of randomness that have been deliberately disregarded here.

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Miroslav VO&CHOVSKk Martin KONRAD, Rostislav CHUDOBA

ACKNOWLEDGEMENTS The present work has been carried out in the framework of the project Simulation of Bond and Crack Behavior of TRC at the Meso Level included in the SFB 532: Textile reinforced concrete: foundation of a new technology sponsored by the German research foundation. The work of the first author was supported by the Grant agency of the Czech Republic under project no. GACR 103/06/P086. REFERENCES 1. Konrad, M. and Chudoba, R., The influence of disorder in multifilament yarns on the bond performance in textile reinforced concrete. Acta Polytechnica, 44, No. 5-6, 2004, pp 186-193. 2. Konrad, M., Chudoba, R. and Kang, B.-G., Numerical and experimental evaluation of damage parameters for textile reinforced concrete under cyclic loading, In ECCM-2006 - I11 European Conference on Computational Mechanics, in print, Lisbon, Portugal, 2006. 3. Chudoba, R., Vofechovslj, M. and Konrad, M., Stochastic modeling of multi-filament yarns I: Random properties within the cross section and size effect. International Journal of Solids and Structures, 43 (34): 2006, pp 413434. 4. Hegger, J., Voss, S., Bruckermann, O., Load-bearing behaviour and simulation of textile reinforced concrete. Materials and Structures, in print, 2006. 5. Schorn, H., Ein verbundmodell fi,ir glasfaserbewehrungen im beton. Bautechnik 80 (3), 2003, pp 174-180. 6. Banholzer, B., Bond Behaviour of a Multi-Filament Yarn Embedded in acementitous Matrix. Ph.D. thesis, Institute for Building Materials Research, RWTH Aachen University, 2004. 7. Lin, Z., Kanda, T. and Li, V. C., On interface property characterization and performance of fiber-reinforced cementitious composites. Concrete Science and Engineering 1, 1999, pp 173-184. 8. Li, V.C., Post-crack scaling relations for fiber reinforced cementitious composites. ASCE Journal of Materials in Civil Engineering 4 (l), 1992, pp 41-57. 9. Kanda, T. and Li, V.C., Interface property and apparent strength of a high strength hydrophilic fiber in cement matrix. Journal of Materials in Civil Engineering 10 (l), 1998, pp 5-13. 10. Phoenix, S. L., Taylor, H. M., The asymptotic strength distribution of a general fiber bundle. Advances in Applied Probability 5: 1973, pp 200-216. 11. Chudoba, R., Konrad, M., Vofechovsk9, M., and Roye, A., Numerical and experimental study of imperfections in the yarn and its bond to cementitious matrix. American Concrete Institute (ACI) Symposium Publication, 2006. 12. Vofechovslj, M., Chudoba, R., Stochastic modeling of multi-filament yams 11: Random properties over the length and size effect. Int. Journal of. Solids and Structures, 43 (34): 2006, pp 435458. 13. Smith, R. L., Phoenix, S. L., Asymptotic distributions for the failure of fibrous materials under series-parallel structure and equal load-sharing. J. Appl. Mech. 48: 1981, pp 75-82. 14. Stang, H., Li, Z. and Shah S. P., Pullout problem: Stress versus fracture mechanical approach. Journal of Engineering Mechanics 116 (lo), 1990, pp 2136-2150. 15. Konrad, M., Jefabek, J. Vofechovsky, M. and Chudoba, R., Evaluation of mean performance of cracks bridged by multi-filament yarns. In Meschke, de Borst, Mang, and Bicanic, editors, EURO-C 2006 Computational Modelling of Concrete Structures, Mayrhofen, Austria, Taylor & Francis Group, London, 2006, pp 873-880. 16. Kaliakin, V. N. and Li, J., Insight into deficiencies associated with commonly used zerothickness interface elements. Computers and Geotechnics 17 (2), 1995, pp 225-252.

Proc. Int. Symp. “Brittle Matrix Composites 8” A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

AN INVESTIGATION INTO DEICER-INDUCED ASR DISTRESS IN CONCRETE Prasada Rao RANGARAJU”, Ketan SOMPURA’), Jan OLEK2) ’) Department of Civil Engineering Clemson University Clemson, SC, U.S.A, e-mail: [email protected], [email protected] 2, School of Civil Engineering Purdue University West Lafayette, IN, U.S.A., e-mail: [email protected]

ABSTRACT Recent investigations into premature deterioration of concrete pavements in some airfields in the United States have suggested the possible role of deicing and anti-icing chemicals in inducing alkali-silica reaction (ASR) in concrete. As a result, a comprehensive research study was undertaken to investigate the role of deicing solutions in causing ASR in mortar and concrete test specimens. This paper presents the findings from the research study conducted to evaluate the influence of potassium acetate deicedanti-icer in causing alkali-silica reaction (ASR) distress in concrete specimens. A test procedure based on modifications to the standard ASTM C 1293 test method was employed in this investigation, wherein the concrete prisms were exposed to deicer solutions during the course of testing. In addition, standard ASTM C 1293 tests were conducted on the same aggregate sources to establish a reference. Expansion of the test specimens was monitored periodically, along with changes in their dynamic modulus of elasticity. pH of the deicer soak solution was also monitored. Visual and scanning electron microscopic (SEM) examinations were conducted at the conclusion o f the tests. Findings from this study indicate that potassium acetate deicer solutions caused aggressive ASR in concrete specimens containing reactive aggregates. In addition, a secondary reaction product, primarily composed of a potassium sulfate phase, was observed in prisms containing both reactive and non-reactive aggregates. Additional research is identified to decipher the precise mechanisms involved in this attack.

Keywords Alkali-Silica Reaction, ASR, deicers, anti-icers, potassium acetate, concrete pavements INTRODUCTION

The United States Federal Aviation Administration (FAA) requires the airports to have an established deicing and anti-icing program for airfield pavements to ensure safe operations in winter conditions. In this regard, a variety of deicing and anti-icing chemicals are used on airfield pavements. The common deicing and anti-icing formulations include potassium acetate, sodium acetate, sodium formate and to a lesser extent urea and glycols. Potassium formate based deicers are more widely used in European nations than in the United States. Recent investigation by the FAA into the condition of some airfield concrete pavements exposed to these new-generation deicers and anti-icers has revealed a trend of rapid and premature deterioration of concrete runways, taxiways and aprons [Scott, J., (NWFAA) Personal Communication]. In a significant number of instances where premature deterioration of concrete was observed, alkali-silica reaction (ASR) was suspected to be the primary cause leading to the distress.

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ASR is a chemical reaction that occurs between certain amorphous and/or poorly crystalline siliceous rocks and alkali hydroxides that are typically present in the pore solution of portland cement concrete. The product of the reaction - ASR gel - is hygroscopic in nature and expands by absorbing moisture. The tensile stresses generated by the expanding ASR gel result in cracking of the surrounding concrete matrix. The mechanism of ASR and other relevant aspects of this reaction have been the subject of 12 International Conferences and thousands of other publications over past 65 years, since its initial discovery by Stanton [I]. Although, several research studies have been conducted in past on the influence of deicing chemicals such as sodium chloride and calcium chloride [2-81, there is little information available in published literature on the effects of other deicing chemicals such as potassium acetate on causing or aggravating ASR in concrete. In light of the growing concerns from premature deterioration of airfield pavements across the country, coupled with the lack of knowledge on the effects of deicing and anti-icing chemicals on ASR, a comprehensive research program was initiated to investigate the effects of these deicing and anti-icing chemicals on portland cement concrete. In this research program, two test methods based on modifications to the standard ASTM C 1293 and ASTM C 1260 test procedures were adopted. Principal modification to these test procedures involved exposing the test specimens to deicer solutions, instead of the environment that was specified in the standard test procedure. The focus of this paper is on the findings from the investigation conducted to evaluate the influence of potassium acetate deicer solution on concrete test specimens using the modified ASTM C 1293 test procedure. The results of the modified ASTM C 1293 test will be compared with the results from the standard ASTM C 1293 test to assess the effects of the deicer solutions on ASR. Data on expansion, dynamic modulus and microstructure of test specimens will be presented, and possible mechanisms behind the effect of deicers will be discussed. Results of standard and modified ASTM C 1293 tests in which a typical reactive aggregate (Spratt limestone) and a typical non-reactive aggregate (IL dolomite) used will be presented. The results from the mortar bar tests using the standard and modified ASTM C 1260 procedure and other aspects of this research program are published elsewhere [9,10]. EXPERIMENTAL PROGRAM

Materials Chemicals: - In this study, a commercial grade liquid potassium acetate deicedanti-icing formulation (KAc) with a concentration of 50% by weight (6.4 molar) was used as soak solution in the modified ASTM C 1293 tests. The pH of the commercial potassium acetate deicer solution was 10.85. A reagent grade sodium hydroxide (NaOH) was used in the standard and modified ASTM C 1293 tests. Agmegates: - The non-reactive coarse aggregate (IL,) used in this study was a quarried dolomite from Illinois, U.S.A (IL). This aggregate has an established history of good field performance and has been used a reference non-reactive aggregate in other laboratory ASR studies Ell]. The reactive coarse aggregate (Spratt) used in this study was a siliceous limestone from Spratt quarry in Ontario Province of Canada. It primarily consists of calcite with minor amounts of dolomite and about 10% insoluble residue. The reactive component of the rock is reported to consist of 3% to 4% of microscopic chalcedony and black chert, which is finely dispersed in the matrix [12]. This aggregate has an established history of being alkali-silica reactive in field structures and has been used as a reference aggregate in numerous ASR studies.

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The non-reactive fine aggregate used in all the ASTM C 1293 test was the standard Ottawa sand from Illinois, U.S.A. This aggregate predominantly contains rounded quartz grains, and has been shown to be non-reactive in previous lab studies [9, 101. Cement: - Two Type I Portland cements - a high alkali cement (HA) and a low alkali cement (LA) were used in this study. The high alkali cement had an alkali content of 0.82% Na2O equivalent with an autoclave expansion of 0.12%. The low-alkali cement had an alkali content of 0.3 1% Na2O equivalent with an autoclave expansion of 0.08%. A more complete chemical composition of these cements is reported elsewhere [ 101. Test Methods Standard ASTM C 1293 Test Method: - In this test method, concrete prisms (75 mm x 75 mm x 285 mm) are prepared with the aggregate in question (coarse or fine aggregate) along with a non-reactive supplementary aggregate (fine or coarse aggregate). The cement to be used in this test is required to have a minimum alkali content of 0.9% f 0.1% NazO,, With the addition of sodium hydroxide to the mix water, the alkali content of the concrete mix is then raised to 1.25% Na20,. by mass of cement in the mix. The concrete prisms are then stored in sealed containers that maintain a 100% relative humidity. The sealed containers are then placed in a 38°C environment. Periodic length change measurements are taken up to one year. A level of expansion greater than 0.04% at one year is indicative of the reactive nature of the aggregate. Modified ASTM C 1293 Test Method: - In the modified ASTM C 1293 test method, the concrete prisms are prepared in a manner similar to the standard test method. However, instead of storing the prisms in a 100% relative humidity environment at 38"C, the prisms were soaked in deicer solution at 38°C. For tests with potassium acetate deicer, the concrete prisms were soaked in a 50% wt. solution of potassium acetate deicer (6.4M). While the concentration of the deicer solution is considerably high, it should be noted that the commercial potassium acetate deicer is applied on bare concrete pavement surface at this concentration during anti-icing operations. For the reference test specimens in the modified ASTM C 1293 procedure, 1N NaOH solution was used as the soak-solution. When using the high-alkali cement in preparing the concrete prisms for the modified ASTM C 1293 test method, the alkali content of the concrete was raised to 1.25% Na20,. by mass of cement by adding reagent grade NaOH as per the standard ASTM C 1293 standard procedure. However, when low-alkali cement was used no additional NaOH was added to the mix water. This modification produced a set of prisms, where the only significant source of alkali to trigger ASR came from external soak solution. This comparative study enabled the assessment of influence of deicer solutions on ASR. Dynamic Modulus of Elasticitv: - The physical distress in concrete prisms was quantified by measuring its dynamic modulus of elasticity (DME). The DME values were determined using the resonant frequency method based on impulse excitation technique (ASTM E 1876-01). A GrindoSonicTMinstrument was used in determining the resonant frequencies of the concrete prisms. The mass and resonant frequency of the concrete prisms was measured at the same time when the prisms were removed from the soak solution for expansion measurements. For sake of simplicity, the dimensions of the concrete prisms were assumed to be constant (i.e., 75 mm x 75 mm x 285 mm) and the effects of gage studs at the ends of the prisms were neglected as it was a common factor for all measurements and would have a relatively small effect on the dynamic modulus of elasticity results. pH of the Soak Solution: - In order to track any changes in the hydroxyl ion concentration of the soak solution, its pH was monitored. The pH measurements were made at the same time when the length-change measurements were made on the concrete prisms. The pH of the soak solution was determined by using an Oakton pH meter with an Accumet low-sodium

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error pH probe. Before every reading, the probe was calibrated using buffer solutions with a pH of 4 7 , 10 and 12.45. Scanning Electron Microscow and Enerm DisDersive X-Rav Analvsis: - Scanning electron microscopy (in back-scattered mode) and energy dispersive X-ray analysis were conducted on polished sections of concrete prisms fiom the standard and modified ASTM C 1293 tests, using a Hitachi S3400N electron microscope. The instrument was operated in a variable pressure mode at a 30Pa vacuum and an accelerating voltage of 2OKeV. The EDX was calibrated using Copper Ka peak. The samples for the SEM investigation were prepared by slicing the prisms using a slow speed diamond saw followed by polishing the samples on a series of magnetic lapping wheels with embedded diamond grit. The final polishing step involved the use of a 0.5 micron diamond paste on a cloth lap. RESULTS AND ANALYSIS In order to identify each combination of aggregate, cement and the soaking condition, a specific notation scheme has been used in this study. For example, Spratt-HA-Control refers to a test in which concrete mix containing Spratt limestone and high-alkali cement were subjected to a 100% relative humidity exposure as required in the standard ASTM C 1293 procedure. Similarly, IL-LA-KAc refers to a modified ASTM C 1293 test, wherein the concrete mixture containing IL dolomite aggregate and low-alkali cement is soaked in potassium acetate deicer soak solution. Results from Standard and Modified ASTM C 1293 Tests on Reactive Aggregate Figure 1 shows the results fiom the standard and modified ASTM C 1293 tests conducted using Spratt limestone aggregate. Also, Figure 1 shows the effect of cement alkali levels on the expansions of concrete prisms observed in the tests.

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Figure 1. Expansion of concrete prisms containing reactive aggregate in the standard ASTM C 1293 test (control) and modified ASTM C 1293 test (1N NaOH and KAc Deicer Soak Solutions). High-alkali cement series: - Based on the guidance provided in the ASTM C 33 specification for aggregates in concrete, any expansion over 0.04% at one year in the standard ASTM C

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1293 test is considered to indicate a reactive aggregate. It can be observed from the standard ASTM C 1293 data shown in Figure 1 under the high-alkali cement series that concrete prisms prepared with Spratt limestone (Spratt-HA-Control) showed expansion well over 0.04% at one year, confirming the reactivity of this aggregate. Data from modified ASTM C 1293 tests in which the concrete prisms with high-alkali cement were soaked in 1N NaOH soak solutions showed virtually identical expansion to those of the “control” prisms in the Standard ASTM C 1293 test. This finding is not surprising in that the alkali hydroxide concentrations in pore solutions of concrete prisms from both the standard ASTM C 1293 test and the modified ASTM C 1293 test with 1N NaOH soak solution can be considered to be practically identical (- 1 Normal). In contrast, the data from the modified ASTM C 1293 test in which the concrete prisms with high-alkali cement were soaked in potassium acetate deicer solution showed more expansion compared to the prisms soaked in IN NaOH solution, or the control prisms from the standard ASTM C 1293 test. The potential reasons behind this effect of the deicer solution are elaborated in the pH and the microstructure investigation section of the results. Low-alkali cement series: - The standard ASTM C 1293 test requires the use of high-alkali cement. Therefore, no control prisms were cast in this study using the low-alkali cement. Data from the modified ASTM C 1293 tests in which the Spratt limestone concrete prisms prepared with low-alkali cement and soaked in 1N NaOH solution showed expansions well over 0.04% at one year. However, the magnitude of expansion of these prisms at one year was lower than those of prisms containing high-alkali cement. Also, the rate of expansion of prisms with low-alkali cement was slower than that of prisms containing high-alkali cement. These findings clearly suggest that the expansion of concrete prisms is affected by both the alkali levels of the cement used and the contribution of alkalis from external environment. Results from the modified ASTM C 1293 tests in which the concrete prisms with low-alkali cement soaked in potassium acetate deicer solution show levels of expansions that were significantly greater than those observed for prisms soaked in 1N NaOH solution. This trend was comparable to the behavior of the Spratt limestone concrete prisms prepared with highalkali cement. These findings suggest that potassium acetate deicer solution has more aggressive effect than 1N NaOH solution on concrete prisms containing reactive aggregate, regardless of the alkali content of the cement.

Standard and Modified ASTM C 1293 Tests on Non-Reactive Aggregate: Figure 2 shows the expansion of concrete prisms prepared with IL dolomite in the standard and modified ASTM C 1293 tests. It can be observed from this figure that the control prisms (IL-HA-Control) showed an expansion of 0.025% at one year. This level of expansion is well below 0.04%, and therefore IL dolomite is considered to be a non-reactive aggregate in the standard ASTM C 1293 test.

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Figure 2. Expansion of concrete prisms containing non-reactive aggregate in the standard ASTM C 1293 test (control) and modified ASTM C 1293 test (1N NaOH and KAc Deicer Soak Solutions). High-alkali cement series: - Results of Modified ASTM C 1293 test in which concrete prisms prepared with high-alkali cement and soaked in 1N NaOH solution and potassium acetate deicer solution did not show any significant expansion. This suggests that the soak solutions did not have any negative influence on the concrete prism. Low-alkali cement series: - Results of the modified ASTM C 1293 test in which the IL dolomite concrete prisms prepared with low-alkali cement showed an unusual trend. While the concrete prisms exposed to 1N NaOH solution did not show any significant expansion at one year (0.032%), the concrete prisms soaked in potassium acetate deicer solution showed significant level of expansion (0.235%) at the end of one year. Investigation into microstructure of these concrete prisms has shown that a secondary reaction product that is rich in potassium sulfate has in-filled cracks in the concrete prisms. This information will be presented and discussed in the SEM-EDX section. Comparing the results of the modified ASTM C 1293 tests on Spratt limestone with IL dolomite, it is apparent that the potassium acetate deicer solution aggravated the level of expansion in concrete containing reactive aggregates. However, no such effect was observed in case of the non-reactive aggregate, in particular, when used with high-alkali cement. While this trend does not appear to be completely valid when low-alkali cement is used in the modified ASTM C 1293 tests with non-reactive aggregate, microstructure investigation of IL-LA-KAc prisms revealed that the expansion in the prisms was not due to ASR. Instead, certain other interactions involving potassium acetate deicer and possibly some sulfate phases present within the hydrated portland cement paste appear to be the main cause. Evidence on this aspect is presented in the microstructure investigation of this paper.

Dynamic Modulus of Elasticity @ME) Figure 3 shows the percent change in DME of Spratt limestone and IL dolomite concrete prisms subjected to the standard and modified ASTM C 1293 tests, relative to their initial DME. In addition, the influence of alkali content of the cement used in the modified ASTM C 1293 test on the percent change in DME is also shown.

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Figure 3. Percent Change in DME of Concrete Prisms Subjected to Standard and Modified ASTM C 1293 Tests It can be observed from Figure 3 that in all the standard and modified ASTM C 1293 tests involving Spratt limestone, there is an initial increase in the DME value up to 60 days, followed by a gradual drop in DME to values less than 100% at 360 days. It appears that the continued hydration of cement and the associated gain in strength and stifhess caused the initial increase in DME values. However, at later ages the distress caused by ASR resulted in significant lose of DME. This affect appears to be more apparent in prisms soaked in potassium acetate deicer solution as compared to other test environments studied. In case of prisms with IL dolomite in the standard and modified ASTM C 1260 tests, there is an initial increase in the DME values of the prisms up to 120 days, followed by a drop at later ages. However, in all the cases with exception of IL-LA-KAc, the DME of concrete prisms at one year is at or above 100% suggesting no severe deterioration in the concrete prisms. In case of IL-LA-KAc, the DME dropped to 75% at 360 days. The reason for the drop in DME values of IL dolomite prisms at later ages is not entirely evident at this point of time. However, evidence from microstructure investigation of polished specimens from E-LA-KAc shows presence of cracks around the coarse aggregate

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particles in the matrix that are in-filled with a potassium sulfate rich phase. More elaborate discussion of these findings will be presented in the section on microstructure observations. It is important to note that the expansion behavior of all the concrete prisms discussed in previous section corresponded well with the observed trends in the DME data, validating the observations. pH of Soak Solution Figure 4 shows the pH of the soak solutions in the modified ASTM C 1293 tests for Spratt and IL aggregates, before immersing the concrete prisms in the soak solution and at the end of one year after soaking the prisms. In addition, the effect of alkali content of cement on the pH of the soak solution was also assessed. The pH of the soak solutions before immersing the prisms is identified in the graph as “plain”, and the pH of the soak solution in which concrete prisms prepared with high-alkali or low-alkali cement are identified as “HA” or “LA”, respectively. It should be noted that the pH of the soak solutions was measured at 38”C, the storage temperature during the course of the tests.

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Figure 4 - pH of soak solution before and after soaking the concrete prisms in the modified ASTM C 1293 test It can be observed from the data shown in Figure 4 that in case of Spratt limestone and IL dolomite, the pH of the IN NaOH solution was identical before and after soaking the concrete prisms. The alkali content of the cement used in the concrete prisms did not have any significant influence. However, in case of potassium acetate deicer soak solution a significant jump in the pH of the soak solution was observed between plain solution and solutions in which concrete prisms prepared with HA or LA cement were stored. The alkali content of the cement in the concrete prism itself did not appear to have an influence on this observation. In a separate study it was found that the “pH jump” observed in potassium acetate deicer solution was caused because of the interaction between potassium acetate and calcium hydroxide in the hydrated cement paste [lo]. It is suspected that the high pH resulting from the interaction between potassium acetate and the calcium hydroxide is triggering deleterious ASR reaction in case of concrete prisms containing reactive aggregates such as Spratt limestone. The consequence of increase in pH of potassium acetate soak solutions on concrete prisms containing non-reactive aggregate (IL-LA-KAc and IL-HA-KAc) is not clear at this time. Although, no deleterious ASR reactions were observed in these prisms, some potassium sulfate rich phase was observed in-filling the cracks. The evidence of this phase is presented in the section on microsh-ucture observations.

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Micro-structural Observations The microstructure observations presented in this paper are limited to sections prepared from Spratt-LA- 1N NaOH, Spratt-LA-KAc, ILHA-KAc, and IL-LA-KAc tests. Figure 5A shows a picture of Spratt limestone concrete prisms soaked in 1N NaOH soak solution and Figures 5B and 5C show an SEM micrograph and EDX spectra, respectively, from the same prism. Figures 5D, 5E and 5F show a picture of Spratt limestone concrete prism soaked in potassium acetate deicer solution, an SEM micrograph and an EDX spectra of ASR gel from this specimen, respectively. Extensive cracking in the Spratt limestone prisms has occurred in both 1N NaOH and potassium acetate deicer solution. The cracks on Spratt limestone concrete prisms are multiple in numbers and form a “pattern” type cracking that is characteristic of ASR distress as seen in Fig. 5D. Cracks in Fig. 5A are not clearly visible due to the presence of a white residue on the surface of the prism. The SEM micrographs shown in Figures 5B and 5E reveal that the cracks in the prisms predominantly go through the reacted coarse aggregate particles. These cracks are either completely in-filled or lined with ASR gel. Typical compositions of the ASR gel in prisms soaked in IN NaOH and potassium acetate deicer solution are shown in Figures 5C and 5F. The ASR gel from the prisms soaked in potassium acetate deicer solution appears to have significantly higher alkali-to-silica ratio as compared to prisms soaked in 1N NaOH solution. Minor amounts of an unknown phase rich in potassium sulfate was also found intermingled with the ASR gel in cracks. Figure 6A shows a picture of IL dolomite concrete prism from the IL-LA-KAc series at the end of one year. While significant cracking in the prism can be observed, it should be noted that the nature of cracking on the IL-LA-KAc is different than that observed on Spratt-LAKAc prism. The distress on the surface of the IL dolomite prism is characterized by a single large crack, rather than a network of cracks that is typical of ASR distress. Figure 6B shows the SEM micrograph from the IL-LA-KAc prism. It is evident from this figures that the crack in the matrix traverses around the coarse aggregate particle. It can be observed in this figure that the crack is in-filled with a reaction product that is rich in a potassium sulfate phase. The composition of the reaction product is shown in Figure 6C. It is not evident at this time, whether the potassium sulfate phase was responsible for cracking or if it has merely precipitated in an existing crack. No ASR gel was detected in this prism. Figures 6D, 6E and 6F show visual and SEM images from concrete prisms prepared with IL dolomite and high-alkali cement and soaked in potassium acetate deicer solution. It is apparent from the figure that no visible cracks or micro-cracks can be observed in this prism, suggesting no deleterious influence of potassium acetate deicers in these prisms. However, small deposits of potassium sulfate phase scattered in the matrix were observed.

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Spratt-LA-KAc

Figure 5. Visual, SEM Images and EDX Spectra from Spratt Limestone Concrete Prisms Soaked in 1N NaOH and KAc Solutions.

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IL-HA-KAC

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Figure 6 . Visual, SEM Images and EDX Spectra from IL Dolomite Concrete Prisms Soaked in 1N NaOH and KAc Solutions. It should be noted that minor amounts of potassium sulfate phase were also observed in some cracks in the Spratt limestone concrete prisms soaked in potassium acetate deicer solution. However, ASR gel dominated the in-filled material in cracks of Spratt limestone prisms. The occurrence of potassium sulfate phase in both Spratt limestone and IL dolomite prisms suggests that the potassium acetate deicer is interacting with some sulfo-aluminate phases in the hydrated cement paste. However, the reason behind the excessive expansion in IL-LAKAc prisms associated with this phase is not clearly understood. Further research is needed to study the possibility of interactions between deicer solutions and other hydration products of portland cement such as ettringite and mono-sulfo-aluminate. CONCLUSIONS Based on the observations of length-change, dynamic modulus of elasticity, and microstructure of concrete prisms, and pH changes in the soak solutions in the standard and modified ASTM C 1293 tests it can be concluded that: 1. Potassium acetate deicer solution is capable of triggering deleterious alkali-silica reaction in concrete specimens that contain reactive aggregates.

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For prisms prepared with a cement of given alkalinity, exposure to potassium acetate deicer solutions caused more distress (expansion and loss in DME) than 1N NaOH solution in the modified ASTM C 1293 test. Significant increase in pH of the potassium acetate deicer solution was observed in the modified ASTM C 1293 tests. Neither the alkali content of the cement nor the type of aggregate used in the prisms had an influence in the pH rise. RECOMMENDATIONS

Before a modified ASTM C 1293 test can be adopted to assess the effect of potassium acetate deicer on its potential to cause ASR, additional research is necessary to understand the origin and effect of the potassium sulfate phase on expansion in the prisms subjected to modified ASTM C 1293 test. Also, further investigation is necessary to understand the interaction between potassium acetate deicer solution and other hydration compounds in portland cement to explain the observed phenomenon in this study. ACKNOWLEDGEMENTS

The support of Innovative Pavements Research Foundation (IPRF)and FAA for this project is gratefully acknowledged. REFERENCES

1. Stanton, T.E. “Expansion of Concrete Through Reactions Between Cement and Aggregate” Proc. ofAmerican Society of Civil Engineers, Vol. 66, 1940, pp. 1781-1 81 1. 2. Diamond, S. “Alkali-Reactions in Concrete-Pore Solution Effects” Proc. of 6th Intl. Con$ on Alkali-Aggregate Reactions, Copenhagen, Denmark, Steen Rostam, 1983, p. 155-167. 3. Chatterji, S., Thaulow, N., Jensen A.D. “Studies of ASR: Part 4”, Cement and Concrete Research, Vol. 17, 1987, p. 777-783. 4. Nixon, P.J., Page C.L., Canham, I., and Bollinghaus, R. “Influence of Sodium Chloride on ASR” Advances in Cement Research, Vol. 1, 1988, p. 99-105. 5. Kawamura, M. and Ichise, M. “Characteristics of ASR in Presence of Sodium and Calcium Chloride” Cement and Concrete Research, Vol. 20, No. 5, 1990, p. 757-766. 6. Duchesne, J., and Berube M.A. “Effect of Deicing Salt and Sea Water on ASR: New Considerations Based on Experimental Data” Proceedings of 10th International Conference on Alkali-AggregateReactions, Melbourne, Australia, 1996, p. 19-23. 7. Sibbick, R.G., and Page C.L. “Effects of Sodium Chloride in Alkali Silica Reactions in Hardened Concretes” Proceedings of 10th International Conference on Alkali Aggregate Reactions, Melbourne, Australia, 1996, p. 822-829. 8. Berube M.A. and Frenette, J. “Testing Concrete for AAR in NaOH and NaCl Solutions at 38°C and 80°C”, Cement and Concrete Composites, Vol. 16, No. 3, 1994, p. 189-198. 9. Rangaraju, P., Sompura K., Olek, J., Diamond S., Love11 J. “Potential for Development of Alkali-Silica Reaction in Concrete in the Presence of Airfield Deicing Chemicals” Proceedings of the 8’hInternational Conference on Concrete Pavements, Colorado Springs, U.S.A, August 14-18, pp. 1269-1288,2005. 10. Rangaraju, P. “Potential for ASR in Concrete in Presence of Airfield Deicing Chemicals” 60% Review Report, IPRF,September 2005, p. 60 (accessed at http://.iprf.org/reportsJune 30,2006)

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11. Touma E. T., Fowler, D.W., Carasquillo R.L. Alkali-Silica Reaction in Portland Cement Concrete: Testing Methods and Mitigation Alternatives, Technical Report ICAR 301-1F, International Center for Aggregate Research, University of Texas, Austin, p. 520,2001. 12. Rogers, C. Multi-Laboratory Study of Accelerated Mortar Bar Test (ASTM C 1260) for Alkali-Silica Reaction, Cement, Concrete anddggregates, CCAGDP, 21(2), 1999, p.191200

Proc. Int. Symp. “Brittle Matrix Composites 8” A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

EXPERIMENTALAND NUMERICAL STUDY OF FROST SALT SCALING OF CONCRETE OguZhan COPUROGLU & Erik SCHLANGEN Delft University of Technology, CITG, Microlab, P.O. Box 5048,2600 GA Delft, The Netherlands, e-mail: [email protected]

ABSTRACT This paper discusses the deterioration in cement based materials due to frost salt scaling. Both experiments and numerical modeling are used to be able to explain the mechanism behind frost salt scaling. The basic mechanism is believed to be the shrinkage and cracking of a thin ice layer on top of the concrete surface. Parameters that are varied in the investigation are the salt concentration in the water layer and ice-layer thickness on the surface. A lattice type model is used to simulate the mechanism in which the material structure is implemented using digital images of the real material. Both experiments and the simulation with the model show that the amount of scaling increases with increasing thickness of the ice layer on the surface. Furthermore it is shown that with the model the well known pessimum effect for salt concentration in the water (most damage at 3% salt) can be reproduced.

Keywords Frost salt scaling, lattice model, fracture, experimental study, ice layer thickness INTRODUCTION The modelling of frost salt scaling (FSS) has been a difficult issue due to its complex physical and chemical mechanisms [ 11. Plain frost action has attracted relatively more attention and thanks to its less complicated mechanism, we increased our knowledge during past couple of decades. The works of Powers, Litvan, Fagerlund, Setzer and many other researchers have drawn the frame of the issue substantially [2]. Unfortunately, similar arguments could not be used for frost salt attack. There have been a number of questions, which could not be answered by a single theory. Due to having insufficient knowledge on this phenomenon, modelling attempts have been restricted to black-box type. However, recently a successful theoretical explanation on the mechanism of frost salt attack was introduced by Valenza and Scherer [3]. The authors propose the mechanism called “glue-spall”. According to this theory the cracking of the icehrine layer is the origin of FSS. They put forward a theoretical explanation for the greater damage of pessimum salt concentration under frost. The principle idea is that following the ice formation on top of the concrete surface, ice starts to shrink due to further cooling. The shrinkage of ice exerts tensile stresses in the ice and causes three consequences depending on the solute concentration of the liquid. These are: 1. Weak salt concentration (0.1%): Due to the ice formation and further cooling of the ice, the exerted tensile stress cannot exceed the tensile strength of ice, so no cracking occurs. 2. Pessimum salt concentration (1-3%): Due to the ice formation and further cooling of the ice, the exerted tensile stress exceeds the tensile strength of ice and breaks the ice, which triggers surface scaling.

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3. Strong salt concentration (10-20%): In this case the ice layer is too soft to exert enough stress to the underlying cementitious material, so no scaling occurs. In the present paper the glue-spa11 theory is simulated with a lattice model. The basics of the model are explained in the next section of the paper. With the model several phenomena that play a role in FSS are studied. The first one is the thickness of the ice layer which is believed to have an important effect on the amount of scaling. Also experiments are performed to study the effect of ice layer thickness. The next parameter that will be discussed in this paper is the pessimum effect. Also here both experimental and numerical simulations are presented. The paper will end with some discussion and concluding remarks.

BASICS OF THE LATTICE MODEL

Heterogeneous materials have complicated fracture mechanisms, which are related to their microstructure. The use of linear elastic fracture mechanics to analytically describe these mechanisms is very hard, since fracture patterns consist of a main crack, with various branches, secondary cracks and micro cracks. Lattice type models are used by theoretical physicists [4] to model fracture mechanisms in various materials. To model concrete fracture these models were introduced by Schlangen and van Mier [ 5 ] . Lattice models are now used quite a lot to model concrete crack patterns, mainly because the simulated cracks are very realistic and resemble to a great detail the cracks observed in laboratory tests and in practice. [5-91. In these models a material is discretized as a lattice consisting of small beam (or spring) elements that transfer forces, as can be seen in figure 1.

Figure 1: a. Lattice of beams, b. definition of forces and degrees of freedom, c. stress strain relation of beam element. The meshes used in this paper are 2D regular triangular meshes. Other options like random meshes are sometimes preferable, see [6]. Each of the beams in the lattice can transfer, in general, normal force (F), shear force (Q) and bending moment (M). The relation among these forces and corresponding displacements for the endpoints (i and j in Figure 1) of a beam can be expressed as follows: EA F=-( ui - Uj) '

I 12EZ

6EZ

Q,= 7 (vi - vj) --(4 I2 6EZ Mi =-(vj-vi)+-(&--)

I2

- 4j)

4EZ

4j

I

2

389

Experimental and numerical study offrost salt scaling of concrete

in which E is Young’s modulus, 1 is the length, A is the cross-sectional area and I is the moment of inertia of a beam element, (u, v) are the translational displacements and @ is the nodal rotation. For a lattice with a regular geometry, the quantities E, A, 1 and I are in principle, equal for all elements. However, these parameters can be varied, either element by element or according to a superimposed microstructure, in order to implement heterogeneity. To construct the system of equations for the complete lattice, each element matrix has to be multiplied by the appropriate rotation matrix and positioned correctly in the system. The final set of equations for the system is of the form: b=Ax

(4)

in which b is the force vector, A is the stiffness matrix and x is the displacement vector. If there are N nodes in the system, then b and x are of length 3N and in general a 3Nx3N matrix. When solving the set of linear elastic equations for a lattice under an applied load, the load vector and stiffness matrix are known and the displacement vector is to be determined by solving eq. 4. One method to solve the set of equations is to use a direct solver which finds the inverse of A by Gaussian elimination. In the simulations presented in this paper the displacement vector x is solved iteratively by minimizing the functional G, which has the dimensions of energy, G = O.SXAX- bx

(5)

The loading in case of the simulation of frost salt scaling is the shrinkage of the ice. The elements in the mesh representing the ice undergo a predefined dilation which is transferred to prescribed nodal forces in the nodes of the elements [lo]. The simulation of fracture proceeds as follows. A linear elastic analysis of the lattice is performed. If the stress in one of the beams exceeds the material strength this beam virtually breaks and is removed from the lattice. Then the stresses in all the beams are recalculated and compared again with their strength, which may result in the next beam to break. SIMULATION OF FROST SALT SCALING MECHANISM A layer of ice forms on top of the material, then the ice shrinks and cracks as the temperature drops. Due to the cracks in the ice, shear stresses are created which cause the concrete to crack and pieces of material to scale off. This is modelled with the lattice model as shown in Figure 2. An arbitrary layer of 3 mm ice on top of 12 mm mortar is modelled. In the horizontal direction periodic boundary conditions are assumed. At the bottom and at the top the nodes are free. The properties in Table 1 are used. Table 1 Material properties used in lattice simulations of concrete with ice layer Material E-modulus [GPa] Tensile Strength [MPa] Aggregate 70 10 ~

Interface

10

~

1

~~

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COPUROGLLU,Erik S C H M G E N

~

Figure 2: Simulation of a FSS action by Delft Lattice Model. The analysis is performed linear elastic. It is assumed that the temperature drops -20 "C and that the difference in thermal dilation between the concrete and the ice is 4.105. In the analysis actually no relaxation of the stresses is used. The crack in the ice, however, starts already at 10% of the applied strain, or in other words at a temperature change of only -2 "C. In practice a lot of the linear elastic stress will disappear due to relaxation. This is also described by Valenza and Scherer [3]. Due to this relaxation the generated stresses for the complete temperature drop of -20 "C will not be enough to crack pure ice, but will be enough to crack brine, as will be further discussed in this paper. The crack continues in the mortar and kinks at a certain distance inside the material. Afterwards it continues until it hits the ice again, and a second crack in the ice is formed. Further continuation of the simulation would create another crack in the ice at another location, and the mechanism would repeat again. Of course also here the mechanism is influenced by the choice of local material properties and relaxation of stresses. However the simulation so far looks promising. INFLUENCE OF ICE LAYER THICKNESS

The effect of ice layer thickness during the frost salt scaling testing is generally prescribed by the standard test methods such as CDF, Scandinavian Slab Test and ASTM C672. In CDF test, the specimen is soaked into the freezing salt solution and the specimens are exposed to 10 mm thickness of ice. In other tests, the thickness of fieezing test liquid is prescribed as 3 mm and 6 mm for Scandinavian Slab Test and ASTM C672 respectively. As can be seen in the standards, there is no commonly used depth of salt solution. This brings the risk of having substantially different experimental results with the different test methods. The authors believe that the thickness of the layer of ice in contact with the testing surface is extremely important since thicker ice formation in contact with material surface would cause more mass scaling when the glue-spa11 action is the main mechanism. After the ice has cracked the ice shrinks further and introduces stresses in the concrete. A thicker icelayer simply generates more force and thus higher stresses in the concrete surface, which should result in more scaling. To study this hypothesis experiments and simulations are performed as discussed below. Experimental study on ice layer thickness A standard (w/c 0.50) mortar mixture with CEM I11 /B 42,5 N HSR LH was prepared according to EN 196. The physical and chemical properties of the cement are presented in Table 2. The reason for using mortar specimens in this study is to minimize the surface heterogeneity and the specimen-to-specimen variation. The mixture was then poured into the

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18~13.5cm cylindrical plastic moulds to the thickness of 35 mm. Then the buckets were vibrated identically for 10 seconds. Finally polyethylene sheets with the diameter of 18 cm were placed on the top surface. Air pores on the specimen surfaces were scraped out by a plastic spatula to obtain a reasonably flat testing surface.

CaO,%

Table 2 Physical and chemical properties of CEM IIIlB 42.5 N HSR LH Physical Strength of standard mortars 44.9 2 days, N / m Z 11.9

SO,,,% SiO2, %

3.38 27.7

7 days, N/mmz 28 days, N / m Z

34.1 53.5

A1203. %

12.0

Blaine, mZkg

376

Chemical

The specimens were cured with water for 1 week and exposed to 3% accelerated carbonation for an additional one week. The aim was to further weaken the surface in order to achieve magnified damage which would make the evaluation and the comparison of the scaling results easier for various ice thicknesses. Finally, prior to the scaling test, the samples were saturated with 3% NaCl solution for 2 weeks. The effect of four different ice thicknesses, 1 mm, 3 mm, 5 mm and 10 mm, was investigated. The freezing thawing cycles of 17&1 hrs freezing at -2&2 "C and 7*1 hrs thawing at 20*2 OC were maintained. 3 cycles were completed for each series and scaled material was collected by filtration after each cycle. Scaled materials were placed in the stove at 110°C and weighed after 24 hrs of drying. As presented in Figure 3, the experimental study revealed the effect of ice thickness on the scaling of the underlying cementitious material. It was clearly observed that for the identical freezing-thawing cycle, salt concentration of freezing liquid and the material properties, higher thickness of ice causes more damage on the material surface. 1 mm of initial depth of freezing liquid gave 74 grams of cumulative mass scaling after 3 cycles while 10 mm caused the scaling of 103 grams. It should be stressed that a 9 mm increase in the initial freezing liquid thickness causes 40% more mass scaling after only three freezing-thawing cycles.

70

I 0

2

4

6

8

10

12

Ice layerthlckness (mm)

Figure 3 Cumulative mass scaling of the identical samples with various ice layer thicknesses (after 3 freezing-thawing cycles).

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It was observed that 10 mm ice layer thickness generates mass scaling with significant number of larger flakes in comparison to the one with 1 mm ice thickness. The typical scaled off particles can be seen in Figure 4.

Figure 4 Typical scaled-off flakes from the mortar specimens having 1 mm (left) and 10 mm (right) thick ice layers on top. Simulation of effect of ice layer thickness In this modelling study, the micromechanicalparameters in Table 1 were used. It was possible to model the effect of ice thickness as well as simulating the damage pattern occurred under the effect of different ice thicknesses. It should be noted that here the mechanical properties of ice represent a plain ice. However the damage pattern depends significantly on the mechanical properties of the material. Hence, the aim is to show the effect of ice thickness rather than to demonstrate a realistic frost salt scaling for the condition given in Table 1. In Figure 5 , the effect of various ice layer thicknesses on the frost salt scaling of the identical mortar specimens is shown. The results imply that apparently the new numerical model is able to simulate the experimental observations. Indeed according to the model, an increase in the depth of ice results severer surface damage under frost salt attack. It might be noteworthy to mention that no damage is observed for 1 mm ice thickness of ice in the modelling results. This is -as mentioned before- simply because of a mismatch between the micromechanical properties of the specimens used in the experimental study and the ones used in the numerical modelling. It is quite possible that the micromechanical values used in the model for matrix and interfacial zones were lower than that of real condition. Another reason could be that the visco-elastic behaviour assumption for ice and mortar is different. The damage pattern created by the different thickness of ice could also be simulated by the model. The size of the scaled flakes in the experiments and the ones generated by the model are observed to be analogues for the corresponding ice thicknesses.

Experimental and numerical study of frost salt scaling of concmte

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Figure 5. Effect of ice layer thickness ( 1 , 3 , 5 and 10 mm) on the frost salt scaling damage magnitude.

INFLUENCE OF SALT CONCENTRATION It has been known for almost four decades that a pessimum salt concentration causes severest surface scaling damage [ 121. However, until recently no researcher succeeded to explain the mechanism behind this phenomenon. Valenza and Scherer [3] proved with the glue-spa11 theory that 1-3% of solute causes ice cracking during freezing which is the main source of surface scaling. In this section, the aim is to present this very well known pessimum effect by experimental and modelling techniques in order to further validate the accuracy of the glue-spa11 theory and its compatibility with the Delft Lattice Model.

Experimental study on pessimum effect The mortar specimens with exact specifications as in the previous paragraph were used in this experimental study. After the carbonation period the samples were saturated by salt solutions of 0%, 1%, 3%, 7% and 12% for two weeks. Frost salt scaling testing was conducted by the corresponding salt solutions with 3 mm thickness. The experimental results seconded the results provided by the previous researchers. In Figure 6 the surface scaling after three freezing-thawing cycles from the experimental study is provided. In the same figure the findings of Verbeck [7] are also presented for comparison. It was noticed that indeed 3% NaCl concentration causes the severest damage compared to the other NaCl concentrations.

3 94

O&uhan cOPUROeL:Lu,Erik SCHLANGEN

I

100,

Qo 80

70

-

60

50

i 4 0 30 20 10 0 0%

2%

4%

6%

8%

10%

12%

14%

salt concentrafion

Figure 6. Effect of NaCl concentration on the frost salt scaling of carbonated slag cement mortars. The small figure is from [12]. Vertical axis denotes the visual inspection rates (1: no scaling, 3: moderate scaling, 5: severe scaling). Modelling of the pessimum effect The aim in this study was mainly to simulate the observations from the experimental study rather than achieving a numerical mass scaling result. In this modelling study the tensile strength of the interfacial transition zones kept very low (0.1Mpa) in order to magnify the scaling damage and enable to differentiate the damage level easily. Another reason for keeping the tensile strength so low is to compensate the internal damage due to the ice formation [1 11 which is not modelled in this study. The mechanical parameters of NaCl ice are different for various solute concentrates. Elastic modulus and tensile strength of the ice at different concentrations were estimated by the following equations provided by Valenza and Scherer [3]; (MPa) = 2.47 - 5 . 1 d ( l - Uj) Oi E = Ei( 2.85 - 1.850,

(6)

(7)

Where ui is the volume fraction of ice, Ei is the elasticity modulus of pure ice, E is the elasticity modulus of NaCl ice and ot is the tensile strength of NaCl ice. The calculated mechanical values of NaCl ice having different salt concentrations are given in Table 3.

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Table 3 Mechanical properties of NaCl ice at -20°C (as used in the model). Tensile Strength (MPa) Solute concentration Elasticity modulus (MPa) 0% NaCl 10.000 2.5 1% NaCl 8460 1.21 3% NaCl 6650 0.48 7% NaCl 4160 0.05 12%NaCl 2230 0.01 The numerical modelling results show that the new numerical model is able to simulate the experimental observations regarding the pessimum salt effect (see Figure 7). According to the assumptions for the mechanical properties, 3% NaCl concentration creates maximum surface damage compared to the other solute concentrations. It was observed that at 1% salt concentration, generated stress was not able to crack the ice completely. This caused an internal damage but this damage is not exactly a mass scaling as can be seen at 3% NaCl. In case of 7% and 12% NaCl solutions, the generated stress was just enough to crack the ice layer but not able to penetrate into the material skin, as also suggested by Valenza and Scherer [31.

1%

3%

7%

12%

Figure 7. Numerical modelling of the pessimum salt concentration effect on the frost salt scaling damage. CONCLUDING REMARKS A new frost salt scaling modelling is introduced based on Delft Lattice Model. The model can demonstrate the frost salt scaling damage in cement based materials in accordance with previously proposed glue-spa11 theory [3]: shrinkage of the ice layer which cools down further, cracks and introduces stresses in the material that lead to scaling. Apparently the new numerical model is able to simulate the frost salt scaling damage as well as the symptoms observed during the experimental observations. The authors believe that this model and its theory fill a gap in understanding the frost salt scaling damage. However it should be strongly stressed that this model (and eventually the others) has to be utilized with realistic ice and material micro properties in order to achieve realistic results. In

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this study the main aim was to evaluate the accuracy of the glue-spa11 theory and the new numerical model by the experimental results. In [113 it is shown that freezing of water in internal pores can cause internal cracking, which lower the mechanical properties of the cement based material. This will increase the probability for scaling. In [13] the effect of carbonation on the change of the mechanical properties is investigated for different cement types. Here it is shown that carbonation changes the microstructure and its mechanical properties. For slag cement pastes carbonation reduces the frost salt scaling resistance, while for ordinary portland cement pastes carbonation has a positive effect. The study also suggest that the thickness of ice layer is extremely important to get consistent results. Although, the test standards prescribe the freezing liquid thickness, extra attention should be paid on this aspect during the experimental studies. An other important aspect which is related to this is the variation of the thickness of the ice layer. This variation is actually the roughness of the surface of the cement based material. This roughness determined the bond strength between ice and surface. Furthermore if the surface has a high roughness additional peak stresses will be introduced in the ice which leads to a higher probability for the ice to crack and thus a higher chance for scaling. Note that the roughness of the surface increases with increasing number of freeze-thaw cycles. Also surface cracking and ice formation inside these cracks will increase the scaling in every new freeze-thaw cycle [ 141. In general it can be concluded that: - if the thickness of the ice layer increases, the force that the ice exerts on the cement based surface also increases, which results in more scaling; - if the stiffness of the ice layer is higher, the force that the ice exerts on the cement based surface increases, which results in more scaling. However, for the latter is has to be kept in mind that the strength of the ice should not be too high, since the ice can only exert stresses afier it has cracked. This is the reason why pure ice generally does not lead to scaling for temperatures down to -20 "C. If the temperature will decrease further also pure ice will crack. REFERENCES 1. Copuroglu, O., Frost Salt Scaling of Cement-Based Materials with a High Slag Content. PhD-thesis, Delft University of Technology, The Netherlands, 2006, pp 188 2. Lindmark, S., Mechanisms of Salt Frost Scaling of Portland Cement Bound Materials: Studies and Hypothesis, Lund University 266 PhD 1999. 3. Valenza, J. J. & Scherer., G.W. (2004). Mechanism for Salt Scaling of a Cementitious Surface. FULEM Spring Meeting 2004, Chicago, USA. 4. Herrmann, H.J. & Row, S. eds, Statistical models for the fracture of disordered media, Elsevier/ North Holland, Amsterdam, 1990. 5. Schlangen, E. & van Mier, J.G.M., Experimental and numerical analysis of micromechanisms of fracture of cement-based composites, Cem. Conc. Composites, 14 (1992) 105-118. 6. Schlangen E, Garboczi E.J., Fracture simulations of concrete using lattice models: computational aspects. Engineering Fracture Mechanics, 1997, Vol57 (213). 7. Jirasek, M. & BaZant, Z.P., Macroscopic fracture characteristics of random particle systems, Int. J. Fracture 69, (1995) 201-228. 8. Bolander, J.E. & Saito, S., Fracture analysis using spring networks with random geometry, Engineering Fracture Mechanics, 61,5-6 (1998) 569-591.

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9. Ince, R., Arslan, A. & Karihaloo, B.L., Lattice modelling of size effect in concrete strength, Eng. Fracture Mech. 70 (2003) 2307-2320 10. Schlangen, E, Koenders, E.A.B. & van Breugel, K., Influence of internal dilation on the fracture behaviour of multiphase materials, Engineering Fracture Mechanics, 2006, Vol 73 (12). 11. Copuroglu, O., Schlangen, E., van Breugel, K., & Fraaij, A.L.A. ,In ICCRRR 2005, Cape Town, South Africa, 2005, pp. 127-133. 12. Verbeck, K. J. a. K., P. "Studies of Salt Scaling of Concrete." Highway Research Bulletin Bull. 150, 1957. 13. Copuroglu, O., Schlangen, E., & Zhu, W., Effect of Carbonation on the Micro-mechanical Properties and Frost Salt Scaling of Cement Pastes: Experimental and Modelling Aspects. In Proceedings "Advances in Concrete Through Science and Engineering", Quebec City (Canada), September 11-13,2006. 14. Valenza, J.J. & Scherer, G.W., Mechanism for Salt Scaling, J. Am. Ceram. SOC.89 [4] 1161-1179,2006.

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and WoodheadPubl., Warsaw 2006

STABILITY OF HYDROSULFOALUMINOSILJCATE COMPOUNDS AND DURABILITY OF AN ARTIFICIAL STONE BASED ON THEM Ekaterina PUSHKAROVA, Vladimir GOTS, Olga GONCHAR V.D. Glukhovsky Scientific Research Institute for Binders and Materials Kiev, Ukraine, e-mail: [email protected]

ABSTRACT The problem of stabilization of structure and features of hydrosulphoaluminate and hydrosulphoaluminosilicate compounds is very novel as it makes for a wide range of waste materials involved to syntheses of the building materials. Coal ashes of different composition and made by different technologies may be successfully used in this process. In this respect, type of additive chosen depends on chemical and mineralogical composition of a waste material used. As have shown carried out earlier researches of binding compositions on the basis of fly ash of bed fluidized combustion and Portland cement, ettringite, a main product of hydrosulphoaluminate binding systems, may be stabilized by directed synthesis of solid solutions based on this phase and subsequent synthesis of hydrogarnet. Additionally, epistilbite-type calcium silicate hydrates were also founded. Stabilization of ettringite and formation of solid solutions based on ettringite by using silica additives such as silica fume and metakaolin was confirmed by SEM-EDX data. Synthesis of above-mentioned solid solutions prevents formation of secondary ettringite making for high service properties such as quick strengthening in all the ages of hardening, high frost resistance and resistance to weathering, corrosion stability etc. There were two binding systems studied: 55-85 %O of fly ash ofbed fluidized combustion and 15-45 % of ordinary Portland cement; (i) 80% of type F fly ash, 20% of ordinary Portland cement, modified with above-mentioned sulfate and (ii) silica additives. For this system, optimal type and composition of additives were determined. Binding systems developed may be applied in hydraulic and road concretes as well as for special application.

Keywords Fly ash of bed fluidized combustion, Solid solutions based on ettringite, Hydrogarnet, Ashcement binding system. INTRODUCTION At the present stage of development of a building industry all over the world the question concerning Portland cement replacement on the mixed and composite cement is actual and it is connected both with economic, and with environmental problems of Portland cement industry. Application of the ash-cement systems activated by sulfate additives, at rather small charges of Portland cement, allows receiving binding compositions which physicalmechanical properties differ from Portland cement binding systems properties a little. At the same time it is known, that ash-cement-sulfate systems are characterized by low parameters of strength at early stages of hardening, and also it is opened a question concerning their durability [1,2] as one of the basic structural elements of the received artificial stone is hydrosulfoaluminate of calcium - ettringite. Its composition, crystals habit, stability of

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Ekaterina PUSHKAROVA, Vladimix GOTS, Olga GONCHAR

structure is caused with properties of an artificial stone which is formed. Formation of ettringite at early stages of hydration in the form of needle crystals promotes creation of primary skeleton of a cement stone structure and increases its initial strength. However after full chemical linkage of gypsum and at presence of aluminates ions in a liquid phase of system occurs its recrystallization in the monosulfatic form. This connected with decrease in strength as a result of phase transformations happen in already hardened cement stone and is accompanied by destructive processes [3]. According to [4] the stability of ettringite crystals substantially depends on modification of initial sulfate of calcium. Unequal thermodynamic stability of crystal structures of hydrosulfoaluminate of calcium received from different kinds of gypsum (gypsum dehydrate, a-and B-semi-hydrate gypsum, anhydrite), defines by different conditions of ettringite crystallization. That is caused by different solubility of noted kinds of gypsum. Crystals of ettringite, received from modification a -and B semi-hydrate gypsum in 28 day find out the tendency to recrystallization and to destruction. This process is promoted by defective crystal structure of calcium hydrosulfoaluminate which results from the increased solubility semihydrate gypsum at high glut. The most stable are crystals of ettringite, which are formed on the basis of gypsum dihydrate and anhydrite. The lowered solubility and speed of dissolution of these substances (compared with semi-hydrated ones) promotes slow recrystallization of big thermodinamic stable crystals in hydrosulfoaluminate of calcium. M.Goria and M.Appiano [5] researched the process of sulfate- aluminate - plasticizer cement hydration and have defined the existence of hydrosulfoaluminate of calcium with presence of silicic acid. Thus, silica additive promotes stable existence of ettringite and its derivatives. Also on a role of silica component for increase of stability of ettringite that is formed at hydration of composite cement, pays attention in works [6-81. But, obviously, the effect from introduction of such additives will depend on their composition and structure. The purpose of work is to research an opportunity to receive a stable stone on the basis of ash filled binding systems modified by silica or sulfate -silica additives which process of hardening provides the directed synthesis in structure of new formation of hydroalumosilicate phases in the form of solid solutions.

MATERIALS AND METHOD OF TESTING As raw materials are applied Portland cement CEM-I, ashes of hydroremoval from Tripolsk energy plant, fly ashes of power station "Zheran", gypsum stone, burning at temperature 45OoC, wastes of ferrosilicon production - silica fume and metakaolin from Gluhovets deposits, received at temperature 800°C. XRD patterns of initial components are presented on fig. 1. For carrying out the researches was used fly ash of bed fluidized combustion from "Zheran" (Poland) power stations which have a following chemical compound: mas. of % SiOz-42.1; F e z0 ~ 7. 4;AlzO3-21.2.; Mn304-0.12; Ti02-0.81; Ca0-13.9; MgO-3.02; so2-7.2; P205-0.23; Na20-1.47; K20-2.13. According to data of X-ray analysis (Fig. 1, curve 1) the fly ash of bed fluidized combustion of power station "Zeran" is presented by and quartz P-SiOZ, anhydrite P-CaS04, oxide of calcium CaO, magnetite Fez03, and also roentgenoamorphous of aluminosilicate phase obtained at dehydration of minerals such as illite, kaolinite, montmorillonite. Ashes of hydroremoval from Tripolsk energy plant according to data XRD presented by glass-like substance which contains a quantity j3-quartz (d =0,430; 0,334; 0,213 nm) and sillimanite A1203.Si02 (d = 0,3364; 0,269; 0,253; 0,220 nm) (Fig. 1, curve 4).

Stabiliiy of hydrosulfoaluminosilicatecompounds and durabiliiy of an artificial stone based on them 40 1

The mineralogical composition of Portland cement (Fig. 1, curve 2) is presented basically by threecalcium silicate 3CaO.SiOz (d = 0,294; 0,274; 0,193; 0,177 nm), threecalcium aluminate 3CaO.Al203 (d = 0,332; 0,304; 0,283; 0,155 nm) and twocalcium silicate 2CaO.SiOz (d = 0,273; 0,260; 0,218; 0,162 nm). As the silica additive is used silica fume which has a specific surface from 2000 up to 3000 m’kg, bulk density - 200...250 kg/m3 and true density - 2200 kg/m3. According to data XRD in composition of silica fume the insignificant quantity of 8-quartz is fixed (d = 0,429; 0,335; 0,283; 0,252; 0,199; 0,182; 0,163; 0,154 nm), oxide of silicon in composition of silica fume have amorphous structure which explains its hydraulic activity at interaction with components of binding substances (Fig. 1, curve 3). As the variant of silica additive is used metakaolin, received after dehydration of kaolin from Gluhovets deposits of Vinnitsa area, under temperature 800°C. Phase structure of metakaolin according to data XRD presented by kaolinit relicts (peaks of low intensity d = 0,712; 0,357; 0,261 nm), 8-quartz (d = 0,424; 0,334 nm) and roentgenamorphous substance. For manufacturing of concretes and solutions on the basis of ash-cement binding compositions modified by sulfate and by silica additives, applied granite aggregate fraction of 5-10 mm, Dniper quartz sand with the gradation factor F, = 1,15. Binding compositions were prepared with compatible grinding in a ball mill of necessary components during 1,5 hours. The specific surface of the received mix made 600 m’kg. Physical-mechanical characteristics of binding substances and concrete on their basis defined according to normative documents of Ukraine, and phase composition of new formations researched by using of complex physical-chemical methods which include X-ray diffraction and scanning electron microscopy with microprobe analysis.

I 60.

50

40

30

20

10

Fig. 1. X-ray grams of initial components: 1 - fly ash of bed fluidized combustion; 2- Portland cement; 3- silica fume; 4 - ,ash of hydroremoval from Tripolsk energy plant RESULTS AND DISCUSSION

Complex research of chemical and mineralogical composition of fly ash of bed fluidized combustion which shows simultaneously pozzolanic (as a result of alumosilicate component presence) and hydraulic properties (due to presence of anhydrite), has allowed to put forward

402

Ekaterina PUSHKAROVA, Vladimir GOTS, Olga GONCHAR

a hypothesis about an opportunity of reception on its basis (at use of optimum quantity of Portland cement) stable artificial stone. For this ash-cement binding system the composition of products hydration is presented by low basic hydrosilicate of calcium and also by solid solutions on the basis of ettringite. At carrying out researches and an establishment of the greatest possible quantity of fly ash of bed fluidized combustion which use will allow to receive binding compositions with service properties which do not differ from properties of an artificial stone on the basis of Portland cement, and are characterized by stability in time, two systems which contained different quantity of ashes, - 45 and 85% (composition 1 and 3 at table 1) have been tested. The specified systems for prevention of processes of hydrosulfoaluminate phase’s recrystallization in the monosulfatic form have been modified by 5% of silica fume, used on replacement of ashes [9]. The analysis of strength development of ash-cement binding systems allow to note changes of stability, and modification of the specified systems by the silica fume additive provides a gain of strength of an artificial stone within 28 days at the used both 45%, and 85% of ashes (table 1). Table 1. The strength development of ash-cement binding substances received with use of fly ash of bed fluidized combustion Commessive strength. MPa, in age Initial raw components Bending strength, MPa

However at later stages (1 year and 7 years) are observed recession of strength in systems with the significant maintenance of a technogenic component (85%), characteristic both for the modified compositions, and for compositions without additives. Noted tendency is a characteristic not only for binding substances, but also for concrete which have been made of the developed systems (table 2). It is authentically caused by insufficient density formed artificial stone and gradual destructions of its structure as a result of the destructive processes connected as with recrystallization of hydrated new formation in time, and with gradual hydration of fly ash of bed fluidized combustion which contains P-CaS04, with formation at late stages of hydration (in case of shortage of Portland cement) monosulphatic hydrosulfoaluminate of calcium. At the same time compositions with the contents of 45% fly ash of bed fluidized combustion are characterized by a stable set of strength at all stages of hardening, the condition for this is made by processes of structurization and phase composition of new formation, including presence of an optimum parity between Portland cement and ashes when at presence concerning a lot of Portland cement, than in the first case, takes place accelerated in time and uniform on volume hydration of ashes with overwhelming formation of low basic hydrosilicate of calcium, ettringite and solid solutions on its basis.

Stability of hydrosulfoaluminosilicatecompounds and durability of an artificial stone based on them 403

Table 2. The strength gain of the developed compositions of concretes on the basis of modified ash-cement binding substances received with use of fly ash of bed fluidized combustion

In accordance with data of X-ray and thermal analysis at early stage of hardening (7 day) the composition of new formation is presented by a portlandite Ca(0H)z (d = 0.193; 0.262; 0.31 1; 0.493 nm), ettringite (d = 0.498; 0.324; 0.277; 0.261; 0.194 nm), CSH(1) (d = 0.307 nm) (Fig. 2. curve.1). At usage of a XRD under small angles (in a condition Cu K a 30/30) in the composition of hydration products it is possible to identify residuum of an ettringite, thus the diffraction reflections of a portlandite Ca(0H)z is missing (fig. 3, curve 1). After 1 year hardening of specimens (according to data of XRD) the composition of new formations is presented by a hydrogehlenite CzASHs (d = 0,261; 0,249; 0,244; 0,238; 0,236; 0,212; 0,207; 0,161 nm) and CaCO3 (d = 0,385; 0,303; 0,249; 0,227; 0,208; 0,191 nm). Also the presence of a scawtite C ~ S i 4 014.2H20 CaC03 (d = 0,599 0,355; 0,303; 0,200 nm) and a (d ~=, 0,355; ~ 0,311; 0,301; 0,245; 0,228 nm) and low basic of hydrogarnets like C ~ A S ~ J H C3AS1,2-1,4H3,6-3,2(d = 0,272 - 0,273 nm) are possible. On the X-ray gram a portlandite Ca(0H)z (d = 0,263 0,193; 0,169 nm) and an ettringite (d = 0,973; 0,561; 0,479 nm) also are arrested (Fig. 2, curve 1). These products based on possible formation of a hydrosulphosilicate like hydroellestadite Calo(Si04) 3(so4)4(OH)2 (d = 0.424; 0.33 1; 0.302; 0.276; 0.269; 0.189; 0.180; 0.169 nm), similar on composition to calcium chondrodite, including ions (so4)?-. In the composition of new formations existence of a hydrosulphosilicate like Cas.(Si(OH)6)3.(S04)3.24H*0 [101 also is possible. Data of electronic microscopy confirms the formation of the specified phases in composition of products of hardening (fig. 4). According to data of electronic microscopy the structure of a sample based on a cementash composition can be circumscribed as very dense, methamic-like, consisting from knitting against each other hydrated phases. The composition of products hydration was represented by the fibrous needle-like shapes of ettringite and its silicate analogue (fig. 5). Considered ashes filled binding systems on the basis of fly ash of bed fluidized combustion which contain of P-CaS04, have been used as modeling system for creation of high-strength binding compositions on the basis of fly ashes in structure of which prevails amorphous substance. As the base system has been used binding composition, which contains 80% of fly ashes and 20% of Portland cement. For modification of this system and improvement of strength gain as modifying additives has been used the anhydrite, received during burning out gypsum stone at T=45OoC, and silica fume or metakaolin. The optimum quantity of additives (sulfate and siliceous) have been picked up with use of mathematical methods of experiment planning and made 8 % (for each additive) from weight of ash-cement mixes. The results of physical-mechanical test of concrete, which were received on the base of ash-cement binding systems, modified by sulfate and silica additive, is shown in the table 3.

P

0 P

Stability of hydrosulfoaluminosilicatecompounds and durability of an artificial stone based on them 405

1

20

Fig. 3.

Fig. 4.

16

12

8

4

I 0

X-ray grams of the patterns based on binding compositions removed under small angles (a condition Cu K a 30 /30): 1, 3 - 45 % of fly ash of bed fluidized combustion and 55 % of Portland cement after 90 and 360 day of curing; 2 , 4 -40 % of fly ash of bed fluidized combustion, 55 % of Portland cement and 5 % of silica fume 90 and 360 day of curing accordingly

Electronic photomicrograph (a) of specimen and results of microanalysis for separate phases (b) (after 7 days of a storage in standard conditions) on the basis of cementash binders of composition 45 % of fly ash of bed fluidized combustion and 55 % of Portland cement

406

Ekatenna PUSHKAROVA, Vladimir GOTS, Olga GONCHAR

separate phases (b) (after 1 year of a storage in standard conditions) on the basis of cement-ash binders of composition 40 % of fly ash of bed fluidized combustion, 55 % of Portland cement and 5% of silica fume

Table 3. The strength gain of the developed compositions of concretes on the basis of

The analysis of the received data (tab. 3) testifies that growth of strength characteristics of concrete on the basis of ash-cement binding substances modified by sulfate and silica additives at all stages of hardening, and strength of concrete on the basis of ash-cement composition modified by only sulfate additive, after 365 days of hardening decreases on 36,6 % compared with compositions on the basis of Portland cement. As obtained data testifies, introduction in composition of binding systems high-dispersive silica additives promotes the stabilization of hydrosulfoaluminate phases in time. Due to introduction of the last in products of hardening binding substances are formed ettringite and solid solutions on its basis of type hydrogarnet 3CaOA1203 1,6SiOz 2,8H20 (d=0,314; 0,270; 0,233; 0,197; 0,167; 0,161 nm) which presence proves the results of microanalysis (fig. 6,) and explains stable strength characteristics of artificial stone in time. Synthesis of hydrogarnet in structure of products of ash-cement-sulfate compositions hydration modified by metakaolin, is accompanied also by formation of hydrosilicate, which contain silica anions, similar to epistilbite (Ca,j(Si(OH)& (SO& 24H20) (d=0,584; 0,399; 0,369; 0,354 nm).

Stability of hydrosulfoaluminosilicate compounds and durability of an artificial stone based on them

Fig. 6.

407

Electronic photomicrographs of specimens (a, c) and results of microanalysis for separate phases (b, d) (after 1 year of a storage in standard conditions) on the basis of cement- fly ash binders of composition, which content as additive of silica fume (a, b) or methakaolin (c, d)

According to data of electronic microscopy (fig. 6, a, b), strength of artificial stone, authentically, is provided due to crystallization of low basic hydrosilicate of calcium on primary formed needle crystals of ettringite, on the other hand takes place some change of a chemical compound of ettringite and formations on its basis of solid solutions which contain silica-anion complexes. It testifies the ability of formed artificial stone to structurally functional adaptation [ 111 under different conditions of service that formation of solid solutions opens an opportunity of some change phase and a chemical compound of new formations, depending on service conditions, without deterioration of physical and mechanical characteristics of received artificial stone. The developed concrete are expedient for using in underground parts of buildings and constructions where the service of sulfate aggression is present, in the bottom layers of road covering at construction of roads, for hydraulic engineering construction and for reception of special concrete.

408

Ekaterina PUSHKAROVA, Vladimir GOTS, Olga GONCZfAR

REFERENCES 1. Alksnis, F.F., Hardening and destruction of gypsum cement composite. Stroyizdat, Leningrad 1988,221 p 2. Mehta, K., Mechanism of sulfate attack on Portland cement. Cem. And Conc. Res., V.13 N93, 1983, pp 401-406 3. Pushkarova, E.K., Nazim, O.A., Pavluk, V.V., Shevchuk, LO., Pavluk, LM., Features of sulfate activation of ash-cement binding systems by different modification of anhydrite. Collection.” Science bulletin of construction”, Kharkov 2003, pp 36-43 4. Larionova, Z.M., Nikitina, L.V.,Garashin, V.R., Phase structure, microstructure and durability of a cement stone and concrete. Stroyizdat 1977,260 p 5 . Goria, G., Appiano, M., 11 Cement0 N93, 1949 N97, 1950 6. Kuznetsova, T.V., KudrJashov, I.V., Timashev, V.V., Physical chemistry of binding materials. M. Vushaya shkola 1989, pp 308-309 7. Ferronskaya, A.V., The theory and practice of application of gypsum cement plasticizer binding substances in construction, Abstract doctoral thesis, G, 1974,47 p 8. Alekseev, S.N., Ivanov, F.M., Modry, S., Shisol, P., Durability of reinforced concrete in chemically aggressive media. M. Stroyizdat 1980,320 p 9. Pushkarova, K.K., Domoslawsky, W., Features of Processes Hydration and Hardening of Binding Compositions based on fluidized fly ash. Proc. of Seventh NCB International Seminar on Cement and Building Materials, New Delhi, India 2000, XL 10. Lukas, W., Hydration of Cement. Cement and Concrete Research, 1976, v.6, N2, pp. 225-233 11. Chernyavskij, V.L., Concrete adaptation. Dnepropetrovsk, Novaya idealogia, 2002, 116p

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

DURABILITY OF AN OVERLAY-OLD CONCRETE INTERFACE: THE ROLE OF A METAL FIBREREINFORCEMENT Quoc-Thanh TRAN, Ahmed TOUMI, Anaclet TURATSINZE Laboratoire MatBriaux et Durabilite des Constructions GBnie Civil INSA-UPS, 135 av. de Rangueil, 31077 Toulouse Cedex e-mail: [email protected]

ABSTRACT The proposed paper focuses on the debonding propagation along an overlay-substrate interface, notably on the damage of the interlocking between the two faces of the interface under static loading. The induced shrinkage length changes of the overlay and the substrate and its influence on the interface debonding are taken into account. The work associates experiment and simulation approaches with the purpose to clarify the role of fibre reinforcement on the interface debonding mechanism. Two types of overlay materials (OM), a fibre reinforced mortar (FRM) and a plain one (PM), are investigated. Direct tensile tests on notched OM specimens were firstly conducted to obtain the tensile strength and the residual normal stress - crack opening relationship. Drying and autogenous shrinkage of OM have been evaluated. The debonding opening - residual normal tensile stress relationship was investigated by static tensile tests. Three-point flexural static tests were then performed on composite substrate-overlay specimens to evaluate their structural behaviour, in particular the durability of the interface. The debonding interface propagation was monitored using a video-microscope with an enlargement of x175. Relying on the identified and quantified parameters, modelling of the above mentioned static tests was carried out by the finite element method using CAST3M code developed in France by CEA (Commission for Atomic Energy). The shrinkage effect was taken into account by using moisture diffusion equations and a relationship between shrinkage strain and water content variation. The model predictions showed a good agreement with the experimental results and proved the important role of fibre capacity to restrain the crack opening by transferring stresses through the crack.

Keywords Repairs, Fibre reinforcement, thin bonded overlays, interface debonding, interlocking, shrinkage INTRODUCTION Cementitious composite materials, such as fibre reinforced mortars (FRM) or fibre reinforced concretes (FRC), are nowadays increasingly used. They are particularly efficient in concrete structure repairs involving the thin bonded overlay technique. The aim of the overlay may be to replace deteriorated concrete, to smooth a damaged surface, and/or to improve the load bearing capacity of a structure by increasing its thickness, or to provide additional cover for corrosion protection. The overlay can be poured in the case of horizontal top surfaces, or can be spread in other cases. Up to now, such repairs have sometimes posed a problem because of their hazardous durability, essentially due to their debonding. There is therefore a demand for better knowledge to predict the behaviour of such repairs and to find solutions to ensure their durability. The work

410

Quoc-Thanh T M , Ahmed TOUMI, Anaclet TURATSINZE

reported here is part of a large programme: Granju [ 11, Turatsinze [2], Sabathier [3], Tran et al. [4-51.. .It focuses on the enhancement of a discrete crack model allowing relevant and efficient prediction of both crack growth and debonding propagation. Especially the role of fibre reinforcement on their durability is clarified in both experiments and modelling.

OVERLAY MATERIALS The control repair material is a mortar. Its mix proportions are given in Table 1. The behaviour of the control mortar is compared with the one of a reinforced overlay: commercially available amorphous metal fibres were used. They come in flexible ribbon form (29 pm thick, 1.6 mm wide and 30 mm long), are stainless and have high tensile strength (2 GPa). They are suitable in the case of thin bonded overlay and are specially convenient in the most aggressive environments. The mix proportion is derived from the control mortar: fibres are added with a content of 20 kg/m3 and superplasticizer dosage was adjusted in order to maintain the same workability. All specimens were tested on the 7' day. Table 1 - Mix proportions of plain mortar Component I Amount (kg/m3) I 500 Cement (CPA CEM 152.5 Rl Superplasticizer Water

3.25 235

THREE-POINT STATIC BENDING TESTS OF COMPOSITE SPECIMENS Composite specimens A composite specimen consisted of an overlay material layer overlaid on a substrate simulating the structure to be repaired. For practical reasons, the substrate was a hollow steel beam. Despite its metal nature that does not reflect really the behaviour of concrete, the advantages of this kind of substrate are the following: The discrepancy of its mechanical behaviour is limited It is reusable and its surface, in contact with the overlay, always keeps the same characteristics In return, it does not present dimensional variations, whereas a concrete substrate absorbs part of the moisture in the newly overlaid mortar and swells To ensure a good bond between the metal substrates and the overlays, the surface of substrates was firstly made rough by milling, then it was overlaid with fresh mortar. After a few days, the hardened mortar overlay was removed from the substrates and another one was cast again. This procedure took several repetitions until a fine bed mortar, uniform and adherent, covered the substrate surface [l-31 (see Fig. 1). In this way, in spite of the metal nature of the substrate, the overlay-substrate interface could be considered as a cement-to-cement contact. Overlay layers were notched at mid-span in order to predetermine and locate a shrinkage crack. Each overlay layer had a length of 500 mm and a width of 100 mm and was overlaid on a

41 1

Durabiliv of an overlay-old concrete inte$ace: the role of a metaljibre reinforcement

700 mm-long substrate. Two sizes of composite specimens were tested. Their dimensional details are given in Table 2. Concerning the curing conditions of specimens, they were placed in a controlled environment at 20°C and a relative humidity of 60% just after casting to investigate shrinkage effects on specimens in 7 days. To avoid an excessive differential length change between the metal base and the overlay, all specimens were covered in plastic film just after casting. They were removed from the moulds after 24 hours and then were wrapped again in plastic films until the moment of testing. In this way, the shrinkage is limited but cannot be totally neglected. Thin bonded layer

\

Steel substrate

Fig. 1 - Details of the treatment of steel substrate surface.

Three-point static bending tests To investigate the behaviour in flexure of the overlay-substrate interface, three-point static bending tests were conducted using the experimental set-up shown in Fig. 2. They were performed in a closed-loop digitally controlled machine with a 50 kN load capacity. Two LVDT(s) (stroke f l mm) were used to measure the deflection of the specimen and the crack opening displacement (COD) of the overlay. For the COD measurement, the LVDT was placed horizontally spanning the notch and located 10 mm below the bottom overlay surface. For the deflection measurement, a midspan point at the level of substrate wing next to the overlay-substrate interface was chosen. concerning the measure of the interface debonding length, a video-microscope with an enlargement of xl75 was used.

Overlay thickNotation

Substrate section ThickHeight Width (mm) (mm) (mm)

ness (mm) I

S5-04 S15-06

I/_i

I I

I

40 60

I

I

.

I

50 150

I I

100 100

I I

I

2.7 3.5

Table 2 - Dimensional details of composite specimens

Fig. 2 - Experimental set-up for three-point bending test on composite specimens.

Quoc-Thanh T M , Ahmed TOUMI. Anaclet TURATSINZE

412

MECHANICAL CHARACTERIZATION Uniaxial static tensile test for overlay materials Uniaxial static tensile tests aim to obtain the tensile strength and the stress - crack opening relationship as input data for the model. They were conducted on 100 x 100 x 120 mm notched prisms as described in [4-51. As results, typical curves for plain mortar and FRM are shown in Fig. 3 and Fig. 4 respectively. Based on the experimental results, the mean stress - crack opening relationship fits the following formulae. For plain mortar:

where o;, is the mean interlocking stress; Rw=2.58 MPa is the tensile strength; w is the crack opening and wlp = 0.15 mm is the crack opening limit. For FRM:

where is the mean interlocking stress; R ~ 3 . 1MPa and 0 ~ 1 . 6MPa are the tensile strength and the stress plateau value of FRM corresponding to the consolidated stage respectively; 0 wlfl . 1 mm are parameters interpreting the beginning, the end of the wo=0.015 mm, ~ ~ ~mm, consolidated stage and the end of interlocking stress activation respectively.

0

005

01

0 15

w (a)

Fig. 3 - Residual tensile stress - crack opening curves for plain mortar

Ow"

02

06

04

08

1'

w

Fig. 4 - Residual tensile stress - crack opening curves for FRM

Young's modulus of overlay material The Young's modulus of FRM was measured through a uniaxial compression test. Four large cylindrical specimens (1 10 mm in diameter and 220 mm in height) were tested on a 3000 kN servo hydraulic testing machine. The faces of the specimens were ground to impose high

12

Durabiliw of an overlay-old concrete interface: the role of a metal fibre reinforcement

413

parallelism. Three Ohmic gauges were glued equidistantly on each specimen to measure the average longitudinal strain. The tests were carried out by controlled loading at the rate of 0.5 MPa /s. The mean value of the Young's modulus on the 7" day is equal to Em = 31000 MPa.

Uniaxial static tensile test for overlay-substrate interface This test aimed to obtain the tensile strength and the residual stress-debonding opening relationship for the overlay-substrate interface. The test is detailed in [4-51. The specimens consisted of two parts: substrate in steel and overlay material. The substrate surface in contact with overlay was processed to provide the same properties as the one related to the flexural test mentioned above. The test was carried out using controlled displacement at a low rate of 2 pdmin. In spite of a minimal loading rate, no test could be controlled successfully until the end of the post-peak phase (see Fig. 5). This demonstrates the brittle character of such an interface. The average value obtained for the interface tensile strength from six tests was Rti = 1.5 MPa. The tensile stress - debonding opening relationship was derived from the measured interface displacement. A simplified formula proposed by Toumi [6] was used:

where w/dis the debonding opening limit beyond which the interlocking totally vanishes. qi and w are the interlocking tensile stress and the debonding opening of overlay-substrate interface, respectively. It was estimated from the interface tensile tests that wld = 10 pm. The corresponding ai- w curve is presented in Fig. 6.

Fig. 5 - Experimental results : tensile stress versus interface displacement.

Fig. 6 - Residual tensile stress - interface debonding opening relationship.

MODELLING OF DELAYED EFFECTS ON MECHANICAL BEHAVIOUR OF OVERLAY-SUBSTRATE STRUCTURES Modelling of drying shrinkage The drying shrinkage was computed using equations [7-91 as follows : Diffusion law:

414

Quoc-Thd TRAM Ahmed TOWMI, Anaclet TURATSIhZE

ac

-= div(D(C).grad((=))

(4)

at

Initial conditions: CO= W - r y B ,

(5)

Boundary Conditions: C = C,, at the drymg surface ac = 0, at the surface in contact with steel substrate an

in which C is the evaporable water content, D is the moisture diffusivity; W , and B, are the water content and the binder content respectively in the mortar formulation. CO is the free water content when evaporation becomes possible (just after demoulding which was performed 1 day after pouring), r is the hydration degree, yis mass proportion of water reacted with the cement; aC / an is moisture gradient at the drylng surface identified by a unit normal n; C, is the equilibrium water content that an element would reach at a given particular environmental condition; C, is the water content of the drying surface, and fis the convective moisture transfer coefficient. The diffusivity D was calculated with the relation derived by Granger et al. [9]:

D(C ) = A .exp( b . C)

(9)

The convective moisture transfer coefficient f can be calculated according to the formula proposed by Torrenti et a1 [lo]:

A strain field is generated by the moisture diffusion from the overlay material. However the resulting length change of the overlay is restrained at the interface by the substrate. A linear relationship between the moisture loss and free shrinkage strain Eh, was used as follows [8]: ~h

= k . AC

(1 1)

According to [5],the required parameters take values given in Table 3. Table 3 - Required parameters for modeling drying shrinkage Co(Kg/m)

131

I I

A(m2/s)

I

b

1.54~10-'~ I 0.05

I

118

I

m4.s-'

5x10'"

I

k [(kglm Y 17.8~ 1 O*

I

Durability of an overlay-old concrete inteface: the role of a metalfibre reinforcement

Modelling of autogenous shrinkage The autogenous shrinkage is a deformation related to the water consumption during the hydration reaction. It is treated as an additional water loss uniform in all the mortar volume and included in relationship (11). The free autogenous shrinkage of overlay material in the present investigation was measured on three prismatic specimens of 40 x 40 x 160 mm. The specimens were removed from the moulds at 21-hour age and the time origin for shrinkage measurement is 3 hours after then to avoid any disorder [l 11. The result is shown in Fig. 7.

1

3

1

2

3 6 5 Twm (days)

415

6

3

I

Fig. 7 - Mean curve of autogenous shrinkage strain versus time of overlay material

Coupling between shrinkage, creep and relaxation Because of creep and stress relaxation, the stresses induced in the overlay by restrained shrinkage would undergo redistribution with time. It has been found that, for this problems, the response can be predicted adequately by using an age-adjusted effective modulus, Eca(t), that is defined as follows [ 121:

in which: E(t) : short-term elastic modulus at time t. to : age (in days) when the shrinkage load (or deformation)is applied. to = 1 day in this investigation. t : elapsed time in days from casting of specimens until the moment of flexural test, t = 7 days. #,to) is the creep coefficient. According to ACI 209R-92 [13], it can be calculated as follows:

where q$, and %, %, yVlsare the ultimate value of creep coefficient in standard condition and correction factors, respectively. Thus the creep coefficient on the 7* day can be obtained and takes the value @= 0.5. ~ ( tis) called the aging coefficient determined from the following equation [14]: ,y(t,to)=

[ E$:d, 1

-

)T

2

--

1

@(t,to)

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Quoc-Thanh T U , Ahmed TOUMI, Anaclet TURATSINZE

in which ER(t,to) and E(ta) are the relaxation function and the Young's modulus at time to respectively. According to Bazant [14], the relaxation function can be determined from the Volterra's integral equation. Thus the aging coefficient x on the 7" day was calculated from Eq. (14) and takes the value x = 0.57. Finally the age-adjusted effective modulus of the overlay material, E,,, on the 7" day could be determined from Eq. (12) and was used to predict the shrinkage effect on the mechanical behaviour of composite structures as presented in the following section.

PREDICTION OF CRACKING GROWTH AND INTERFACE DEBONDING PROPAGATION IN COMPOSITE ELEMENTS Modelling fundamentals Calculations were performed with the CAST3M code developed in France by CEA (Centre for Atomic Energy). To manage crack or debonding propagation, it was decided to control the propagation from the stress state calculated at the first node beyond the crack tip or beyond the debonding tip as sketched in Fig. 8 [3-61. Such a method avoids controlling the propagation by the stress state calculated at the tip of the crack or of the debonding, where the strength theory analysis predict stress singularity (infinite value). In this way, the crack tip or the debonding tip is propagated to the next node when the stress state at the controlling node no longer satisfies a stability criterion (which is dependent on the material or on the interface). The interlocking is taken into account by closing forces imposed between the nodes facing each other along the crack surfaces or along the debonded zone. The literature shows that, in the case of cement-based materials, debonding is initiated by tension perpendicular to the interface. Therefore, in the present model, the closing forces were assumed to depend on the normal relative displacements between each pair of facing nodes, perpendicular to the crack plane or to the debonding plane.

Fig. 8 - Finite element calculation: controlling node of crack or debonding propagation.

Implementation In this section, the three-point static bending tests on the overlay-substrate specimens described in the previous section were simulated using the discrete crack model mentioned above. Moreover the shrinkage effect on the mechanical behaviour of the composite structures was also taken into consideration. Thus the calculation procedure consists of two stages. In the first stage, the numerical simulation of stress build-up in the repair system due to restrained shrinkage was performed. The free shrinkage strain E, included drying and autogenous one, was determined through Eqs. (4-11) and experimental data as presented in Fig. 7. An equivalent stress field a i n the overlay was computed according to the Hook's law:

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in which E,, is the age-adjusted effective modulus of the overlay material on the 7'h day and is given by Eq. (12). The calculated stress fields CT were then transformed into equivalent nodal forces to simulate the shrinkage effect. In the second stage, these nodal forces were superimposed to the mechanical loading. To model the cracking behaviour of the overlay material and the overlay substrate interface, the required mechanical properties were presented in Table 4. Table 4 - Experimental mechanical properties of the overlay materials and the interface interface Young's modulus Tensile strength c w relationship

Em = 3 1 GPa

Em = 3 1 GPa

Rtp= 2.58 MPa Eq. (1)

RB = 3.1 MPa Eq. (2)

Rti = 1.5 MPa Eq. (3)

Results and discussion Experimental and numerical results for the three-point static bending tests on the overlaysubstrate composites are shown in Figs. 9-13. It can be seen that the model prediction is in a good agreement with the experimental results. Figs. 9-10 allow comparison of interface debonding propagation for both studied composite geometries (S15-06 & S5-04). It can be seen that for the same geometry of specimens, the load required to initiate the interface debonding in the case of FRM is always higher than the one obtained in the case of plain mortar (= 30%). Moreover, at a given load level, the interface debonding length in the case of plain mortar is always higher than the one obtained in the case of FRM. Figs. 11 and 12 show that the same conclusions can be drawn concerning the COD kinetic. Indeed, fibre reinforcement ensures structural continuity through the crack, thus controls the COD and the debonding propagation in the first stage. Concerning the load versus deflection curves (Figs. 13-14), one can notice that there is no significant difference between two cases of overlay materials. It can be explained by the fact that with the point chosen to measure the deflection (on the substrate portion), this parameter depends essentially on the relative high rigidity of steel substrate and is slightly influenced by the interface degradation.

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CONCLUSION The present work has focused on debonding propagation along the interface between a cementbased thin bonded overlay and a substrate. The evolution of interlocking between two crack surfaces through the overlay as well as two faces of the debonding interface under static loading were thus investigated. The shrinkage effect on the mechanical behaviour of repair structures was taken into account. Based on the cohesive crack concept, a model has been developed to study the influence of major parameters on the durability of the bond of such overlays. The relevant parameters were identified by a series of tests and then introduced into the model. Then an appropriate procedure of calculations was conducted to predict the mechanical behaviour of plain mortar and fibre reinforced mortar as repair materials. The numerical predictions show that the proposed model is efficient to predict cracking and interfacial fracture behaviour. The following conclusions can be drawn.

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1. The cracking of the overlay and debonding propagation along an interface is governed by the interlocking phenomenon. 2. Fibre reinforcement appears as a solution to improve the durability of thin bonded cementbased overlay. It delays the crack opening growth and debonding propagation. The ongoing work is dedicated to numerical and experimental investigations on concrete substrates repaired by the thin bonded overlay technique and the first results are promising.

REFERENCES 1. Granju, J.L., ‘Debonding of thin cement-based overlays’ ASCE Journal of Materials in Civil Enginering 32 (2) (2001) 114-120. 2. Turatsinze, A., Farhat, H. and Granju, J.-L. ‘Influence of autogenous cracking on the durability of repairs by cement-based overlays reinforced with metal fibres’, Materials and Structures, (36) 2003 673-677. 3. Sabathier, V., Granju, J-L., Bissonnette, B., Turatsinze, A., Tamtsia, B., ‘Thin bonded cementbased overlays: Influence of initial bond defects and fibre reinforcement’, in Seventh International Symposium on Brittle Matrix Composites, Woodhead Publishing ltd, Poland, October, 2003. 4. Tran, Q-T., Toumi, A., Granju, J-L., ‘Interface between an old concrete and an overlay’, Third international conference on composites in construction, CCC2005-France,July, 2005. 5. Tran, Q. T., Toumi A., Granju, J.-L, ‘Experimental and numerical investigation of the debonding interface between an old concrete and an overlay’, Materials and Structures, Online publication: 2005-10-05. 6. Toumi, A., Bascoul, A., Turatsinze, A., ‘Modeling of fatigue crack growth in concrete subjected to mode I crack opening’, FRAMCOS-4, edited by R. De Borst, J. Mazars, G. PijaudierCabot, J.G.M. Van Mier, (A.A. Balkema Publishers, 2001) 637-640. 7. Rahman, M.K., Baluch, M.H., Gadhib, A.H., ‘Simulation of shrinkage distress and creep relief in concrete repair’, Composites Part B: Engineering 31 (6-7) (2000) 541-553. 8. Feron, C., ‘Etude des mecanismes de generation de contraintes et de fissuration par retrait g&nC dans les structures B base de materiaux cimentaires’, Phd Thesis of INSA Lyon, 17 June 2002. 9. Granger, L., Torrenti, J-M., Acker, P., ‘Thoughts about drylng shrinkage: experimental results and quantification of structure drylng creep’, Materials and Structures 30 (1997) 588-598. 10. Torrenti, J.M., Granger, L., Diruy, M., Genin, P., ‘Modelisation du retrait du beton en ambiance variable’, Revue Frangaise de Genie Civil, 1 (4) (1997) 687-698. 11. Granju, J-L., Sarkis, M., Amaud, M., Escadeillas, G., ‘Temps zero de reference pour les mesures de retrait’, Materials and Structures 37 (2004) 449-455. 12. Carino, N-J., Clifton, J-R., ‘Predictionof cracking in reinforced concrete structures’, Technical report NISTIR 5634, National Institute of Standard and Technology, Gaithersburg, MD 20899. 13. ACI 209R-92, (reapproved 1997), ‘Prediction of Creep, Shrinkage, and Temperature Effects in Concrete Structures’, reported by ACI committee 209. 14. Bazant, Z-P. ‘Prediction of concrete creep effects using age-adjusted effective modulus method’, ACI Journal, (1972) 212-217.

Proc. Int. Symp. “BrittleMatrix Composites 8“ A.M. Brandt, YC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

PLENARY INVITED PAPER

ON SCIENCE ASPECTS OF SIMULATING PRACTICAL MANIFESTATIONS OF CONCRETE Piet STROEVEN and Jing HU Faculty of Civil Engineering and Geosciences, Delft University of Technology Stevinweg 1,2628 CN Delft, The Netherlands e-mail: [email protected]

ABSTRACT Materials science aspects are discussed that govern manifestations of concrete produced by computer simulation (=compucrete). Stochastic heterogeneity is linked up with scatter in a descriptor of material structure or of performance. Sample size leading to an acceptable scatter level in the descriptor defines the size of the representative volume element (RVE) that is homogeneous as to the very parameter. The RVE size increases with the degree of configuration-sensitivity of the descriptor of structure, or with the degree of structure-sensitivity of the descriptor of performance. In cases of finite degrees of sensitivity, the descriptor’s probability density function (histogram) is skew to the left. This causes biases due to stochastic heterogeneity effects in sub-RVE experimental or computer-simulation designs. Physical modelling is performed differently by RSA-based systems, like HYMOSTRUC 3D, and by concurrent algorithm-based systems, like SPACE. Manifestations of concrete produced by RSA-based systems provide only a realistic representation of material composition, whereas the representation of reality also encompasses material configuration in case of concurrent algorithm-based systems. Only the latter type can be used therefore to study structure-sensitive performance characteristics (or configurationsensitive features).

Keywords Computer simulation, stochastic heterogeneity, configuration, structure sensitivity, concrete INTRODUCTION Concrete is a composite material with a complex structure. It is widely used for a very wide range of applications, among which most of the infiastructural facilities. Hence, impact of the material on society is large. Nevertheless, the material is manifesting durability problems on large scale, requiring major investments for repair and renovation. Hence, thorough research efforts are urgently needed to prevent such problems continuing in the future. Growing computer simulation capabilities offer a promising perspective in combination with limited experimental verification testing for developing economic strategies to approach such technological problems as an alternative for time consuming and expensive systematic experimental investigations. But like the old wisdom: “the quality of the research is controlled by the quality of the sample”, we are confronted in such new developments by a similar “law”, i.e., “the quality of the research is controlled by the quality of the compucrete”. Compucrete constitutes the computer-made manifestation of concrete under investigation, whereby manifestation is defined as schematization of reality with operational potentials. The schematization involves a practical and scientifically sound simplification of reality. This has nothing to do with the alchemical pursue for finding gold or an elixir for extending men’s life

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in a research design not based on the scientific knowledge of today, sometimes used for characterizing modelling approaches in concrete technology [l, 21. Contrary, the quality of the compucrete should be cared for in a similar fashion as that of the sample in experimental research. This paper will go into the details of the scientific framework that, when respected, will guarantee the quality of manifestations of concrete for objected research purposes. Different manifestations of concrete will be demonstrated having different operational capabilities. Misunderstanding as to the limitations involved in the various schematizationsof reality incorporated in the compucrete will inevitably lead to biased predictions, and uneconomic solutions to the problems of today. To be honest, this is indeed a problem of today. CONCRETE AND COMPUCRETE

When it is pursued schematizing concrete’s material structure in a physical model, here referred to as a manifestation of the concrete (=compucrete), it must be obvious that the degree of schematization should be functional, causal or operational. It should allow offering a framework for reliably estimating the material features we are interested in. When targeting different features of material structure, it could be of economic relevance to use different manifestation. Or in simpler words: when looking for how much of a certain phase (particles, pores) we are confronted with, a simple problem, the manifestation can be dramatically simpler, than when the configuration, or dispersion of particles or pores is at issue; research into the latter problem is far more demanding for the degree of schematization of reality incorporated in the compucrete. This points, first of all, to assessment of the different categories of concrete’s features that could be subject of interest. To do so, we introduce the well-known concept of stochastic heterogeneity [3,4]. This is not a material property, as suggested sometimes in research. Instead, it has to do with sampling. We make or take (e.g.,by coring) samples of the concrete material body, mostly in the form of simple shapes and in building codes prescribed sizes. These samples can be weighed to determine volumetric density. A large number of similar samples will yield a group of data on density of the material. A Gaussian-type of histogram is obtained when frequency of occurrence is plotted. For an infinite number of observations, this would transform into a smooth-shaped probability density function. The histogram’s mode or average value is an unbiased estimator of material density. Larger samples (specimens) will lead to narrower histograms, and thus lower amount of scatter (Fig. 1-left). Stochastic heterogeneity is associated with scatter of the same descriptive parameter (here, density) of material structure. When scatter is of acceptable low level for the researcher’s purposes, the sample is defined as homogeneous in the descriptive parameter (again, here density), and the specimen is said to be of representative size. In other words, we deal here with a representative volume element (RVE)of concrete’s feature (density) under investigation. Since density is a so-called composition parameter (only reflecting how much of the material is at issue), the RVE is of only modest size; say four to five times the maximum grain size [5]. Again, starting with a large number of specimens, we want to collect information on the two-dimensional nearest neighbour surface-to-surface spacing (NND) distribution among the aggregate grains (since we expect this a prime factor governing shrinkage crack evolution in which we are, supposedly, interested). As in the previous case, we want engineering information derived from this large group of identical specimens. For that purpose, we could take

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Fig. 1. Smoothed absolute frequency histograms of composition descriptor (e.g.,density, left), and of configuration one (e.g., spacing distribution, right) of “large” and “small” specimens with respect to structural dimensions, say, maximum grain size. Number of specimens is 400. average values and compose histograms, as probability density distributions. The histograms will be skewed to the left, however (Fig. 1- right). And scatter will exceed that for volumetric density considerably when based on specimens of equal size. As a consequence, the sample size has to be increased far more to amve at an acceptable scatter level for defining the sample of representative size, and homogeneous as to the average surface-to-surface spacing. Of course, this size would be insufficient for defining the surface-to-surface distribution function (i.e., the full range from small to large distances) homogeneous. To do so, the surface-tosurface distances should be collected in histograms and the various histograms compared until differences would be acceptable. The average surface-to-surface spacing is a descriptive parameter of moderate configuration sensitivity; the distribution hnction of this descriptive parameter is highly configuration-sensitive. A decline in configuration-sensitivity is accompanied by a decline in the size of the associated RVE and thus of the level of homogeneity. Observed at the same sample, the skewness of the probability density histograms increases with the degree of configuration sensitivity. The relevance of the foregoing is coming from the structure-sensitivity of most of the properties we are interested in, i.e., they are influenced by the configuration of material structure. Practical consequences are the relatively large specimens necessary to avoid biases. Such biases would be the result of stochastic heterogeneity, or, as it is referred to in fracture mechanics, of “the” size effect. This is illustrated in Table I. The level (ie., the largest dimensions) of the microstructure governs the RVE size involved. Changing maximum grain size would cause a modification in the size of the RVE. In sub-representative experimental or computer simulation designs (a quite common situation in concrete technology), the effect of stochastic heterogeneity on structure-sensitive properties (such as shrinkage cracking) could only be eliminated in a comparative set up (e.g., evaluating the effect of grain size on the property at issue) by keeping the ratio of RVE to sample size constant [4], a precaution seldom taken by researchers in concrete technology [6]. The relevance as to the computer simulation methodology, the major topic of this paper, will be made explicit in what follows. Table I. Spacing parameters of aggregate grains in compucrete

92.6 46.3

0.1083 0.1062 0.0989

0.075 1 0.0993 0.1640

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COMPUTER SIMULATION METHODOLOGY

We are confronted in concrete technology on meso-level with packed particles in the form of aggregates, and in the fresh state on micro-level with cement particles, eventually blended by those of a mineral admixture. This is the prime target of physical computer simulation strategies. In the latter case, hydration is simulated in a second stage. Generally, particles are assumed spherical. Yet, recent studies have demonstrated that particle shape can exert significant effects on packing density in the jammed state, relevant for concrete aggregates [7]. Moreover, the efficiency of cement blending by finer grained mineral admixtures will probably also be affected by shape, because it at least partly relies on the internal migration capacity of the finer blend through the grain skeleton of the packed cement particles [8]. The available physical computer simulation systems for packing of particles in concrete technology can be placed in two distinct groups. Common strategy is to apply an RSA (random sequential addition) procedure for placing the particles inside a container. Fundamentally different strategies in solving the particle overlap problem that occurs during this procedure are followed, however, in the two groups. The systems solely based on RSA-procedures reject particles that overlap with earlier generated ones, and continue the RSA procedure. The number of rejections will increase dramatically at increasing densities, which makes the operation quite time-consuming. Maximum attainable density is found lower than 40% [7]. But of major concern is that “any relation between these RSA packings and an experimental granular packing is at least tenuous” [7]. To be more specific, RSA procedures cause a suppression of particle clustering because accidental overlap in the simulation will lead to new positions but more remote from their potential clustering location. The final result is a model material (compucrete) in which particles are more evenly distributed than in nature. HYMOSTRUC 3D system, developed at Delft University of Technology [9], can be referred to as an example, but other RSA systems are basically similar [lo]. The strategy of solving the overlap problem in the second group is based on so-called concurrent algorithms. Static as well as dynamic implementations have been realized. The system used in [7] is of the first type, whereas the SPACE system, developed at Delft University of Technology [l 11, is of the second type. In both cases, a more dilute system is produced in the first stage (for which RSA procedures can be employed), whereupon the container size is gradually reduced. The static system locally shifts positions of overlapping particles during mechanical contraction, the dynamic system makes use of a dynamic (Newtonian) simulation mechanism (also reflecting production conditions), in which the forces that are added to the particles can be manipulated so that repulsion, attraction, or gravitation effects can be included [12]. Jammed states can be produced (at even higher density than in the static case), quite accurately corresponding to experimental values obtained on compacted aggregates with different compositions [131. Particle clustering is properly simulated by the SPACE system; mean values of NNDs are 0.236 pm (left) and 0.262 pm (right) for the patterns in Fig. 2. Even in this relatively dilute two-dimensional system, the total number of (near) contacts between particles is 8 for the SPACE structure, and zero in the random case. Differences can be expected much more dramatic when based on dense 3-D grain packing. More generally speaking, the configuration of the packed particles forms a (far) more realistic schematization of reality than can be obtained by RSA-based systems. Of come, these distinct capabilities of RSA- and concurrent algorithm-based systems to schematize reality as to particle configuration in the fresh state will exert direct influences in more matured states on configuration, either of unhydrated cement grains (e.g., relevant for self-healing capacity), or of pores, and thus on the pore-depercolation process (relevant for durability issues) [141.

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Fig. 2. Two-dimensional “particle” packing 1: i mixing algoritlhm (left) and by RSA procedure (right) for 40% area coverage (particle size be:tween 1 and 10 pm, container size 50 pm).

MODELLING Most of the world around is structured, and details of the structure play a crucial role in their resistance to decay. This is also the case when discussing men-made concrete. Of course, the relative amounts of components in composites as concrete are of eminent importance. Yet, it is the phase’s co-operative action that really gives extra dimension to survival capabilities of the composite. In scientific terms, we can state that most performance capabilities of concrete we are interested in are (to different degrees, though) depending on material configuration. So, basically the schematization of reality reflected by compucrete should be quite sophisticated. Only when the configuration-sensitivity of the phenomenon we are interested in is proven low, we may study a less sophisticated manifestation of concrete. Otherwise, seeing the operational capabilities of the different manifestations of concrete, only a concurrent algorithm-based approach (e.g.,SPACE) can be expected providing reliable data. RSA-based systems can do the job only when the low sensitivity is evidenced by the concurrent algorithm-based system. In summary, HYMOSTRUC 3D (and other RSA alternatives) should be employed only when the property is of near-composition nature. An illustrative porosimetry examde is uresented in Figs. 3 and 4.based on the fresh cements of Fig. 2.

Fig. 3 . Simulated (dynamic: left; random: ria structures in the equally - it) . pore - - long- matu-ed 2D cements of Fig. 2.

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Fig. 4. Cumulative porosity distribution (lefi) and pore size distribution (right) for the hydrated cement pastes (material structure visualised in Fig. 3), of which the fresh states are generated by SPACE (denoted as ‘dynamic’) and by RSA procedures (denoted as ‘random’). Fortunately, structure-sensitive processes (appealing to material configuration) have the tendency to reveal declining sensitivities. The first crack strength will be highly depending on material configuration, but configuration-sensitivity will decline as damage increases. E.g., ultimate tensile strength was found not dramatically influenced by the particle size distribution (grading) of the aggregate (volume fraction being similar), as shown by Fig. 5. The sensitivity to grading was found negligible for fracture energy, Gf [15], as suggested by General Fracture Mechanics Report [ 161. This could have been expected, since the surface profile (tortuosity) of the leading crack as a major parameter has been analytically demonstrated mainly governed by composition, i.e., by volume fraction of the aggregate [ 171.

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Fig. 5. Cracking in tension observed in FE simulation of SPACE-generated 2D compucrete (left). Load-displacement diagrams reveal marginal effects of grading (A and B) (right) [ 151. The same tendency is found in studies of porosity. Dispersion of pores in matured concrete is governed by the configuration characteristics in the fresh state [ 181. Even total porosity manifested a low sensitivity on configuration; this gave rise to about 10% higher porosity values in SPACE simulations than in RSA based ones [ 191, as shown by Fig. 4(left). It can be expected that the evolution process of gradual de-percolation of the initially fully connected pore space will reveal declining configuration sensitivity also. This is evidenced by quite

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comparable values of the de-percolation threshold (end of the de-percolation process) provided by Bentz & Garboczi [20], Ye [21], and Chen [22], despite distinctly different characteristics of the de-percolation process. De-percolation starts relatively early during the hydration process due to the independence of the pores in [20]. De-percolation starts extre-mely late during hydration in Ye’s approach, because of the relatively even cement particle distribution in the fresh state. Particle interaction is incorporated in the SPACE system. As a consequence, a much wider range of three-dimensional surface-to-surface spacing values of the hydrating cement grains is leading to more gradual de-percolation evolution than in H Y MOSTRUC 3D system [23]. The various curves in Fig. 6 , however, seem to roughly tend for similar cases to the same threshold value, reflecting its low configuration sensitivity.

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f

The more even spatial distribution of the cement particles in the fresh state produced by RSA-based systems (reflected by the steep decline in connected porosity) will also significantly affect the characteristics (amount, size distribution, spatial dispersion) of the unhydrated cement paste in matured concrete [24]. Hence, SPACE manifestations of concrete should be considered on the scientific grounds presented herein more appropriate than H Y MOSTRUC 3D, or similar RSA systems, for studying the “self-healing” capacity of concrete, unless in the unlikely case that these investigations reveal low configurations-sensitivity of the phenomenon. The Interfacial Transition Zone (ITZ) in concrete is considered a key phase in controlling mechanical and durability properties of concrete (despite some debate [25]). One might consider the ITZ phase as a separate one, and use compucrete to investigate its properties. However, even a composition property as local density is governed by particle migration to and from the ITZ (in conformity with the idea that the ultra fine silica fume particles added to the cement would predominantly densify the ITZ, in agreement with disproportionately improved strength and durability of concrete). This is indeed what happens as demonstrated for (rice husk ash) blended Portland cement mixtures [26]. Migration of particles is a long-range process, leading to gradients in volume fractions of the different particle sizes extending much deeper in what is normally considered bulk paste. So, particle grading (as on engineering level

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obtained by sieving) is changing over distances away from the aggregate grain’s surface exceeding significantly the ITZ thickness for density [27]. Hence, HYMOSTRUC 3D and similar RSA-based systems (as well as analytical approaches based on the same assumption [28]), cannot provide the relevant structural details of the ITZ, not even the compositional ones. This conclusions has far-ranging implications for computer simulation approaches, because Chen has revealed the significant influences of boundary conditions (cement pocket between aggregate grain surfaces, “bulk” paste, and mixed situations) exerted on even total porosity (supposedly an approximate composition parameter), but very dramatic effects of rigid surfaces on the de-percolation process of pores [22,23]. Fig. 7 visualizes the latter.

r

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Thickness of the removed layer bm)

Fig. 7. Connected porosity is restricted to a narrower zone (0.5 pm) in the ITZ of SPACEgenerated compupaste (w/c=0.3; Blaine numbe~300)between aggregate grain surfaces (container walls) than total porosity (limited to zone of 2 pm) at about 75% degree of hydration (DOH) [21]. CONCLUSIONS

Each descriptive material parameter (of structure or of performance) has its own scale of stochastic heterogeneity. This implies that RVEs in which the material can be considered homogeneous for the given descriptor are of different sizes, too. Basically, the RVE for material composition is relatively small (say, 4 to 5 times the maximum structural parameter, such as grain size), but the RVE for a highly configuration-sensitive parameter will be significantly (up to one order of magnitude) larger. Depending on the degree of configuration-sensitivity, RVEs will have intermediate dimensions for various descriptors of material structure or material behaviour. Unbiased estimates for composition of material structure or structure-insensitive material behaviour can be obtained even in sub-representative sampling designs by averaging over repetitions. Evasion of biases in sub-representative experimental or computer-simulation approaches can only be achieved when configuration- or structure-sensitivity are involved in comparative studies whereby the ratio of RVE and sample dimensions are kept constant (so, adaptations are required when maximum grain size or hydration time are changed). The probability density function of the descriptor (represented by histogram of data from repetitions)

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will be skew to the left. Violating the required adaptation of RVE size will add effects of stochastic heterogeneity (size effect in fracture mechanics) to the structural influence studied. RSA-based computer-simulation systems (like HYMOSTRUC 3D) generate a manifestation of concrete (=compucrete) in which the schematization of reality is unbiased as to material composition only. So, composition and structure-insensitive properties can be studied by such systems. Concurrent algorithm-based systems (like SPACE) generate a manifestation of concrete in which the schematization of reality involves a realistic representation of configuration of material structure. Hence, such systems can be employed for studying the full ranges of configuration-sensitive features of material structure and of structure-sensitive material properties. Processes in materials (cracking, pore de-percolation) manifest declining sensitivities to material configuration, so that approximate estimation of advanced stages in such processes (at ultimate, or post ultimate deformational states, or at de-percolation threshold, respectively) may be obtained by RSA-based systems, too. Even composition (volume fraction of pores, or particles) of part of the material system (like the ITZ) can be influenced by configuration aspects of the whole system, so that extreme care should be bestowed on selecting an RSA-based system for application to seemingly compositional problems. The RSA-based system can safely be employed when results obtained by concurrent algorithm-based system provides evidence of the composition nature of the descriptor of material structure or of material performance that is targeted. REFERENCES 1. Maier, G., Lettieri, M.A., Piola, In situ mechanical characterization of dam concrete and stress state by dilatometric measurements and inverse analysis. In. Carpinteri A, Mai YW, Tirchie, R.O., Ferro, G. (eds.), Proceedings of 1 lth International Congress on Fracture, Turin, Italy, March 2005, available on proceedings CD 2. Breugel, K. van, Modelling of cement-based systems - the alchemy of cement chemistry, Cement and Concrete Research, 3419 (H.F.W. Taylor Commemorative Issue), 2004, pp 16611668 3. Freudenthal, A.M., The Inelastic Behaviour of Engineering Materials and Structures. Wiley, New York, 1950 4. Hu, J., Stroeven, P., Local porosity analysis of pore structure in cement paste. Cement Concrete Research, 3512,2005, pp 233-242 5. Stroeven, P., Stroeven, M., Size of representative volume element of concrete assessed by quantitative image analysis and computer simulation. Image Analysis & Stereology 20 (supplement l), 2001, pp 2 16-220 6. Bisschop, J., van Mier, J.G.M., Effect of aggregates on drylng shrinkage microcracking in cement-based composites. Materials & Structures, 35,2002, pp 453-461 7. Williams, S.R., Philipse, A.P., Random packings of spheres and spherocylinders simulated by mechanical contraction, Physical Review E, 67,051301,2003, pp 1-9 8. Stroeven, P., Hu, J., Stroeven, M., Exploitation of particle migration mechanism to promote economy and ecology in concrete technology, Key Engineering Materials, 302, 2006, pp 1925 9. Breugel, K. van, Simulation of Hydration and Formation of Structure in Hardening Cement-based Materials. PhD thesis. Delft University Press, Delft, 199 1

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10. Roelfstra, P.E., A Numerical Approach to Investigate the Properties of Numerical Concrete. PhD Thesis. EPFL-Lausanne, Lausanne, 1989 11. Stroeven, M., Discrete Numerical Model for the Structural Assessment of Composite Materials. PhD Thesis. Delft University Press, Delft, 1999 12. Stroeven, P., Stroeven, M., SPACE approach to concrete's space structure and its mechanical properties. Heron, 4614,2001,pp 265-289 13. Stroeven, P., Stroeven, M., Assessment of packing characteristics by computer simulation. Cement and Concrete Research, 29, 1999, pp 1201-1206 14. Stroeven, P., Chen, H., Stroeven, M., On connectivity of porosity in model cement paste. Proceedings of Brittle Matrix Composites 8, Warsaw, to be published in 2006 15. Stroeven, M., Askes, H., Sluys, L.J., A numerical approach to determine Representative Volumes for granular materials. In: Mang HA, Rammerstorfer FG, Eberhardsteiner J (eds.), Proceedings of 5" World Congress on Computation Mechanics, Vienna, 2002, proceedings published on CD 16. RILEM TC QFS "Quasi-brittle fracture scaling and size effects"- Final Report. Materials and Structures, 37,2004,547-568 17. Stroeven, P., A stereological approach to roughness of fracture surfaces and tortuosity of transport paths in concrete. Cement Concrete Composites, 22,2000, pp 331-341 18. Hu, J., Stroeven, P., Depercolation threshold of porosity in model cement; approach by morphological evolution during hydration. Cement Concrete Composites, 27/1, 2005, pp 1925 19. Stroeven, P., Hu, J., Modelling in Concrete Technology - Balancing between Science and Alchemy? Cement and Concrete Research, 2005, submitted for publication 20. Bentz, D.P., Garboczi, E.J., Percolation of phases in a three-dimensional cement paste microstructural model. Cement Concrete Research, 21, 1991, pp 324-344 21. Ye, G., van Breugel, K., Fraaij, A.L.A., Three-dimensional microstructure analysis of numerically simulated cementitious materials. Cement Concrete Research, 33, 2003, pp 2 15222 22. Chen, H., Stroeven, P., Ye, G., Stroeven, M., Influence of boundary conditions on pore percolation in model cement paste. Key Engineering Materials, 303,2006, pp 486-492 23. Chen, H., Numerical Modelling on ITZ Microstructure and Its Influence on the Effective Elastic Property and Diffusivity of Concrete. PhD Thesis, Delft University of Technology, 2006, to be published 24. Stroeven, M., Stroeven, P., SPACE system for simulation of aggregated matter; application to cement hydration. Cement Concrete Research, 29, 1999, pp 1299-1304 25. Diamond, S., Huang, J., The interfacial transition zone: reality or myth? In: Katz A (editor), The Interfacial Transition Zone in Cementitious Composites. E & FN Spon, London, 1998, pp 3-39 26. Stroeven, P., Stroeven, M., Reconstruction by SPACE of the interfacial transition zone. Cement Concrete Composites, 23,2001, pp 189-200 27. Bui, D.D., Hu, J., Stroeven, P., Particle size effect on the strength of rice husk ash blended gap-graded Portland cement concrete. Cement Concrete Composites, 27/3,2005, pp 357-366 28. Zheng, J., Mesostructure of Concrete - Stereological Analysis and Some Mechanical Implications. Ph.D. Thesis, Delft University of Technology, Delft, 2000

Proc. Int. Symp. “BrittleMatrix Composites 8” A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

RHEOLOGY OF FIBER-REINFORCED CEMENT SYSTEMS USING A CUSTOM BUILT RHEOMETER Katherine KUDER’, Nilufer OZWRT’, Edward Mu3,Surendra SHAH4 ‘SeattleUniv.; Civil and Env. Eug.; Seattle, WA 98122, USA, email: kuderk(d,seattleu.cdu 21stanbul Technical Univ., Civil Eng., Ayazaga Kampusu, Istanbul, 34469, Turkey, email: owurtnilk2itu.cdu.Q Formerly at Center for Advanced Cement-Based MaterialsMorthwestem Univ.; currently at James Hardie Research; 10901 Elm Avenue; Fontana, CA 92337, USA, email: [email protected] Center for Advanced Cement Based Materials, Northwestem Univ., 2145 Sheridan Road, Evanston, IL., 60208, USA, email: s-shahkilnorthwestem.edu

ABSTRACT Fiber reinforcement can be used to enhance the mechanical performance and durability of cement-based materials. However, incorporation of fibers can have an adverse effect on the fresh state properties of these materials. To use fiber reinforcement effectively, the effect that fibers have on the rheology of cement-based composites must be understood. In this work, a custom designed and built parallel plate rheometer is used to evaluate the fresh state properties of stiff cementitious systems. Previous work showed that this rheometer produces reasonable results and that the parameters obtained with the rheometer compare well with other established values. In this work, the rheometer is used to study a variety of stiff cementitious systems. The effect of water and sand content on the rheology of cement paste is determined. Finally, the fresh state characteristics of stiff steel fiber-reinforced cement paste is evaluated and conclusions are drawn about the influences of the reinforcement on flow behavior.

Keywords Rheology, fresh state characteristics, fiber reinforcement, parallel plate rheometer

INTRODUCTION Fiber reinforcement is used in cementitious composites to improve the brittle nature of these materials and has been shown to enhance both mechanical performance and durability’”. However, inclusion of fibers can compromise the fresh state properties of these materials, which in turn can affect the efficacy of the reinforcement. Fibers can make cement-based systems difficult to mix and place, possibly leading to poor fiber dispersion or excessive voids in the matrix. To use fibers effectively, the influence that they have on the fresh state, or rheological, properties of the materials that they reinforce must be known. Fresh state characteristics are often described using rheometers, which can be used to study how a material flows, or deforms, under an applied stress. Two categories of rheometers exist: commercially-available and custom-designed and built. Commercially-available rheometers were originally developed to test polymer systems. These rheometers allow for different geometries, such as the parallel plate, coaxial cylinder and vane configurations, to be used with

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Katherine KUDER? Nilufer OZYURT, Edward Mv, Surendra SHAH

one machine by simply changing fixtures. However, sample sizes must be small and the torque capacity is low, limiting testing to relatively fluid cement paste and mortar systems, typically without fiber reinforcement. In response to these limitations, a number of researchers have developed rheometers for evaluating the rheology of highly fluid concrete systems. Custom designed and built rheometers include the BMI.,, CEMAGREF-IMG, BTRHEOM, parallel plate rheometer developed at the University of Illinois at Urbana-Champaign, IBB and Two-Point rheometer6-'*,which allow for the inclusion of larger particles, such as coarse aggregate.

As materials technology continues to advance, more complicated cement-based systems are being developed. New cement mixtures can be quite stiff with low watedcement ratios and fiber reinforcement. To be able to tailor these materials for their intended applications, the fiesh state processing characteristics need to be known. In the present research, a parallel plate rheometer is designed and built to evaluate the flow of highly stiff fiber-reinforced systems. Previously, results from the rheometer were verified and an experimental procedure was developed that yielded reproducible results. In this work, the rheology of a variety of cement systems is evaluated using the newly-developed rheometer. The effect of water content and sand addition on the flow behavior of cement paste systems is studied. In addition, the influence of steel fibers on the rheology of stiff fiber-reinforced systems is evaluated. CUSTOM DESIGNED AND BUILT PARALLEL PLATE RHEOMETER The objective of the work was to develop a rheometer that could evaluate the rheological behavior of stiff fiber-reinforced cement paste and mortar systems. With this in mind, a number of considerations were made while selecting the rheometer configuration. First, an adequate gap size was required so that fibers and sand particles could be included in the mix. Next, the occurrence of plug flow and wall slip, testing artifacts that can lead to underestimated rheological parameters, needed to be minimized. Finally, for the rheometer to be able to shear stiff materials, a high torque capacity motor was needed. Figure 1 presents the parallel plate rheometer that was built13, which consists of two round plates that have a diameter of 254 mm. The occurrence of plug flow is minimized with this geometry since the shear rate varies along the radius of the plates. Wall slip is reduced by including square grooves (6.3 x 6.3 x 2.5 mm ) that are machined onto the plates. The gap between the two plates is adjustable, allowing the inclusion of fibers and sand particles in the mix. The upper plate is rotated by a high torque capacity motor that is capable of approximately 20 N-m of torque, enabling the shearing of stiff materials. Material is prevented from flowing away during the test by a plexi-glass wall is included that surrounds the bottom stationary plate. This plate introduces a frictional effect that will influence the rheological measurements. To reduce this effect, a large diameter to gap ratio was used (typically greater than 10). In addition, in a previous work, the effect of this wall was evaluated and a procedure to minimize its influence was determined. This procedure was used in this work and is described in the Experimental Program section. The rheometer is attached to a data acquisition system for continuous monitoring during testing.

433

Rheology offiber-reinforced cement systems using a custom-built rheometer

6J

Figure 1. Parallel plate rheometer: (a) rheometer, (b) plates with square grooves and (c) velocity distributions

EXPERIMENTAL PROCEDURE Mixes were prepared in a planetary mixer and then placed into the rheometer. A rigorous mixing procedure was used: First, the cement, sand and fibers (if applicable) were placed in a Hobart mixer and mixed for 1 minute at the low speed. Next, water was added and the material mixed for 1 minute at the low speed and then there was a 30 second rest. Finally, the materials were mixed for 2 minutes at the high speed. This procedure was selected because it was found to disperse the sand and fibers well. The cement was LaFarge Type I. Short steel fibers, produced by Bekaert, which had a length of 6 mm and a diameter of 0.16 mm, were used. River sand, with a maximum particle diameter of 3mm, was used in the oven-dry condition. After each mixture was prepared, it was placed into the rheometer. To obtain a level surface and a uniform void system, a normal compression force of 2.36 kg was applied on the top of the sample as it was vibrated for 10 seconds. This procedure was found to be necessary to achieve good repeatability 1 3 . The gap height between the plates was kept between 10 and 13 mm. This range of gap heights was found to minimize the influence of the wall (most prevalent at large gap heights) and the influence of the teeth (most prevalent at small gap heights). At least three replications were made for each parameter investigated.

Katherine KUDER, Nilufer OZYURL Edward Mu. Surendra SHAH

434

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Figure 4 presents the rheological measuring protocol, which is similar to the one proposed by Geiker et. a1 14. At each shear rate, torque measurements were recorded for 20 seconds. The torque at each shear rate was then obtained by averaging the values that corresponded to the equilibrium region. To ensure that a steady-state condition was reached, the data from the highest rotation rate was neglected. Torque and rotation speed were converted to shear stress and shear rate. An example of data obtained using this procedure is shown in Figure 3 and demonstrates that there is a linear relationship between shear stress and shear rate. The wellknown Bingham relationship was used to describe the flow and is given by:

where TO is the Bingham yield stress, describing the stress needed to initiate flow, po is the Bingham plastic viscosity, which is the resistance of the material to flow, and T and Z; are the shear stress and shear rate, respectively. The yield stress was determined from the resultant shear stress versus shear rate data for the slowest two shear rates, since the yield stress corresponds to the shear stress at a shear rate of zero.

435

Rheology offiber-reinforced cement systems using a custom-built rheometer

400

~

0

5

10 15 20 Shear rate (d)

25

30

Figure 3. Example of shear stress vs. shear rate behavior with Bingham fit to data

RESULTS Figure 4 presents the Bingham parameters as a function of water content. Three different neat pastes were studied, with w/c = 0.30, 0.35 and 0.45. As is expected, both the yield stress and viscosity decrease as the w/c increases.

Figure 4. Effect of w/c on (a) yield stress and (b) viscosity of neat cement paste with standard deviations

Figure 5 presents the effect of sand addition on the rheology of cement paste with a wlc = 0.45. Sand was added at 10 and 30% of the cement weight. Addition of sand has a small influence on the yield stress, but has a significant effect on the viscosity, which increases by more than 4 times, from 0 to 30% sand.

Katherine KUDER, Nilufer OZYURT, Edward W,Surendra SHAH

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Finally, the rheology of steel fiber-reinforced cement paste was determined. Two watedcement ratios, 0.30 and 0.35,with three fiber volume contents, 1, 2 and 4%, were studied. Figure 6 and Figure 7 show the effect of steel fiber volume on the yield stress and viscosity of the cement pastes, respectively. For both w/c, the yield stress decreases until a critical volume fraction is reached, and then increases. The viscosities appear to decrease until reaching a critical point, however, the decrease for the stiffer matrix, with w/c = 0.30, is much greater. At a fiber dosage of 4%, the yield stress and viscosity are similar for both w/c.

Figure 6. Effect of fiber content on yield stress for w/c = 0.30 and 0.35 with standard deviations

437

Rheology of3ber-reinforced cement systems using a custom-built rheometer

1 *- 0 . 3 1. 0.35

0 0

1

2

3

4

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Figure 7. Effect of fiber content on viscosity for w/c = 0.30 and 0.35 with standard deviations

DISCUSSION AND FURTHER ANALYSIS The results presented here for the steel fiber-reinforced cement pastes are not what are generally expected. It is hypothesized that the initial decrease in Bingham parameters is due to the thixotropic nature of the cement paste. The stiff steel fibers might increase the amount structural breakdown that occurs during mixing, thus initially reducing the yield stress and viscosity of the material. However, once a certain critical volume fraction of fibers is reached, the mechanical interlocking, or entangling, of the fibers could be dominating the flow behavior. This phenomenon of mechanical fiber interactions has been observed with fiber reinforced polymers [25]. Thus, at lower fiber dosages, the properties of the cementitious matrices are dominant, which is why the yield stress and viscosity are lower for w/c = 0.35 than 0.30. After the critical volume fraction, the absolute values of the Bingham parameters are much closer, regardless of w/c, suggesting that the fiber interlocking is governing the behavior. It is also interesting to note that the critical volume fraction appears to be higher for the w/c = 0.30 than the w/c = 0.35, possibly indicating that a stiffer material undergoes more structural breakdown before reaching the critical point. To isolate the effect of the mechanical interaction of the fibers, the rheology of a Newtonian fluid was evaluated. Rheological measurements were determined for a commercially available honey (Sue Bee Clover Honey) with varying dosages of the steel fibers investigated previously. The results from these experiments are shown in Figure 8. From 0-2%, the change in viscosity is small; however, between 2 and 4%, a large increase in the viscosity is observed. Thus, there

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Katherine KUDER, Nilufer OZYURI: Edward W,Surendra SHAH

does appear to be a point at which the mechanical interlocking of fibers dominates the flow behavior for both Bingham and Newtonian liquids.

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CONCLUSION

A parallel plate rheometer was designed and built to evaluate the rheology of stiff fiberreinforced cement pastes. Previous work verified that the rheometer worked properly. In the present study the rheology of a variety of cement paste systems was studied, including stiff steel fiber-reinforced cement pastes. As expected, the research showed that the yield stress and viscosity of cement paste systems increased as the water content was reduced and as the sand content was increased. In addition, the influence of steel fibers on the rheology of stiff cement pastes systems was studied. Contrary to what was expected, the Bingham rheological paraemeters decreased as the fiber content increased until a critical volume fraction was reached. This trend is explained by a coupling effect between the structural breakdown of the material, which occurs at low fiber volumes, and the mechanical interlocking of the fibers, which occurs at high volumes. ACKNOWLEDGEMENTS

The authors gratefully acknowledge the financial support of Lafarge and WR-Grace. This work was also funded by NSF PATH Grant no. CMS-0122045. The second author also would like to

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acknowledge the financial support of TUBITAK (The Scientific and Technical Research Council of Turkey) and ITU (Istanbul Technical University). Fibers were provided by Bekaert. REFERENCES 1. 2. 3. 4. 5.

6. 7. 8. 9. 10. 11. 12. 13.

14.

Nawy, E.G., Fundamentals of High-Performance Concrete. 2nd ed. 2001: John Wiley & Sons, Inc. 464. Balaguru, P.N. and S.P. Shah, Fiber-Reinforced Cement Composites. 1992, New York: McGraw-Hill Inc. Lawler, J.S., D. Zampini, and S.P. Shah, "Permeability of Cracked Hybrid FiberReinforced Mortar under Load". ACI Materials Journal, 2002.99(4): p. 379-385. Voigt, T., K.B. Bui, and S.P. Shah, "Drying Shrinkage of Concrete Reinforced with Fibers and Welded-Wire Fabric". ACI Materials Journal, 2004. lOl(2): p. 233-241. Banthia, N., S. Mindess, and J. Trottier, "Impact Resistance of Steel Fiber Reinforced Concrete". ACI Materials Journal, 1996. 93(9): p. 472-479. Coussor, P., Rheologie des Boues et Laves Torrentielles - Etudes de Dispersions et Suspensions Concentrees. 1993, L'Institut National Polytechnique de Grenoble, et Etudes du Cemagref. p. 418. Wallevik, O.H. and O.E. Gjorv. "Development of Coaxial Cylinder Viscometer for Fresh Concrete". in Properties of Concrete, RILEM Colloqium. 1990. Hanover: Chapman and Hall. Hu, C., et al., "Validation of BTRHEOM, the New Rheometer for Soft-to-Fluid Concrete". Materials and Structures, 1996. 29: p. 620-631. Struble, L.J., U. Puri, and X. Ji, "Concrete Rheometer". Advances in Cement Research, 2001. 13(2): p. 53-63. Beaupre, D., Rheology of High Performance Shotcrete, in Civil Engineering. 1994, University of British Columbia: British Columbia. Banfill, P.F.G. and G.H. Tattersall, The Rheology ofFresh Concrete. 1st ed. 1983, Boston: Pitman Advanced Publishing Program. 356. Tattersall, G.H. and S.J. Bloomer, "Further Development of the Two-Point Test for Workability and Extension of Its Range". Magazine of Concrete Research, 1979.31: p. 202-2 10. Kuder, K.G., et al. "The Rheology of Fiber-Reinforced Cement Paste Evaluated by a Parallel Plate Rheometer". in Advances in Concrete Through Science and Engineering. 2004. Northwestern University, Evanston, IL, USA: International Union of Laboratories and Experts in Construction Materials (RILEM). Geiker, M.R., et al., "The Effect of Measuring Procedure on the Apparent Rheological Properties of Self-Compacting Concrete". Cement and Concrete Research, 2002.32: p. 1791-1795.

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, V.C. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

THE INFLUENCE OF CEMENT PASTE VOLUME IN MORTAR ON THE RHEOLOGICAL EFFECTS OF THE ADDITION OF SUPERPLASTICIZER Jacek GOLASZEWSKI Silesian University of Technology, Faculty of Civil Engineering Akademicka 5,44 - 100 Gliwice, Poland e-mail: [email protected]

ABSTRACT The results of investigation into the influence of polyether type superplasticizer addition on rheological parameters and its changes with time of different in cement paste volume fresh mortars are presented and discussed in the paper. Rheology results have been evaluated according to the Bingham model, which describes the rheology with the parameters: yield value and plastic viscosity. Rheological parameters were measured using Two-Point Workability Test. Because of the similar nature of rheological behaviour of fresh mortar and concrete, presented in paper relationships for mortars can be also used to state the rheological properties of fresh concrete. The obtained results show that the character and range of superplasticizer influence on rheological parameters of mortar is strongly influenced by cement paste volume. Volume of cement paste is also a significant factor affecting range and direction of changes of the rheological parameters of superplasticized mortars with time. The results show, that the character of influence of superplasticizer on rheological parameters of fresh cement paste and on fresh mortar may significantlydiffer each from other. These differences increase with decreasing volume of cement paste in mortar. On the ground of obtained results the empirical relations joining the rheological parameters of mortars with superplasticizercontent were established. Analysis of results covers also comparison between influence of cement paste volume on rheology of mortars with and without superplasticizer. It is concluded that effects of superplasticizer addition on rheology of fresh mortar and concrete should be also considered in respect to cement paste volume. It is difficult to unequivocally predict effects of superplasticizer content on rheology of mortar or fresh concrete basing only on tests made on cement paste, and not making allowance on degree of aggregate filling by cement paste in this mortar or concrete.

Keywords Rheology, Workability, High -range water reducers, Mortar INTRODUCTION When using superplasticizers it is possible to produce flowing and self compacting mortars and concretes of low wlc and, in effect, to produce mortars and concretes of high strength and durability. More effective use of superplasticizersrequires a creation of the systematised data base concerning their influence on rheological parameters of fresh mortars and concrete depending of composition of these materials and of their components properties [l-31. It is well documented that fresh mortar and concrete exhibit the Bingham viscoplastic behaviour according to the formula:

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Jacek GOLASZEWSKl

where z (Pa) is the shear stress at shear rate (Us) and z (Pa) and qpl(Pas) are the yield value and plastic viscosity respectively [4-lo]. The physical interpretation of yield value is that of the stress needed to be applied to a material in order to start flowing. When the shear stress is higher then yield value the mix flows and its flow resistance depends on plastic viscosity. In a recent years, a lot of studies concerning the influence of superplasticizers on fresh cement paste, mortar and concrete rheology were done. It was stated, that the effects of superplasticizer addition on rheology of these materials depend mostly on cement and superplasticizerphysicochemical properties, superplasticizer content, time of superplasticizer addition, and on w/c [ 1-4,6,9,11]. However, obtained relationships of superplasticizers influence on rheological properties of cement paste, mortar and fresh concrete not always are unequivocal. Superplasticizersreduce both the yield value and the plastic viscosity of cement paste [2]. They also reduce the yield value of mortar and fresh concrete, but either reduce, not change or increase plastic viscosity of these materials [6,9,11]. Analysis of so far executed studies indicates, that different effects of superplasticizeraddition on rheological properties of mortars and fresh concrete may be related to the cement paste volume content in these materials. The methodology and the results of investigation into the influence of superplasticizer content on the rheological parameters and its changes with time of different in cement paste volume fresh mortars are presented and discussed in the paper. It is worth to notice, that it was stated in [4,5,10,11] that the character of rheological behaviour of mortar and concrete is similar, and that mortars can be considered to be a model concrete. Thus, the relationship presented in the paper relate both to mortars and to fresh concrete.

EXPERIMENTAL Testing program In the research the influence of cement paste volume in mortar and superplasticizercontent on rheological properties of mortars was investigated. Cement paste volume in mortar was expressed in the terms of factor of uncompacted sand filling by cement paste - we.Methods of computing the factor cpdpare presented in details in existing literature [6,12,13]. Factors taken into consideration and its levels are shown in Table 1. Cement type, sand type and grading, and wlc ratio ( = 0.35)were kept constant. For comparison between influence of cement paste volume on rheology of mortars with and without superplasticizer, the influence of cement paste volume and wlc on rheological properties of mortars without superplasticizer was also investigated. The factors taken into consideration in this part of research and their levels are shown in Table 2; cement type, sand type and grading were kept constant. Waterlcement ratios were selected so as to obtain similar value of g for mortars of certain sandcement ratio both with and without superplasticizer. Rheological parameters of cement pastes filling mortars were not determined because of its high gravitational and centrifigal separation. Measurements of rheological parameters of fresh mortars Rheological parameters of mortars were measured according to the two-point workability test methodology using Viskomat PC rheometer. The origin and principles of the two-point workability test are presented in [4]. Viskomat PC and its measuring element are presented in Fig 1; it is in detail described in [6,9]. In general, in the two-point workability test rheological parameters of fresh mortar are measured by applying a given shear rate and measuring the resulting shear stress. Because of the Bingham nature of rheological behaviour

The influence of cement paste volume in mortar on the rheological effects of the addition of supeplasticizer

443

Table 1. Program of testing of cement paste volume and superplasticizer content influence on rheological properties of mortars Factor qdp 1.24 1.89 2.52 3.77

Superplasticizer content 2.0,2.5, 3.0,3.5 1.0, 1.5,2.0,2.5 0.5, 1.0, 1.5,2.0,2.5 0.5, 1.0, 1.5,2.0,

WaterIcement

Sandcement

0.35

311 2/ 1 1.511 1/1

Table 2. Program of testing of cement paste volume and w/c ratio influence on rheological properties of mortars without superplasticizer Factor qdD 1.46 - 1.68 2.22 - 2.48 2.85 - 3.30 4.28 - 4.78

Waterlcement 0.44,0.47,0.50,0.53 0.44,0.47,0.50,0.53 0.47,0.50,0.53,0.56 0.50,0.53, 0.56

Sandcement 311 211 1.511 111

of fresh mortars, the measurements should be taken at no less than two considerably different shear rates. The rheological parameters are determined by regression analysis according to the relation:

where T is the shear resistance of sample measured at rotation rate N and g (N.mm) and h (N.mm.s) are constants corresponding respectively to yield value z,,and plastic viscosity qpl. By suitable calibration of rheometer it is possible to express g and h in fundamental units. According to [5] in the Viskomat PC rheometer zo = 7.9 g and qpl = 0.78 h, but all results are given below in terms of g and h.

Materials and mixes CEM I 32.5 cement and polyether based superplasticizer were used for the investigations. Their main properties are given in Tables 3 and 4. The sand used was PN EN 196:1996 CEN model sand (2 mm max.). The mix proportions of tested mortars are presented in Tables 1 and 2. The superplasticizerdosage relate to the total mass of liquid superplasticizer. Mortar mixing and testing procedures The mixing procedure reminded in accordance with EN 196-1:1994; superplasticizer was added with water. Rheological parameters of mortars were measured at 10, 30 and 60 min after the end of mixing according to the procedure shown in Fig 1. Table 3. Properties of cement CEM I 32.5

Table 4. Properties of superplasticizer Major constituent polycarboxylate acid

Density Concentration [g/cm3] [%] 1.06 40

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Jacek GOtASZEWSKI

Fig. 1. Viskomat PC, its measuring element, and measuring procedure

RESULTS AND DISCUSSION Obtained relations of influence of superplasticizer content on rheological parameters of fresh mortars different in cement paste volume (in factor qdp)are presented in Fig. 2. The analysis of variance ANOVA of influence of cement paste volume (factor wP) and superplasticizer content on rheological parameters of mortars are presented in Tables 5 and 6. It can be seen, as one can expect, that cement paste volume in mortar (factor 9dp),superplasticizer content, and interaction of these factors significantly influences parameters g and h of mortars and its changes with time. Increasing content of superplasticizer, likewise increasing cement paste volume (increasing factor wP) causes a non-linear decrease in g of mortars. Initially increasing superplasticizer content causes fast decrease of parameter g until a certain minimum. Further increasing of superplasticizercontent no longer changes parameter g. Such nature of changes in g does not depend on cement paste volume in mortar (on factor wP). Nevertheless, cement paste volume in mortar (factor qdp)influences strongly range of changes in value of g caused by increasing superplasticizercontent. As may be seen fiom obtained results, for mortars with small cement paste volume changes of g when superplasticizers quantity increases are great and reduce as cement paste volume in mortar increases. Simultaneously, the lesser is cement paste volume in mortar (factor wP), the higher superplasticizer addition is necessary to obtain mortar of given value of g. It is also worth to notice, that if cement paste volume in mortar is small, obtainment of mortar of similar value of g as for mortar with high content of paste is difficult, even if quantity of added superplasticizeris very large. Value of g of all tested mortars, independently on cement paste volume (factor wP) and superplasticizer content, increases with time. The range of increase of g with time is lesser when superplasticizercontent andor cement paste volume in mortar (factor wP) are larger. For example, increase in value of g of mortars of wP= 1.89 with 1.0, 1.5, 2.0, 2.5, and 3.0% addition of superplasticizer within 50 minutes is correspondingly of 31.4, 15.3, 10.7, 7.4, and 5.8 N.mm. It means that it increases in comparison to initial value of g by correspondingly of 96, 86, 79, 69 and 61%. For mortars of wP = 1.89, 2.52, and 3.77 with 1% addition of superplasticizer increase in value of g within 50 min is correspondingly of 31.4, 5.8 and 1.9 N.mm. It means that it increases in comparison to initial value of g by correspondinglyof 96,54 and 32 %. Nature of the influence of superplasticizer content on h of mortars, in opposition to character of influence on g, depends on cement paste volume in mortar (on factor qdP).

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Table 5 . Analysis of variance ANOVA of influence of cement paste volume in mortar and superplasticizer content on g of mortars after 10,30 and 60 min after end of mixing

Source of variation A: Cement paste volume B: SP content AB

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In the case of mortars of larger cement paste volume = 2.52 and 3.77) increase of superplasticizer content initially slightly increases value of h and next, after reaching some maximum value, value of h decreases. Range of such changes increases as cement paste volume in mortar increases, but generally is not too great and only in limited extent influences rheological properties of mortars. In the case of mortars of lower cement paste volume (qdp= 1.24 and 1.89) increase of superplasticizer content causes large (larger if cement paste volume in mortar is lower) increase in the value of h. Maximum value of h is reached when superplasticizer content is very high - further addition of superplasticizer may develop only small decrease in value of h. Maximum value of h and superplasticizer content when such value is reached, depends on cement paste volume in mortar. The lower the cement paste volume in mortar, the higher maximum value of h reached, and it occurs at higher superplasticizer content. It should be mentioned that h of mortars with a given superplasticizer content in general increases when cement paste volume in mortars is decreased. Use of smaller cement paste volumes facilitates obtainment of mortars of higher value of h. It is essential in case of self compacting mortars

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Direction of changes of h of mortars with time depends first of all on volume of cement paste in mortar and, in the case of mortars with low cement paste volume, also on quantity of superplasticizer. In the case of mortars with larger cement paste volume (of 'pdp = 2.52 and 3.77), the value of h increases as a time passes. It should be mentioned that scope of such increase to a lesser extend depends on superplasticizer content. For example, increase in the value of h within 50 min for mortars of qdp= 1.89 with 1.0, 1.5, 2.0, 2.5, 3.0% superplasticizer content is correspondingly of 4.2, 3.9, 3.8, 3.4 and 2.8 N.mm.s, what means that its increase is in comparison to initial value of h correspondingly of 29,28,22,20 and 18%. In the case of mortars with lower cement paste volume (of = 1.24 and 1.89) direction of h changes with time depends on superplasticizer content. Initially, when superplasticizer content is low, value of h decreases with time. Range of such decrease is lower when superplasticizer content increases. After exceeding some specific superplasticizer content (the higher, the lower cement paste volume in mortar) value of h of mortars shows increase with time. For example, increase of the value of h within 50 min for mortars of = 2.52 with

wP

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The influence of cement paste volume in mortar on the rheologikal effects of the addition of supeplasticizer

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1.O, 1S,2.0, 2.5, and 3.0 % superplasticizer content is accordingly of -3.0, -3.3, -0.8, 4.0, and 6.1 N.mm.s what means change in comparison to the initial value of h by correspondingly of -17, -15, -3, 14 and 22%. It should be mentioned that in case of mortars of low cement paste volume (of (pmip = 1.24), even in case of 3% superplasticizer content, value of h decreases as time passes. Presented above results are unfortunately insufficient for formulation of general relations connecting the values of g and h parameters and superplasticizer content and factor of uncompacted sand filling by cement paste However, the regression analysis of obtained test results shows that changes of g and h of mortars with changing superplasticizer content may be described using the following relations:

wP.

where: SP - superplasticizer content; A,, B,, Ah, Bh, c h - material constants depending mainly of cement paste volume in mortar and sand graining (factor (pdp), of properties of cement and superplasticizer, and of time. Values of material constants and correlation coefficients for tested mortars are presented in Table 7. Obtained relations of influence of w/c on rheological parameters of fresh mortars different in cement paste volume (in factor qdp)are presented in Fig 3. The analysis of variance ANOVA of influence of cement paste volume in mortar (factor qdP)and w/c on rheological parameters of mortars are presented in Tables 8 and 9. It can be seen, that cement paste volume in mortar (factor (pdp), w/c, and interaction of these factors significantly influences g and h of mortars and its changes with time. Both rheological parameters of mortars decrease as cement paste volume in mortar (factor qdp)and wlc increase, and range of such changes is larger when cement paste volume in mortar is lower. As time passes, both g and h increase independently of w/c and cement paste volume. Only mortars with low paste volume (factor qdp)and low w/c indicate slight decrease in value of h with time. Range of g changes with time decreases, while range of h changes increases if w/c increases. Comparing relations for mortars with and without superplasticizer, one may say that nature of increasing superplasticizer content influence on value of g is analogical as nature of increasing w/c influence, while nature of superplasticizer content influence on h is other than nature of wlc influence. Increase of w/c causes decrease in the value of h, while increasing superplasticizer content causes initially distinct increase in value of h and then, after exceeding specific superplasticizer content, value of h decrease. At a given value of g, value

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The influence of cementpaste volume in mortar on the rheological effects of the addition of superplasticizer

449

Table 8. Analysis of variance ANOVA of influence of cement paste volume in mortar and w/c ratio on g of mortars after 10,30 and 60 min after end of mixing g after 10 min

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AB

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sig.

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of h for mortar without superplasticizer is always distinctly lower than the value of h for mortar with lower w/c, but with superplasticizer addition. Nature of influence of increasing wlc on rheological parameters may be explained by simultaneous increase of cement paste fluidisation and increase of aggregate’s grains dissipation resulting from cement paste volume increase. Increase of superplasticizer content in mortar of given wlc causes de-flocculation of cement grains and increase of free water volume in mortar, but without essential influence on aggregate’s grains dissipation. Larger volume of free water causes both fluidisation of paste and reduction of capillary cohesion in mortar [ 6 ] .These effects cause decrease in value of g of mortar. The nature of such changes may be interpreted as apparent increase of cement paste volume in mortar. Decrease of capillary cohesion does not influence considerably h of mortar [6].Therefore changes in value of h when superplasticizer content is increased depends on changes in cement paste rheological properties and free water volume. Clarification of such changes mechanism requires further studies. Greater aggregate’s grains concentration in mortar explains, why for similar value of g mortars with superplasticizer of low wlc are characterised by greater value of h than for mortar without or with lower superplasticizer quantity but of greater wlc. Changes in rheological parameters with time of mortars of given w/c and superplasticizer content depend on cement paste volume in this mortar and are greater when its volume is lower. Basing on [6] is possible to say, that in the same time occur effects of decrease in free water volume in cement paste and of gradual decay of superplasticizer action in result of cement hydration, and of increase of capillary cohesion in mortar. In mortar with lower paste volume, and therefore lower grade of dissipation of aggregate’s grains decrease of free water quantity leads to quicker increase of capillary cohesion of mortar, and therefore to quicker increase in the value of g. CONCLUSIONS Range, and in the case of h parameter, also nature of superplasticizer content influence on rheological parameters of mortars depend on cement paste volume in mortar. Increasing superplasticizer content causes fast decrease in value of g of mortars until a certain minimum. The minimum value of g possible to obtain due to increasing superplasticizer content increases with decreasing cement paste volume in mortar. Value of g of superplasticized mortars increases with time and the range of this increase is clearly lesser when cement paste

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Jacek GOLASZEWSH

volume in mortar is higher. Increase of superplasticizer content first causes increase in value of h of mortars. In the case of mortars with high cement paste volume such increase is low, its range increases as cement paste volume in mortar decreases and in the case of low cement paste volume mortars it is very high. In the case of mortars with high cement paste volume value of h reaches specific maximum value and then, as superplasticizer content further is increased, its value decreases. Such effect is weak or does not occur in case of mortars with low volume of cement paste. Value of h of mortars with higher cement paste volume increases with time, scope of such changes to a lesser degree depends on superplasticizer content. In the case of mortars with lower cement paste volume direction of changes in value of h with time in such mortars depends on superplasticizer content. If superplasticizer content is low, value of h of mortars decreases with time. After exceeding a particular superplasticizer content, larger when quantity of cement paste in mortar is lower, value of h of mortars increases with time. Nature of influence of increasing superplasticizer content and increasing wlc on the value of g of mortars is similar. In the same time such influence on the value of h may be distinctly different. Due the above, it is not possible to set out such composition of mortars without and with superplasticizer that for similar value of g will be characterised by the same value of h. Effects of superplasticizer addition on rheology properties of fresh mortar and concrete should be considered in respect to cement paste volume. It is difficult to unequivocally predict effects of superplasticizer content on rheology of mortar or fresh concrete and its changes with time basing only on tests made on cement paste, and not making allowance on degree of aggregate filling by cement paste in this mortar or concrete. REFERENCES [ 11 Aitcin, P-C. High Performance Concrete, EF&N SPON, London, 1998. [2] Ramachandran, V S. Concrete Admixtures Handbook. Properties, Science and Technology. 2nd Edn, Noyes Publications, Park Ridge, USA, 1995. [3] Neville, A M. Properties of Concrete (in Polish), Polski Cement, Krakbw, 2000. [4] Tattarsall, G H, Banfill, P F G.The Rheology of Fresh Concrete, Pitman Books Limited, Boston, 1983. [5] Banfill, P F G. The rheology of fresh mortar. Magazine of Concrete Research. Vol. 43 (154), 1991. pp 13-21. [6] Szwabowski, J. Rheology of mixes on cement binders (in Polish), Wydawnictwo Politechniki Slqskiej, Gliwice, 1999,239 pp. [7] Ferraris Ch. F.: Measurement of the Rheological Properties of High Performance Concrete: State of Art Report. Journal of Research of the National Institute of Standards and Technology, Vol. 104, No. 5, 1999,461 - 478. [8] Larrard de F.: Concrete Mixture Proportioning. A scientific approach. E&FN SPON, London and New York 1999. [9] Golaszewski, J. Adjusting of fresh concrete workability using superplasticizers, (in Polish), Wydawnictwo Politechniki Slqskiej, Budownictwo z. 99, Gliwice, 2003,2 16. [10]Banfill P.F.G.: The rheology of fresh cement and concrete - a review. Proceeding of 1lth International Cement Chemistry Congress, Durban, South Africa 2003, 50 - 63. [ 111Golaszewski, J. Rheology of mortars and rheology of fresh concrete. Cement Wapno Beton, No. 1,2006, 17-28. [12] Szwabowski, J. Rheology of self compacting fresh concrete (in Polish). TV Sympozjum Naukowo - Techniczne Reologia w Technologii Betonu, Gliwice, Poland 2002,61-75. [13]Kuczyhski, W. Technology of concrete. Part 2. Concrete proportioning (in Polish). Arkady, Warszawa 1972.

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, l?C. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

THE RHEOLOGICAL PROPERTIES OF FRESH STEEL FIBRE REINFORCED SELF-COMPACTING CONCRETE Tomasz PONWEWSKI The Silesian University of Technology, Department of Building Processes Akademicka 5,44- 100 Gliwice, Poland,e-mail:[email protected]

ABSTRACT In this paper the methodology and test results of the investigation are presented the influence of steel fibres on rheological and mechanical properties of Steel Fibre Reinforced Self-Compacting Concrete (SFRSCC) are discussed. The rheological parameters of SFRSCC - behaves as a Bingham body, where the rheological parameters yield value g and plastic viscosity h were determined to using a new kind of rheometer BT2 to mortar and concrete mix research. The mechanical parameter of SFRSCC - the cube compressive strength& is presented as well. In the research, an experimental verification of and significance of an influence: volume fraction of fibres, fibres factor, lengths and the shape of fibres on rheological properties of SFRSCC was investigated. In this paper the results obtained for the mixes with three types of steel fibre shapes are presented. Concrete mixtures are proportioned to provide the workability needed and the required properties in the hardened concrete. The length of fibres does not have the significant influence on yield value g and plastic viscosity h of SFRSCC. The significant influence of the length of fibres on plastic viscosity h of tested hooked steel SFRSCC was only observed. The rheological properties of SFRSCC from workability point of view are better for SCC than with other types of fibres.

Keywords Steel fibres, rheology, self-compacting concrete, workability INTRODUCTION Technology of self-compacting concrete allows the shaping structure of engineering objects in quicker and safer way than in cases of concrete with traditional properties. The technological operations of concrete elements forming, in case of self-compacting concrete, are considerably simplified and the end results allow to expose hardened concrete structures in more extended ways [l]. One modification of considered concrete is to add to its volume various kinds of fibres as dispersed reinforcement [2]. This is not a new issue in the technology of concrete, however, in case of the self-compacting concrete, it provides an area of research. Problems resulting from using such modified concrete mixes were determined based on carried out tests of workability of fresh self-compacting concrete mix modified with steel fibres in rheological context. Technological problems in applying self-compacting concrete modified with steel fibres as dispersed reinforcement is the subject of the present article. Analysing of the influence of fibres on workability and durability of the parameters of concrete is one of new tendencies in research of self-compacting concrete [3-71. The research upon of the influence steel fibres of various geometric parameters were presented to determine the impact of its volume fraction, the length and the shape on rheological and mechanical properties of self-compacting concrete. The general tendency of the improvement of hardened self-compacting concrete characteristics with the increase of contents of fibres in

452

Tomasz PONIKEWSKI

its volume, makes workability of these concrete mixes worse during forming. The current problem, also in case of self-compacting concrete modified with steel fibres, is a technological difficulty of its production and carrying out of technological processes in concrete works. It compels to recognise the real nature of workability and to determine the impact of added fibres on phenomena taking place in fresh and hardened self-compacting concrete. ASSUMPTIONS AND METHODOLOGY OF RESEARCH

Results of workability tests of self-compacting cement mixes modified with steel fibres in rheological context are presented in this paper. Testing carried out with method of rheometrical of workability test (RWT) were conducted with a rheometer for mortars and concrete mixes - BT-2. (Fig. 1).

a) b) Fig. 1. Rheometer BT-2 to determine rheological parameters of concrete mixes a) general view of the apparatus during the measuring procedure; b) readout and the verification of results of the rheometrical measurement. The approximation of measurement results conducted by two-parameter Bingham rheological model and three-parameter Hershell-Bulkey model were done. It allowed to determine two basic rheological parameters - yield value g and plastic viscosity h. The values were determined by two-parameter model. Composition of the tested self-compactingmixture is presented in table 1.

The rheological properties offish steel fibre reinforced self-compacting concrete

453

The concrete mix was modified - variable kinds and volume fraction of steel fibres were used. Steel fibres were selected out of a large number of fibres available on the market. Results of testing of self-compacting mixtures modified with eleven kinds of steel fibres are presented in this article. In the first step tests were carried out for the variable volume fraction of fibres in the matrix. In the second step avariable level of the fibre reinforcement was examined for the fibre factor FF [8] taking geometric parameters of fibres into consideration (length Lfand diameter dr> as well as fibre volume fraction V , in the mixture, according to the Taking the level of the fibre reinforcement into consideration in following pattern (V,*Lf/df). testing (FF)allows to determine the influence of each parameter that characterise the used diffused reinforcement on workabilty of self-compacting mixtures in rheological context in more reliable way. In first step, tested fibre volume fraction (5)in the concrete mixture was 0.5-1.0-1.5-2.0 % which corresponds to the 39.25-78.50-1 17.75-157.00 kg/m3 contents. In second step, the levels of variability (FF) was considered 0.2-0.4-0.6-0.8, which corresponds to fibre mass that is subject to slenderness of fibres, as presented in table 2.

Geometric characteristics of tested fibres and fibre volume fraction in concrete mixture according to the level of fibre reinforcement were presented in table 2. The shape of fibres due to the variability of their geometry is an additional factor influencing test results but overlapping with considered remaining variable parameters of fibres. RESULTS OF TESTS AND DISCUSSION Properties of self-compacting mixtures modified with steel fibres were tested to determine rheological parameters measured with RTU method. On the basis of pre-examinations, determining relationship between the time and the flow diameter measured with Abram’s cone method, an estimated self-compacting limit was determined for tested mixtures with steel fibres, according to the assumption: flow time T5o = max 9 seconds and the flow diameter R = min. 600 [mm]. The above mentioned assumptions of the self-compacting limit were obtained for maximal yield value g on the level 600 [Nmm]. Any plastic viscosity level (h) as a limit for self-compacting mixtures with steel fibres was unambiguously determined. Figures 2 and 3 present the influence of type and volume fraction of straight steel fibres on rheological parameters of self-compacting mixtures - yield value g and plastic viscosity h value.

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a> b) Fig. 5. Influence of kind and fibre factor of wavy steel fibres on: a) yield value g, b) plastic viscosity h value Conditions of self-compacting was obtained for all considered wavy fibres in the whole range of variability of volume fraction. In case of the factor (FF), self-compacting limit for all wavy fibres was level 0.6. Influence of the hooked steel fibres on rheological parameters of self-compacting mixtures was presented in figures 6 and 7. The increase of yield value g and plastic viscosity h value along with the increase of hooked fibres volume fraction in self-compacting mixtures was shown. In this research group, addition of 64x030 fibres to the mixture resulted in the biggest increase of g and h parameters and limitative fulfilling of self-compacting condition for volume fraction 0.5. Similar parameters were obtained for fibres 30~0.5.Addition of 60~0.65 fibres to the mixture resulted in the smallest increase in the g parameter. Condition of selfcompacting was obtained for volume fraction close to 1.0%. All tested concrete-mix with hooked fibres, except of the discussed above 64x030 fibres, fulfilled self-compacting limit within the whole range of fibre reinforcement. On the basis of carried out tests it is possible to feature estimated brackets of properties of self-compacting mixtures with steel fibres of various geometrical parameters and volume fraction. Table 3 presents limit of properties of self-compacting for variable volume fraction of fibres and fibre factor together with fibres weight quantity.

~

Tomasz PONIKE WSKl

456

1

Hooked fibre5 2000 1800

1200

5 1000

7

800 0 600

x3OxO.5

400 200 n

0,O

0.5 1,0 1,5 2.0 Flbre volunn Fraction [Q

0.0

2.5

0,5

1.0

1.5

2.0

2.5

Flbn volume Indion l%l

b) a) Fig. 6. Influence of kind and volume fraction of hooked steel fibres on: a) yield value g, b) plastic viscosity h value I

Hooked f l b w

2000

Hookedflbrer

1800

p 1600 E 1400 1200 2 1000

4 7 0 E’

+50

800 600 400

x 0.45

+Mx0,5

tBox 0.80

200 0,o

0,2

0,6

%re fictor

08

1,0

0,O

0.2

0.4

0.6

0.8

1.0

Flbn hctor

~~

a) b) Fig. 7. Influence of kind and fibre factor of hooked steel fibres on: a) yield value g, b) plastic viscosity h value Table 3. Limit of DroDerties of self-compacting for variable volume fraction (Vj) and fibre

Lack of self-compacting effect of mixtures in the whole considered first step of research of added straight fibres 13~0.16was shown. Total self compacting effect of mixtures modified

The rheologicalproperties offesh steelfibre reinfoxed self-compacting concrete

457

with steel fibres was stated in case of two types of wavy fibres 30x0.7 - 50x1.0 and straight fibres 6~0.16.For these two types of fibres the tests were not carried out in the second step of research. There is a lack - insignificant though - of consequence in results of tests. Straight 13~0.16fibres in the first step of research have not indicated any self-compacting effect within the total research area, however in the second step these properties were kept up to FF value 0.4 i.e. for 42 kg/m3. Wavy fibres 50x1.0 fibres in the first step have indicated selfcompacting effects within the total research area i.e. maximum 157 kg/m3, however in the second step these properties were not kept for FFvalue 0,4 i.e. for 125.6 kg/m3.Hooked fibres 30x0.5 were the last incorrect case. In the first step they indicated self-compacting properties for V, = 1.0% i.e. at most for 78.5 kg/m3, however in the second step, self-compacting properties were indicated within the whole considered research area i.e. even for 104.7 kg/m3. Any impact of the length of fibres on changes of rheological parameters of the considered modified mixtures were not ambiguously determined. It is possible on the basis of fig. 8 to determine the impact of the kind of steel fibres on the value of yield value g of self-compacting mixtures and compressive strength &, for the volume fraction 2 %, and for the fibre factor 0.8.

I

C o l i p r e Z &th

f:wPa]trVp&h

'O

1I

10 20 30 40 50 60 Gampreriw atnngth f, CPa] for F, = 0.8

70

a) b) Fig. 8. Influence of steel fibres on yield value g and compressive strengthf,, a) for volume fraction 2%, b) for fibre factor 0,8 Addition to self-compacting mixture of steel fibres indicates an increase of yield value g for all considered kinds of fibres but compressive strength increased only in two fibre addition 50~0.45and 30~0.7.Fibres 25~0.40for V,= 2.0%were characterised by the biggest increase of yield value g at the invariablef, value. It confirms the necessity of carrying out broader and more reliable research on the influences of steel fibres on self-compacting and concrete mechanical properties. Wavy fibres 30x0.7 were characterised by the smallest increase of yield value g i.e. smallest deterioration of workability at an unambiguous increase of compressive strength. CONCLUSIONS The presented results of testing on self-compacting concrete modified with steel fibres show that the influences of fibres addiction worsen workability of fresh mixture and increase compressive strength of hardened SFRSCC made out of self-compacting mixtures. To keep the self-compacting effect of mixtures modified with steel fibres, the volume fraction of 2.0% seems to be recommended to ensure its maintenance. This in not however the case with all of the fibres taken under consideration. The number of possibilities to apply steel

45 8

Tomasz PONIKIE WSKl

fibres to ensure self-compacting effects increases along with the decrease of fibres volume fraction but simultaneouslyprobability to improve mechanical properties drops down. Problems occur with homogenous filling of concrete volume with the added fibres and the required technological processes for this type of concrete make the keeping homogenous structure even more difficult. Pomped SFRSCC should be delivered directly to a forming place, with the limiting of horizontal relocation of mixtures within formed concrete structure. The slenderness and volume fraction of steel fibres in the mixture worsens its workability but improves strength parameters, though not for all fibres. Keeping the homogeneity of steel fibres during the process of self-compacting concrete forming is the current research problem. It seems recommendable to carry out broader research to determine influence of steel fibres on properties of fresh and hardened self-compacting concrete based on variability of so called fibre factors. Taking workability under consideration, it seems to be proper to add shorter fibres with higher volume fraction into concrete mixture. This should ensure homogeneity of formed concrete structure. The influence of added fibre shapes, important from fibres anchorage energy in selfcompacting concrete matrix, has not been unambiguously determined in the research. Currently, the author conducts research of relationship between energy to draw fibres out of the concrete matrix and fibres geometric parameters, as well as research of the influence of real distribution of diffised reinforcement on concrete compressive strength parameters. It is necessary to remember the diversified shape of tested fibres together with their diversified slenderness. It is recommendable to carry out additional research to eliminate overlapping of variable factors.

REFERENCES 1. Kaszyliska M.: Mix design of the self-compacting concrete, in: Proc. Int. Symp. 'Brittle Matrix Composites 7', A.M.Brandt, V.C.Li, I.H.Marshal1, Warsaw, 2003,33 1-338. 2. Brandt A.M.: Applying fibres as reinforcement in the concrete elements, Conference: Concrete on the Threshold of the New Millennium, Cracow, Poland, 2000,433-444. 3. Ambroise J., Rols S., Pera J, Properties of self-levelling concrete reinforced by steel fibres, High Performance Fibre Reinforced Cement Composites (HPFRCC 3), Maim, W E M publications PR06, Cachan, 1999. 4. Griinewald S., Walraven J.C., Rheological measurements on self-compacting fibre reinforced concrete, Third Int. Symposium on SCC, Reykjavik, Edited by Wallevick and Nielsson, RILEM publications, PRO 33, Bagneux, 2003. 5. Barraghn B., Zerbino R., the R. ghetto, Soriano M., de la C. Cruz, Giaccio G., Bravo M.: Development and application of steel fibre reinforced self-compacting concrete, 6th RILEM Symposium on Fibre-Reinforced Concretes (FRC) - BEFIB 2004, Varenna, Italy, 457 - 466. 6. Ding Y., Thomaseth D., Niederegger Ch., Thomas A., Lukas W.: The investigation on the workability and flexural toughness of fibre cocktail reinforced self-compacting high performance concrete, 6th RILEM Symposium on Fibre-Reinforced Concretes (FRC) BEFIB 2004, Varenna, Italy, 467 - 478. 7. Griinewald S., Performance-based design of self-compacting fibre reinforced concrete, PhD-thesis, Technische Universiteit Delft, Netherlands, 2004. 8. Groth P., Nemegeer D., The use of steel fibres in self-compacting concrete, First Int. Symposium od SCC, Stockholm, Edited by Skarendahl and Petersson, RILEM publications PR07, Cachan, 1999.

Proc. Int. Symp. “BrittleMatrix Composites 8” A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

CORRELATION BETWEEN BENDING RESISTANCE OF EPOXY COMPOSITE SPECIMENS MAINTAINED IN WATER AND IN PETROLEUM Tatiana LYASHENKO”, Vitaly VOZNESENSKY” Alexandra DOVGAN”, Igor PODAGELIS” I ) Odessa State Building and Architecture Academy PO Box 76, Main Post Office, 65001 Odessa, Ukraine, e-mail: [email protected] ’) Vilnius Gediminas Technical University 11 Sauletekio ave., LT-2047 Vilnius, Lithuania, e-mail: [email protected] ABSTRACT The relation between criteria of resistance of modified epoxy composites in two media is analysed. Computational experiments on the fields of these criteria in coordinates of composition, described by experimental-statistical models, are used. In the absence of modifiers (furfural and zeolite) coefficients of resistance correlate positively to a high degree, and increase in water resistance due to optimisation of mineral kamework implies the gain in petroleum resistance. As the dosages of both modifiers are increased, the correlation weakens, disappears, and becomes negative. At maximal modification the maximisation of one resistance index may lower the other.

Keywords Epoxy composite, furfural, zeolite, experimental-statistical model, Monte-Carlo

INTRODUCTION There exist at least three important roles [l] in research and development, in manufacturing process, and during service life of composite materials, that analysis of correlation between material quality criteria could perform. Firstly, it is indispensable when arranging an express-control of material properties, in particular, based on non-destructive tests. Secondly, it allows a number of criteria by which composite material should be designed and optimised to be reduced. Thirdly, correlation analysis can be helpful in revealing the conditions (composition and process parameters) under which the mechanisms of structure formation or destruction change. However, it is not infrequent that the circumstances and the results of applying this rather fine statistical tool to restricted volume of experimental data (obtained usually in resources consuming experiment) prove to be incorrect. The potential of correlation analysis for materials science would increase if experimental data were enriched with information mined in computational experiments. Such virtual experiments [2] are carried out on the fields of material properties (Y) in coordinates of composition and process parameters (vector x). The fields can be described by experimentalstatistical (ES) models built on results of real physical experiment, fulfilled preferably to

certain optimal design. These means of computational materials science were used, in particular, to analyse the relation between the indices of resistance of modified epoxy composites in two media. CONDITIONS AND RESULTS OF THE EXPERIMENT When developing [2-31 protective epoxy composite (Ukrainian patent 5408, 2005) a number of quality criteria were determined for 18 compositions, in correspondence with optimal nonsymmetric 18-point design of experiment. Four composition factors separated into two groups were varied. Dosages of organic modifier (furfural), XI = Xl.0 f AX1 = 5 f 5, and fine grained mineral modifier (zeolite), X2 = 8 f 8 M.P. (for 100 mass parts of epoxy resin cdbfacro))at 18 M.P. of hardener), comprised the group “Brittle matrix modifiers”. Quantities of quartz sand, X3 = 175 f 125, and diabase powder, & = 70 f 20 M.P., defined the second group, “Mineral framework”. Among other properties bending strength of dry material (R), the strength after holding in water (Rw) and in petroleum (Rp) for 180 days were determined for each composition allowing corresponding resistance coefficients (Kw = Rw/R, Kp = Rp/R) to be also estimated. Experimental values of these resistance indices for 18 compositions are generalised in Table 1 where characteristicsof variation series for Kw, Kp, and their ratio are given.

Table 1 Characteristicsof experimental data (over 18 compositions) Notation Average Y Standard deviation SY Maximal value Ym, Minimal value Y,i,

Kw 1.029 0.057 I 1.174 0.932 . . .~ I 0.242

I I

KP 0.962 0.044 1.067 0.881 0.186

I

1

I I

I

KPKW 0.937 0.051 1.020 0.863 0.158

I

Eighteen pairs of experimental values have shown rather weak positive correlation between indices of resistance in water and petroleum. Correlation coefficient r{Kw, Kp} = 0.44 (significant at 5% risk, non-linear model increasing coefficient of determination less than by 0.02). This non-trivial result has shown that it could be difficult to design the composites resistant concurrently to water, petroleum, and industrial runoffs. This has given impetus to more in-depth analysis of the relation between two measures of resistance, in two media. FIELDS OF RESISTANCE CRITERIA IN COORDINATES OF COMPOSITION For more sophisticated analysis of correlation between KWand Kp it has been necessary to divide the pair data into several populations. Such assemblages of the sufficient size can be obtained in computational experiments on the local fields of KW and Kp at various combinations of field factors (called gradient factors [2]) and levels of other factors that would change the local fields (changing factors). Built on results of physical experiment non-linear 4-factor ES-models (1-2) describe all the variety of local fields of resistance coefficients in composition coordinates (conventionally

Correlation between bending mistance of epoxy composite specimens maintained in water and in petroleum

46 1

normalised as xi = (X-Xo.i)/AXi ). The models (with only significant coefficients at 10% risk and generated [2] experimental errors s,{Kw} = 0.009 II se{Kp} = 0.019) are structured by factor subsystems: block (a) presents effects of modifiers at medium levels of mineral framework factors, block (b) characterises the influence of the filler and sand at medium levels of modification, with block (c) evaluating the synergism of modification and framework parameters. K, = 1.02 f

0 ~ ~ + 0 . 0 1 ~ ~ ~ - 0 . 0+10~ .1 0~ 21 ~ ~ ~ ~ (4

+ 0.01~2+ 0 . 0 5 ~ 2

-0 . 0 2 ~ ~ ~ ~

+ 0.02XIx4 + 0.02x2x3 (2) The full fields (1-2) of two properties, over whole region of all composition coordinates under study, have the following values of main generalizing indices, practically coinciding with their analogues in experiment (Table 1): = 1.12 (at X I = x2= +I, x3= -1, %= -0.44) and Kp.max= 1.07 (at X I = +0.71, x2= +1, x3= +1, xq= +0.18), absolute increases DK, = 0.27 and DKp = 0.19. Unlike random experimental values, the levels of full and local fields merely calculated by ES-models present deterministic values. In this case the fields are referred to as “modeldeterminate”. Shown in Figure 1.a-b are the local model-determinate fields Kw(x3, xq) and Kp(x3, xq) for non-modified compositions (at lower levels of modification factors, X I = x2= -1). This is the case for which the procedure of transforming the model-determinate fields to the random fields, and thus obtaining random values of properties for correlation analysis, is outlined. The previous version of the procedure [4] has been improved so that prediction properties of the models could be properly accounted for.

COMPUTATIONAL EXPERIMENT The algorithm allows for transformation of the model-determinate fields of two properties, specifically Kw (x3, xq) and Kp(x3, x4) defined by (1-2) at fixed XI= x2= -1, into random fields with Monte Car10 method. Random coordinates x3, x4 are generated, uniformly distributed inside the square [+I, f l ] of normalised factors (contents of sand and diabase), corresponding to N =18 random granular compositions of “Mineral framework” (N equals the number of compositions investigated in physical experiment). For each composition (generated point x) the values of Kw and KP are calculated by the models and random prediction errors AKw and AKp are determined, accounting for the errors of physical experiment and of the approximation. The inclusion of random errors brings the conditions of computational experiment closer to the circumstances of potential real experiment.

Tatiana LYASHENKO, Vitaly VOZNESENSSKI:Alexandra DOVGAN, Igor PODAGELIS

462

1

1

1

x3

1

x3

Figure 1. Model-determinate fields (a, b) of resistance coefficients in framework coordinates for non-modifiedmatrix, (c, d), and random fields (e, f ) respective prediction variance functions

Correlationbehoeen bending reskttznce of epoxy composite specimens maintained in water and inpetroreUm

463

At each point x = ( x3, xq) the random error A(x) = se. tx. [d(x)]O involves experimental error (above-indicated for both properties), generated random value (from standard normal distribution), and square root of prediction variance function. The d-function is defined by design of experiment the models has been built on and by structure of corresponding model, from which all insignificant effects have been eliminated. Shown in Figure 1.c-d are variance functions d(x3, x4) for resistance levels of non-modified composites (without furfural and zeolite) estimated by models (1-2). Added to predicted model-determinate values at N random points in the region of property field the prediction errors give N random levels of the property, thus presenting one sample realisation of the random field. The representation of random fields in Figure l.e-f is the result of 200 realisations, with Monte Carlo participating twice in building random field model - when simulating the point of the field and its level. One realisation of the pair of random fields Kw ( ~ 3 ,xd) and Kp(x3, x4) gives one estimate of r{Kw, KP} exhibited through variations in “sand + diabase” system (and so denoted as 1-34). Then M such realisations would produce M estimates (for any numerical characteristic of the fields and any measure of their correlation). This enables frequency distribution f(r34) of correlation coefficient to be obtained (as the result of computational experiment at one point in the space of “Modifiers”). Experience shows and special computational trials confirm that M 2 200 would suffke for the credible conclusions. The empirical distribution of r34 obtained after 200 realisations of paired random fields Kw(x3, x4) and Kp(x3, xq) at X I = x2 = -1 is shown in Figure 2 (f - relative frequency of r values). The distribution centre is characterised by median value Me of r34 equal to 0.66.The quantile qo5 estimates the lowest level of correlation coefficient (at conventional risk of 0.05). With 90s = 0.44being considerably greater than zero, strong positive correlation between two resistance indices of non-modified composites, when varied are framework factors, may be thought of as practically certain.

-0.85

-0.55

-0.25

0.05

0.35

0.65 r34WwKpJ

Figure 2. Two distributions of r34{K~,KP} at different conditions of modification Represented in Figure 2 is also one more distribution of r34{K~.KP} analogically obtained for different conditions of modification. The distribution polygons reflect the changes in {KwKp}-relation (revealed in computational experiment) when going to other content of modifiers, specifically, to maximal dosages of furfural and zeolite (8 and I6 m.p. respectively, X I = x2= +l). It should be noted that uniform random variations of framework composition (random combinations of sand and diabase dosages), varying the levels of Kw and Kp on

Tatiana LYMHENKO, Wtaly VOZNEENSSKI:Alexandra DOVGM, Igor PODAGELIS

464

which their correlation is measured, could not be responsible for these changes. The positive correlation between Kw and Kp weakens with increase of the content of modifiers. At medium dosages (about 4 parts of furfural and 8 of zeolite, XI=XZ=O)the correlation disappears and then transforms into negative one. At upper levels of modification factors the critical quantile q95 of r34-distribution equals -0.17, revealing the significant negative correlation. This means, in particular, that increase of water resistance of the composites could be accompanied by reduced petroleum resistance. Since described above virtual trials can be performed, and estimates of distribution of r34{Kw, Kp) can be obtained, for any conditions of modification under study (at any XI, xz), the computational experiment has been carried out on the square of modification factors by “full factor experiment” design of second order (“32”). The results have allowed the secondary non-linear ES-model to be built describing the effects of modification parameters on correlation between Kw and Kp of the composites with various mineral frameworks.

THE FIELDS OF r{Kw, Kp} IN COORDINATES OF COMPOSITION When modeling the estimates of correlation coefficient Fisher transformation, z = Arc tg r, is used so that the values of r calculated by the model would fall within the interval between -1 and +1 [l, 41. Thus built has been incomplete cubic model (3) for r34{KW, Kp} (on average values of the obtained distributions), in dependence of modification parameters (normalized to XIand XZ). 234

= 0.352 - 0.430~1-o.082xi2- 0 . 1 0 1 ~ 1+~02 .

-0.436~2 f

The model

describes the field of r34{Kw,

Figure 3. Coefficient of correlation between resistance criteria of composites with various mineral frameworks versus dosages of furfival (XI)and zeolite (XZ)

154~1~~~

+ 0.102x,x22

0x22

(3)

in “Modifiers” coordinates shown in Figure 3, the inverse Fisher function being used when mapping. A maximum of 0.687 (XI = -0.5 and x2 = -1) practically coincide with r34 for non-modified composites, a minimum of -0.414 corresponds to upper levels of the content of both furfural and zeolite. The difference in r34 over the extent of modification conditions is quite great, Dr34 = 0.687 - (-0.414) = 1.101, i.e., more than half of the possible range. To define the boundaries of a zone where null-hypothesis on true correlation coefficient p34 {Kw, KP) must be admitted (that linear relation between Kw and Kp is not revealed), the models for 5-th and 95 percentiles of rdistributions (905 and q95), analogous to (3), areused. Determined by these models with Monte Car10 method has been “no correlation” corridor - the dosages of modifiers at which r34{Kw, Kp} would

Kp}

Correlation between bending resistance of epoxy composite specimens maintained in water and in peholeum

465

be between -0.32f0.02 and +0.36f0.04 (in Fig. 3 the surface of r34 between these levels rounded to f0.3 is not shaded). On these results at least three conclusions could be made about correlation between the coefficients of resistance in water and petroleum of the epoxy composites with different mineral framework. Firstly, with no such modifiers as furfural and zeolite in the composition (XI= x2 = -l), Kw and K p correlate positively to a high degree, and maximisation of water resistance, achieved with certain ratio between framework grains, is attended by increase of resistance in petroleum. Secondly, with combined increase of contents of both modifiers, the correlation weakens up to utter disappearance. Thirdly, at the small zone (about 2% of {XI,x~}-square) close to maximal modification the correlation becomes negative, and maximisation of one resistance index may lower the other. In another computational experiment the correlation coefficient r12 {Kw, Kp} was estimated, when varying uniformly the factors of “Modifiers” (simulating random combinations of furfural and zeolite dosages), at various fixed levels of the framework parameters. Modelled on the results the field of rIZ{KW, K p } in mineral framework coordinates is displayed in Figure 4. Almost 80% of grain compositions do not show significant correlation between the coefficients of resistance in water and petroleum. Only as polymer layers in inter-grain space thinner with increased filling, the correlation begins to change. If therewith such thinning of the layers is accounted for the content of sand, the tendency of 1-12 to grow up to significant positive level (corresponding to = 2% of grain compositions) comes to light. However, if increased is the content of filler, r l 2 { K ~ K, p } reduces to critical level. At about 20% of the field region the gradients of Kw and Kp would point in different directions. Additional information about the influence of composition on correlation between the indices of composite resistance in water and petroleum was mined with computational experiments in other pair coordinates. Analysed and compared with relationships considered , above (Fig. 3-4) were I - ~ ~ { K wKp} versus quantities of furfural (XI) and sand (xg), together with rI3{KW,KP} in zeolite (xz) and filler (xq) coordinates, represented in Figure 5. This has allowed the following conclusions. Firstly, changing the coordinate from “zeolite” (XZ, Fig. 3) to “sand” (x3, Fig. 5.a) does not change the form of 2-dimensional field of correlation coefficient. Consequently, the increased content of furfural is truly responsible for attenuation of correlation between water resistance and petroleum resistance of the epoxy composites. Secondly, the joint -1 / influence of fine-grained diabase and zeolite (Fig. 5.b) amounts to effect of Figure 4. Coefficient of correlation the latter only. This confirms the between resistance criteria of composites hypothesis [5] that zeolite is not merely at varied conditions of modification a filler, but specific mineral modifier vs. dosages of sand (x3) and diabase (x4) of epoxy matrix.

Tatiana LYMHENKO, Wtaly VOZNESENSSKI:Alexandra DOVGAN. Igor PODAGELIS

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(a)

_-

,\

\ ~

x2

XI

Figure 5. The fields of correlation coefficients r24{Kw, Kp} in coordinates “furhral, sand” (a) and r13 {Kw, Kp} in coordinates “zeolite, diabase” (b)

CONCLUSIONS The results of computational experiments on correlation analysis of water resistance and petroleum resistance of modified epoxy composites serve as the logical basis of the necessity for compromise optimisation [3,6] of material composition by resistance criteria. The compositions found have been used in discharge structures at zones of filling and repair stations at some highways in Lithuania, 20% increase in service life of such structures being forecasted. REFERENCES 1. Lyashenko, T.V., Vonesensky, V.A., Modeling and analysis of varying correlation between properties of brittle matrix composites. In: Proc. Int. Symp. Brittle Matrix Composites 5, A.M. Brandt, V.C. Li, and I.H. Marshall eds., Woodhead Publ. Ltd. -BIGRAF, Cambridge-Warsaw 1997, pp 417-426 2. Vonesensky, V., Lyashenko, T., ES-models in Computational Building Materials Science (in Russian), Astroprint, Odessa 2006, 116 pp 3. Lyashenko, T., Voznesensky, V., Dovgan, A., Podagelis, I., Sharshunov, A., Design of repair compositions for concrete by workability and durability criteria with methods of computational materials science. In: Proc. Int. Conf. Life Cycle Assessment, Behaviour and Properties of Concrete and Concrete Structures, Brno 2004, pp 300-305 4. Lyashenko, T., Vonesensky, V., Krovyakov S., Analysis of water effect on fracture toughness in cement-based composites using computational materials science methods. In: Proc. Int. Symp. Brittle Matrix Composites 6, A.M. Brandt, V.C. Li, and I.H. Marshall eds., Woodhead Publ. Ltd. - ZTUREK RSI, Cambridge-Warsaw 2000, pp 2 10-2 19 5. Vonesensky, V., Lyashenko, T., Dovgan, A., Gara, A., Analysis of the property fields to prove the specific role of zeolite in epoxy composite (in Russian). In: Bulletin of Odessa State Building and Architecture Academy, 15,2004, pp 54-6 1 6. Lyashenko, T., Voznesensky, V., Boiko, S., Shtakelberg, D., Analysis of concrete property fields and search for the best compositions using Monte Car10 method. In: Proc. Int. Symp. Brittle Matrix Composites 7, A.M. Brandt, V.C. Li, and I.H. Marshall eds., Woodhead Publ. Ltd. -ZTUREK RSI, Cambridge-Warsaw 2003, pp 351-358

Proc. Int. Symp. '%BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

INORGANIC GEOCEMENT MATRIX IN COMPOSITE MATEFUALS Pave1 V. KRIVENKO, Mykola A. MOKHORT V.D. Glukhovsky Scientific Research Institute for binders and Materials Kiev National University of Civil Engineering and Architecture Kiev, Ukraine, e-mail: [email protected] ABSTRACT The purpose of the present research was the determination of the influence of composition of geocements on properties and structure of artificial stone. The introduction of various classes of modifiers in composition of geocements allows to adjust directionally the processes of structure formation, technological and operational properties of geocements matrices and materials on their basis. The investigations and also long natural tests of materials have allowed to draw a conclusion about high durability and properties of such materials.

Keywords Durability, forecasting of durability, geocements, metakaoline, processes of structure formation, rigid inorganic matrixes, soluble glass, zeolite-like new formations INTRODUCTION Beginning from 1957 in Kiev has been developing a new direction in the binders, the background of which was the discovery of binding properties of the alkali metals [l]. The alkali metals compounds act not only as activators of hardening but are responsible for the formation of main structural elements of the alkaline cements - zeolite-like compounds of different types [2]. Just these hydration products, analogs to natural zeolites have been discovered in ancient concretes (Ancient Greece, Rome, Egypt, Syria) [3,4]. Durability of the ancient concretes and similarity of their structure with the alkaline cement concretes allowed to predict their high durability. Effectiveness and high performance properties of the alkaline cement concretes are supported by over 40-year experience of service of the constructions. The paper covers practical application of the geocement- based systems, referred to the I class of alkaline cements [2], in construction, in building materials production, as well as other industries. The geocements are themselves inorganic polymers, combining properties of both inorganic compounds (high strength, thermal resistance) and organic ones (corrosion resistance, adhesion to different materials, etc.) and are analogies to natural alkaline aluminosilicates of the zeolite and feldspathoid types [24]. MANUFACTURE AND USE OF THE GEOCEMENT- BASED COMPOSITES Over last 12 years, the scientists of the V.D.Glukhovskii SRI for Binders and Materials have developed and tried on pilot- and commercial scale a range of geocement- based materials, since geocement is as one of the most advanced types of contemporary materials exhibiting

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Pave1 V. KRIVENKO, Mykola A. MOKHORT

high efficiency, ecological compatibility, safety and durability. Types of the developed geocement- based materials are given schematically in Fig. 1.

Fig. 1. Types of geocement- based composites Heat insulating, heat insulating refractory and constructive materials with mineral fibres, expanded rocks and gas geocements High temperature resistance of the geocements and high adhesion to different materials predetermined the development of a wide range of heat insulating and heat insulatingconstructive [5], heat insulating refractory [6], constructive materials [7] with mineral fibres and different fibre- reinforced materials, as well as waste products of their production [8], expanded rocks[9], organic fabrics [lo], as well as gas geocement concrete with mineral fillers [ll], the technical characteristics of which meet or exceed those specified by the requirements of existing standards. Application of geocements allows to replace organic binders in almost all classes applications of such materials as fibre- reinforced heat insulating and heat insulating refractory mats, boards and blocks, different plastics (with strength up to 120 MPa), perlite and vermiculite boards, as well as to utilize with high efficiency the wastes products of fibres production. The replacement of traditional binders in gas concrete allows to considerably enhance temperature resistance of these materials (up to 80OOC) and utilize the ashes of heat power stations with high efficiency. Fibre-reinforced heat insulating materials with mineral wool, superfine, thin and thickened basalt fibres. These materials may be produced by any method known in the technology of production of fibre- reinforced heat insulating materials, thought the most advantageous were found to be two methods, they were: spraying in a fibre- settling chamber and a “wet” process in which water pulp is used. Density of these materials varies within 90-400 kg/m3 at binder content between 2 and 30 mass %, heat conductivity at T=298K being 0.044-0.056

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W/(m'K), compressibility (at P=0,002 MPa) varying between 2 and 35%, temperature resistance of the ready products - within 600-850 "C depending on fibre type, sorption moisture content and water absorption of the materials falling beyond the standard specified requirements. Heat insulating refractory materials with mullite- silica fibres. At mean density 310-420 kg/cub.m, these materials exhibit flexural strength- up to 0.4 MPa and the lower heat conductivity at T= 600°C (0.18-0.22 W/(m'K)), as compared with commercial products, the lower loss on ignition (0.01-1 %), have the increased content of A1203+Si02 (97-98%) and possess the higher temperature resistance (1300°C). Heat insulating materials with perlite and vermiculite. A variety of products with different expanded rocks has been developed. Inorganic nature of these materials and high temperature resistance of the geocement determine their high fire resistance and fire safety. At density varying from 150 to 300 kg/m3, these materials have compressive strength 0.26- 1.0 MPa, flexural strength - 0.15- 0.3 MPa, heat conductivity - 0.062 - 0.09 W/(m'K). Constructive materials with basalt rovings have density 850- 1000 kg/m3, surface density 2.6-9.85 kg/sq.m, binder content 42-49 mass %, organic matters content of the final product 0-10 mass %., flexural strength- 28-30 MPa, maximum use temperature 250-500°C. They are characteristic of considerably lower density at enhanced strength compared with the asbestoscement materials, and equal or considerably higher maximum use temperature (the maximum use temperature of the asbestos- Portland cement materials is 25OoC and they may be used for replacement of temperature resistant asbestos- cement materials with siliceous additives at working temperatures up to 500°C as well. Constructive materials with glass- and basalt non-woven materials and fabrics have density 900-1800 kg/cub.m, surface density 4.4-9.0 kg/m2, binder content 21-38 mass %., organic matters content of the final product and 5-12 mass YO,flexural strength 94-130 MPa, maximum use temperature 300-4OO0C, maximum temperature of short-time exposure 400500°C. Constructive high temperature materials with mullite- siliceous fibres exhibit density 1600-1700 kg/m3, binder content - up to 25 mass %, flexural strength - up to 11 MPa, maximum use temperature 1000°C. Constructive materials with carbon fabrics have density 2200 kg/m3, binder content 50 mass %, organic matters content of the final product- 11 mass %, flexural strength 110 MPa, maximum use temperature 300°C. Constructive and heat- insulating- constructive materials with expanded rocks have density 640-1600 kg/cub.m, binder content- up to 25 mass %, compressive strength 1.0-15.0 MPa, flexural strength 2.8-10.0 MPa. The majority of them is heat resistant (residual strength at firing at T= 800°C is up to 225 %). Constructive materials with jute fabrics have density 1000-1250 kg/cub.m, flexural strength 30-45 m a , Young's modulus- 2-14x10' Pa, temperature resistance- up to 105°C. Heat-insulating-constructive gas geocement concretes exhibit compressive strength 2-5 MPa, density 400-550 kg/cub.m, the much higher temperature resistance- up to 800°C compared with the known-in-the-art gas cement concretes, allowing to use them not only in construction, but for insulating high -temperature industrial equipment [ 111.

Abrasion resistant composites and instruments The abrasion resistant composites with high thermo-mechanical characteristics and the instruments on them have been developed by modification of the geocement by silicon carbide [12], partially interacting with the geocement with the formation of a protective aluminosilicate film on the surface of the grains, determining largely the composite's thermoresistance. The material of which the instruments were produced commercially in

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Ukraine, is characteristic of compressive strength 47-50 MPa, flexural strength- 23-25 MPa, breaking strength 5-7 MPa, thermoresistance- 25 heat cycles "8OOOC -water". Wood- based materials The use of the geocements in wood- based materials allows not only to replace toxic organic resins, but to improve considerably fue resistance [13]. The commercial production of the door leafs with the geocement- based adhesive was launched in Ukraine. The strength of the tenon joints was over 0.96 MPa, strength of the adhesive joint of the wood fibre- reinforced boards with wood in non-uniform breaking - over 2.8 MPa, strength of the adhesive joint of the wood in cleavage parallel to the fibres - over 8 MPa. The attempts have been taken to use the geoecement-based adhesives instead of organic ones in the production of plywood.. The splitting strength along the adhesive joint in testing plywood was over 2,2 MPa. One more example of the use of geocements in the production of wood-based materials may be geocement- based chip wood aggregate concrete and wood particle boards [ 141. Geocement- based wood particle/ chip wood boards are characteristic of densities from 500 to 1500 kg/m3, flexural strength- 8-20 MPa, swelling less than 1% for 24 hrs, binder content- up to 40 %. The use of geocements in the production of the wood particle boards allows to replace completely the organic binders and to involve in the production of up to 70 mass % of hardwood waste products. Chip wood aggregate geocement concretes exhibit density 600-800 kg/m3, compressive strength- 5- 7 MPa, low water absorption - below 60 mass %. The content of chip wood in the concrete is 38-40 mass %. Application of the geocements in metallurgy High thermal resistance of the geocements allowed to develop a range of products for metallurgical industry: pipes for feeding metal melt, sieves for melt filtration, heat insulating feeders, etc. [ 151. The use of the geocements allows to replace high-cost ceramic products and to exclude a high- temperature firing. Also, the developed geocement- based moulding and core sands allowed to replace high toxic organic adhesives in the foundry [16]. The developed mixes are characteristic of high green strength (up to 0,25 MPa) and strength in a dry state reaching 7,5 MPa, minimal strength of the moulding sand after the metal is fed, high reclamation of moulding sand (8090 'YO) and meet the requirements to moulding sands. Geocement- based protective coatings Intumescent fireproofing and expanded heat insulating coatings have been developed on the base of geocements. Intumescent fireproofing and expanded heat insulating coatings [ 171. The formation of cellular- pore structure of the expanded layer is reached by introducing a granular material, which was preliminary prepared from a reactive mix by a technology of production of expanded soluble glass, into a coating composition, together with a complex mineral filler. The coatings exhibit: coefficient of volume increase 15-23, compressive strength up to 4,6 MPa, water absorption- below 10 mass %, adhesion to metal 0,78-0,84 MPa, to ceramics - 4,8-5,0 MPa, the expanded coating is characteristic of the coefficient of heat conductivity h=0,04 1-0,065 W/(m'K). The tests and introduction at the enterprises of Ukraine showed their efficiency in replacing traditional compositions made with organic components. Mineral corrosion resistant coatings based on geocements [181 are characteristic of compressive strength after hardening within a temperature range 20-150°C- 50-1 10 MPa, changes in mass and volume after boiling in 2% milk acid solution fall within/ beyond the ranges 1,5-2 'Yo, whereas in 1-3% solutions of sulphuric acid - falling within 2-2.5 'YO at

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coefficient of resistance 0,85-0,97 for organic mediums and 0,7-0,78 for sulphuric acid. The use of autoclave treatment allows to produce a material, which is able to resist/ withstand exposure of these chemicals of increased concentrations (5- 10 %), the coefficient of resistance being 0,98-1,OO after attacks of the organic chemicals and 0,8-1,0 - of sulphuric acid, and volume changes not exceeding 2-2,s %. Adhesion of the coating to concrete is 3-8 MPa. The geocement- based composites of the enhanced acid resistance have been introduced as a twolayered coating of the external lining of the sewage waters disposal canal. The coatings are in service in solutions (concentrations below 5 'YO) of sulphuric acid, pickling solutions and solutions of heavy metal salts. Geocement- based fireproofing coatings for wood, wood- based materials and fibrereinforced boards are characteristic of resistance in open flame with T= 800-1000°C for 10-20 minutes [ 191.

Geocement-basedcompositions for repairing and restoration A variety of geocement- based materials for restoration and repairing needs has been developed: primers and putties based on the compositions, the bases of production of which are reported in [20], inorganic anchoring binding systems [21], systems for repairing gas permeable members of machines working at high temperatures. These materials are characteristic of strength 24-70 MPa, high water resistance and strength gain and inconsiderable linear shrinkage deformations. Inorganic geocement- based adhesives One of the most effective application of the geocements is their use in bonding various materials: mineral fabrics, fibre- reinforced heat insulating articles and paper, concrete, brick and ceramics , wood with concrete and gypsum, for metal, expanded polystyrene, plastics, etc. [ 19,22-231. The adhesive joints exhibit strength in non-uniform pulling-out 3-15 MPa, high frost (over 100 cycles) and water resistance. The pilot- scale tests held in Ukraine and Finland showed that the bonded materials, thanks to layers of the adhesive, enhance summary strength and rigidity (for such materials as fibre- reinforced boards, paper, mineral fabrics, expanded polystyrene), and exhibit considerably higher fire resistance. PHYSICAL-CHEMICAL RESEARCHES Research of influence of initial composition of geocements and type of treatment on processes of structure formation and linkage of alkaline oxides The research of physical-mechanical properties of an artificial stone has shown, that the compression strength of samples depends on a mode and temperature of processing and changes from 60 MPa after hardening at 20°C during 28 day in damp conditions up to 80 MPa after drylng at 100°C during 6 hours. The rontgenograms of samples of an artificial stone (at hardening at temperatures 20100°C) show the presence of X-ray amorphous phase in structure of an artificial stone, and also not reacted initial materials, about what the peaks of quartz testify (d=0,429; 0,261; 0,169; 0,150 nm) (fig. 2, curves 1-3). For determination of phase composition of an artificial stone on the basis of geocement systems the autoclave treatment, which provides increasing of a degree of crystallization of hydrated phases without change of its phase composition was used. After autoclave treatment composition of new formations of non-modified geocement (Na2OA12034Si021OH20): mainly analcime (d=0,564; 0,345; 0,292; 0,268; 0,250; 0,241 nm) (fig. 2, curve 4). The change of composition (from Na20A12034Si021OH20 to (0,72Na20+0,28K20)A1203

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4SiO~lOHz0)by introduction of KOH result in synthesis of new formations like analcime (d=0,564; 0,345; 0,292; 0,268; 0,250; 0,241 nm) and zeolite P (d=0,713; 0,505; 0,442; 0,4108; 0,352; 0,332; 0,319; 0,317; 0,267 nm) (fig. 2, KP. 5). The introduction of Ca-inclusive additive (slag or OPC) result in formation of, except analcime, amicite (d=0,564; 0,422; 0,314; 0,272 nm), garronite (d=0,710; 0,501; 0,410; 0,316; 0,267 nm), gismondine (d=0,723; 0,569; 0,496; 0,425; 0,330; 0,3 18; 0,269 nm) (fig. 2, curve 6).

The linkage of alkaline ions For definition of depth of hydratation processes and the interactions of components in geocement at structureformation a method of the chemical analysis of interstitial liquid of an artificial stone was used. The linkage of alkaline ions in structure of geocement depends on a type of treatment and temperature, and most intensively this process passes at a microwave treatment. Fig. 2. X-ray spectra of geocement after various conditions of hardening: 1-28 days of hardening at 20°C; 2-28 days of hardening at40"C; 3-6 h of hardening at 100°C; 4-6 h of hardening at autoclaving (17Of2"Cy 0,s MPa), geocement with ratio of oxides (NaZO/A120+1; 5-6 h of hardening at autoclaving (17 M" C , 0,8 Mpa), geocement with ratio of oxides (0,72NazO+ 0,28K20)/A1203=1; 6- 1 h of microwave treatment (P=140 watt, T=160°C); 76 h of hardening at autoclaving (17Ofl0C, 0,8 MPa), geocement with Cainclusive additive.

A- Analcime ( N A & L ~ S I ~ O O ~ ~ 12H20); Acc - Analcime-C (NA(SIZAL)O&O); P-Zeolite P (N~s.~ALs.~SIIO Q.-~ ~ ~ Z ~ ~ H Z ~ ) Quartz SIOz; Am -Amicite (KzNAz&sb~1~5Hzo); GT Garronite (NACAz5(SIloAL6032) 14HzO); Gs - Gismondine (CmSIz084Hz0) 60

50

40

-

30

20

20

With increasing of hardening temperature from 20°C up to 100°C decreases the value of pH of interstitial liquid from 11,2 up to 9,7. The introduction in composition of geocement of the Ca-containing additive reduces the value of interstitial liquid from 10,l up to 9,3. It is possible to explain it to that the introduction of such additive results in linkage of a part of

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unbound alkali in new formations like amicite (d=0,564; 0,422; 0,314; 0,272 nm), garronite (d=0,710; 0,501; 0,410; 0,316; 0,267 nm), gismondine (fig. 2, curve 6). So in hardening process of geocement there is a linkage of alkaline ions, and most intensively this process passes in the first 1-7 day. The introduction in composition of geocement of the Ca-containing additive accelerates the linkage of alkaline ions in structure of insoluble new formations. The linkage of Na-ions in time at structureformation of geocement. Starting from initial concentration in system can be told, that in the age of 2 day the samples are characterized by linkage of 84.99% of free Na-ions in case of the unmodified geocement and 89.79% - at introduction in composition of geocement of the Ca-containing additive. After 7 day hardening, accordingly, 91.26 YOand 92.36 % of Na ions are bonded already. Then process of linkage of Na-ions in time is considerably slowed down, especially in the unmodified geocement. So, after 28 day the unmodified geocement is characterized by presence in interstitial liquid of 8,57 % of of free Na-ions (from total in system), and modified geocement - only 1.2%. The introduction of the Ca-containing additives in composition of geocement allows to speed up the process of linkage of Na-ions owing to formation of zeolites - amicite, garronite, gismondine. The process of linkage of K-ions in time at structureformationof geocement. Starting from initial concentration in system can be told, that in the age of 2 day the samples are characterized by linkage of 95% of free K-ions in case of the unmodified geocement and 96% - at introduction in composition of geocement of the Ca-containing additive. After 7 day hardening approximately 99% of K ions are bonded. Then process of linkage of K-ions in time is considerably slowed down, especially in the unmodified geocement. So, after 28 day the unmodified geocement is characterized by presence in interstitial liquid of 1,25% of of free Ka-ions (from total in system), and modified geocement - only 0,85%. The microscopic research The microscopic research of a sample of an artificial stone on the basis of geocement has allowed to establish presence of submicrocrystalline structure, which is characterized by presence of microcrystals by the size no more than 3,5 microns seldom located in a amorphous matrix. A low roughness of a surface and continuous strong connection of submicrocrystalline and amorphous phases allow to assume strength and durability of such structure, which by virtue of the nature can only in a small measure fall under influence of the external destroying factors. The electron microscopic research of an artificial stone on the basis of the unmodified geocement after autoclaving (the fig. 3) also has shown, that in structure of geocement in the majority are formed the crystals of hexagonal form photography geocement which can be related to analcime. (3810X)

of stone

QUESTIONS OF DURABILITY

The structure of geocement gluing compositions represents microconcrete that has allowed to assume as a basis by development of mathematical model of durability of a geocement artificial stone the dependence which reflects influence of raw materials, and also influence of structural characteristics, technological, operational and special factors on development constructive or destructive phenomena in time at formation of an artificial stone on a basis of clinker-containingbinding systems [25]. Thus the characteristics and factors are expressed as

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the factors, promoting to influence on the base (accepted as average (standard)) durability of the ordinary heavy weight Portlandcementconcretefll = 60 years). At the comparative analysis of durability of geocement and OPC systems and interpolation of the mentioned above dependence with reference to peculiarities of composition of geocements, physical and chemical peculiarities of composite building of hydroaluminasilicate matrixes and development of its structure in time, and also the analysis of long-term tests and control (more than 12 years) have allowed to interpret dependence [31] by following mathematical expression: Y = YB.DUR~.DUR~.DUR~.ABCDEFGHLTKL, where: YB - the base durability equal to 60 years (accepted by analogy with OPC concrete); DUR1, DUR2, DUR3 - factors of increasing of durability of a geocement stone in comparison with OPC stone. Depending on exploitation conditions of glued connection one of them is accepted as significant, values of two others are accepted equal 1,O; A, B, C, D, E, F, G,H, I, J, K, L - factors of influence of quality of initial components, technological, operational and special factors on development of constructive or destructive phenomena in glued connection on the basis of geocement compositions. DURl - factor of increasing of durability of a geocement gluing composition in comparison with OPC composition by criterion of frost resistance (1,O-2,O). DUR2 - factor of increasing of durability of a geocement gluing composition in comparison with OPC composition by criterion of atmosphere resistance (1,O-13). DUR3 - factor of increasing of durability of a geocement gluing composition in comparison with OPC composition by criterion of corrosion resistance (0,9-1,8), accepted on the basis of the data of the comparative analysis of corrosion of geocements and sulfateresistant portland cement in organic environments and a solution of a sulfuric acid, and also comparative oil resistance. A - factor of quality of metakaoline (0,8-1,l). B - factor of quality of materials for preparation of a liquid phase (soluble glass, Na and K hydrooxides, silicafume) (0,9-1,2). C - factor of quality ofbinder preparation (1,O-1,l). D - factor of quality of a gluing composition preparation (0,9- 1,l). E - factor of a time interval of use of a ready composition (0,8-1,l). F - factor of open drylng time (0,9-1,l). G - factor of thickness of a layer of glue line (0,9-1,2). H - factor of temperature influence of an environment at facing (0,9-1,l). I - factor of a temperature regime of glued connection hardening (0,9-1,l). J - factor of conditions of gluing work (0.9-1.O). K - factor of personnel training (0,7-1,l). L - factor of quality of surfaces (0,8-1,2). The presented model of estimation of durability allows to calculate the minimal and maximal service life of glued connections of, for example, concrete and inorganic facing material, using, accordingly, the minimal and maximal values of factors. The durability of glued connection the concrete - inorganic facing material, estimated on the given model, can make 10-407 years. CONCLUSIONS

The experience gained from small- and large- scale industrial uses of the geocements in construction and other industries showed high technical and economical advantages of the geocements and the materials based on them.

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Results of researches of phase formation and hardening of geocements with a ratio of oxides (Na,K)20/A1203=1,Si02/A1203=4,H20/A1203=10-15 shown, that the structure of new formation is mainly composed of hydroaluminosilicates like analcime, zeolite P and garronite. The introduction of calcium to reactionary mixture leads the to the formation of zeolite P and analcime, and also hydrosilicates Ca with structure of ksonotlite and girolite types. The introduction of additional quantity of Si02 in combination of geocement leads to dominance zeolite-like formations with increased contents of Si02: minerals of a Nashabasite-gmelenite, fogasite and mordenite types. It was established, that at hardening of a geocement stone within 28 day there is a linkage of 93-99 % Na and K ions. The introduction of the Ca-containing additives in composition of geocement allows to speed up the process of linkage of Na and K ions owing to formation of zeolites - amicite, garronite, gismondine. On the basis of conducted researches (atomic force microscopy) it was established that in geocements described by a H20/A1203ratio lower than 10, hardening process passes through stages of formation of hydroaluminosilicates: of amorphous, submicrocrystalline and crystalline structure. In geocements with H20/A1203>10 the stage of formation of submicrocrystalline structure is very poorly expressed. The nucleation of large crystals thus happens immediately in the amorphious phase, that results to a significant retardation of hardening and crystallization processes, and also to the deterioration of the structure and properties of geocement. It was established, that the durability of geocements, estimated on the mathematical model, can make up to 400 years.

REFERENCES: 1. Glukhovsky V.D. The Soil Silicates.GosstroiizdatPublish, Kyiv, 1959, pp 127 (in Russian) 2. Krivenko P.V. Alkaline Cements: Terminology, Classification, Aspects of Durability. In: Proc. 10” Int. Congr. on the Chemistry of Cement, Inform Trycket AB, Gothenburg, Sweden, 1997, Vo1.4, 4iv046, pp. 6 3. Davidovits J. Geopolymeric reaction in archaeological cements and in modem blended cement. In: Proc. of the Conf. Geopolymers, 1988, Vol.1, pp 93-0105 4. Krivenko P.V., Pushkaryova E.K. Durability of the Slag Alkaline Cement Concrete. Budivel’nik Publish., Kiev, 1993, pp 224 5 . Mokhort M.A. Relationship between structure, composition and properties of the fibrereinforced heat-insulating composites based on geocements modified by organic polymers. In: Proc. I1 Int. Conf.”Alkaline Cements and Concretes”, Oranta Ltd. Publish., Kyiv, 1999, pp. 195-207 6. Kryvenko P.V., Bbrodko O.A., Mokhort M.A. The mullite-silica fibre-reinforced heatinsulating refractory materials made from alkaline aluminosilicate binder. “Budivnytstvo Ukrainy”, Kyiv, 1996, N6, pp 3 1-34 (in Ukrainian) 7. Mokhort M.A., Tsibulya Yu.L. Geocement Composites based on basalt Fabric And alkaline Aluminosilicate Binder. In: “Cement Combination for Durable Concrete”, Proc. 6‘h Int. Cong. “Global Construction: Ultimate Concrete Opportunities”, Thomas Telford Publish., London, 2005, pp 287-292 8. Mokhort M.A. A method of utilization of non-stretched basalt fibre. “Budivnytstvo Ukrainy”, Kyiv, 1998, N6, pp 31-32 (in Ukrainian) 9. Krivenko P.V., Petropavlovskii O.M., Mokhort M.A., Resenchuk V.M. Heat-insulating and heat-insulating-constructional materials based on alkaline aluminosilicate binder and natural expanded materials. In: Proc. Scientific-Technical Seminar “Effective Heat Insulation and

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Hydraulic Sealing of the Buildings and Structures and Their Protection in Case of Fire in the Process of Erection and Renovation”, State Scientific Research Institute of Building ConstructionsPablish., Kyiv, 1999, pp 26-32 10. Kryvenko P.V., Petropavlovskii O.M., Mokhort M.A., Vozniuk G.V. A geocementbased composite with jute fabric. “Budivnytstvo Ukrainy”, Kyiv, 2001, N2, pp 2-24 11. Krivenko P.V., Mokhort M.A., Kovalchuk G.Yu. The structure and properties of the geocement-based gas concrete. In: Proc. Sixth Int.Conf. on Failure, Durability and Retrofitting, Singapore, 2000, pp 149-156 12. Gontchar V.P. The studies on thermo-mechanical characteristics of the alkaline aluminosilicate binders modified by a silicon carbide. In: Proc. Second Int. Conf. ”Alkaline Cements and Concretes”, Oranta Ltd. Publish., Kyiv, 1999, pp 303-3 12 13. Baranovskii A.V. The use of alkaline aluminosilicate adhesive in the production of woodbased materials. News of the Academy of Construction, Kyiv, 1997, Vol. 3, pp 44-48 14. Mokhort M.A. Slag alkaline and geocement concretes on Basis of organic filler with high durability. In: Proc. Int. Conf. “Innovations and Developments in Concrete Materials And Construction”, Thomas Telford Publish., London, 2002, pp 597-605 15.Krivenko P.V., Brodko O.A., Vozniuk G.V. Ecologically safety moulding and core sands based on alkaline aluminosilicate bond with high degree of reclamation for foundry. Khimichna Promislovost Ukrainy, Kyiv, 1996, N25,pp 22-25 (in Ukrainian) 16.Voznyuk G.V. Environmentally friendly composites based on geocements for foundry. In: Proc. Second Int. Conf. ”Alkaline Cements and Concretes”, Oranta Ltd. Publish., Kyiv, 1999, pp 237- 247 17. Kryvenko P.V., Pushkaryova K.K, Sukhanevich M.B. The development of physicochemical bases of directed synthesis of inorganic binders in the system NazO-Al203-Si02HzO for obtaining ecologically friendly intumescent materials. “Budivnytstvo Ukrainy”, Kyiv, 1997, N2, pp 46-49 (in Ukrainian) 18. Kryvenko P.V., Popel G.N. Corrosion resistant coatings based on alkaline aluminosilicate binders. In: Proc. Fourth CANMET/ACI Int. Conf. on Durability of Concrete, Sydney, Australia, 1997, pp 241-253 19. Kryvenko P.V., Petropavlovskii O.M., Mokhort M.A., Vozniuk G.V. Alkaline cement for glueing building facing materials. In: Proc. Int. Symp. “Non-Traditional Cement & Concrete”, Bmo, Czech Republic, 2005, pp 568-582 20. Kryvenko P.V., Petropavlovskii O.M., Mokhort M.A., Popel G.N. Physico-chemical fundamentals of obtaining the composite materials from geocements for restoration. In: Proc. of the Annual Int. Seminar RUR’98 “Restorations, Renovation and Urboecology”, Odessa, Belgorod -Dnestrovskii, 1998, pp 260-268 21. Kryvenko P.V., Mokhort M.A., Popel G.N. The geocement-based grout mix for anchoring application. In: Proc. of the Annual Int. Seminar RUR’98 “Restorations, Renovation and Urboecology”,Odessa, Belgorod -Dnestrovskii, 1998, pp 226-227 22. Krivenko P.V., Mokhort M.A., Vozniuk G.V. Inorganic cements for building materials. In: Beitrage zur Baustoff-Forschung 2001 F.A. Finger-Institut f?ir Baustoffkunde, Gutenberg Druckerei Weimar, Weimar, 2001, pp 222-228 23. Kryvenko P.V., Petropavlovskii O.M., Mokhort M.A. Industrial Uses of Geocementbased Materials in Construction and Other Industries. In: Proc. of Int. Conf. “Geopolymer 2002”, Melbourne, Australia, 2002, pp 11 24. Mokhort M.A. Einflulj von Ausgangzustand und Erhartungsbedingungen bei der Verfestigung von Geozementen. In: Tagunsbericht 15 Int. BaustofBagung “Ibausil’’, Weimar, Bundesrepublik Deutchland, 2003, Band 1, pp 1-0137 - 1-0143 25. Gordon S.S. Forecasting of durability of reinforced concrete. “Concrete and reinforced concrete ‘I, Moscow, Russia, 1992, N26, pp 23-25

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

PROPERTIES OF CLINOPTILOLITE BASED AUTOCLAVED COMPOSITES Wlodzimierz MOZGAWA, Waldemar PICHOR AGH University of Science and Technology, Faculty of Materials Science and Ceramics Mickiewicza 30,30-059 Krakow, Poland e-mail: [email protected], [email protected]

ABSTRACT Among several available methods of removing heavy metal cations from wastewater, the sorption method with use of zeolite appears to be encouraging. Clinoptilolite, which is a naturally occurring zeolitic pozzolana, shows ion exchange properties and is mainly used for immobilization of very different wastes, e.g. heavy metal cations. The investigated Polish natural zeolite deposit contains about 20-25 wt. % of clinoptilolite and considerable amounts of different phases: quartz, smectite, metal oxides, and others. Before application as a matrix for metal uptake from wastewater the zeolite should be concentrated and pretreated as to increase its sorption capacity. Very important problem after ion sorption and immobilization is solidification of zeolite in durable form like mortar, concrete or another. Sorption of heavy metal cations (Pb(II), Cr(III), Cd(II), Zn(I1)) from aqueous solutions on clinoptilolite was studied using atomic absorption spectrometry (AAS) and FT-IR spectroscopy. The results of physical properties investigation of the clinoptilolite based composites obtained during autoclaving process with different amount of lime at 180°C are shown. The comparison between results obtained for materials containing initial clinoptilolite and zeolite after sorption process are presented. The paper describes results of structural studies (IR spectroscopy, XRD analysis and SEM observations) of zeolites on which sorption of various metal cations has been carried out. Obtained results indicate that the proposed method of solidification the clinoptilolite based materials after heavy-metal uptake is promising and applicable in simple way in Autoclaved Aerated Concrete technology or production of Calcium Silicate blocks.

Keywords: Zeolite, clinoptilolite, immobilization, autoclaved materials, heavy-metals

INTRODUCTION Constant trapping of heavy metals cations in the zeolites' framework structure is one of the possible ways of cation immobilisation - the process of their inactivation and making them not dangerous. Chemical immobilisation is the most effective procedure of trapping the cations, which are harmful for human beings and natural environment. Heavy metal cations can be immobilized by zeolites by two mechanisms: ionexchange and chemisorption [ I]. The ion-exchange involves substitution of ions present in zeolite crystalline lattice by metal ions from the solution [2]. The type of cation (the position of the cation in the selectivity as well as the cation concentration in the solution will determine the ion-exchange efficiency. Ion-exchange properties of zeolites are due to the presence of not compensated negative charges, which originate from heterovalence substitution (Si4+by A13+in tetrahedral positions) as well as the presence of surface functional groups (mainly OH- and H2O) in crystalline lattice discontinuity points. As the result of lattice negative charge compensation by metal cation outer-sphere or inner-sphere complexes

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are formed. Stability of the outer-sphere complexes is very low because in the binding of the zeolite with the metal ion complexed in such a way intermolecular forces dominate [3]. In contrast, the second mechanism of metal ions immobilization, i.e. chemisorption, results in the formation of stable inner-sphere complexes [4]. This is because the functional groups (mainly OH-) form strong chemical bonds with metal ions outside the hydration envelope [11. In zeolites, ion-exchange processes generally dominate over chemisorption. The process of chemical immobilisation is based on the exchange of the alkali and alkaline earth metals cations by the heavy metals cations. This is connected to new chemical bonds formation and deformations of starting structures. Such changes can influence the IR spectra. The possible application of clinoptilolite in the immobilization process is due to specific features of this zeolite. Clinoptilolite shows well examined ion exchange properties [5,6]. The experimental selectivity series for the metal cations suggest the possibility of using of clinoptilolite as an ion-exchange trap. Alkali-activated clinoptilolite and other zeolite based materials are potentially effective matrices for the immobilization of hazardous wastes compared to alkali-activated cement paste [7]. Available known deposits of this mineral and easy methods of its exploration give way for broad use of clinoptilolite [8]. In the present work we show the results of investigation the physical properties of clinoptilolite based composites obtained during autoclaving process with different amount of lime at 180°C. The comparison between results obtained for materials containing initial clinoptilolite and zeolite after process of ion sorption are presented. The work also describes results of structural studies (IR spectroscopy, XRD analysis and SEM observations) of composites based on zeolites on which sorption of various metal cations has been carried out. Obtained results indicate that the proposed method of solidification the clinoptilolite based materials after heavy-metal uptake is promising and applicable in simple way in Autocalved Areated Concrete technology or production of Calcium Silicate blocks. EXPERIMENTAL Clinptilolite crystallizes in monoclinic C2/m group with the following unit cell parameters: a=17.62.&,b=17.91& c=7.39.& p=116"16'. Theoretical composition of this zeolite is given by (Kz,Naz,Ca)3[A1&30072]'24H20 formula but it can contain some Fe(III), Mn(II), Ba(II), Sr(I1) and Mg(1I) ions. The contents of non-tetrahedral cations: K(I), Na(1) and Ca(1I) as well as AVSi ratio can change in different minerals. Clinoptilolite is an example of the group of zeolite in which SBU (Secondary Building Unit) is built of 4-4-1 complex [9]. It contains 4-, 5-, 8- and 10-membered rings in its structure.

Preparation of clinoptilolite samples Mineral samples containing about 20-25% of clinoptilolite were concentrated by separation process which allowed to obtain maximum content of zeolite material. The concentration procedure included grinding, drylng at 60°C for several days, mixing with water (10 hours, mechanical stirrer) and material decantation. The process was repeated several times. On the basis of X-ray phase analysis the approximate 90% clinoptilolite content was determined. The samples contained also some amount of quartz and insignificant amount of montmorillonite. Clinoptilolite was transformed into sodium form to increase the efficiency of ionic exchange. The zeolite was activated by aqueous solutions of 2M NaCl. Then the Pb(II), Cd(II), Cr(I1I) and Zn(I1) ions were introduced into zeolite structure from 11 mM/dm3 water solution of Pb(N03)2, Cd(N03)2'4HzO, Zn(N03)2'6Hz0and Cr(N0&'9HzO. After ion-exchange process the zeolites were washed with distilled water. In a typical sorption experiment, a suspension of the zeolite in water (20 g/dm3) was shaken with the appropriate metal salt solution for 24

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hours at 25°C in a rotary shaking thermostat. The suspension was then centrifuged at 10000 rpm for 10 min. Basing on the previous studies of the influence of time on the amount of the metals sorbed it was found that 24 hours is a sufficient period of time to reach the equilibrium state between the metal ions in the solution and sorbed on Clinoptilolite [7]. Preparation of composites Pre-treated clinoptilolite was mixed with different amount of Ca(0H)z (analytically pure) in vibration mill for 10 minutes. The amounts of added Ca(0H)z were 8, 12, and 16% of Ca(0H)Z which correspond to the 6, 9 and 12% of CaO, respectively. After homogenization process the cubic shape samples, 20x20~20mm were produced by compacting the damp powder at 15 MPa. Samples were hydrothermally curing under saturated steam pressure at 180°C for 12h. In all cases during samples' forming and curing the distilled water was used. Measurements Infrared spectra were measured on a Bio-Rad FTS-60 spectrometer. Spectra were collected after 256 scans at 4 cm-' resolution. Samples were prepared by the standard KBr (MIR) pellets method. The X-ray structural analysis was performed with Philips X-ray difkactometer X'Pert system using CuK, radiation. The SEM pictures were measured on JEOL 5400 scanning microscopy with microprobe analyser LINK ISIS (Oxford Instrument). Atomic absorption spectrometry (AAS) was used to determine the concentration of metal cations in the solution after the sorption experiments (Philips PU-91OOx). Additionally, pH of these solutions was controlled. The physical properties of composites were obtained by standard method on the cubic samples.

RESULTS AND DISCUSSION Results of heavy metal immobilization processes (AAS results) are presented in Table 1. Total amount of heavy metal ions sorbed on clinoptilolite is a measure of its sorption capacity. Of all the metals studied, sorption capacity decreases in the following order: Cr>Zn>Pb>Cd, but the sorption capacity of chromium are about four times grater then other cations. Table 1. The sorption of heavy metal cations on clinoptilolite Sorption [mM/kg]

The samples obtained in autoclaving process before investigation of physical properties were dried at 105'C. The result of bulk density are shown in Fig.1. As it was expected, the differences between bulk density of compacted samples were in error range, so the influence of sample composition are rather insignificant. Mean value of bulk density of composites is about 1,40 g/cm3.

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Content of Ca(oH),,Wh

Fig. 1. Bulk density of autoclaved samples due to Ca(OH)2 content

In Fig.2 the result of investigation the compressive strength are presented. Effect of lime content is visible for samples with reference clinoptilolite and for clinoptilolite with introducing Pb(II) and Cr(I1I) ions. In those cases the compressive strength of sample with 8% Ca(OH)2 is about 15% smaller than compressive strength of samples with greater amount of lime. For samples with clinoptilolite treated by other cations solutions no significant effects of Ca(OH)2 content were observed. Introduced into clinoptilolite metal cations have influence on compressive strength of composites obtained by autoclaving process. Except of the Pb(II) the effect is small and rather negative. The decrease of recorded compressive strength of samples are visible especially for Zn(II), where the drop of such parameter reaches almost 20%. 20 7

8

16

Content of Ca(OH),, wt%

Fig.2. Compressive strength of autoclaved samples due to Ca(OH)2 content The IR spectra of materials after autoclaving process containing initial clinoptilolite and after sorption process are shown in Fig.3. Four groups of bands on the spectra are present: 1. The bands connected with the internal Si-O(Si) and Si-O(A1) vibrations in tetrahedra or alumino- and silico-oxygen bridges (the range of 1200-400 cm-' ), 2. The bands due to the presence of zeolite water (the range of 1600-3700cm-'), 3. The bands due to pseudolattice vibrations of structural units (the ran e of 500-700 cm-'), 4. The bands connected with COs vibrations (the range 1350-1500cm- ).

P

Inorganic gwcements matrix in composite materials

48 1

Introduction of non-tetrahedral cations into alumino-silicate framework can change their IR spectra in the range of pseudolattice vibrations located at about 700-500 cm-'. I

Fig.3. IR spectra of composites with 12% of Ca(OH)2 On the basis of the MIR spectra of zeolites after sorption in the region of 4000-400 cm-' it is difficult to obtain directly clear information. All the spectra are quite similar. The essential changes in IR spectra caused by heavy metals sorption are observed in the range of the pseudolattice vibrations, i.e. 700-500 cm-' as it has been shown in the previous study [lo], whereas the remaining bands in the IR spectra are almost unchanged. The bands in the range of 715-650 cm-' due to the vibrations of four- and five-membered rings in the zeolite structure, have been analysed in more details.

Fig.4. IR spectra of clinoptilolite after sorption of cations in the range of 715-650 cm-'

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In this range, the bands at about 693 and 676 cm-’ occurred. The latter band is sensitive to changes in the type and the amount of non-tetrahedral cations. This band is connected with symmetrical stretching vibrations of Si-0 bond existing in the 4- and 5membered rings of the zeolite [ 111. The change of integral intensity of this band can give information on the amount of heavy metal cations sorbed in the zeolite structure. In Fig. 4, the IR spectra of clinoptilolite after sorption of cations in the range of 715-650 cm-’ are presented. For all the spectra, the baseline correction has been carried out. Precise estimation of intensity changes is possible after decomposition of the spectra into component bands [lo]. Systematic change in the intensity of the 676 cm-’ band (with respect to the 693 cm-’ band) after different cations introducing is noticeable. Thus, the band at 676 cm“ can be considered as an “indicator” whose intensity changes after heavy metal sorption. However, to estimate precisely the amount of ions sorbed, another analytical method should be used (e.g. AAS). However, the observed changes are much smaller and show different course of metal cations sorption than the AAS results. In Fig.5 the XRD pattern of samples are presented. The patterns show that recorded peaks originate fiom clinoptilolite and incorporated into clinoptilolite grains of quartz. Only weak peaks of tobermorite are visible. No peaks of Ca(OH)z were recorded that means the reactions of lime and clinoptilolite is effective. C - Climptilolite Q Quartz T -TObefmoriteI I A

-

Q

,

10

‘

I

20

.

,

-

30

1

40

.

1

50

.

1

60

2 theta, deg

Fig.5. XRD patterns of composites with 12% of Ca(OH)*

In the Fig.6 - 8 the SEM observation of composites are presented. The composites with in different manner modified clinoptilolite have quite similar textures. In the SEM pictures grains of clinoptilolite bonded by C-S-H phase are visible.

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Fig.6. SEM observations of sample with reference clinoptilolite and 12% of Ca(OH)2, magn.750~(lea), 7 5 0 0 ~(right)

In the Fig.6 is presented SEM observation of sample with Clinoptilolite without introduced heavy metal cations. The C-S-H phase was produced by pozzolanic reaction has rather compact form, but a larger magnifications show that C-S-H phase locally crystallized in needle-like shape forms or tobermorite. This suggests that in that system the hydrothermal reaction may lead to produce bulk paste which provide to the strength of composite.

Fig.7. SEM observations of sample with 12% of Ca(OH)2 and clinoptilolitewith introduced the Cd(I1) ions, magn. 1 0 0 0 ~(left) and Pb(1I) ions, magn. 7 5 0 0 ~(right)

Fig.8. SEM observations of sample with 12% of Ca(OH)*and clinoptilolitewith introduced the Zn(I1) ions, m a p . 7 5 0 0 ~(left) and Cr(1I) ions, magn. 7 5 0 0 ~(right)

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In the Fig. 7 and 8 are presented the result of observations the composites with modified clinoptilolite. The overall view of bulk paste is similar to the previous observation, but in cases of clinoptilolite with introduced heavy metal cations into structure the region of tobermorite formation are more discontinuous and needle-like shapes of C-S-H phase except the samples with chromium modified clinoptilolite are faulted. In case of Cr(III) ions the amount and fibrous shape of C-S-H phase are very similar to the unmodified clinoptilolite.

CONCLUSIONS According to obtained results we formulate the following conclusions: 1. Clinoptilolite is very promising materials to obtain tough composites by hydrothermal reactions with lime, 2. Incorporation of metal cations causes changes in the IR spectra of zeolites in the range of the pseudo-lattice vibrations. This process results in the increase in the intensities of the band at 674 in the spectra. 3. The influence of introduced heavy metal ions into structure of clinoptilolite have rather secondary effect.

ACKNOWLEDGEMENTS This work was supported by Polish Ministry of Scientific Research and Information Technology grant no 3 T08D 039 26.

REFERENCES 1. Jenne, E. A., Adsorption Models. In Jenne, E. A. (ed), Adsorption of Metals by Geomedia: Variables, Mechanism and Model Applications”, Academic Press, San Diego, pp. 11-36 (1998). 2. Inglezakis, V.J., Loizidou, M.D., Grigoropoulou, H.P, 2002. Equilibrium and kinetic ion exchange studies of Pb”, C?, Fe3+and Cu” on natural clinoptilolite. Water Research, 36, (2002) 2784. 3. McBride M.B., Chemisorption and precipitation reactions. In: M.E. Sumner (ed.) Handbook of soil science. CRC Press, New York, B265 (2000). 4. Godelitsas, A., Transition metal complexes supported on natural zeolitic materials: an overwiew. In: Misaelides P. et al (Eds.). Natural Microporous Materials in Environmental Technology. Kluwer Academic Publishers Printed in the Netherlands, 271 (1999). 5. Ciciszwili G.W., Andronikaszwili T.G., Kirov G.N., Filizowa L.D., The Natural Zeolites (in Polish). WNT, Warsaw 1990. 6. Vasylechko V.O., Gryshchouk G.V., Kuz’ma Yu.B., Zakordonskiy V.P., Vasylechko L.O., Lebedynets L.O., Kalytovs’ka M.B., Adsorption of cadmium on acid-modified Transcarpathian Clinoptilolite. Microporous and Mesoporous Materials, 60 (2003) 183. 7. Mozgawa W., Application of Clinoptilolite for immobilization of heavy metal cations (in Polish). Papers of the Commission on Ceramic Science, Polish Ceramic Bulletin Polish Academy of Science - Krakbw Division, Polish Ceramic Society, Ceramikdceramics 66 (2001) 756. 8. Ortega E.A., Cheeseman C., Knight J., Loizidou M., Properties of alkali-activated Clinoptilolite. Cement and Concrete Research 30 (2000) 1641.

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9. Breck D.W., Zeolite molecular sieves, John Wiley & Sons, NY, London, (1974) 10. Mozgawa W., Bajda T., Spectroscopic Study of Heavy Metals Sorption on Clinoptilolite, Physics and Chemistry of Minerals, 3 1 (2005) 706. 11. Mozgawa W., Bajda T., Application of Vibrational Spectra in the Studies of Cation Sorption on Zeolites, Journal of Molecular Structure, 792-793C, (2006) 170.

Proc. Int. Symp. "Brittle Matrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

ADVANCED MODELS FOR THE SIMULATION OF ANISOTROPIC DAMAGE IN FIBRE AND TEXTILE REINFORCED CERAMICS Werner HUFENBACH, Robert BOHM, Albert LANGKAh4P Institute of Lightweight Engineering and Polymer Technology (ILK) Technische Universitat Dresden D-0 1062 Dresden, Germany e-mail: [email protected] ABSTRACT The high lightweight potential of fibre and in particular textile reinforced ceramics for high performance applications can only be used optimally, if the structural components as well as the composite material itself are designed according to the acting loads. The basic requirement for this purpose is the use of validated design and dimensioning concepts that primarily consider the failure and damage behaviour as well as the stress and deformation state. For many practical problems, the induced stresses can be calculated relatively well by means of existing analytical, numerical and experimental methods. In contrast, for the realistic description of the damage behaviour of ceramic composite structures, in particular for those with textile reinforcement, validated models barely exist. At the ILK, first physically reasonable failure criteria have been developed and successfully applied for fibre and textile reinforced ceramics. On the basis of extensive multi-axial fracture tests with unidirectional fibre reinforced ceramics, these novel fracture mode related failure criteria could be confirmed in the practically relevant (9z&tress-plane. , Starting from the advanced fracture criteria, new approaches for the description of the failure behaviour of woven ceramic composites have been developed and verified. Thereby, it could be shown that such ceramic composites could be subdivided into two classes with respect to the failure behaviour. For the simulation of the anisotropic damage behaviour of textile reinforced ceramics, adapted material models have been developed at the ILK. These models realistically describe the different fracture modes as well as the anisotropic damage phenomena under consideration of the existing textile architecture, the matrix system and the fibre-matrix-interface.

Keywords ceramic composites, textile reinforcement, failure criteria, damage models.

INTRODUCTION Recently, reinforced ceramic lightweight materials attract high attention for high temperature and tribology applications under complex working loads. At the same time, a targeted reinforcement with continuous fibres or textiles clearly enlarges the range of application of this novel material group. To use the existing property potential of continuous fibre reinforced ceramics as well as of textile reinforced ceramics for the high performance sector, structural components made of this new material needs to be designed according to the acting loads [ 131. Therefore, a basic requirement is the adoption of assured practical oriented calculation

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methods and design concepts which have to consider the acting stress and strain state as well as the failure and damage behaviour in a realistic way [4,5]. For many practical problems, the induced stresses could be calculated already quite well with the available analytical, numerical and experimental methods. In contrast, for the realistic description of the damage behaviour of ceramic composites, especially for those with textile reinforcement, validated models barely exist. For the predominant number of fibre and textile reinforced ceramic composites, a suitable estimation of the failure behaviour need to include a fracture mode related damage analysis because distinct nonlinearities occur due to micromechanical effects (pseudo-plasticity) [ 6 ] .For the gathering of this non-linear stressstrain-behaviour due to specific damage phenomena, a couple of models based on the foundations of continuum damage mechanics have already been developed [7-91. In fact, these models successfully describe the non-linear deformation incidents but they do not consider the essential influence of different fracture modes. Such fracture mode related damage models are most suitable for an assured evaluation of the failure behaviour of complexly loaded textile reinforced ceramic structures. PHYSICALLY BASED FAILURE CRITERIA As an origin for a realistic failure analysis, uniaxial and multi-axial fracture tests with fibre reinforced as well as with textile reinforced ceramics have been conducted. These experimental investigations with ceramic composites show that the failure behaviour depends on different parameters like the fibre-matrix-combination,the structure of reinforcement or the manufacturing technology. Thereby, concerning their failure, the observed ceramic composites could be subdivided into two major groups (Fig. 1) [2]:

Group 1: Group 2:

Ceramic composites that show a uniform fracture plane which proceeds parallel to the warp or the weft direction of the reinforcement Ceramic composites that show a very brittle fracture in tension with a low ratio of tensile strength to shear strength; in the compression domain, a fibre buckling failure occurs with a heavily jagged fracture distribution

Group 1: (e.g. M30-Sic)

Group 2: (e.g. T300-Sic)

Fig. 1:Failure behaviour of different woven ceramic composites

Group 1 Group 1 contains all ceramic composites that show a largely uniform fracture plane like carbon fibre reinforced woven composites with a Sic-matrix that have been fabricated with the CVI-technique. A realistic description of the failure phenomena in this case could be undertaken according to the so-called action plane related fracture criterion. The introduced

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489

coordinate system and the corresponding fracture angles for the criterion are indicated in Fig.

2. 1 (warp direction)

eft direction)

Fig. 2: stresses in the fracture plane related coordinate system with the associated fracture angles The examination of the experimental fracture tests leads to the following considerations regarding an improvement of the specific fracture hypothesis: The ratio of warp yams and weft yarns of the textile reinforced ceramic composite could not be “demerged” so that a validation based on unidirectional single layers with respect to fibre failure and inter-fibre failure is not possible. The failure occurs in a plane parallel to the warp yarn direction or the weft yarn direction. Thereby, the fracture is characterised by the angles 0 , ~and em. On the basis of these observations, the fracture conditions for woven ceramic composites of group 1 could be formulated. Thereby, two failure cases need to be distinguished. Initially, for the formulation of the fracture conditions, a coordinate system (n, p, t) that is carried along with the action plane is introduced analogue to the action plane related failure criterion of HASHINPUCK (Fig. 2). For woven composites of group 1, the fracture angle 0 , ~always takes the value of 0” or 90”.Therewith, the following quadratic relations for the tension and the compression domain come into consideration as fracture conditions for in-plane loading:

with RY’ denoting the fibre parallel tensile strength, R,, as the shear strength, R E ) as the resistance in the action plane and p l Zas well as pu denoting the slope parameters. For bidirectionally reinforced composites, these fiacture conditions are valid with the following assumptions for the tensile strength and the shear strength:

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Werner HUFENBACH, Robert BoHM, Albert LANGKAMP

RY) = Rit) und R12= R,, .

(3)

For the predominant number of ceramic composites with woven reinforcement, this assumption is feasible because the material behaviour in warp direction and in weft direction only marginally differs. Group 2 A failure behaviour that is characteristic for the second group could be identified in liquidly siliconised (LSI) fibrous ceramics. By means of the performed fracture tests, it could be concluded that a very brittle failure together with a low ratio of tensile strength to shear strength exists. The reason for this high brittleness could be found in the manufacturing process. Within the liquid silicon infiltration process, the silicon carbide matrix arises from the reaction of the molten silicon with the matrix carbon of the preform material. It is assumed that during the infiltration process also the carbon fibres react with the elementary silicon so that damage of the fibre structure is produced. This damage is notably heavily distinct at high degrees of siliconising. Compared to other manufacturing technologies, it causes a distinct reduction of the typical pseudo-plasticity of fibrous ceramics and for this reason a further increase of brittleness. The specimens that failed under tensile loading show a fracture plane that is oriented perpendicular to the first normal stress and that is independent of the orientation of the woven reinforcement. Under compressive loading, a failure due to instability of the fibres in the wake of microcracks frequently occurs before a combined failure due to perpendicular stresses and shear stresses could appear. The evaluation of the fracture tests with LSI-ceramics show that the failure behaviour could be described as a first approximation with an extension of the well known normal stress hypothesis with directional dependent strengths:

with (T, as the absolute value of the highest normal stress. Thereby, the parameter functions R:+)(cp)and R/-) (9) denote the directional dependent tensile strength and compressive strength. The fracture condition (4)is mathematically easy to handle. However, it requires an extensive experimental effort regarding the determination of the directional dependent strengths. Above all, the normal stress failure is also influenced by the interaction of further principal stresses. Such coherences could by proved by means of compression-inner pressure-tests on liquidly siliconised tube specimen. Here it is particularly remarkable that also stresses that are not acting in the fracture plane contribute to failure. SIMULATION OF THE GRADUAL DAMAGE BEHAVIOUR

For a practical design of fibre and textile reinforced ceramics by means of the finite element method, a description of the damage behaviour on the basis of continuum damage mechanics

49 1

Advanced models for the simulation of anisotropic damage in fibre and textile reinforced ceramics

(CDM) is required [lo]. The theory of CDM is based on the assumption that micromechanical damage processes could be specified on the macroscopic scale by smearing the damage phenomena in a definite volume. In contrast, it may be assumed that the micromechanical incidents are considered reasonably exact by means of the aforementioned failure criteria. Then, the damage mechanics approach conduces to the description of the non-linear stressstrain-behaviour after the initial failure. The proposed damage model uses two damage parameters for the characterisation of material degradation, D1 and 4. These damage parameters describe the decrease of stiffness in the principal axes of orthotropy of the researched fibre and textile composites. Additionally, a further damage parameter D, is introduced that describes the change of shear stiffness during increased loading. For the general case of a three-dimensional stress state, the effective stress then results in

a = Do.

(5)

The matrix D transforms the CAUCHY-stresses of the damaged material to the effective stresses of the undamaged smeared material according to the principle of strain equivalence by LEMAITRE.In most cases, stresses in thickness direction could be neglected so that a twodimensional stress state exists

l 1-0,

o

0

1-0,

0

0

0

1

0 1

1-Ds

By inverting the compliance matrix in the damaged state

g=S,,D,

(7)

with SOdenoting the compliance matrix in the initial state, the effective stiffness matrix in the damaged state results in

with d = 1- (1 - DJ(1 -D2)Vp2V;,.

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Then, the constitutive law for the ceramic composite reads

Damage threshold In the framework of continuum damage mechanics, the introduced fracture conditions act as damage threshold functions according to the yield surface in the theory of plasticity. Thereby, the domain of linear deformation is bordered by the fracture surfaces of the different failure modes. The threshold function r which describes the elastic domain during the damage process is a function of the effective stress5,The failure surfacesf; and& describe the failure surfaces in the tensile and compressive domain. For composites of group 1 it is obtained

The effective stresses acting in the failure plane can be calculated by means of transformation conditions from the stresses in the global or the material adapted coordinate system. Damage evolution laws In the literature, already several different damage evolution laws for different composite materials have been published [6-lo]. The main problem in the application of these evolution laws for the description of the damage progress and the decrease of the stiffhesses respectively is the possibility of an assured experimental validation of the associated material parameters. On the basis of experimental tests with different textile reinforced ceramics, a first approximation for an easily manageable damage evolution law has been chosen. Thus, the stress-strain-curve as well as the damage evolution law will be subdivided into four parts:

c

o

if

E, <E,

if

Eg >E,

In the first domain no damage occurs and the stress-strain-curve is linear. After E~ > E, matrix nonlinearities due to microcrack growth arise which are accordingly considered by the damage functionD, ( E ~ ) .After the occurrence of mesoscopical inter fibre failure, the damage function is extended toD,(&,) + D,(E~). Depending on the structure of textile reinforcement and the combination of fibre and matrix, a level of saturation could appear. Then, in the

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particular layer no further inter fibre failure will occur. Further cracks then only arise in other layers until the catastrophic failure of the whole structure is detected.

CONCLUSION The realistic failure criteria and damage models developed at the ILK for fibre and textile reinforced ceramic composites allow a better and more load adapted use of the high lightweight potential of modem structures made of ceramic composites. The possibility of evaluating different fracture modes by means of adapted fracture criteria grants an identification of the highest loaded parts of a structure as well as an estimation of fracture consequences. Therewith, an appropriate foundation for the analysis of the successive failure and damage implications in textile reinforced composite structures is available.

REFERENCES

f& die Praxis. Miinchen, Carl Hanser Verlag, 1995 [2] LANGKAMP,A.: Bruchmodebezogene Versagensmodelle fiir faser- und textilverstarkte Basisverbunde mit polymeren, keramischen und metallischen Matrices, Dissertation, TU Dresden, 2002 [3] KROLL,L.; HUFENBACH, W.: New proof of laminate design by a physically based failure criterion. Proc. of the 10th International Conference on Composite Materials (ICCM lo), Whistler (Kanada), 14.8.-18.8. 1995, Vol. I, pp. 715-722 [4] HUFENBACH, W.; KROLL, L.; LANGKAMP, A.; HOPKEN, J.: Non-linear deformation behaviour of anisotropic fibre reinforced ceramic. Proceedings of the European Conference on Spacecraft Structures, Materials and Mechanical Testing, Braunschweig, 4.-6. November 1998, S. 411-417 A.: Physically based failure [5] HUFENBACH, W.; KROLL,L.; HOPKEN,J.; LANGKAMP, criterion for long-fibre reinforced ceramics. In: High temperature ceramic matrix composites, Hg. W. Krenkel, R. Naslain, H. Schneider, Weinheim: Wiley, 2001 [6] GASSER,A., LADEVEZE,P., POSS, M.: Damage mechanisms of a woven SiC/SiC composite: modeling and identification. Composite Science & Technology 56 (1996), pp. 779-784 [7] CMUS, G; GUILLAUMAT, L.; BASTE,S.: Development of damage in a 2D woven C/SiC composite under mechanical loading: I. Mechanical Characterization. Composite Science & Technology 56 (1996), pp. 1363-1372 [8] EL BOUAZZAOUI, R.; BASTE,S.; C w s , G.: Development of damage in a 2D woven C/SiC composite under mechanical loading: 11. Ultrasonic Characterization. Composite Science & Technology 56 (1996), pp. 1373-1382 [9] ~~ATZENMILLER,A.; LUBLMER,J.; TAYLOR,R.L.: A constitutive model for anisotropic damage in fiber-composites. Mech. Mater., Vol. 20, 1995, No. 2, pp. 125-152. [lo] HUFENBACH, W.; LANGKAMP, A.; BOHM, R.: Erweiterte Modelle zur Simulation der Schtidigung von faser- und textilverstarkten Keramiken. 15th DGM-Symposium Verbundwerkstoffe und Werkstoffverbunde, Kassel, 6.-8.4.2005, pp. 435-44 1

[ 11 PUCK,A.: Festigkeitsanalyse von Faser-Matrix-Laminaten - Modelle

Proc. Int. Symp. "Brittle Matrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

MODELLING OF METALLIC INTER-GRANULAR LAYERS IN POLYCRYSTALLINE CERAMICS Eligiusz POSTEK'), Tomasz SADOWSKI*),Christophe DENIS3) "The University of Leeds, School of Earth and Environment, Institute of Geophysics and Tectonics, Leeds, LS2 9JT, UK e-mail: [email protected] 2, Faculty of Civil and Sanitary Engineering, Department of Solid Mechanics, Lublin University of Technology, ul. Nadbystrzycka 40,20-618 Lublin, Poland e-mail: [email protected] 3)Univer~ity of Pierre et Marie Curie, Paris 6, CNRS U M R 7606, Laboratory of Computer Sciences LIP6,8 rue du Capitaine Scott, 75015 Paris, France e-mail: [email protected] ABSTRACT The aim of the paper is to present a constitutive model for the case of uniaxial tension of the polycrystalline materials, including the inter-granular metallic layers that create its internal structure. The quasi-static deformation process of the material comprises elastic deformation of brittle grains, elasto-plastic deformation of intergranular layers and additional deformation due to micro-porosity development in layers. A Representative Volume Element (RVE)was analysed taking into consideration an initial internal structure of the material obtained f?om SEM photographs

Keywords Ceramics, interfaces, finite strains

INTRODUCTION A typical application of polycrystalline materials is the fabrication of cutting tools. The tools are working in such severe conditions as high dynamic and temperature loadings. An exemplary two-phase material used for them may consist of elastic grains and ductile interfaces. An example of SEM image showing grains, interfaces and their idealization are presented in Fig l a and Fig lb, respectively. The grains can exhibit anisotropic behaviour. The interfaces are thick enough not to be treated as only contacting adhesive layers. Our interest will focus on the behavior of the relatively thick intergranular layers which affect performance of entire sample. The interface material has different grades of porosity.

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Eligiusz POSTEK, Tomasz SAD0 WSm,Christophe DENIS

MATHEMATICAL FORMULATION Incremental equation of equilibrium The problem is elasto-plastic with the assumption of large displacements, [1-31. We consider nonlinear terms of the strain tensor. The virtual work equation is of the form

Figure 1. SEM image of a polycrystal (a), idealization (b). gradient decomposition into elastic and plastic parts (c)

where S and E are the I1 Piola-Kirchhof stress tensor and Green Lagrange strains, f, t and u={u,v,w} are body forces, boundary tractions and displacements. All of the quantities are determined at time t+At in the initial configuration. To obtain the above equation at time t+At in the configuration at time t the relations [4], [S], are used

Now, we apply incremental decomposition to the quantities in the equation above: strains, stresses, displacements and forces f+Af

E=:E + AE, “ $S=: S

+ AS,

u + Au,

f+Af~=‘

‘+” f=‘f +AS, ‘+&t=‘t +At

(4)

Since the I1 P-K tensor at time t in the configuration t is equal to the Cauchy stress tensor the stress decomposition is of the form

Modelling of metallic inter-granular layers in polycrystalline ceramics

497

Then, we employ the following strain increment decomposition into its linear and nonlinear parts in the following form -

AE = Ae + Aq, , Ae = a u , Aq = A(AU')AU'/ 2

(6)

whercAul_is the vector of the displacement increment derivatives w.r.t. Cartesian coordinates and ( A , A ) are the linear and nonlinear operators, [2].

-

ax

-

A=

a

0 - 0 aY a 0 0 -

az

a a

o

ay ax a 0 -a az ax a o - - a az

AunX 0 0 Av, 0 0 Awvx 0 0 0 Auoy 0 0 AvSy 0 0 Aw,, 0 0 0 Auqz 0 0 Avaz 0 0 Awq2 (7) AuaY AusX 0 AvVy AvnX 0 AwvY AwtX 0 0 AunZ AuVy 0 AvSz Any 0 Awfz AW,~ Au~= 0 A u - ~ AvvZ 0 Avox AwaZ 0 Aw~,

ay

Substituting the described relations, into the virtual work equation, Eqn 3, and assuming that the equation is precisely fulfilled at the end of the step we obtain the following incremental form of the virtual work equation

Employing the finite element approximation Au = NAq and A d = B',Aq, where N is the set of shape functions and Aq is the increment of nodal displacements and considering the following set of equalities

where :T is the Cauchy stress matrix

we obtain the following discretized form of the virtual work equation

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Eligiusz POSTEK, Tomasz SAD0WSKL ChristopheDENIS

Now, we will deal with the constitutive model and employ the linearized constitutive equation, in fact with the stress increment, AS.

Finite strains When considering the finite strains effect [6], [7], the gradient F=a(X+u)/aX is decomposed into its elastic and plastic parts, F =F'FP ,Fig. lc. To integrate the constitutive relations the deformation increment AD is rotated to the un-rotated configuration by means of rotation matrix obtained from polar decomposition F = VR = RU , Ad = R~+lADR,+, , then the radial return is performed and stresses are transformed to the Cauchy stresses at n+l, an+l= R n + l ~ ~ + I R The ~ t stresses I. are integrated using the consistent tangent matrix [8] and the integration is done in the un-rotated configuration as for small strains.

Stress updating procedure To integrate the constitutive relations we exploit the relations given above using the integration for the un-rotated configuration and the backward integration rule. The algorithm arises f?om the depicted above relations for rotated and un-rotated strain rates and Cauchy stresses. The outline of the integration scheme is given below. 0

Compute deformation gradient

.Compute polar decomposition

0

Compute deformation increment over the step

0 Now, we take the elements of the strain increment A d and obtain the AD' and perform rotation of the increment of spatial deformation to the un-rotated configuration

Ad' = R;,AD'R:,, 0 Then, we perform integration of the small strains constitutive model using backward Euler integration rule (predictor - corrector) where the stresses depend on the history, this is reflected by the stresses at time t and internal variables a,.

Modelling of metallic inter-granular layers in polycrystalline ceramics

0

499

Transform the stresses to the true Cauchy stresses at t +At .

The integration in the un-rotated configuration is performed using a consistent tangent formulation, [81.

Constitutive model. The constitutive model is the Gurson Tvergaard model [9-111 with the yield function as follows

where o Mis the Mises stress, omis the mean stress, B is the Mises stress in the matrix,fis the void ratio and ql,q 2 ,q3 are the Tvergaard coefficients. The stress integration algorithm comprises the elastic trial stress (predictor) and the corrector. It conforms the radial return algorithm. The algorithm can be derived basing on [ 121. The elastic trial stress are of the form =am+ D

The deviatioric and the volumetric stresses are of the form

The increment of the plastic strains can be obtained fiom the normality condition aF = d;l--.

aa '

Further, the plastic strains increment and the unit normal vector are of the form

The volumetic and deviatoric plastic strains are as follows

The stress at the end of the step after performing radial return

Eligiusz POSTEK, Tomasz SADOWSU, Christophe DENIS

500

Since the increment of plastic strains is 1

Asp' = -AcpI + Acqnm+l 3 the stress at the end of the step may be expressed as follows

- KAc,I

G,,,+~ =

- 2GAcqnm+,

Finally the updated stresses are of the form T

3GAE

= om+,- KAE,I -+Si+l qm+l

where the (T) indexed values are the trial stresses.

NUMERICAL RESULTS The mechanical properties of the polycrystal consisting of elastic grains (tungsten carbide) and metallic interfaces (cobalt) are as follows: grains; Young's modulus 4.lxlO"Pa and Poisson's ratio 0.25, interfaces: Young modulus 2.1x101'Pa, Poisson's ratio 0.235, yield limit 2.97xlO"Pa and small hardening modulus 1.0xlO'Pa. The dimensions of the sample are lOOxlOOxl0 pm.The scheme of the RepresentativeVolume Element (RVE) is given in Fig. 2.

(4

(b)

(c)

Figure 2. Mesh of representativevolume element (a), interfaces (b), considered joint (c).

501

Modelling of metallic inter-granular layers in polyciystalline ceramics

I

d

P

W - I . ,

, , > o . 6 ~ a7 0.72 0.74 0.76 0.78

,

as

.

,

0.82 0.84

. c

,

o.&

0.~8 0.9

lnnd factor

lnad farlN

(a) Figure 3. Parametric study: porosity, displacements at the centre of the loaded face (a) and equivalent plastic strains in the considered joints (b) versus loading factor. U W

0.W

ams

a

0.m

’-$ a m

4

0.m

amis

6

k

o

a

J

om

,

,

aoz

,

oai

,

ow

,

pas

POm1tv

(4

, OM

,

om

, a08

1

oor

O w a r J , , o

om

urn

, om

,

aw

,

OM

Pomihl

,

om

,

om

, aos

c

ow

(b)

Figure 4. Maximum displacement of the midpoint of the loaded face (a) and maximum equivalent plastic strain (b) in the considered interface joint. The sample is discretized with 48894 elements and 58016 nodes. The sample is fixed on one side and loaded with the uniform pressure of 400 MPa on the other one. There is imposed symmetry condition in the bottom of the sample. Since the grains are elastic the sample fails due to large plastic strains occurring in the elasto-plastic interfaces, in particular in joints [13, 141. However, in this case, we focus our attention on the additional deformation and additional plastic strains occurring in the interfaces due to existence of initial porosity therein. We will consider a “control node” at the midpoint of loaded face of the sample and a joint between 4 grains (Fig. 2b and Fig. 2c) where the plastic strains start to appear early. The sample is loaded until 89% of the total load. The parametric study is performed for the values of porosity in the interfaces such as: 0.005,0.01,0.05,0.09 (Fig. 3). The displacements and plastic strains strongly depend on the porosity of the material. This is presented in Fig. 3 and Fig. 4. We start to follow the equilibrium path at 70% of the total load when the plastic strains begin to develop for lowest value of porosity. The horizontal displacement of the chosen node increases nonlinearly and at the end of the process (89% of the total load), for porosity 0.09 is 15.3% higher than for the material with porosity 0.005. The dependence of the plastic strains development in the chosen joint on the porosity of the material is dramatically strong. The equivalent plastic strains are 362% higher for the material with the

502

Eligiusz POSTEK, Tornasz SAD0WSKJ Christophe DENIS

high porosity (0.09) then for the material with the small porosity (0.005). The dependences of the maximum displacement and the maximum equivalent plastic strain on porosity at the end of the process are shown in Fig. 4a and Fig 4b, respectively. The dependences for the considered two-phase material are nonlinear and strongly dependent on porosity. Let us have a closer look at the qualitative effects of the porosity on the behaviour of the composite. The displacement fields are presented in Fig. 5 and Fig. 6. The sample containing interface material of low porosity is presented in Fig. 5 while the sample with the high porosity interface material is given in Fig. 6. The difference of the behaviour of the two samples is significant. The pictures of the displacement fields of both samples (Fig. 5a and Fig 5b) are qualitatively almost the same at the beginning of observations. The "control displacements" are almost the same (Fig. 3a). The plastic strains are starting to develop in the interface material of low porosity and are still small in the interface material of high porosity (Fig. 3b). When observing the surface of the samples (Fig. 5a and Fig. 6a) we may notice only spots of plastic strains.

Ir,

-1

%RE%&

Ir,

""I

y?s%4 L o w -

(a) (b) Figure 5. Displacement fields at load factor 0.7 (a) and at load factor 0.89 (b), porosity 0.005.

(a) (b) Figure 6. Displacement fields at load factor 0.7 (a) and at load factor 0.89 (b), porosity 0.09.

Modelling of metallic inter-granular layers in polycrystalline ceramics

503

Figure 7.Equivalent plastic strain distribution at load factor 0.7 (a) and at load factor 0.89 (b), porosity 0.005.

(b) load factor 0.7 (a) and at load factor 0.89 (b), 0.09.

Figure 9. Equivalent plastic strain distribution in the interface at load factor 0.7 (a) and at load factor 0.89 (b), porosity 0.09.

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Eligiusz POSTEK, Tomasz SAD0 WSU, Christophe DENIS

The picture changes at the end of the loading process. Both samples undergo plastic deformation. It can be concluded analysing Fig. 3 and Fig 7b and 8b. The last two figures show the distribution of the equivalent plastic strain on the surface of the samples at the end of the loading process. The plastic strains are developed in the interfaces. Coming back to the displacement fields we may notice discontinuities in the fields. The discontinuities are much more distinct in the sample with high porosity interface material (Fig. 6b) than in the sample containing low porosity material (Fig. 5b). The discontinuities qualitatively manifest the sliding of the grains because of the plastic deformation in the interfaces. The development of plastic deformation can be observed better in Fig .9a and in Fig. 9b. FINAL REMARKS The influence of porosity of the interfaces material was investigated. It has been found that the behaviour of the two-phase composite strongly depends on the porosity of interfaces in the material. Amount of porosity changes the qualitative behaviour of the two-phase material. High porosity values cause earlier appearance of the slips in the interfaces and it is anticipated that decreasing of failure loads, as well. Summarising, we have shown that the composite structure is particularly sensitive to the development of porosity in the ductile interfaces, namely, the equivalent plastic strains in the interfaces increase rapidly with the increase of porosity. It implies that the composite is very sensitive to imperfections in the interfaces as well. ACKNOWLEDGMENTS

T. Sadowski and E. Postek are currently supported by Polish Ministry of Education and Science - grant SPB, decision No 65/6.PR UE/2005-2008/7. The support of the Civil & Computational Engineering Centre at UWS is appreciated. REFERENCES 1. Owen, D.R.J., Hinton, E., Finite Elements in Plasticity: Theory and Practice, Pineridge Press, Swansea 1980 2. Bathe, K.J., Finite Element Procedures, Englewood Cliffs, New Jersey, Prentice Hall, London 1996 3. Kleiber, M., Incremental finte element modelling in non-linear solid mechanics, Polish Scientific Publishers, Warsaw, Ellis Horwood, Chichester 1989. 4. Malvern, L.E., Introduction to the Mechanics of Continuous Medium, Englewood Cliffs, New Jersey, Prentice Hall, London 1969. 5. Crisfield, M.A., Non-linear Finite Element Analysis of Solids and Structures, John Wiley, New York 1991. 6. Pinsky, P.M., Ortiz, M., Pister, K.S., Numerical integration of rate constitutive equations in finite deformations analysis, Computer Methods in Applied Mechanics and Engineering, 40, 1983, pp. 137-158 7. Peric, D., Owen D.R.J., Honnor, M.E. A model for finite strain elasto-plasticity based on logarithmic strains: Computational issues, Computer Methods in Applied Mechanics and Engineering, 94,1985, pp. 101-118

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8. Simo, J.C., Taylor, R.L., Consistent tangent operators for rate independent elastoplasticity, Computer Methods in Applied Mechanics and Engineering, 48, 1985, pp. 101-1 18 9. Gurson, L., Continuum theory of ductile rupture by void nucleation and growth. part I yield criteria and flow rules for porous ductile media, Journal of Engineering Materials and Technology, Transactions of ASME, 99, 1977, pp. 2-15 10. Tvergaard, V., On localization in ductile materials containing spherical voids, International Journal of Fracture, 18, 1982, pp. 237-252 11. Tvergaard, V., Influence of voids on shear-band instabilities under plane strain conditions, International Journal of Fracture, 17, 1981, pp. 389-407 12. Simo, J.C., Hughes, T.J.R., Computational Inelasticity, Springer, New York-London 1998 13. Postek, E., Sadowski, T., Hardy, S.J., The mechanical response of a ceramic polycrystalline material with inter-granular layers. In: Proc. “VIII International Conference on Computational Plasticity, COMF’LAS VIII”, E. Oiiate, D.R.J. Owen eds. Barcelona Sept. 2005, CIMNE Barcelona 2005 14. Sadowski, T., Hardy S., Postek, E., A new model for the time-dependent of polycrystalline ceramic materials with metallic inter-granular layers under tension, Material Science and Engineering A, 2006, pp. 230-238

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, V.C. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

A NEW MICROMECHANICS BASED PREDICTIVE METHOD FOR POROUS CERAMICS BEHAVIOUR UNDER COMPRESSION

Sylwester SAMBORSKI I), Tomasz SADOWSKI 2, of Applied Mechanics, Faculty of Mechanical Engineering Lublin University of Technology 36 Nadbystrzycka St., Lublin 20-618, Poland, e-mail: [email protected] 2)Department of Solid Mechanics, Faculty of Civil and Sanitary Engineering Lublin University of Technology 40 Nadbystrzycka St., Lublin 20-6 18, Poland, e-mail: [email protected] ')Department

ABSTRACT A new research method of the mechanical characteristics estimation and damage assessment for porous ceramics under compression was hereby proposed. Micromechanical approach concerns initial material structure parameters before loading. Phenomenological modelling fmds the relation between the stresses and the strains. SEM analysis of initial material structure and quasi-static compression of ceramic specimens with unloading and multiple reloading gives a lot of material data for modelling. A detailed analysis of material constants and damage assessment is thus possible.

Keywords: Porous ceramics, crack growth, damage evolution INTRODUCTION Nowadays ceramic materials have wide range of applications, for example as thermal coatings in steel or cement fabrication furnaces, ceramic layers in combustion engines, structural parts of military equipment etc. It is justified with their excellent mechanical and thermal strength. However, the main disadvantage of ceramics application is its brittleness [ 11. Thus, classical assumption that ceramics is a linear-elastic-brittle is no longer valid. The global behaviour of the material under loading is non-linear. It is necessary to investigate internal structure of the material and degradation process of material to predict properties. For these reasons we propose a new method of damage assessment in a polycrystalline ceramics with initial structure consisting of: grains, boundaries, pores, inclusions etc. The idea of the method results from the observation of the loading-unloading-reloading process of the material, which is connected with the analysis of the strain stage. Experimental results are compared with theoretical modelling of ceramics behaviour by application of micromechanical and phenomenological approach. Examination of brittle materials behaviour in the process of controlled deformation, estimation of the global material stiffness tensor by deformation analysis and current anisotropy assessment (dependent on damage evolution) are the main advantages of the proposed method. Modelling of material damage development is made both by micromechanical and phenomenological approach. The micromechanical

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Sylwester SAMBORSKI, Tomasz SADOWXKI

approach requires determination of the initial material structure by Scanning Electron Microscopy (SEM). The phenomenological approach covers experimental testing of the material under quasi-static compression with observation of several repeated cycles of loading, unloading and subsequent reloading, what enables a detailed analysis of the strain stage changes and further - evolution of elastic material properties as well as damage assessment.

MODELLING OF MATERIAL BEHAVIOUR IN COMPRESSION

Micromechanicalapproach At this level of polycrystallineceramics modelling one must pay special attention to structure characterization. SEM observations provide necessary information describing the structure morphology, i.e.: grain sue distribution, grainshape, n boundaries texture, pores sue distribution, the shape of pores, P crack patterns (lunking), the preferable way of crack growth (intergranullaror transgranullar). The main material parameters in micromechanical modelling are resistances to crack propagation: along grain boundaries and through grains. In micromechanical modelling we analyse a representative part of the material. For simplicity we consider a two dimensional model and so called Representative Surface Element (RSE). It contains significant number of grains, representative for the whole sample. The exemplary structure of RSE is shown in Fig. 1. In the case of porous polycrystalline ceramics it consists of polygon grains and boundaries with pores initially spread inside grains and along the grain boundaries. Pores are assumed to be spherical, with the radius d4),for each q-th pore . Pores are stress concentrators and in the prospect - the fracture origins [S]. It is related to differences of thermal expansion coefficient for various crystallographic directions of the neighbouring grains, which leads to microcracking during cooling phase of fabricationprocess. With the progress of the loading process cracks can initiate in the specimen from pores and spread along straight segment of grain boundaries. This is due to fact that surface fracture energy of grain boundaries ( y,) is lower than the surface fracture energy of grains ( yg). In most cases one can assume, that:

Afier load increase, the number of straight cracks N, gets higher. Subsequent crack propagation is related to local stress state ( q j )and is connected with the crack kinking in order to expand along the next segment of grain boundary (wing crack initiation). In theoretical modelling it is possible to estimate some privileged directions of cracks propagation.

A new micromechanics based predictive method for porous ceramics behavior under compression

509

Figure. 1. Representative Surface Element (RSE) Due to the small value of the overall material strains we can decompose the total strain tensor E~ into several parts [3]. Such approach gives us the possibility to describe separately: elastic behaviour, porosity presence and crack development during loading, as follows:

Let us define local strain state at material point by tensorial functions describing appropriate induced by cracks 2 ; and 2 ; due to deformations: elastic $, porosity dependent ,:?t dislocation slip (in semi-brittle ceramics, for example MgO). We are now able to estimate macroscopic strain state performing averaging procedure over the RSE:

for purely elastic deformation,

to describe pore induced strains,

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Sylwester SAMBORSKI, Tomasz W OWSKI

for crack influence on the material response and:

as plastic strains. Bearing in mind, that A describes a surface area of the RSE. A$’, A:’, 4:’ are the areas of the: q -th pore, s -th crack and v -th grain with dislocation slip, respectively. N , is the number of pores, N , - the number of cracks and N,,, - the number of grains with plastic effect in the RSE.We do not take into account evolution of porosity during loading.

Phenomenologicalapproach The basic phenomenological relation, expressing the macroscopic strains as a function of applied stress is:

for i, j, k, 1, m,n=l, 2, 3. Sgu is the compliance tensor, which reflects the damage of material structure due to load increase. W~ is the damage tensor of rank two, p stands for porosity and @ - for plastic effects. Coupling the relations (2) to (6) with (7) we may compute the components of the compliancetensor.

Loading-Unloading-Reloadingprocess and damage description The scheme of loading-unloading-reloading process is shown in Fig. 2. The first step is loading to a certain force 4 , corresponding to the stress a?’.Then, we start unloading to almost zero load. After that - reloading to a force F, = 4 + AF , unloading, reloading to a force F3 = F, + AF and so on, until the rupture load is reached. The force increment AF (or stress increment AaL) depends on the expected material strength to failure. As a consequence of the presented loading scheme, mechanical features of ceramics may be estimated at each stage of the process. It is possible to estimate experimentally or theoretically calculate the following characteristics: a initial Young modulus ( E L ) ,which is a function of material porosity, a unloading moduli ( E E ) , for each stress value a(,.). For the purpose of damage descriptionwe introduce a scalar damage parameter ( W ) [2]. This feature characterizes the damage state of the material in the cross-section perpendicular to the loading axis. Thus, the longitudinaldamage can be defined as the function of the initial density of the material and current level of loading:

c$“

A new micromechanics based predictive method for porous ceramics behavior under compression

5 11

Figure 2. Scheme of loading-unloading-reloadingprocess Calculation of unloading moduli E$) was done on the basis of the total longitudinal strains ( E ? ' ) and the corresponding permanent strains ( E:(")), i.e. remaining after total unloading for each stress level dn). One can calculate the difference between the total and permanent strains:

as Then it is easy to estimate the unloading modulus E k ) for the stress level o(")

Additional information about the material damage and it is mechanical features evolution may

5 12

Sylwester W B O M U , Tomasz SAD0 WSKl

be obtained by the measurement of the circumferential strains, specimen, Fig. 3. The similar plot to Fig. 2 can be obtained for transversal strains

(

possibility of defining transversal damage of the material q p,

&))

E$),

of the cylindrical

E$).

Thus, there is the

in the similar manner

to (8):

where ET(p) is the Young modulus in transversal direction and E$ is the corresponding value after total material unloading. That of' means, we can generalize the scalar damage parameter to second order damage tensor two as follows:

0

{mq}={;;

"12}={

@22

w , o 0

}

(12)

@L

The value of E$ ,which is necessary to calculate q , can be obtained by considering the following formula: (4

E g ( p , ct)) =- v p

A$)

'

(13)

0

where: A€$') denotes difference between total strain E?) and corresponding Figure 3. Strain measurement with gauges permanent strains E$") after unloading of the material in the transversal direction. Here vp is the Poisson's coefficient of initially porous material.

of'

EXPERIMENTAL PROCEDURES

Material samples Two sets of ceramic samples were made by powder sintering in the Institute of Electronic Materials Technology in Warsaw. There were four sets of alumina (A1203) specimens of the porosities ( p ) from about 3% to 38%. The samples of magnesia (MgO) have from 9,8% to almost 21% of pores. All specimens were cylinders of the average dimensions 14 x 5Omm (see Fig. 3).

5 13

A new micromechanics based predictive method for porous ceramics behavior under compression

Measurement A system of four strain gauges (“Vishay” EA-06-240LZ) was glued on each specimen. They were delivering continuous data of axial and circumferential strains. Each gauge worked in quarter bridge, monitored by the “ESAM Traveller” system consisting of 8-channel bridge coupled with the PC and supervised by the adequate software. Samples were loaded by the universal testing machine “Zwick 2100”. The force increment (AF) was correlated to material porosity and the expected rupture load. Straining velocity was of the order of 100 pdmin. Experimental outcomes in the following section. RESULTS AND DISCUSSION

The Young modulus of fully dense material was estimated to be 410GPa for A1203 and 316,4GPa for MgO (cf. [l]). The plots included below show the evolution of unloading moduli in axial and transversal direction: E$ and E g , respectively. Figure 4 presents the influence of the initial porosity on the values of unloading moduli for MgO at the very moment of rupture. As one can see, there is significant difference between the moduli measured in both directions. Namely, unloading modulus in transversal direction is always lower. Moreover, the higher initial porosity, the bigger difference between moduli. It is reflected by macroscopis fracture scheme, i.e. samples are axially splitting at the last stage of loading process (Fig. 6; cf. also [4]).

O J

a

I 12

P [%I

1

1 16

20

24

Fig. 4. Unloading moduli at the moment of rupture for magnesia vs. initial porosity The behaviour of alumina seems to be similar. However, the values of moduli are higher and the distribution of porosity is wider.

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Sylwester SAMBORSKI, Tomasz SADOWSKI

1 , I I I I 1 04 ’ P 0 10 20 30 40 Fig. 5 . Unloading moduli at the moment of rupture for alumina vs. initial porosity

[“/.I

Fig. 6. Axial splitting of ceramic sample in compression Consecutive figures contain information about damage at the moment of rupture. In accordance to the relations given in preceding sections the values of damage tensor are

A new micromechanics based predictive method for porous ceramics behavior under compression

5 15

connected with unloading moduli E , and Em. Figure 7 and 8 show evolution of damage parameters in loading direction (@))and in perpendicular direction (q').

8

12

16

20

24

D .-I

[%]

Fig. 7. Damage parameters at rupture for magnesia vs. initial porosity

A1203

0,400

0,200

r* I

0,000 0

' -I

01

~

0

t

- --Q-- - - 10

cc____---

I 20

1

30

40 ~

Fig. 8. Damage parameters at rupture for alumina vs. initial porosity

P [%]

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Sylwester SAMBORSKI, Tomasz SADOWSKI

Again, axial splitting of the samples is reflected in bigger values of transversal damage: @) is always higher than It is worth noting that samples of both ceramics, which have higher initial porosities, are clearly more “damage tolerant”. Namely, they can withstand higher damage until they fail. For example, in case of MgO (Fig. 7): @)=0,15 at p = 10%

6).

and

yo” = 0,70 at p = 21%. The alumina exhibits analogous properties (Fig. 8). CONCLUSIONS

The aim of our current research is to develop a new research method. This method gives the possibility of assessment of influence of material porosity, and the current damage state on the mechanical strength of ceramics. Comparison of theoretical modelling with results of experiments will enable a quick assessment of strength of ceramic parts of construction, that are subjected to ftequent load changes. Experiments show considerable influence of initial material porosity on mechanical strength of ceramics. ACKNOWLEDGEMENT

This work was funded partly by: 1) State Committee for Scientific Research (Poland) within the years 2004-2006 as a research project, grant 3 T08D 027 26 and 2) Ministry of Education and Science (Poland) grant No 65/6.PR UE/2005-2008/7. REFERENCES

1. Davidge R.W., Mechanical Behaviour of Ceramics. Cambridge University Press, Cambridge 1979 2. Lemaitre J., A Course on Damage Mechanics. Springer-Verlag, Berlin 1996 3. Moss W.C., Gupta Y.M., A Constitutive Model Describing Dilatancy and Cracking in Brittle Rocks. J. Geophysical Research, 87, 1982, pp. 2985-2998 4. Sadowski T., Mr6z 2.: Deformation process of low MgO ceramic cylinder subjected to compressive loading. In Proc. Int. Symp. “Brittle Matrix Composites 3”, A.M. Brandt and I.H. Marshall (eds.), Warsaw 1991, Elsevier Applied Science, London 1991, pp. 366-376 5. Sammis C.G., Ashby M.F., The Failure of Brittle Porous Solids under Compressive Stress States. Acta Metallurgica, 34, 1986, pp. 5 11-526

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

THE INFLUENCE OF MODIFICATION OF THE STRUCTURE OF SILICATE MATERIALS ON THEIR PROPERTIES AFTER NON-AUTOCLAVED HARDENING Elena S. SHINKEVICH, Yevgen S. LUTSKIN, Olga P. GNYP, Aleksandr A. KOICHEV, Juliya V. DOTSENKO Odessa State Academy of Building and Architecture Didrihsona 4, 65029 Odessa, Ukraine, e-mail: [email protected]

ABSTRACT The introduction of modern equipment in practice as the speed mixer-activators allows to realizing the phenomenon of mechano-chemistry in the production of silicate materials on the base of lime-silica binding substance. The activating of mixture of raw materials as slip defmed the transition from autoclave treatment to thermo-moisture one of silicate materials, the optimization of the composition and the of hardening conditions in the conditions of thermo-moisture treatment on the basis of experimentally-statistical modeling ensured the receipt of materials with the required properties. The possibility of the practical realization of reserve of structure mineral substances for the energy consumption of production silicate materials are proved experimentally. The analysis influence "mixture-technology-structure-properties" has been fulfilled on experimental-statistic models. The changing of properties of silicate materials under the influence of surface of mineral additive, of hardening conditions and content of gypsum addition have been estimated. Correlation analysis allows to receive new information about the influence of the factors of structure and the technology on a degree of correlation the between structure and the properties of building materials. Optimal compositions and the regimes of hardening are recommended for the receipt of articles of a different functional purpose.

Keywords Lime-silica binding substance, slim, non-autoclave silicate materials, experimental-statistic models, walls articles, coefficient of heat conductivity, computer Science-Materials, activation, optimization INTRODUCTION The conservation of energy problem is actual for building industry. The development and the introduction of resources conservation technologies are favour reduction of power imputes during the production of articles. The receipt of articles with the improved heat-physical properties cuts down the losses of heat on the stage of exploitation, too. Hydrosilicates calcium imparts the compressive strength properties to the silicate materials and is forming during the process of interaction lime with silica in the conditions of the saturated water steam in the autoclaves. The higher temperature (175-200°C) and the higher pressure (0.8-1MPa) of the autoclave treatment are needed for the intensification of physical and chemical processes in the system "Ca0-Si02-H201', at first, for the creation of the solutions oversaturated by the ions Ca2' and Si2-.The synthesis of hydrosilicates calcium is possible and thermodynamically advantageous under the ordinary conditions: normal temperature and pressure (T=2W2"C and P=O. 1MPa) with the use of active forms of silica [13.

The possibility of the receipt of silicate materials under normal temperature and pressure allows speaking about the substantial potential reserves of their production. THE BASIC THEORETICALPOSITIONS

The joint chemical and mechanical activating is used for decreasing temperature and pressure of thermal treatment in this research. The mechanical activating is carried out in the speed mixer-activator. The introduction of modern equipment in practice as the speed mixeractivators allows to realizing the phenomenon of mechano-chemistry in the production of silicate materials on the base of lime-silica binding substance [2]. The higher activity of the components of the mixture of raw materials as slip after the treatment in the mixer-activators is involved the deliverance of atoms, ions and electrons from CaO and Si02 and the formation defects and dislocations on the surface of crystalline substances. During the process of activation liquid phase is saturated by the ions Ca2' and SiOY, the dissociation of the molecules of water is accelerated with the formation of the active groups H' and O H and the conditions for the synthesis of hydrosilicates calcium at low temperatures and atmosphere pressure are created. The activating of the mixture of raw material as slip defined the transition from autoclave treatment to thermo-moisture one of silicate materials, the optimization of the composition and the of hardening conditions in the conditions of thermo-moisture treatment (TMT) on the basis of experimentally-statistical modeling ensured the receipt of materials with the required properties. The compositions and the method of the preparation of the activated mixture, which provides the receipt of material in the conditions of TMT with lower bulk density and heat conductivity (in comparison with autoclave silicate materials) and the standardized compressive strength, water-resistance and frost-resistance are worked out [3]. The method of the preparation of the mixture of raw material as slip provides the optimum time of the activating and the certain sequence of dosage in the mixer-activator of water, the binding materials (lime, ground quartz sand, the mineral addition), unground quartz sand, the additions of superplastiphicationand gypsum. The chemical activating is carried out by the introduction chemically active silica into the binding substance. The porous addition is introduced instead of the ground quartz sand. It cuts down expenditures for the grist substantially. The optimum quantity of mineral addition is determined during the researches. The deterioration of the effective toughness of the mixture during the process of the activating allows to introduce the mineral addition instead of quartz sand without the increase of water-hard materials ration [4]. It compensates the negative influence of the higher water requirement of the mineral addition to frost-resistance. The use of mineral addition with a different specific surface is a next peculiarity. The mineral additions are traditionally powdered to the specific surface of binding materials or used them with their natural ultradispersed specific surface. However, the mineral additions with low chemical activity are able to influence on the structuroformation and the properties of materials greatly. EXPERIMENT

The dependence of influencing of values of specific surface of mineral addition, of the content of gypsum addition and of the of hardening conditions on the complex of the parameters of the modified structure and the properties with the purpose of the recommendation of optimum

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technological decisions of production of silicate materials after non-autoclave hardening are the subject of this research.

Experimental-statistic modeling The forming of informative base for the optimization of technological decisions is provided by the use of the methods of the experimentally-statistical modeling (ES-modeling). The multifactor experiment was carried out with the use of the mathematical theory of planning of the experiment. Two comparable complexes of six factor experimentally-statistical models (ES-models) are computed. ES-models describe the dependence "mixture-technologyproperties" and "mixture-technology-structure" [5]. It allows the dependence "contenttechnology-structure-properties"to study. The ultimate compressive strength (Rb),frost-resistance (F) and the coefficient of heat conductivity (A) are resulted in this research. The next characteristics of structure are researched: the content of hydrosilicates calcium of a different mineral composition, the degree of their hydration, general, opened and closed kinds of porosity, their numeral correlations, the relative size and distribution of pores according to the sizes. The change of the properties and the modification of the structure were carried out at the expense of variation quantity and quality composition of the mixture of raw materials and the of hardening conditions. The specific surface of mineral addition was fixed at followin levels as three factors of the composition of mixture: ul=SsP1=350 m2kg, u2=SSp2=425 m k g , u3=S8,3=5O0 m2/kg. These factors are bound by the linear dependence: u1+2)2$2)3=1. The of hardening conditions were regulated by two technological factors: the duration of preliminary ageing ( T ~ , ~in. ) the normal conditions at the temperature of 20+2"C - from 0 till 12 hours (factor &) and the duration of thermo-moisture treatment ( T ~ T from ) 10 till 18 hours (factor Xs) at the temperature of the isothermal ageing 80 ...85°C and the relative humidity 100%. The gypsum addition changed from 0.0 to 5.0 percents from the mass of dry components (factor &). It allowed to compare the compositions without gypsum and with a different content of gypsum. The addition of gypsum was added for the adjustment of reological properties of the mixture of raw materials as slip. There are normalizing factors x, by formula: xi=(Xi-X~i)/Xi,where -1Sxi51. The experiment were camed out by constant water-hard of materials ration (W/H). As a result of the realization of experiment six-factor ES-models, describing the change of the characteristics of the structure and the properties under the influence of the specific surface of mineral addition, of the gypsum addition and of the of hardening conditions are calculated. The content of the gypsum addition, the of hardening conditions form the space as a cube on the diagrams. The influence of specific surface of mineral addition is described by the three-cornered diagrams which moved on the field of the three-dimensional cubic diagram of the technological factors for the search of optimal decisions during the analysis [6].

8

Compressive strength The change of compressive strength under the influence of specific surface of mineral addition, the content of gypsum addition and the of hardening conditions are described by the model (1).

520 Elena S . Shll?KEmH, Yevgen S.LUKWlN Olga P.GNYR Aleksandr A. KOICHEK Juliya V.DOTSENKO

Under the influence of all six factors the relatively change of compressive strength 6%, calculated according to the complete model equal (1) - 6Rb=Rbm&,-6.4. It shows the high influences of this technological parameters on the compressive strength. According to complete model (1) the model with optimum values of specific surface of mineral addition is calculated which is graphically shown on fig. 1. Isosurface of the compressive strength inside the cube is calculated according to the complete model at the optimum values of specific surface of mineral addition providing the individual maximums of compressive strength in the factor space of each of three-comered diagrams. The maximums of compressive strength for compositions with gypsum and without it differ 4 times.

- 0,027V13

-0,072v1X5 hOv3~5

- 0,354 X26

+0,394Vlx6 +O,481v3%

- 0,011X5X6

The role of gypsum in the forming of compressive strength changes in depending of the values of specific surface of mineral addition. So, the maximal values Rb'-{cg=O%}=10.9MPa got for the compositions without gypsum and &""{cg=5%}=22.5MPa addition of gypsum.

Fig. 1. Changing of compressive strength under the influence of hardening conditions and content of gypsum addition (isosurfaces are inside the cube); hardening conditions (isolines on the square diagrams); specific surface of mineral addition (isolines on the triangular diagrams).

The influence of modifcation of the structure of silicate materials on their properties after..

52 1

The compressive strength is increase at the expense of the gypsum addition (5%) on the large specific surface of mineral addition (S,l) - 2.25, and on shallow one (Ssp3)- 2.7 times. And according to the joint influence between the volume of specific surface of mineral addition and the content of gypsum addition on the large specific surface of mineral addition (S,l) - 2.5 and on shallow one (Ssp3)- 3 times. When the fixed volume of specific surface of mineral addition, the joint influence of gypsum addition and the of hardening conditions are provide, the compressive strength changes on the shallow specific surface of mineral addition (Sspl) - 3 times and on large one (Ssp3)- 2.5 times. The calculation showed that joint influence on the compressive strength of three groups of factors: the specific surface (Sspl, Ssp2,Ssp3),the of hardening conditions (& and Xs) and the content of the gypsum addition (&) are equivalent. Every group of factors is able to provide the increase of the compressive strength more than twice. The analysis of modeling results of the compressive strength and the results of physical and chemical researches allowed to formulate the suppositions about the reasons of such influence of compounding-technological factors on the compressive strength and another properties.

Mineralogical and chemical content of hydrosilicates calcium The following physical and chemical researches carried out: the electronic microscopy, the differential-thermal, X-ray and chemical analysis. There are hydrosilicates calcium in a structure: hillebrandite B, its variety: hillebrandite C and a foshagite with different morphology and different quantities [7]. The content of minerals is calculated according to the ES-models - “mixturetechnology-structure”. The content of hillebrandite B changes from 4 to 51%, hillebrandite C - from 7 to 33%, foshagite - fiom 0 to 18%. The general porosity changes from 30 to 40%, the ration of opened porosity and general one changes by 35%, the capillary water-absorption - by lo%, the ration of opened porosity and closed one - 5 times.

RpI. MPa

Fig. 2. Correlations between compressive strength and hillebrandite C2SH(B) content for Different specific surface of mineral addition: SS,,=350, S,,2=425, S,,3=500 rn’ikg.

The correlation analysis is camed out for the estimation of the connection of compressive strength with the characteristic of the structure [ 5 ] . The correlation connection

522 E l m S . SHLVKEVITCH Yevgen S. LUTXLh! Olga P.GNYI: Aleksandr A. KOICHEK Juliya V. DOTSENKO between the compressive strength and hillebrandite differs substantially for a different values of specific surface of the mineral addition: on the large specific surface of mineral addition (S,l) r = -0.28, on the middle (S,z) r = -0.26 and on shallow one (S,3) r = 0.72 /fig.2/. The maximal value of the compressive strength (E20MPa) is determined by the maximal content in the hard phase of mineral foshagite, the content in the equal parts of hillebrandite B and hillebrandite C.

Frost-resistance The change of frost-resistance is shown as the diagrams on fig. 3. Three surfaces inside a cube correspond to the grades of the frost-resistance F15,25,35. The of hardening conditions have less influence by 2.5 times on a frost-resistance, than specific surface of mineral addition.

Fig. 3. Changing of frost-resistance under the influence of different factors, of Fig. 1.

The 'oint use of mineral addition with the values of specific surface SsP1=350and SSp3=500mk g in the equal parts ( s s p ~ + s ~ 3 =s,1=sv3=0.5) l; favours the increase of frostresistance. The use of mixture with a different specific surface of mineral addition lowers the loss of the mass when freezing and thawing out from 10 to 3%. The content of gypsum addition in an optimum quantity 3% in the mixture favours also the increase of frostresistance.

1 .

Heat conductivity The change of the coefficient of heat conductivity is graphically shown as the diagrams on fig. 4. The coefficient of heat conductivity can change 3.0 times depending on the specific surface of mineral addition, the content of gypsum addition and the of hardening conditions: from 0.43 to 1.3 Wt/mK. The mineral addition with specific surface Ssp1=350 m2kg is preferable for the

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523

A

Fig. 4. Changing of maximum and minimum values of coefficient of heat conductivity under the influence of harening conditions and content of gypsum addition (isosurfaces are inside the cube); hardening conditions (isolines on the square diagrams); specific surface of mineral addition (isolines on the triangular diagrams).

providing of minimum value of heat conductivity coefficient for the compositions without gypsum, and with a gypsum one - with S,,~=500m2/kg/fig. 4.1. The values of the specific surface of mineral addition, which provide the minimum of heat conductivity, don't coincide with the values of the specific surface, which provide the maximum of compressive strength and frost-resistance. The bulk density 1450-1650 kg/m3 of the materials which content mineral addition is 15-20% lower than the bulk density of silicate materials without any mineral additions. The walling materials with such values of heat conductivity coefficient and the density are classified as heat effective. CONCLUSIONS The analysis of ES-models characteristics of the structures and the properties showed that the mineral additions with the set specific surface are the modifiers of the structure, allow to regulate the kinetics of formation of hydrosilicates calcium and the properties of silicate materials [8]. According to ES-models, the properties of materials can be regulated within wide limits: the compressive strength from 4.9 till 22.5MPa, the coefficient of heat conductivity from 0.46 till 0.99 Wt/m.K, the frost-resistance from 15 till 35 cycles. The analysis of the connection of "composition-technology-structure-property" allows to manage the properties of silicate materials.

524 E l m S. m v I T C H , Yevgen S. LUlXKlh! Olga P. G m Aleksandr A. KOICHW Juliya V DOTSENKO

Fig.5. Optimization of heat effective silicate materials. Using the results of ES-modeling the rational compositions and the conditions of hardening are recommended for the receipt of articles of a different functional purpose /fig.5/: - the stones of the walling facial classes B 10 and B 15 with the frost-resistance F35, the coefficient of heat conductivity less than 0.7 Wt/mK, the coefficient of softening more than 0.85 and the density1550-1650 kg/m3; - the stones of walling ordinary classes B5 and B7.5 with the frost-resistance F15 and F25, the factor of heat conductivity less than 0.56 WT/mK and the density 1450-1550 kg/m3; - the hollow blocks of walling ordinary classes B7.5 with the frost-resistance F>15, the factor of heat conductivity less than 0.56 WT/m.K and the density 1100-1200 kg/m3. The use of modem technological equipment and the methods of mathematical modeling, including the elements of computer Material-Science, during the analysis of the results of researches, allowed to realize the idea of the receipt and the regulation of the properties of the silicate materials after non-autoclave hardening. According to the analysis of ES-models of the structure and the properties kinetic model (physical) was offered, which in a differential form describes the basic stages of structuroformation and the change of their duration depending on the composition of limesilica binding substance, pH environment and the values of the specific surface of the components. REFERENCES

1. Mtschedlov-Petrosyan, O.P., Babuschkin, V.I., The thermodynamics of the hardening process of cement. 4th Int. Symp. Chem. Cem., Washington 1960, v. 1, pp 533-544. 2. Barabash, I.V., Solomatov, V.I., Shinkevich, E.S., Paster, N., Dyrikova, S. The mode of production of concrete mixture. SU 1761731Al C04B40/00 // 15.05.1992. 3. Shinkevich, E., Sidorova, N., Lutskin, E., Sidorov V., Politkin, S. 2004. Raw Mix for Obtain Modified Silicate Materials and Method of Its Prepare. Declared patent # 64603 A, 7 C04B28/20, Ukraine.

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4. Shinkevich, E., Lyashenko, T., Barabash, I., Sherbina, S., Voznesensky, V., Experimentalstatistical modeling the effect of multi-fractional filer on reological indices of compositions. In: Proc. of 5* European Rheology Conference, Ljubljana 1996, pp104-105. 5. Lyashenko, T., Voznesensky, V., Ivanov, Y., Modeling the influence of mix proportions on correlation between destruction pace and thixotropy of suspensions. Book of Abstracts: 3d Int. Meeting of the Hellenic Society of Rheology, Patras 2001, pp 76. 6. Shinkevich, E., Lutskin, E., Tchesskii, Yu., Bondarenko, G., Researches and mathematic modelling properties of cellular silicate compositions. In: Proc. of the 2ndInt. Symp. NonTraditional Cement & Concrete, Brno 14-16 June 2005, pp 148-153. 7. Shinkevich, E., Lutskin, E., Sidorova, N., Vinogradskiy, V., Politkin, S., Modified nonautoclave hardened siliceous materials. Structure, mixture, properties. In: Proc. of the 2ndInt. Symp. Non-Traditional Cement & Concrete, Brno 14-16 June 2005, pp 141-147. 8. Shinkevich, E. Modelling of structure and properties of modified silicate compositions. Proc. of the 42ndInt. Seminar on Modelling and Optimization of Composites, Odessa, 24-25 April 2003,24 p.

Proc. Int. Symp. “Brittle Matrix Composites 8” A.M. Brandt, ?!C. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

R-CURVES FROM EQUIVALENT ELASTIC CRACK APPROACH: EFFECT OF STRUCTURAL GEOMETRY ON FRACTURE BEHAVIOUR Viclav VESELY, ZbynEk KERSNER Institute of Structural Mechanics Faculty of Civil Engineering, Brno University of Technology Vevefi 33 1195,602 00 Bmo, Czech Republic e-mail: vesely.vl @fce.vutbr.cz, [email protected]

ABSTRACT The paper deals with the effect of structural geometry on the description of fracture behaviour of quasi-brittle materials by equivalent elastic crack approach. Particularly it is focused on the geometry effect on an R-curve shape. The effect is investigated in the paper using characteristics of stress constraint at the crack tip, for which the theory of two-parameter fracture mechanics is employed. The paper is conceived as a study of the dependence of an R-curve on stress constraint described here by biaxiality factor B. The methods by Baiant & Kazemi and Ouyang, Mobasher & Shah for determination of size independent R-curve is taken into account in the study. For both methods, R-surfaces are constructed and determination of “true” R-curve appropriate to structural size, geometry and material from the R-surface is explained. Finally, the structural analysis from determined R-curves is performed.

Keywords Equivalent elastic crack, R-curve, constraint effect, geometry, structural analysis. INTRODUCTION The fracture behaviour of quasi-brittle materials can be well described by an R-curve. However, several drawbacks are linked to the R-curve approach. The major handicap is the dependence of R-curves, except for material parameters, on the size of the structure as well as on its geometry (the shape of the structure and the loading setup) [3, 5, 111. The former difficulty has been eliminated by techniques based on size effect that determine the R - c w e as an envelope of G-curves for a series of structures of a certain category with different sizes [2, 101. Nevertheless, the later has not been overcome in a satisfactory manner yet. Partial solution has been brought recently by resistance surface concept proposed by the authors, which is a synthesis of the R-curve approach (based on equivalent elastic crack) and twoparametric linear elastic fracture mechanics (LEFM) [ 12, 131. Authors apply tools of two-parametric LEFM to solution of the problem of a dependence of R-curve on structural/specimen geometry. The R-surface concept extends the twodimensional space of R-curve, known as the dependence of the crack growth resistance R on the length of equivalent elastic crack extension da, to a three-dimensional one. For the extension of the R-curve into the R-surface, the parameter of constraint of stress (deformation) at the equivalent elastic crack tip is used (T-stress or equivalently biaxiality factor B). The Rsurface is created as a 3D plot of a set of R-curves constructed for structures of certain category

528

Vblav VESELf, Zbynek KEMNER

Proposed approach can be exploited in a structural analysis. Prediction of fracture behaviour of particular specimen or structure is performed by using so-called “true” R-curve that exactly corresponds to the material parameters, size and geometry of the structure. This R-curve is extracted from the R-surface as a section determined by constraint conditions related to investigated case. This kind of a section can be then considered as the “true” Rcurve appropriate to the structure of given material, size, shape and loading setup. So far, the R-surface concept was formulated by the authors for R-curves determined in two different ways [12, 131: For R-curves constructed by means of transformation of loaddeflection diagram into R-curve by using unloading compliance method [e.g. 3, 51 and for Rcurves from analytical method based on Baiant’s size effect law - Baiant-Kazemi method [2, 11. Only the later has been proven to produce R-curve, and consequently R-surface which is structural size independent. In proposed paper, the authors investigate another sizeindependent formulation of R-surface which employs the Ouyang-Mobasher-Shah method [10, 111 for determination of R-curve, and compare it to R-surface constructed by using Baiant-Kazemi method. In the paper, the utilization of R-surface concept is shown on consecutive numerical example in which structural analyses are performed for a set of three-point bending beams with different constraint conditions. For each particular notched beam, the “true” R-curves from the R-surface and subsequently the load-deflection diagrams are constructed. All results presented in the paper were gained via specialised calculation procedures prepared within a commercial mathematical package. CONSTRAINT EFFECT Two-parameter fracture mechanics comes out from Williams series [14] which approximates the stress and displacement fields near the crack tip. Two-parameter LEFM takes into account the first two terms of the series instead of one term in the classical fracture mechanics approach. The first term of the Williams series is related to K-factor whereas the second one corresponds to the T-stress [7]. The next terms tend to zero with increasing distance from the crack tip, and therefore they are neglected. The stress tensor q,in a certain point near the crack tip can be expressed as

where r and Bare the polar coordinates of the point andf;i(9 is known function of polar angle B(the origin of coordinates is assumed to be at the crack tip) and & is Cronecker delta. Two-parameter fracture mechanics is successfully applied to the field of elastic and elastic-plastic fracture, e.g. [8]. More detailed description of crack tip stress field can explain causes of the fact, that structures with the same values of K-factor can exhibit considerably different fracture behaviour. Significance of such a phenomenon, called constraint effect, depends on the difference of stress multiaxiality of compared structures. Stress multiaxiality arisen at the discontinuity tip due to loading is induced and affected by the structural geometry. The constraint of stress (deformation) near the crack tip caused by the stress multiaxiality (affected by structural geometry) has essential influence on fracture behaviour of investigated structure. From above mentioned facts, it is obvious, that there is a relation between the constraint at the crack tip and the structural geometry. Since the constraint effect can be characterised by values of T-stress or equivalently biaxiality factor B, it is suitable to use these quantities also

R-curvesfrom equivalent elastic crack approach: effect of structural geometry on fracture behavior

529

as characteristics of geometry for capturing the effect of structural geometry on its fracture behaviour. This feature is employed in presented paper. The non-dimensional biaxiality factor B used in the paper is defined as [9]

R-CURVES FROM SIZE EFFECT

The methods for R-curve determination exploiting size effect are considered to be the only ones which provide the R-curve independent on structural size. According to these techniques, the R-curve is constructed as an envelope of G-curves for a series of structures of a certain category with different sizes. 0

In the Baiant-Kazemi (BK) method [2, 31 the size independent R-curve is created by the G-curves for geometrically similar structures (one geometrical configuration with constant ratio of crack length to structural depth, i.e. = uo/W= const., different sizes W). On the other hand, the Ouyang-Mobasher-Shah (OMS) method exploits G-curves of structures with the same initial crack length (one geometrical configuration with constant crack length UO, different sizes

w>.

Both methods are analytical and enable to construct the R-curve from the knowledge of two basic material fracture characteristics of equivalent elastic crack models: the fracture energy Gfand the effective crack extension cfof Baiant's size effect model [2, 31 or equivalently the effective fracture toughness KI," and the critical crack tip opening CTOD, of Jenq-Shah twoparameter model [4, 111. The function of geometry of investigated configuration is another input to both analytical methods. The calculation of points of the R-curves is not described in the paper. Detailed information can be found in original papers proposing the methods, i.e. [2] for BK method and [lo] for OMS method, or in monographs [3, 5, 111. For given aforementioned inputs of the methods one can observe the change of R-curve shape with the change of geometrical characteristics of the structure. If an appropriate geometrical characteristic is chosen as a parameter p , the parametric study of the R-curve shape can be performed. The formulation of the BK method assigns the relative initial crack length a to be the parameterp. For the OMS method, the initial crack length ao, eventually in proportion to c~ as parameterp is suitable. Numerical example For illustration, the parametric study for testing configuration of three-point bending (TPB) notched beam for both methods is shown in Fig. 1. Since the axes of the graphs are nondimensional (standardised by the fracture characteristics Gf and cf)the relevant shape of Rcurve is between the initial point [O,O] and the starting point of R-curve plateau [ l,l].

The value of parameter of BK method varies from 0.02 to 0.9, which corresponds to practical limits of correctness of geometry function (see Fig. la). Particularly for shorter cracks (approximately to = 0.3) the geometry effect is rather strong. The parameter uo/cfof OMS method varies from a value close to 0 to a value tending to co. The former limit of the parameter generates the linear increasing part of R-curve below

530

Vdclav VESELt Zbynek KEMNER

the plateau; the later in fact shifts the start of the R-curve plateau to the beginning of the R-curve.

1.2

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Figure 1: Comparison of geometry effect of initial crack length on the shape of R-curve for TPB geometry: a) BK method, b) OMS method R-SURFACES

Both methods for determination of size independent R-curve are based on equivalent elastic crack approach. According to this approach, the assessment of stability of a crack and afracture process zone ahead of its tip in a real quasi-brittle body is transformed to the assessment of stability of an effective sharp crack in a elastic (brittle) body, whereas the length of the effective crack is assumed longer than the real one to ensure the equivalence in fracture behaviour of the real and the effective structure. As such, this approach enables to evaluate the parameters of the stress field near the equivalent elastic crack tip. The stress and displacement fields can be expressed in the form of Williams series, as it was introduced above. Classical equivalent elastic crack models take into account only the first term of the series which corresponds to the stress intensity factor KI, equivalently the crack driving force G, and neglect the other terms. The R-surface concept, in accordance with twoparameter LEFM, employs the second one as well. The reason for expanding of the quasibrittle fracture mechanics to next parameters consists in the necessity of capturing and/or eliminating the effect of geometry of the structurehpecimen on its fracture behaviour. The Tstress (or equivalently the biaxiality factor B) which is appropriate to the second member of the Williams series can serve as the measure of constraint and consequently the characteristics of structural geometry. The R-surface concept extends the two-dimensional space of R-curve to a threedimensional one [12, 131. The characteristic of constraint at the equivalent elastic crack tip is used as the third dimension. It is usually the biaxiality factor B for the advantage of nondimensionality. The R-surface is created from R-curves determined for reasonable range of the parameter p according to distinct size independent method. To each point of the R-curve a value of biaxiality factor B is assigned from the function B ( a ) according to actual relative crack length a.The function B ( a ) for typical testing configurations can be found in literature, e.g. [7, 91, or can be constructed based on FE analysis of stress or displacement fields near the crack tip [ 6 ] .

R-curvesfrom equivalent elastic crack approach: effect of structural geometry on fvacture behavior

53 1

Numerical example The R-surface (according to both methods) is determined for the three-point bending notched beam of following dimensions: depth W = 80 mm, breadth B = 80 mm, length L = 480 mm, span S = 400 mm, crack length a varies within the range of validity of geometry function. In Fig. 2 and 3, the construction of R-surface from appropriate set R-curves displayed in 3D space is illustrated for method BK and OMS, respectively. In graph d) of both figures, the 3D view of the R-curves constructing the R-surface is shown. Three individual projections of the R-surface are added for clarity into graphs a), b), and c). In this example, the R-surface is created in space of axes R and Au standardized by Gf and cfi respectively. Following dissimilarities of both methods in this example should be emphasized: The R-surface according to BK method is created by extending the R-curves of the entire range of parameter p , as it was used in the previous example (see Fig. la), into the 3D space. The parameter p of R-curves participating on the construction of R-surface for given example vary only within p = Adcf = 31.25 (for lower limit of validity of geometry function = 0.02) andp = 0.694 (for = 0.9 which is close to its upper limit 0.92). Approximations (wire models) of the R-surfaces constructed from R-curves according to both considered methods are displayed in Fig. 4. The shape of constructed R-surfaces indicates a considerable geometry effect. For BK formulation particularly the part of the surface formed by short crack R-curves (up to a = 0.3) is worthy of attention. The OMS method shows rather strong geometry effect on the entire R-surface. GEOMETRY APPROPRIATE R-CURVES The R-surface may signalize the geometry effect on fracture behaviour of quasi-brittle material. However, it is not directly applied for particular structural analysis, but serves as a space from which the “true” R-curve is extracted according to actual size and geometry conditions (i.e. constraint conditions) for analysed case. Under the term “true” R-curve, the Rcurve is assumed which exactly corresponds to the material parameters (introduced to the Rsurface calculation by material characteristics Gf and q),size (introduced by the dimensions of the structure, e.g. W) and geometry of the structure (introduced by the function of geometry and the initial crack length a0 or its relative value a). This “true” R-curve is extracted from the R-surface (which already involves the material and size parameters) as an intersection of the R-surface and function of B (which brings into the concept the effect of structural geometry). The construction of the ‘‘true” R-curve is explained in following example.

Numerical example The “true” R-curves of the structure described in previous example with relative initial crack length a = 0.1, 0.3, 0.5 and 0.7 made of concrete with fracture parameters Gf= 61 Jm-* and cf= 0.05 m are determined by using both considered methods. For the construction of the “true” R-curves, the R-surfaces from the previous example are used. After the transformation of the surfaces into the absolute coordinates (by employing values of Gf and cf), they are crossed by a cylindrical surfaces ruled by the function B ( a = a/W) appropriate to given size W ( W = l.6cf) and given relative notch length a.The situation is illustrated in Fig. 5.

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R-curves from equivalent elastic crack approach: effect of structural geometry on fracture behavior

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Figure 5: Extraction of the “true” R-curve appropriateto the material, size and geometry of given structure from corresponding R-surface: a) BK method, b) OMS method size W (indexed as W = 1.6q in the graphs), and ii) with the part after the critical point appropriate to infinitely large structure (denoted as inf. W).All R-curves end at a certain point emphasized in the graphs by the empty diamond sign which indicates the attainment of the upper limit of geometry function a = 0.92, i.e. the ligament of size of 8% of W remains uncracked at this point. To all graphs, the R-curve plateau for infinitely large structure is appended (starting at point [0.05,61]). STRUCTURAL ANALYSIS FROM R-CURVES The application of R-curve concept in the fracture mechanics of quasi-brittle materials, with aforementioned limitations, is possible in material quality assessment, where the R-curve is considered as a material property. The other significant field of use of the concept is the prediction of structural failure. The R-curve itself straightforwardly describes the aspects of the fracture process associated with the crack propagation under loading, i.e. the behaviour of the crack. Nevertheless, also the behaviour of the entire structure can be predicted from the Rcurve. Most common is the transformation the R-curve into the load-displacement curve performed by employing LEFM. Numerical example The load-deflection diagrams of TPB specimens of varying notch lengths are calculated from R-curves determined above. In graphs b), d), f) and h) in Fig. 6 the P-d curves are drawn corresponding to R-curves from graphs a), c), e) and g), respectively. CONCLUDING REMARKS In the presented paper, the comparison of Baiant-Kazemi method and Ouyang-MobasherShah method, both for size independent R-curve determination, was introduced. The application of the technique capturing the effect of structural geometry on the fracture behaviour by means of parameter of constraint at the equivalent elastic crack tip was shown. This approach is referred to as the R-surface concept. The prediction of load-displacement diagrams from R-curves was performed.

535

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Figure 6: Comparison of R-curves determined by considered methods (left column) and corresponding load-deflection diagrams constructed from the R-curves (right column) Results of presented numerical example indicate considerable differences between both considered methods; whether in the R-curves and consequently the R-surface shapes or in the predicted load-displacement diagrams. Moreover, for the Baiant-Kazemi method there exists

536

VPclav VESELt Zbynek KERSNER

a discrepancy between the R-curve calculated according the method and the “true” R-curve corresponding to the real progress of the fracture. The significance of this discrepancy depends on the shape of the R-surface, where the particular hcture process proceeds. Unfortunately, the authors have accomplished no comparison of calculated result with experimental tests yet. Therefore, presented research may be regarded as a case study so far. Of course, it is intended to replenish this study with experimental data. ACKNOWLEDGEMENTS This outcome has been achieved with the financial support of the Ministry of Education, Youth and Sports, project No. lM680470001, within activities of the CIDEAS research centre. REFERENCES 1. Baiant, Z. P., Gettu, R. & Kazemi, M. T. (1991) Identification of nonlinear fracture properties from size effect tests and structural analysis based on geometry-dependent Rcurves. International Journal of Rock Mechanics, Mining Science & Geomechanics Abstracts, Vol. 28., NO. 1,43-51 2. Baiant, Z. P. & Kazemi, M. T. (1990) Determination of fracture energy, process zone length and brittleness number from size effect, with application to rock and concrete. International Journal of Fracture, 44,111-13 1 3. Baiant, Z. P. & Planas, J. (1998) Fracture and size effect in concrete and other quasibrittle materials. Boca Raton: CRC Press 4. Jenq, Y. S., Shah, S. P. (1985) A two-parameter fracture model for concrete. Journal of Engineering Mechanics, 111,1227-1241 5. Karihaloo, B. L. (1995) Fracture mechanics and structural concrete. New York Longman Scientific &Technical 6. KnBsl, Z. (1995) Evaluation of the elastic T-stress using a hybrid finite element approach. International Journal of Fracture, 70,9-14 7. KnBsl, Z. & Bednu, K. (1998) Two-parameter characterization in fracture mechanics. Engineering Mechanics 5(3), 133-142 8. KnBsl, Z., Bednu, K. & Radon, J. (2000) Influence of T-stress on the rate of propagation of fatigue cracks. Physical Mesomechanics 3(5), 5-9 9. Leevers, P. S. & Radon, J. S. (1982) Inherent stress biaxiality in various fracture specimen geometries. International Journal of Fracture, Vol. 19,311-325 lO.Ouyang, C., Mobasher, B. & Shah, S. P. (1990) An R-curve approach for fracture of quasi-brittle materials. Engineering Fracture Mechanics, Vol. 37, No. 4,901-916 11. Shah, S. P., Swartz, S. E. & Ouyang, C. (1995) Fracture mechanics of structural concrete: aplications of fracture mechanics to concrete, rock, and other quasi-brittle materials. New York: John Wiley & Sons, Inc. 12.Vesely, V. & KerSner, Z. (2004) Resistance surface concept for concrete fracture. In proceedings of FRAMCOSJ Conference, Vail Colorado, 407414 13.Vesel9, V. (2004) Parameters of concrete for description of fracture behaviour. PhD Thesis, FCE BUT Brno, Brno 14.Williams, M. L. (1957) On the stress distribution at the base of stationary crack. ASME Journal of Applied Mechanics, 24, 109-1 14

Proc. Int. Symp. "Brittle Matrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

FRACTURE TOUGHNESS AT SHEAR (MODEII) OF CONCRETES MADE OF NATURAL AND BROKEN AGGREGATES Grzegorz GOLEWSKI', Tomasz SADOWSK12 Faculty of Civil and Sanitary Engineering Technical University of Lublin Nadbystrzycka 40,20-618 Lublin, Poland, e-mail: [email protected], [email protected]

'

ABSTRACT Hardened concrete is a brittle material characterised by non - homogeneous internal structure. In such composites destruction under the influence of external loads is caused by: discontinuities created by technological defects or local differences of mechanical properties of concrete components. In the vicinity of discontinuities local concentration of stresses occur. They can cause sudden defects propagation which lead to degradation of the whole element in the brittle ffacture process. In the present work the basic type of microcracks, which develop in concrete composites under shear, has been discussed cracks which have appeared in the aggregate- mortar contact area and, defects occurring in cement matrix. The paper presents discussion of basic properties of mineral aggregates (natural and broken) with specification of properties affecting their brittle damage. In order to describe cracks development at shear experiments were carried out for concrete made of two types of aggregates: natural gravel aggregates (G) and broken limestone aggregates (L). To determine the influence of grain-size distribution of coarse aggregate for each concrete series, two types of optimal proportions of different - sized aggregates were used with a maximum grain size up to 8 mm (serie I) and up to 16 mm (serie 11). The tests regarding estimation of macroscopic crack resistance were curried out for I1 fracture mode (at shear). During experiments the critical values of stress intensity factors Kffc and the fracture work J1fc were determined. For each sample load - displacement curve Pcf) at loading point and acoustic emission signals (EA) were recorded. It was determined that the values KII,and J,, were higher for the concrete L. In order to evaluate the differences of the internal structure of the investigated materials, the microscopic observation of the fracture surface was carried out. The microstructural observations were performed using the scanning microscope LEO 1430VP applying two electron detectors: SE (secondary electrons) and BSE (backscattered electrons).

Keywords Fracture toughness, crack, gravel and limstone aggregate, shear, microscopic observation, acoustic emission INTRODUCTION Due to its internal structure concrete is a composite in which cement paste (a continuous phase), is treated as a brittle matrix, whereas aggregate and the remaining concrete ingredients (a scattered phase), are inclusions inside this matrix. Inclusions usually create a reinforcing structure and affect positively its strength and durability, but they are also used for other purposes e.g. in order to improve its thermal insulation or to decrease composite weight [13.

538

Gaegorz GOLEWSfl, Tomasz SAD0 WSKI

Hardened concrete is a brittle material characterized by non - homogenous structure. Concrete brittleness is of decisive importance as regards material strength and toughness in the places where its structure has defects. Close to concrete discontinuities such as notches, air voids, cracks, internal material corrosion etc. local stress concentrations can occur which, under external load, cause sudden propagation of the damage. Dynamic damage growth can result in the destruction of the whole element. Such a phenomenon is called brittle fracture. Usually it is a dynamic process characterized by catastrophic results, it is irreversible and it occurs without previous plastic strain. In order to prevent such a sudden destruction of the construction it is necessary to have deep knowledge about damage development processes and fixture in construction materials. The knowledge of these processes allows to affect the improvement of concrete quality of which constructions are made, to evaluate their defects and to define the reasons for their occurrence. Damage mechanics and fhcture mechanics constitute a tool to understand the process of material degradation and their practical use can add to obtaining the composites characterized by the highest quality, durability and working reliability possible [2].

CHARAKTER OF DAMAGE OF CONCRETE COMPOSITES

Concrete is a multi-phase and non-homogenous material. Therefore there can be many reasons for defects and many places for their development. First damages usually occur in the concrete matrix or in the aggregate - mortar interfacial transition zone. They can occur due to e.g. nonhydrated cement grains, pores occurring in all composite phases or the amount and type of the aggregate used. According to J. Peng and others [3] the basic types of micro-cracks which occur in deformed concrete for the case of two-dimensional analysis are: a) cracks which occur in the aggregate-mortar contact area: stable cracks (characterized by constant length) which do not develop towards the central part of the cement matrix, cracks which tend to develop towards the central part of the cement matrix, cracks developing accross aggregate grains b) cracks which occur in the cement matrix itself. While describing the structure of microcracks in deformed concrete it also important to determine: the defects locations, types of crack, direction of cracks and number of defectss. In the case of the two dimensional analysis one can differentiate straight line type cracks or wingtype cracks when the direction of straight line type cracks [4-81 is changed. The occurrence of combined types of cracks in concrete resulted in using various theoretical models of micro-cracks to create a model of behaviour of this type of materials. It results in the necessity of carrying out experimental research based on the I and II fracture modes. The analysis according to the II fiacture mode (shearing) is of particular importance - since it makes possible to introduce the type of a closed crack whose edges slide against each other perpendicularly towards to the crack front and parallel to the crack plane. As regards concrete this model is of particular significance due to its low shearing resistance and high sensitivity to this type of stress. The way of loading the element according o the 11 fhcture mode was presented among others in the following work [9-lo].

539

Fracture toughness at shear (Mode Il) of concretes made of natural and broken aggregates

PROPERTIES OF MINERAL AGGREGATES As coarse aggregate for concretes, mineral aggregates are usually applied. They can be divided into natural and crushed ones. Natural aggregates are formed as a result of breaking rocks caused by natural forces, whereas crushed aggregates are formed as a result of crushing rocks due to intended human actions. The aggregate quality is determined by the conditions under which they were formed and mineral composition of rocks from which they originate. Moreover, their physical characteristics such as their: size, shape and type of grain surface are also of significance [ 111. These properties are decisive as regards its adhesion and the composition of aggregate-mortar interfacial transition zone and have fundamental influence on brittle fracture processes in concrete [ 12-14]. The grains shape of natural aggregates is usually spherical, oval, rounded or irregular, whereas that of crushed ones: elongated, stocky, flat, plate - shaped etc. The texture of natural aggregates grains is smooth and that of broken ones rough or crystalline [111.

AIM AND SCOPE RESEARCH

Component of concrete mix portland cement CEM I class 42.5 R from Oiarow Cement Plant pit sand, fraction 0-2 from Markwzow n. Lublin gravel, fraction 2-8 (G1 and G2) from the deposits in Sobolewo n. Suwalki gravel, fraction 8-16 (G2) limestone grit, fraction 2-8 (L1 and L2) from the deposits in T r ~ ~ k a w i n. c aKielce limestone grit, fraction 8-16 mm (L2) water from municipal water supply system superplasticizer Arpoment P (1.5% of cement mass)

Quantity [ks/m3] 352 676 1207 (GI), 482 (G2) 725 (G2)

141 5,28

In the present work the results of empirical research which describe the damage development of gravel concretes (G) and limestone concretes (L) were presented. The research on determining macroscopic fracture mechanics was carried out according to the I1 fracture mode (shearing). During the experiments carried out basic parameters of fracture mechanics were defined i.e. critical stress intensity factors K,lC and unit works of failure JII,. Four types of construction concrete made of gravel aggregate and limestone aggregate underwent testing. In order to determine the influence of coarse aggregate grain size distribution, for each type of concrete two types of optimal proportions of different - sized aggregates were used, with the maximum, grain diameter up to 8 and up to 16 mm. The compositions of concrete mixes were listed in Table 1. All concrete mixes were made up to the consistence of V2 and w/c ratio = 0.40. The times measured using the Vebe apparatus according to PN-EN 206-1 [15] and sand points for individual mixes were contained in Table 2.

540

Grzegorz GOLEWSKI. Tomasz SAD0 WSKl

An attempt was made to select the optimum proportion of different - sized aggregates in such a way as to be contained in the most advantageous area between limiting grain size distribution curves. The recommendations used were based on the German standard DIN 4226-1 [16]. An example of the grain size distribution curve for L1 concrete was presented in Fig. 1. After being formed the samples were thickened by approximately 30 s on the vibrating table. From each mix samples were made for auxiliary and basic research. In the tests of concretes properties the following were used: 12 cubes (150 mm edges) to assess compressive strength and tensile strength by splitting (6 for each type of tests), 9 cylinders - 150 mm diameter and 300 mm high to perform compressive strength tests (3 cylinders) and to determine the coefficient of elasticity for concrete (6 cylinders). For principal experiments 8 cubes 150 mm edge were formed for the purpose of determination macroscopic hcture toughness at shear.

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Fig. 1. An example of a grain-size distribution curve for concrete - batch L1. All samples were removed after 2 days from being embedded in concrete and then they were kept for 14 days under extremely wet conditions and for another 14 days under laboratory conditions. After 28 days of curing the tests of concrete properties and basic tests were carried out. Concrete strength characteristics were defined using a hydraulic press type ZDlOO. The values obtained were presented in Table 2. Table 2. Characteristics of concrete mix and concrete.

Fracture toughness at shear (Mode II) of concretes made of natural and broken aggregates

54 1

TESTS OF MACROSCOPIC FRACTURE TOUGHNESS AT SHEAR

In the tests to evaluate concrete fracture toughness, a method of loading samples according to the I1 fracture mode was used. For the experiments cubes with two initial cracks were used. The target crack sizes were obtained by embedding in concrete the cubes being formed two 4 mm steel sharpened flat bars. The sizes of the samples used and the way of loading them were shown in Fig. 2. The size of the initial crack was selected in such a way as to hlfil the condition 0,3 < a/W I0,5 [17], where a represents the depth of the initial crack and W the total height of the sample.

Fig. 2. Sampleused for fracture toughness using the II fracture mode. The tests presented were carried out using a strength machine ZWICK ZlOO with a computerised results recording system. During the experiments,at the same time for each sample the dependencies load - deflection of the applied force Pdf) and acoustic emission signals (AE). were recorded. AE measurements were carried out in an automatic way using a one-channel acoustic emission analyser EA-IFTR equipped with a memory card EA-100 wit a computerised registration of the results. Based on the dependency P f l destruction curves were obtained for each sample. Using these graphs and AE distribution the value of a critical force PQ was determined which caused the developmentof the initial crack. An example of a destruction curve was shown in Fig. 3.

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542

Grzegorz GOLEWSkX Tomasz SAD0 WSKl

Critical stress intensity factors KLI,were determined according to the dependence determined by J. Watkins (1) [18].

where: PQ is the value of critical force which begins of the initial crack development and is identified on destruction graphs as a small inflection of a curve of its extreme (Fig. 3), b - sample height above the initial crack, B - sample thickness, a - length of the initial crack. Load - deflection curves P@obtained in the research were used to define the unit work of failure JII~,. These calculations were performed according to the dependency contained in the American standard ASTM E 820-01 (2) [ 191. J,

=

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where: A - energy absorbed in the sample up to the moment of the initial crack growth. This energy can be calculated by finding the area (i. e. by taking the integral) underneath the curve Pfl to the critical force point PQ;B and b are the same as in formula (1). RESEARCH RESULTS AND THEIR ANALYSIS

During the tests carried out various fracture mechanisms were observed. The L1 and G1 samples usually failured at the moment of the initial crack growth (Fig. 4a). At that moment a sudden force decrease and a dynamic initial crack propagation was observed. The character of destruction of these samples had features of quasi-plastic deformations. In a few samples of L2 and G2 a sudden total destruction was observed (Fig. 4b) i.e. at a significant value of critical force PQa complete shearing of samples took place along the initial crack plane. In one of the examined cubes one could see the existence of two twin branched cracks which were initiated at the top of the initial crack.

Fig. 4 The ways of sample failure were observed in the following tests: a) extension of initial crack; b) sudden total destruction.

In Table 3 the calculated values of critical stress intensity factors K1lCand unit unit works of failure values were presented.

Fracture toughness at shear (Mode II) of concretes made of natural and broken aggregates

543

Table 3. Average values obtained in testing fracture mechanics parameters

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Type of concrete L1 G1 L2 G2

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Based on the results obtained one can say that the type and the grain - size distribution of the aggregate used have a direct influence both on strength characteristics and fracture mechanics parameters. Analysing Table 3 it was stated that the parameters tested had higher values in the case of L concretes. It is also observed that the fracture mechanics parameters increase in both gravel and limestone concretes together with the size of grains of the coarse aggregate used. MICROSTRUCTURAL TESTING OF CONCRETE In order to assess the differences in the structures of the concretes tested, a microscopic observation of sample fragments obtained from destruction zones during fracture toughness testing was carried out. Using a scanning electron microscope to assess the morphology of concrete fractures [20-211 allows to estimate the manner of material destruction and provides an answer to the question what type of microcrack arose in composite. The microstructural testing was carried out using a scanning microscope LEO 1430VP at magnification from 100 to 5000 times. In order to carry out the analysis of fracture concrete surfaces two electron detectors were used SE (secondary electrons) and BSE (backscattered electrons).

Fig. 5 An example of the microstructure of the examined concretes: 1-limestonegrains, 2microdamages in contact surfaces, 3-cavities in the place of separation gravel grains from the matrix a) view of microdamages in the contact areas between the matrix and limestone grains; b) view of cavity after removing gravel grain fiom the matrix.

544

Grzegon GOLEWSKI, Tomasz SAD0 WSKl

Fig. 5 presents characteristic pictures of the fixture surface microstructure of sample for the examined concretes. On the destruction surface of limestone concrete (L) one could see the cracks on the contact surfaces which did not cause the separation of the aggregate grains from the matrix (Fig. 5a). These damages had a local character. The length of microcracks in relation to the grain size was small. No propagation of defects inside the matrix was observed. Such a picture of fractures is an evidence of good properties of the filler used which implied the occurrence of compact and tight layers in the aggregate - mortar contact area. For gravel concretes (G) cavities after removing coarse aggregate from the mortar (Fig. 5b) were observed. Usually in such places in the cement matrix the straight line or winged type cracks were visible. FINAL REMARKS In the course of the tests carried out it was found out that concretes manufactured with the use of

limestone crushed aggregates were characterized by a higher crack resistance than the concretes containing natural uncrushed gravel filler. The highest increase of values were observed while comparing unit unit works of failure JII, with reference to G1 series of concretes to the advantage of L1 concretes - 42.2% and G2 concretes to the advantage of L2 - 53.4%. Such good results were obtained for L series of thanks to the angular shape of limestone grains and their rough texture. Additionally, a positive influence on the increase in the hcture toughness of L concretes could have been achieved by the physical and chemical composition of carbonate rocks from which limestone is obtained. These properties of aggregate grains implied its much better adherence to the mortar than the rounded gravels characterized by smooth texture. During the experiments higher values of hcture mechanics parameters were noted as the gmin sizes of the aggregate used increased. The most significant differences occurred at percentage comparison of L1 concretes to the advantage of L2 one. The increase in fracture toughness depends mainly on the aggregate - mortar contact layers and the inclusions of large limestone aggregate grains (up to 16 mm) lead to forming compact and tight structure just in these areas of the composite. In the case of gravel concretes the increase of K I I and ~ JII~ for G2 concretes in relation to G1 was insignificant and it was caused mainly by small adherence of these aggregates to the mortar which was presented in Fig. 5b. In the tests to determine the critical force the use of an acoustic analyzer of acoustic emission was helpful - it let define precisely and univocally the critical forces PQ . It can be stated that analyzing concrete properties using the II fracture mode can be used to assess the durability and safety of the working constructions subjected to shear e.g. the analysis of concrete and reinforced concrete beams in the support areas. Moreover, the use of the results presented in the work can give rise to the optimization of the following: design processes of concrete mixtures curing of concrete - process technology in such a way as to obtain the material characterised by the minimum number of initial defects which ,thanks to the increased crack resistance, affect in a positive way the structure reliability. The results obtained are also of crucial significance for the theoretical description of microcracks in concrete when their increase takes place at the mixed fracture mode and when both I and 11crack mode are considered [22]. In such a case there are changes in the direction of the microcracks development, which is caused by non - homogenous material structure and it is necessary to take into account the influence of shear.

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ACKNOWLEDGEMENTS The Authors are currently supported by the Polish Ministry of Education and Science, grant No 65/6, (PR/UE/2005/7).

REFERENCES 1. Brandt A. M., Kasperkiewicz J and others: Methods of diagnosis of concrete and high performance concretes based on the structural research. Institute of Fundamental Technological Research Polish Academy of Sciences, Warszawa 2003. (In Polish). 2. Van Mier J. G. M.: Fracture processes of concrete. Assessment of material parameters for fracture models, CRC Press, Boca Raton, New York, London, Tokyo, Florida 2000. 3. Peng J., Wu Zhimin, Zhao G.: Fractal analysis of fracture in concrete. Theoretical and Applied Fracture Mechanics, vol. 27, no. 211997, 135-140. 4. Sadowski T.: Modeling of semi - brittle MgO ceramic behaviour under compression, Mechanics of Materials, vol. 18, 1994, 1- 16. 5. Sadowski T.: Mechanical response of semi - brittle ceramics subjected to tension compression state. Part I: Theoretical modeling, International Journal of Damage Mechanics, vol. 3, 1994,213-233. 6. Sadowski T.: Mechanical response of semi - brittle ceramics subjected to tension compression state. Part 11: Description of deformation process, International Journal of Damage Mechanics, vol. 4, 1995,293-317. 7. Basista M., Gross D.: The sliding crack model of brittle deformation: An internal vaoable approach. Itemational Journal of Solids and Structures, vol. 35, no. 5,64998,487-509. 8. Yu S, Feng X.: A micromechanics-based damage model for microcrack - weakened brittle solids, Mechanics of Materials, vol. 20, no. 1/1995, 59-76. 9. Ayatollahi M. R., Aliha M. R. M.: Cracked Brazilian disc specimen subjected to mode I1 deformation. Engineering Fracture Mechanics. 2005, Vol. 72, No. 4, s. 493-503. 10. H. W. Reinhardt, J. Ozbolt, S. Xu, A. Dinku: Shear of structural concrete members. Advanced Cement Based Materials, vol. 5, no. 3-4/1997,75-85. 11. Neville A.: Properties of concrete. Fourth edition. Polish cement, Krakdw 2000. (In Polish) 12. Prokopski G., Halbiniak J.: Interfacial transition zone in cementitious materials. Cement and Concrete Research, vol. 30, no. 4/2000, 579-583. 13. Elsharief A., Cohen M. D. Olek J.: Influence of aggregate size, water cement ratio and age on the microstructure of the interfractial transition zone. Cement and Concrete Research, vol. 33, no. 11/2003, 1837-1849.

14. Zheng J. J., Li C. Q., Zhou X. Z.: Thickness of interfacial transition zone and cement content profiles around aggregates, vol. 57, no. 7/2005,397-406. 15. PN-EN 206- 1:2003 Concrete Part 1: Specifications, properties, production and conformance. (In Polish) 16. DIN 4226-1: Zuschlag fur Beton; Zuschlag mit dichtem Gefbge, Begriffe, Bezeichnung und Anforderungen. 17. Dixon J. R., Strannigan J. S.: Determination of energy release rates and stress-intensity factors by the finite element method. Journal of Strain Analysis vol. 7, 1972. 18. Watkins J.: Fracture toughness test for soil-cement samples in mode 11. International Journal of Fracture, 23, 1983, 135-138. 19. ASTM E 1820-01: Test Method for Measurement for Fracture Testing. American Society for Testing and Materials, Philadelphia, 1996.

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20. Soroushian P., Elzafraney M.: Qualitive analogy between performance loss and microstructural manifestation of different damaging effects on concrete. Magazine of Concrete Research, vol. 55, no. 512003,419-427. 21. Scrivener K. L.: Backscattered electron imaging of cementitious microstructures: understanding and quantification, vol. 26, no. 8/2004,935-945. 22. Sadowski T. (ed).: Mutiscale modelling of damage and hcture processes in composite materials. CISM courses and lectures no. 474,2005, Springer, Wien New York, 309.

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ.. Warsaw 2006

ON FRACTURE ENRGY OF CONCRETE FOR SHORT-TIME LOADING IN TENSION Janusz R. KLEPACZKO Laboratory of Physics and Mechanics of Materials Metz University Ile du Saulcy, 57045 Metz, France, e-mail: [email protected]

ABSTRACT In the first part of this paper recent experimental results performed in tension are discussed showing a very high rate sensitivity of the failure stress o-,,when a threshold strain rate is exceeded, approximately 1.0 s-'. Quantitative analyses of the rate effect on the failure stress are carried out by application of a local cumulative failure criterion [ l , 21. In the second part an analysis of different energy components during rapid material separation is outlined. It appears that not only the failure stress oFbut also the fracture energies defined in different ways are very sensitive to the local strain rate. Finally, some quantitative analyses of the failure stress and energies are performed for concrete in tension at high loading rates.

Keywords Concrete, high loading rate, failure stress, failure energy.

INTRODUCTION Experiments performed during recent years have clearly shown that the failure stress OF in tension substantially increases above the threshold strain rate -1 .O s-'. Great interest in the study of dynamic behavior of concrete observed in recent years is caused by development in LPMM-Metz a new experimental technique based on the phenomenon of spalling, [3].This method to determine the strength of concrete in tension up to -120 s-' has become more common among different research laboratories, for example [4-61. The experimental technique developed in LPMM for dynamic tension test of brittle or quasi-brittle materials is based on the concept of Hopkinson bar as a measuring device, [7]. The principle of the arrangement based on only one Hopkinson bar, which is shown in Fig. 1, are as follows: a striker bar which is accelerated by a gas launcher impacts coaxially an instrumented Hopkinson bar in tight contact with a concrete cylindrical specimen. The striker, the bar and the specimen are of the same diameter 40 mm. The lengths are respectively: 120 mm, 1000 mm and 120 mm. The striker and the measuring bar are made of the same hard aluminum alloy 6060-T5 having very close mechanical impedance in comparison to concrete. After impact of the striker on the measuring bar an incident compression wave fi ( t ) is generated, this wave propagates along the bar up to the interface badspecimen. Because a slight difference in the mechanical impedances of the measuring bar and the specimen, the incident wave is partly reflected back as tension wave OR 6) and partly transmitted into the specimen as compressive wave oT(t).This transmitted compressive wave is reflected from the free end of the specimen as a tension wave.

Janusz R.KLEPACZKO

548

..

+-

P

i 1000 mm

specimen

Measuring bar

80 mm

projectile

Fig, 1 Arrangement for dynamic tensile test by spalling, Jl,J2 and 53 are SR stations. A superposition of the transmitted compression wave into specimen with the reflected tension wave, which propagates in the opposite direction, generates a fast increasing tension stress causing fiacture at the specific distance from the specimen free end. Possibility to use different striker lengths (80 mm, 120 mm 160 mm) and to apply different impact velocities enables to achieve relatively high local loading rates of spalling. However, application of high amplitudes of the incident wave leads to a consecutive multiple spalling, for example Landon and Quiney [9]. In such case the chronology of spalling failures is very important. When a multiple spalling occurs only the first one was analyzed (only the first spalling is recorded by the measuring bar). In order to find the chronology of spalling, as well as to observe development of failure, a coupled arrangement of six fast CCD cameras has been applied [8]. It is important to note that the specimen wave history of loading can be exclusively controlled by the Hopkinson bar data. In order to determine the whole wave history the Hopkinson bar called also the measuring bar has been instrumented with three SR gage stations, Fig. 1, with adequate voltage supply units,potentiometer circuits and three wide band amplifiers. More details on measurements are given elsewhere [lo]. The experimental technique developed in LPMM-Metz is nowadays more frequently used in different laboratories, and the result is always the same: the failure stress of concrete increases substantially when the strain rate is higher than -1.0 s-I. RATE SENSITMTY OF CONCRETE FAILURE AT HIGH LOADING RATES

Although a detailed discussion of the high rate sensitivity of strength for concrete has been published earlier in Proceedings of BMC-7, [ 111, here some fundamental results are outlined anew. The LPMM experimental data were obtained up to strain rates -120 s" in tension for wet and dry mini-concrete MB 50, [lo], they are shown in Fig2 in the double logarithmic scale of the relative failure stress @IF) versus strain rate. The DIF means the Dynamic Intensification Factor, defined as the ratio of the current failure stress at specific strain rate to the quasi static failure stress at strain rate -10-~s-'. It is clear, at strain rate -10' s-' the DIF of the wet concrete reaches value -10. That means that the dynamic strength is ten times then the quasi-static one. For the dry concrete behavior is similar, at strain rate -10' s-' the DIF reaches value -8.

549

On fracture eneqg of concretefor short-time loading in tension

. 2

10

loo

log strain rate ( U s )

Fig.2 Comparison of relative tensile strength (log DIF) for wet and dry micro-concrete MB50 versus logarithm of strain rate, [lo]. The concrete that has been tested in LPMM denoted by MB 50 had the maximum aggregate size 2.0 mm, relatively high cement dose CPA HP and the ratio of the watedcement was 0.5. The technology in preparation of the concrete was optimized to assure a high level of homogeneity of all components. All doses and mechanical characteristics of the concrete tested in LPMM are given in [lo]. The stress rates determined by the numerical analyses of the records vary between 800 GPds up to 5000 GPds, corresponding to strain rates varying from -20 s-' to -120 s-'. In some figures in this paper the ratio of the dynamic strength to the quasi-static strength, which is the DIF is applied. Figure 2 clearly indicates a high rate sensitivity of the failure strength versus strain rate. The threshold strain rate can be found for log(D1F) = 0, that is DIF = 1. The DIF is lower for the dry concrete MB50 because the quasi-static strength in that case is higher than that for the wet one, respectively -5.0 MPa and 4 . 0 MPa. Another results of experiment obtained via spall technique, being very similar to that presented in Fig.2, were reported in [4]. The data have been completely reanalyzed and they are shown as log@IF) versus logarithm of strain rate in Fig.3. The results reported in [4] were obtained for a concrete with the quasi-static compression strength OF = 34.1 MPa and the quasi-static tension strength 2.69 MPa. This concrete can be classified as a mediumstrength material. The results presented in Fig.3 can be again approximated by a straight line. It can be shown that the best linear fit can be obtained with the slope m I= 0.743. The threshold strain rate is estimated as &, I= 5.17 s-'. More recent results reported in [5] have been completely reanalyzed and the outcome is presented in Fig.4. Again, the spall technique was applied. A concrete tested had an estimated compressive quasi-static strength OF = 35 MPa and the maximum aggregate size -8.0 mm. In addition to the smooth cylindrical specimens several tests were carried out with notched specimens. The diamonds in Fig.4 mean the data for smooth cylindrical specimens, the squares mean the notched specimen data. The quasi-static strength in tension was given as aF0 = 3.24 MPa.

Janusz R.KLEPACZKO

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LOG (DIF) VS. LOG STRAIN RATE

Fig.3 Logarithm of relative tensile strength (DIF) for concrete versus logarithm of strain rate, reanalyzed after [4].

LOG(DIF) VERSUS LOG. OF STRAIN RATE I 0.8 0,7

0,6 IL

E B 99

0,5 0.4

0.3

t

1 - 1 1

Fig.4 Reanalyzed results of experiments for concrete reported in [5], here are presented in the double logarithmic scale of relative tensile strength log(DIF) versus logarithm of strain rate.

On fracture energy of concrete for short-time loading in tension

55 1

A relatively narrow range of strain rate and dispersion of the data limit a more detailed quantitative analysis, but a trend is clear, the DIF increases as a function of strain rate. In Fig.4 the reanalyzed data after [5] are plotted in the double logarithmic scale: log(DIF) versus log€. Although the dispersion of experimental points is substantial a linear fit of the form log(DZF) = u + m log€ seems to be a good approximation with a = -0.1 and m = 0.346 . The rate sensitivity of the failure strength was discussed in detail in [ll]. Here only the rate sensitivity defined as m will be analyzed.

This is a commonly used definition of the so-called logarithmic rate sensitivity. Therefore the slope of the data shown in Fig.4 for log D = 0 determines the threshold strain rate for u = log& the threshold strain rate is estimated for that case as kTH= 0.79 s-I. This value is much lower than obtained by analysis of data shown in Fig.3 where kTH= 5.17 s-’. In general, the threshold strain rate varies from 1.0 s-l to10 s-I. It seems that for wet and poor quality concretes the threshold strain rate is closer to 1.O s-I. THE CUMULATIVE FAILURE CRITERION According to experimental data presented in Figs 2-4 it is convenient to interpret an increase of strength when the loading time is increased by a cumulative criterion of failure, [ 11. Many processes on the micro-, meso- and macro-levels, involving an increase of the DIF, can be analyzed as source of damage accumulation in dynamic loading, including material separation during spalling. Thus, variety of possible damage processes in different materials causes that there is no universal failure criterion applicable for semi-brittle materials like concrete and rocks. In formulations of failure criteria applicable at high loading rates or short time loading the approach based on cumulative rate-dependent damage has many advantages. One of advantages is it’s a simple implementation in computer codes, for example [3]. A possibility in formulation of time-dependent failure criterion is application of specific approach to material separation by taking into account an assistance of thermal vibration of lattice. It is assumed that the thermal vibration of atoms always assist in material separation. The concept is not new and was long time ago. It is out of this contribution to discuss details of such processes. It is assumed that in addition to the assistance of thermal vibration of atoms the micro-crack shielding contributes to additional time delay during the material separation, [ 11,121. Specific form of the failure criterion based on Markov’s processes and Yokobori’s stress-dependent activation energy has been proposed by Klepaczko in 1990, [ 1,131. It may be mentioned that the cumulative criterion based on the time accumulation of rate-dependent defects leads to the logarithmic rate sensitivity m as the meaningful constant. A general form of the criterion is given by

where OFO, t , and a(()are the material constants at constant temperature. The constant tCois the longest time of loading when the failure stress OF( t,, ) = 0.0 approaches the quasi-static

552

Janusz R.KUEPACZKO

failure stress oF0, that OF = OFO for tc = t , . The absolute temperature and the activation energy are respectively T and AGO, k in the Boltzman constant. Thus, the transition point fiom low to the high rate sensitivity is defined. The exponential a depends on the absolute temperature T and is directly related to the activation energy of the material separation AGO with k being the Boltzman constant. When a proportional loading is assumed, that is the case of the spa11 experiment with concrete, the critical stress at failure is related to the strain rate O,

=E t t ,

(3)

The criterion after integration can be written in a simplified and explicit form T = const

(4)

where E is the Young's modulus of the material and t is the local strain rate. Assuming that UFO is the quasi static failure stress the dynamic intensification factor DIF,denoted fiuther by D, can be introduced into Eq.(4), then

[

(;o)]k&A

D = (i+a)t,, -

(5)

In the limit for D = 1 the strain rate threshold tmcan be determined, then

In order to analyze the experimental data in the double logarithmic scale, log D = f(logb), Eq.(5) is transformed into logarithmic form, then on obtains (7) Of course, the first term in Eq.(7) is a material constant and the relation is reduced to the simplified form logD = a + mlogt where m=- 1

a+1

Therefore the logarithmic rate sensitivity m is directly related to the cumulative model of failure (2). The logarithmic rate sensitivity can be explicitly expressed by the physical quantities, then

On fracture energy of concretefor short-time loading in tension

1

m=kT

or

m=

+1

553

kT AGO+ kT

Since concrete is a composite all quantities discussed within the fkamework of the cumulative failure criterion should be understood as mean quantities. The logarithmic rate sensitivity is therefore related to the ratio of energies, kT is the mean energy of the atom vibrations and AGO is the mean energy of material separation. The material constants obtained after experiments of spalling for MB 50 concrete are cited after [ 141 in Table 1. Those values were determined by the method of the linear regression. Material constants tCO [PSI 4300) AGO[eV]

Wet MB 50 52.3 0.912 0.0237

Dry MB 50 88.0 1.267 0.0328

When the values of exponent a for the dry and wet concrete are introduced into Eq.(5) the result is that the dynamic strength increases with strain rate to the power close to % (more exactly 0.441 for the dry and 0.523 for the wet concrete. Fig.5 shows comparison of experimental results and prediction by the cumulative criterion in the form of Eq.(7). Predictions for the both wet and dry concrete fit very well to the experimental data.

I 10

wetmicmamrete dry micro-concrete uiterion wet micro-ooncrete

- - criteriondrymicrc-concrete

I

*

9 81 ~

7l 6c

5l i

~

4,

i1

31

I

I 2' 1

I

I

I

10

100

strain rate (log, 1Is)

Fig.5 Data for wet and dry MB 50 micro-concrete compared with prediction by the cumulative criterion, after [141.

554

Janusz R.KLEPACZKO

Although a more detailed quantitative analysis has been performed for the concrete MB 50 tested in LPMM it is clear that the cumulative criterion can be applied for other reanalyzed experimental data shown in this study, Fig.3 and Fig.4. For this reason the cumulative failure criterion seems to be a universal tool in application to variety of quasi-brittle materials and it offers a possibility for application in numerical codes. SOME ENERGY CONSIDERATIONS It is well known that the energy is a principal variable in all branches of fracture mechanics. Every material has certain ability to accumulate energy, elastic or plastic, before fracture or failure. The principal energy components involved in fracture processes are the stored elastic energy, the energy dissipated during loading (plastic or related to damage) and in the dynamic processes the kinetic energy, for example the energy of micro-cracking, may intervene. Therefore a relation is a permanent search to relate failure criteria to those three forms of energy. The total energy related to the fracture processes is then composed roughly with those three parts plus the thermal energy related to the atom vibrations of a solid, WT = kT. In the isothermal processes of fracture the total energy can be divided into

Where WE, GF and wk are respectively the elastic energy, the energy dissipated or released during failure and the kinetic energy involved in the process of material separation. Assuming that the contribution of the kinetic energy is small the quasi-brittle materials break as shown schematically in Fig.6. Where the critical stress is reached when the elastic energy in a solid reaches the critical value WE = WF. The fracturing releases the energy. The elastic energy is released in dynamic processes usually by unloading elastic waves and vibrations, the energy dissipated in separation is the area indicated by GF.

tc =-EC E

tc

ts

TIME

Fig.6 Scheme of loading and fracturing of quasi-brittle solid. In dynamic failure processes in quasi-brittle materials, as it was shown by Griffith the released energy GF is much lower than the elastic energy WF. Therefore the main indicator

On fracture energy of concretefor short-time loading in tension

555

when a solid breaks is the maximum elastic energy being stored just before failure. The energy balance can be written as follows

In dynamic loading, and especially in the case of spalling, the time periods shown in Fig.6 , tc and ts ,where respectively they are the critical time and the time of the final separation of a specimen, are very short. For example, in experiments performed at LPMM tc varies from -10 p s to -50 ps. A quasi-static approach to material separation transferred to dynamic processes by a simple analogy leads very frequently to wrong assumptions and conclusions. The dynamic failure differs to that in quasi-static conditions. Simply in dynamic separation not one but many sites on penny-shaped cracks are activated and the time of separation is much shorter than expected by analogy to slow material separation. The reason is shown in Fig.7 for two hypothetic situations, that lower and higher loading times.

Fig.7 Rapid separation of semi-brittle material, a) - longer separation time, b) shorter time of separation, i.e. lower and higher strain rate. Because many penny-shaped cracks propagate with high velocities, close to -0.3 CR,where CR is the velocity of Rayleigh waves, CR = Cr , they shield each other causing the “rate effect”. The micro-cracks posses as well its inertia and they carry a small kinetic energy amplifylng the “rate effect”. In conclusion, physical sources of the rate effects in semi-brittle materials are dzflerent than in metals and polymers. Assuming that the changes of Young’s modulus at different strain rates are negligible and the micro-damage does not change much the tangent modulus E, = do/dE. In fact those two effects compensate each other, the simplified estimation of fracture energy is given by 1 W, (i.)= -o,i. tc 2

or

1 2 W, (i.)= -0, (i.) 2E

Substitution for oF(i.) ,Eq.(4), the explicit failure energy is obtained

(13)

Janusz R.KLEPACZKO

556

Since m

= (l+&

the relation for the failure energy can be rewritten to the following form

WF(€)=A €

2m

Where A is the material constant involving combination of constants assumed in the failure criterion (2)

Because the logarithmic rate sensitivity m depends on the activation energy AGO,Eq.(lO), the fracture energy is to some extent also related to temperature. At T = 0 the logarithmic rate sensitivity reduces to mo = (AGO)-'.The values of the failure energy calculated after Eq.(14) seem to be reasonable in comparison to the experimental data reported in the literature. In Table 2 all values of the exponent 2m obtained in this analysis are assembled. SOURCE LPMM LPMM [4]

[51 This study

MATERIAL Wet concrete MB50 Dry concrete MB50 Concrete Concrete Mean value

2m 1.046 0.882 1.486 0.692 1.0265

It is interesting to note that that the exponent defining the rate sensitivity of the failure energy is close to unity. It means a "viscous" variation of the failure energy with strain rate. In that case Eq.( 14) simplifies to the following form

2E This relation can serve as a preliminary estimation of the failure energy at different strain rates. It is important to note that the ''viscosity" constant depends on the quasi-static constants and a, which can be assumed in within an expected range. Of course, the failure energy WF(&) is calculated for a unity of volume. This is not the case of the energy released during separation GF, Fig.6. The energy GF is usually given as the specific energy per unit of the hctured surface, units [J/m2]. Estimation of this energy for the case of dynamic failure depends on a model assumed in a particular analysis. In [5] a standard definition of the specific fracture energy was applied to determine GF via the spa11 tests. The fracture energy GF is the integral of the force F over the crack opening S, then 1 GF=- SF(8)dS A , where A is the cross sectional area of specimen. In this paper the impulse transfer was applied to determine GF. This solution was possible because at the specimen free end an

0nfi;acture energv of concretefor short-time loading in tension

557

accelerometer was attached. Therefore the velocity of ejected specimen fragment could be measured as a function of time, in that case the total energy of separation is given by GF = AI8, where Al is the impulse transfer. It was found that the mean specific fracture energy increases from GF= 288 J/mz at the mean strain rate 33 s-' to GF= 334 J/m2 at 67.8 s-' . It was shown again that not only the critical failure stress OF and the critical energy WFdepends on the rate of loading but also the specific energy of material separation GFis rate dependent. CONCLUDING REMERKS

The review of experimental results presented in this paper clearly demonstrates that experimental technique based on the spa11 phenomenon and developed at LPMM-Metz for concrete testing in dynamic tension yields useful data to be implemented into different applications, including computer codes. The results obtained by different laboratories which applied the same or similar experimental arrangements confirm a very high rate sensitivity of the failure stress which occurs above the strain rate threshold, typically from -1.0 s-' to -10 s-I. The same trend is observed concerning the failure energy WF.Also the specific energy of material separation GF seems to be rate dependent. The local cumulative failure criterion discussed in this paper enables to estimate not only the failure stress OF as a function of strain rate but also the failure energy WF.The criterion was already applied in numerical codes. Further research in this area would be very useful if the decisive variables, that are the strain rate or the loading rate, could be extended to a maximum possible range for the same material, from quasi-static rates to high available ones. REFERENCES 1. Klepaczko, J.R. Dynamic crack initiation, some experimental methods and modeling. In: Crack Dynamics in Metallic Materials, J.R. Klepaczko ed. Springer Verlag, Vienna 1990, p.255. 2. Klepaczko, J. R., and Brara, A., An experimental method for dynamic tensile testing of concrete by spalling, International Journal of Impact Engineering, 25,2001, p.387. 3. Brara, A., Camborde, F., Klepaczko, J. R., Mariotti C., Experimental and numerical study of concrete at high strain rates in tension, Mechanics of Materials, 33,2001, p.33. 4. Wu, H., Zhang, Q., Huang F., Jin, Q., Experimental and numerical investigation on the dynamic tensile strength of concrete, International Journal of Impact Engineering, 32, 2005, p.605. 5. Schuler, H., Mayrhofer, Ch., Thoma, K., Spa11 experiments for the measurement of the tensile strength and fracture energy of concrete at high strain rates, International Journal of Impact Engineering, 32,2006, p.1635. 6 Weerheijm, J., Van Doormaal, J.C.A.M., Tensile failure of concrete at high loading rates: New test data on strength and fracture energy from instrumented spalling tests, International Journal of Impact Engineering, 33,2006, in print. 7. Davies, R.M., A critical study of the Hopkinson pressure bar, Phi. Trans., A240, 1948, p.375.

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8. Faure, L., and Klepaczko, J. R., A Fast Video Setup with CCD Cameras, Technical Report, ISGMP-LPMM, Project GDR 972 “Impact on Materials”, Metz University, France, 1996. 9. Landon, J. W. and Quinney, H., Experiment with the pressure Hopkinson bar, Proceedings Royal Society of London, series A : Math. Phys. Sci., A103, 1923, pp. 622-643. 10. Klepaczko, J. R., and Brara, A., An experimental method for dynamic tensile testing of concrete by spalling, International Journal of Impact Engineering, 25,2001, p. 387. 11. Klepaczko, J.R. On a very high rate sensitivity of concrete failure at high loading rates and impact. In: Proc. Int. Symp. “Brittle Matrix Composites 7”, A.M. Brandt, V.C. Li and I.H. Marshall, Wodhead h b l . Ltd, Cambridge and Warsaw 2003, p. 1. 12. Denoual, C., and Hild, F., Dynamic fragmentation of brittle solids: a multi-scale model, European Journal of Mechanics, Nsolids, 21,2002, p. 105. 13. Klepaczko, J.R., Statistics of the transition from brittle to ductile fracture and a new local fracture criterion. In: Proceedings of DYMAT Meeting: “Dynamics of Damage and Fracture”, CEG, France, 1990, p. 1. 14. Brara, A., Klepaczko, J.R., Experimental characterization of concrete in dynamic tension, Mechanics of Materials, 38,2006, p.253.

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, FC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

DYNAMIC RESPONSE OF CONCRETE AT HIGH LOADING RATES A NEW HOPKINSON BAR DEVICE Ilse VEGT and Jaap WEERHEIJM Delft University of Technology Faculty of Civil Engineering and Geosciences P.O.Box 5048,2600 GA Delft, the Netherlands E-mail: [email protected] ABSTRACT Explosion scenarios in tunnels, the potential hazards from storage of high energetic materials and terrorist attacks have become important safety issues. The mechanical response of concrete structures exposed to impact loading can only be predicted with proper material modelling that includes the rate effect of concrete. The influence of moisture on the rate effect of concrete is studied with a Split Hopkinson Bar. The results are presented in this paper. A new testing device for very high loading rates is developed. The development of the device and a new measurement technique is presented. The fmt results on concrete are promising.

Keywords Concrete, dynamic tensile loading, rate effects, high loading rate (1000 GPds), test development, Modified Split Hopkinson Bar, moisture effects. INTRODUCTION Last years, explosion scenarios in tunnels and the potential hazards from storage of high energetic materials have received significant attention. Moreover, due to terrorist attacks, structures that can withstand explosions have become a major concern. Knowledge about the response of concrete structures to explosive loading is required for reliable safety assessment and the design of protective structures. Concrete is a rate-dependent material, which means that the mechanical properties depend on the applied loading rate. The rate dependency of concrete complicates the prediction of the mechanical response of concrete structures exposed to explosive loading even more. Therefore, the response can only be predicted properly with material models that include this rate effect. With the current knowledge on computational modelling, force- and stress distributions can be calculated in concrete structures under complex dynamic loading conditions. However, material models that properly account for the true dynamic material response are still in their infancy. Therefore, these material models are the weak link in advanced finite element calculations. Reliable test data, that support material modelling, are only available to a limited extend. Data on the rate effect of concrete at very high loading rate are scarce. Furthermore, the mechanisms

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behind the rate effect are not completely understood. Without the proper experimental data and the insight on the cause of the rate effect, numerical models cannot be validated. In a combined experimental and numerical research programme the rate effect of concrete at different loading rates and the mechanisms behind the rate effect are studied [11.

RATE EFFECT OF CONCRETE The dynamic material properties under tension are important parameters for the response of concrete structures under impact loading. The tensile strength and fracture energy have a major influence on the failure mode and the residual bearing capacity of the structure. Also, the rate effect of concrete is more pronounced under tensile loading than in compression. Therefore, the current research is focussed on tension. Research conducted on the rate effect of concrete under tensile loading conditions show that beyond a certain threshold the rate effect becomes more pronounced [2, 3,4]. Consequently, the rate dependency of concrete in tension can be subdivided into two regimes; the regime with moderate rate effects for loading rates in the range from lo4 GPds (static loading rate) up to 15 GPds, and the regime with extensive rate effects for loading rates beyond 15 GPds (Fig. 1). Loadlng rate efteet on

:o*

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Figure 1. Rate dependency of concrete under uniaxial tensile loading. ~

Micro-inertia effects in the fracture process zone mainly cause the rate dependency of concrete in the high regime [2,4]. The moisture content is assumed not to be dominant in this regime. In the moderate regime, on the contrary, the moisture content is believed to play an important role in the strength increase of concrete [5 - 81. The free water in the micro-pores is assumed to exhibit the so-called Stefan-effect causing a strengthening effect in concrete with increasing loading rate. Although the influence of moisture in the moderate regime has been proven by several different researchers [5 - 81, the mechanisms at meso- and microlevel are still not completely clear. In order to be able to include the rate effect of concrete in a physically true material model, the meso- and micro mechanic behaviour of concrete as influenced by the humidity level should be known. To investigate this behaviour an experimental study is carried out in the moderate regime. The results from these experiments are summarked in this paper. Also, earlier research

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was mainly focussed on the determination of the strength at medium loading rates. Data on the dynamic fracture energy or on the rate effects at higher loading rates are very limited. To be able to conduct research on the tensile strength and fracture energy at very high loading rates (1000 GPds) a new Hopkinson Bar device was developed, i.e. the second subject of the current paper.

EXPERIMENTAL SET-UP Loading rate To determine the tensile dynamic properties of concrete, experiments are conducted at three different loading rates. To study the behaviour under moderate loading rates the gravity driven Split Hopkinson Bar (SHB) at the Stevin laboratory of the Delft University of Technology (DUT) is used. Static tests with a loading rate of lo4 GPds are conducted as a reference. The behaviour of concrete at very high loading rates is studied with a so-called Modified Split Hopkinson Bar device (MSHB), which can reach loading rates up to 1000 GPds. Concrete specimens The specimens used for the experiments are cylindrical specimens with a diameter of 74mm. The length of the specimens for the static and the SHB tests is lOOmm and for the MSHB tests 240mm. The cylindrical specimens were drilled out of concrete cubes, to make sure that effects that will occur at the edge zones (sides) of the moulds are not included in the specimens. The concrete cubes are demoulded after one day and placed in a wet environment of 95% RH for 14 days, to avoid dehydration. Next, the cubes are subjected to a controlled environment of 50% RH and 20°C for another 14 days. At an age of 28 days, the specimens are drilled out of the cubes. The composition of the concrete that is used for the experiments is similar to the concrete used in earlier research [2] in order to be able to compare the results from the experiments with results obtained in the past. For the cement a Portland cement (CEM I 32.5R) is used. The composition of the concrete is shown in Table 1. The mechanical properties of the concrete are determined at 28 days and 42 days (with the exception of the Young’s modulus, which is only determined at 42 days). The compressive cubic strength and tensile splitting strength are determined following the recommendations EN 12390-3, EN-12390-4 and EN-12390-6. The static Young’s modulus is obtained by prism pressure tests. The mechanical properties of the concrete are summarized in Table 2. Table 1. Composition of concrete. Mix A

Aggregate kg/m3 1810

Cement kg/m3 375

Water cement ratio

Air %

0.5

2.5

Table 2. Mechanical properties of concrete. 28 days 42 days

Cube strength MPa 48.8 52.4

Tensile splitting strength MPa 3.6 3.6

Young’s modulus GPa 35.7

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Moisture content To study the influence of the moisture content on the rate effect of concrete, the cylindrical specimens are subjected to four different conditions for approximately 14 days (after being drilled out of the cubes): 50% RH and 20°C (‘normal’); immersion in water (‘wet’); drying in an oven of 35°C (‘dry35’); drying in an oven of 105OC (‘dryl05’). Most of the specimens are kept in the final conditions until testing. However, the specimens that are dried at 105°C are taken out of the oven at 42 days and wrapped in a plastic bag, to avoid any moisture absorption and to be able to cool before being tested. Approximately 42 days after casting the experiments are carried out and the strength and fracture energy are determined. The influence of moisture on the rate dependency of concrete will not only be studied in the moderate rate regime but also in the high regime. The experiments under very high loading rates are not carried out yet. The experiments in the moderate regime are conducted and the results are summarized in the next chapter. EXPERIMENTS AT MODERATE LOADING RATES (SHB) Equipment Static and SHB tests The Split Hopkinson Bar used for the tests at moderate loading rates consists of two vertical cylindrical aluminium bars (0 74mm)between which the concrete specimen (length 1OOmm) is glued. The tensile stress wave is generated with a drop weight, which slides along the lower bar and hits an anvil at the end. In the applied set-up the strain-time relation is measured with strain gauges at the upper bar, while the deformations are measured directly on the notched specimens with LVDT’s (Linear Variable Differential Transducers). Therefore, the measurements have to be synchronized to obtain the stress-displacement curve. The strength and fracture energy can be determined from this stress-displacementcurve. The static loading device used for the reference tests is also situated in the Stevin laboratory. The experiments are deformation controlled. During the experiment the deformation and the force are measured and the load-deformation curve can be immediately obtained. Results SHB The data that has been collected on the tensile strength and fracture energy under medium loading rate conditions at four different moisture levels are presented in Table 3. Table 3. Average strength and fracture energy for SHE3 and static tests. Normal Dry (35°C) Dry (105°C) Wet

fi static

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MPa

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5.58 6.40 5.04 6.35

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1 ft,

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Gf,SHE Nlrn

Gf, SHE 1 Gf, ataic

120 130 135 80

120 300 265 165

1.o 2.3 2.0 2.1

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The ratio between the dynamic strength and the static strength is often used as an impression of the rate dependency. As can be seen from the data all moisture levels show an increase in strength when the loading rate increases. The ratio between the dynamic and the static strength is obviously higher for the wet concrete. This indicates that the wet concrete is more rate dependent than the concretes that are kept under normal or dry conditions. The ratios presented in Table 3 are similar compared to results obtained by other researchers [3, 5 - 81. Differences found in the absolute values might be explained by the lower temperature of drying (and therefore the higher moisture content), the difference in drying time and moisture gradient and the difference in specimen geometry. Split Hopkinson Bar tests conducted with specimens dried at 5OoC have resulted in strength values and dynamicktatic strength ratios that do not match the trends observed in the results presented in the current paper. This indicates that the drying procedure greatly influences the results on the rate dependency of concrete [9] and should be taken in consideration when comparing results with literature. MPa

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Figures 2 and 3. Stress displacement curves for static tests (Fig. 2) and Split Hopkinson Bar tests (Fig. 3) for different moisture levels. Not only the strength of the concrete was determined, the stress-displacement curves were also determined (Figs. 2 and 3). The area under the stress displacement curves represents the fracture energy. The static fracture energy of the dry and normal concrete is comparable. The fracture energy of the wet concrete is lower, due to the lower strength and the lower bearing capacity in the second part of the curve. Figure 13 shows the stress-displacement diagrams for the Split Hopkinson Bar tests. When the static and SHB stress-displacement curves (Figs. 2 and 3) are compared, it is observed that the tails of the curves seem not to be influenced by the higher loading rate. For all four moisture contents the rate effect seems to be most pronounced for the ascending branch and the first part of the descending branch. The peaks of the SHB stress-displacement curves of the wet and dry concrete are much wider than the normal concrete. The formation of propagating micro-cracks causes a growth in damage of the material, and therefore in the energy absorption, until finally the formation of a macro-crack takes place. The stress-displacement curves will be combined with microscopic analysis of the fracture zones and will form the basis of the material model.

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DEVELOPMENT OF MSHB FOR HIGH LOADING RATES MSHB set-up The loading rate that can be generated with the SHB set-up at the DUT is about 50 GPds. For the very high loading rates (up to 1000 GPds) a new Modified Split Hopkinson Bar set-up is developed at the TNO Defence Security and Safety in Rijswijk. The feasibility was demonstrated by the TNO prototype test set-up [lo]. Based on the results of the feasibility study and the promising perspectives, TNO decided to build a small and large scale test set-up (with diameters of respectively 74 and 30Omm). The Modified Split Hopkinson Bar (MSHB) is based on a different principle than the SHB; the principle of spalling. The small MSHB set-up consists of a steel bar with a length of 2m and a diameter of 74mm (Fig. 4). The rod is placed horizontally and supported by strings. The rod can move freely in horizontal direction and the wave propagation in the rod will therefore hardly be influenced by the supports. The shock wave is introduced into the rod by detonating an explosive charge at one end of the rod. At the other end of the rod, the concrete specimen (diameter 74mm, length 240mm) is attached to the steel rod with a plaster that has almost no strength. The plaster will break when tensile stresses are introduced, making sure that only the compressive part of the impact wave travels into the concrete specimen. The concrete specimen is first loaded in compression and will fail in tension due to the reflected tensile wave (spalling). The principle of the large MSHB set-up with a diameter of 300mm is the same as the smaller set-up (Fig. 5).

Measuring equipment of MSHB The measuring equipment of the Modified Split Hopkinson Bar is based on the equipment of the Split Hopkinson Bar. The strain-time relation is measured at the steel bar to check the amplitude, energy and reproducibility of the loading pulse induced by the explosive charge. The transmitted pressure pulse, the wave propagation and reflection process are recorded with a number of strain gauges distributed along the notched specimen (Fig. 6). The applied load can be derived from the strain measurements on the specimen. The resulting stress at the failure zone (notch) can be determined using the uniaxial wave theory to quantify the wave interaction process.

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Figures 6 and 7. Detail of strain gauges at the concrete specimen and detail of the new deformation measuring device after a test. To be able to determine the stress-displacement curve it is necessary to measure the deformation of the notch directly. Similar to the SHB of the DUT, the measured deformation must be combined with the stresses in the notch section to obtain the desired stress-displacement curve. The fracture energy can then be established from this curve. Other researchers [l 11 that have used similar experimental methods as the MSHB, have obtained the fracture energy by determining the energy captured in the part that is blasted from the specimen due to the spalling mechanism. However, with this method it is only possible to get a value for the fracture energy and not a stress-displacement curve. The aim of the current research is to develop a method to measure the deformation directly and derive the load-displacement relation. The methods and displacement gauges used for the Split Hopkinson Bar cannot be used in the case of very high loading rates. In spite of their small own mass, the LVDT’s start to vibrate due to the high accelerations. The vibration of the LVDT’s disturbs the measurements in such a way, that no useful data can be obtained. Therefore, the new measuring device is required to be almost weightless to avoid problems with vibration. The device also has to be linear within its measuring range and be able to obtain reproducible results. The newly developed measuring device consists of a strain gauge glued on a supporting material. The ends of the supporting material are glued onto the specimen, leaving a certain unglued area around the notch. The supporting material should behave linear and be able to stretch for a certain distance before it breaks. Different strain gauges and different supporting materials were tested under static loading conditions to determine which would perform the best. Eventually it was decided that strain gauges with a length of 30mm will be used and that a synthetic foil will be applied for support (Fig. 7). When the strains of this combination of strain gauge and supporting material are plotted against the imposed static deformation, a linear curve is obtained up to failure. The new measuring device is also tested under high loading rate conditions. For the purpose of these tests an aluminium specimen was used instead of a concrete specimen. The measuring device was placed over the joint aluminium specimen - steel bar. LVDT’s were also placed over the joint, to compare the results of the new device with the original method. The tests showed that the new device behaves similar as the LVDT’s but without the disturbing vibrations (Fig. 8).

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Figure 8. Comparison new devices (Displ-strl and Displ-str2) with LVDT's (Displ 1 to 3).

The low frequency wave on the displacement curve in Figure 8 is no disturbance but is caused by the up en down travelling loading pulse in the aluminium specimen. The time difference between two successive peaks of the wave is the time needed for the wave to travel up en down. The test results of the strain gauges on the foil support are promising. Nevertheless, the new measurement system should be carefully calibrated dynamically before conducting full test series, to make sure that the behaviour of the new measuring system is also linear under high dynamic loading conditions. Data from first MSHB tests After the tests with the aluminium specimen, a limited experimental test series with concrete specimens was performed to check the feasibility of the set-up and the new measuring device for concrete. Experiments were conducted to check whether the transmitted loading pulses into the concrete were reproducible, the loading pulse was uniaxial and the behaviour of the concrete was elastic under the incoming pressure wave and the returning tensile wave after passing the notch.

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Figures 9 and 10. Strains of three tests (Fig. 9) and Compressive strains at W/R4 and R5 (Fig. 10).

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Dynamic response of concrete at high loading rates a new Hopkinson bar device

The experiments are reproducible, which is shown in Figure 9. The strains in the concrete specimen at a distance of 65 mm from the joint at the steel bar (R3) are compared for three different tests (Fig. 9). As can be seen, the strains in the concrete are almost equal. The loading pulse is observed to be uniaxial. The loading pulse in the concrete specimen is determined at 65 mm from the joint. The strains are measured at two sides of the specimen, R3 and R4. As can be seen from Figure 10, the strains at R3 and at R4 are approximately the same. Figure 10 also shows the strains at a distance of 130 mm from the joint (R5). The compressive strains do not decrease during the travelling of the wave, which indicates that the behaviour of the concrete is elastic. The behaviour of the concrete under the returning tensile wave (the tensile stresses that pass the fracture zone) can be assumed to be elastic, since the strains of the returning wave are never larger than the strength at the notch. Furthermore, because the cross-section at the notch is reduced the stresses due to the returning wave are smaller in the specimen than in the notch-section.

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Figures 1 1 and 12. Examples of determined stresses at the notch section (Fig. 11) and deformation measurement with strain gauges glued on a foil support (Fig. 12).

The above mentioned experiments show that the tests can be reproduced, that they are uniaxial and that the concrete is not damaged due to the compressive wave. However, the question remains if it is possible to determine a stress-displacement curve with the combination of the direct deformation measurements and the strains measured at the concrete specimens. Examples of the deformation measurements and the strains at a distance of 24 mm from the notch are given in Figures 11 and 12. The stresses are calculated using the dynamic Young’s modulus which is derived from the longitudinal wave velocity. When the stresses at the notch are determined using the uniaxial wave theory and these measurements are combined, a stress-displacementcurve can be determined (Fig. 13). However, as can be seen from Figure 13, the stress-displacement curve could not completely be determined; the tail of the curve is missing. This is caused by the fact that the strain measurements at the position 24 mm from the notch are not only composed of the initial compressive wave and the returning tensile wave, but also of the secondary compressive wave returning from the free end of the concrete specimen, which was first joint with the steel bar through plaster. This makes it almost impossible to determine the real strains at the notch for the performed tests.

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Figure 13. Example of a stress-displacementdiagram. The conclusion is that a stress-displacement curve can be established with the method as described before. However, the curve could only be determined until the time when the secondary waves are starting to interact with the returning tensile wave. When the length of the specimen and the positioning of the strain gauges on the concrete specimen are adjusted properly, these problems will not occur and a full stress-displacement curve can be determined. CONCLUDING REMARKS

In a combined experimental and computational research programme the mechanisms of the dynamic response of concrete are studied. The aim is to understand and model the true, dynamic behaviour of concrete in tension for moderate as well as high loading rates. Therefore, the experimental part focuses on determining the stress-displacement relation for static, moderate and high loading rates. Furthermore, the influence of moisture on the rate effect of concrete is studied. The results of the static and Split Hopkinson Bar experiments presented in this paper show a clear influence of the moisture level on the rate dependency of concrete. The wet concrete exhibits a more pronounced rate effect in tensile strength than the dry and normal concrete. An important finding in the study on the effect of moisture content in the moderate rate regime is the dominant effect of the drying procedure. When comparing results the influence of the drying procedure on the absolute values of the static and dynamic strength should be taken into consideration. The other result reported in this paper is the possibility to derive experimentally the stressdisplacement curve at very high loading rates. So far, this was not possible. By direct measurement of the displacement in the fracture zone and by combining this displacement with the stresses at the fracture zone, it proves to be possible to construct a stress-displacement curve.

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REFERENCES 1. Vegt, I., Weerheijm, J., Pedersen, R.R. and Sluys, L.J., Modelling of impact behaviour of concrete- an experimental approach. In G. Meschke et a1 (eds.): Computational modeling of Concrete Structures, Proceedings of EURO-C 2006, 27-30 March 2006 Mayerhofen, Taylor & Francis group London, 2006, pp 45 1-458. 2. Weerheijm, J., Concrete under impact tensile loading and lateral compression. Thesis. TNO Rijswijk, 1992. 3. Cadoni, E., Albertini, C., Labibes, K. and Solomos, G., Behavior of plain concrete subjected to tensile loading at high strain rate. In R. de Borst et a1 (eds.): Fracture Mechanics of Concrete Structures, Swets & Zeitlinger Lisse, 2001, pp 341-347. 4. Brara, A. and Klepaczko, J.R., Experimental characterization of concrete in dynamic tension. Mechanics of Materials 38,2006, pp 253-267. 5 . Cadoni, E., Labibes, K., Albertini, C. and Berra, M., Strain rate effect on the tensile behaviour of concrete at different relative humidity levels. Materials and Structures 34, 2001, pp 21-26. 6. Ross, A.C., Jerome, D.M., Tedesco, J.W. and Hughes, M.L., Moisture and strain rate effects on concrete strength. ACI Materials Journal 93 (3), 1996, pp 293-300. 7. Rossi, P., van Mier, J.G.M., Boulay, C. and le Maou, F., The dynamic behaviour of concrete: influence of free water. Materials and Structures 25, 1992, pp 509-514. 8. Rossi, P., Strain rate effects in concrete structures: the LCPC experience. Materials and Structures Supplement March, 1997, pp 54-62. 9. Vegt, I., Weerheijm, J. and Schlangen, E., Influence of moisture content on the dynamic behaviour of concrete. Subjected to ECF- 16 Conference July 3-7 2006, Alexandroupolis, Greece. 10. Weerheijm, J., Doormaal, J.C.A.M. and van de Kasteele, R.M., Development of a new test set-up for dynamic tensile tests on concrete under high loading rates. Part 2: Results and evaluation of test series B and C on concrete. TNO Rijswijk, 2003. 11. Brara, A. and Klepaczko, J.R., Fracture energy of concrete at high loading rates in tension. International Journal of Impact Engineering (In Press), 2006.

Proc. Int. Syrnp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

MICROSTRUCTURE ALTERATIONS UNDERLYING ELECTROCHEMICAL PROCESS OF CHLORIDE-INDUCED CORROSION Dessi A. KOLEVA, Jing HU, Piet STROEVEN Faculty of Civil Engineering and Geosciences, Delft University of technology, Stevinweg 1,2628CN Delft, The Netherlands e-mail: [email protected]

ABSTRACT The electrochemical processes of chloride-induced corrosion are intimately associated with structural alterations that take place at different interfaces in reinforced mortars, such as the interface between mortar and aggressive media, between cement paste and aggregate, and the one between steel reinforcement and mortar matrix. Scanning electronic microscopy (SEM)techniques render possible quantitative characterization of the composite microstructureat various interfaces, including structural morphology of corrosion and hydration products as well as the pore structure. The correlation between electrochemical impedance spectroscopy (EIS) measurements and microstructure investigationsallows breaking down the electrical properties of reinforced concrete to micro-level interface microstructures. The results indicate that different components in the EIS modelling concept correspond well to the specific interfaces. This will provide significant implications for computer simulation of the corrosion process and modelling of concrete performance in aggressive environments.

Keywords EIS, cathodic protection, pore structure, SEM image analysis, chloride-induced corrosion INTRODUCTION Reinforced concrete is a macroscopic heterogeneous composite material aggregated at different structural levels. Hence, multi-phase interfaces are involved in concrete structure and concrete behaviour. The steel reinforcement is embedded on meso-level in supposedly homogeneous concrete bulk. However, on a lower structural level (micro-level), the concrete bulk is consisted of cement paste and aggregate (sand) particles, with air voids and macro pores dispersed in the cement paste matrix. Further, cement paste can be decomposed into unhydrated cement, hydration products, and capillary pore structure. The latter is generally assumed to have significant relevance to permeability and other transport phenomenon in concrete technology. Corrosion of embedded steel reinforcement in concrete represents a great concern with respect to the durability and safety operation of concrete structures. When sufficient aggressive ions (e.g., sulphate and chloride from seawater or chloride from de-icing salts) have penetrated to the reinforcement or when the pH of the pore solution drops to low values due to carbonation, the protective film is destroyed and the reinforcement steel is depassivated. Corrosion prevention and protection techniques have been a focus of interest for decades in the field of civil engineering. Various protective methods, including epoxy-coated steels, overlays, membranes, impregnations or inhibitors, are used to prevent corrosion in new structures. The present authors have carried out an extensive experimental program on

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chloride-induced corrosion in reinforced concretes and on efficiencies of cathodic protection techniques for corrosion prevention in concrete structures. The electrical properties (material behaviour) of reinforced mortar and concrete in aggressive environment are monitored by electrochemical methods, specifically by electrochemical impedance spectroscopy (EIS). On micro-level, material microstructure at aforementioned interfaces and the bulk is investigated by scanning electronic microscopy (SEM). Most relevant to the corrosion process is the interface between the steel reinforcement and the mortar (concrete). Chloride-induced corrosion is a highly localized attack, leading to macro-cell corrosion situation (corresponding to very high local penetration rates of aggressive substances). Localized corrosion is a result of chloride penetration through the concrete cover, resulting in interaction of chloride ions at the steel-paste interface. The corrosion products are mainly host of iron oxichlorides and iron oxyhydroxides complex. It should be noted that only part of the chlorides, i.e., the free chlorides, are responsible for corrosion initiation at certain threshold values. The rest of the chlorides are chemically bound, forming chloro-aluminate complexes, e.g. Friedel’s salt (3CaO.Al203.CaC12.10H~O). In this case, chloride ions incorporate in the cement hydrates, thus forcing portlandite to release OHso as to reach equilibrium with the alkali ions in the pore solution. Another part of the chloride ions, remain free (water soluble) in the pore solution, e.g. in the form of NaC1. The processes of chloride penetration and corrosion products formation depend on a wide variety of factors, including the concrete quality, the pH value of pore solution, concentrations of alkali ions in concrete and the aggressive environments. On the other hand, chloride penetration and corrosion process will induce structural changes in the material in terms of various aspects, e.g., modifling the pore structure and interfacial structure, inducing cracking at the steel-paste interface (as a result of volume expansion of corrosion products). These microstructure alterations are underlying the evolution of electrical behaviour of reinforced concrete during the corrosion process. EIS is a technique working in the frequency domain. The basic concept is that an electrochemical interface can be modelled as a combination of passive electrical circuit elements, i.e., resistance, capacitance and inductance. When an alternating voltage is applied to these elements, the resultant current is obtained by using the Ohm’s law. In a simple way, impedance can be thought of as the resistance of a circuit to an alternating waveform [13. EIS can provide information on a number of parameters, such as the presence of surface films, bulk concrete characteristics, interfacial corrosion and mass transfer phenomena. However, detailed interpretation of the EIS results can be a difficult task and the equivalent circuit varies with steel condition. By combining different components of EIS with specific interface microstructures, this contribution is expected to provide more fundamental and detailed knowledge of the corrosion process in the reinforced concrete system. This knowledge will provide significant implications for computer simulation of the corrosion process and modelling of concrete performance in aggressive environments.

EXPERIMENTAL Three groups of reinforced mortar specimens (40 mm in diameter and 150 mm long) were cast according to standard procedures, using ordinary Portland cement CEM I 32.5 (with cement-to-sand mixing proportion of 1:3 and water cement ratio of 0.6). The relatively poor cement quality and high water cement ratio are designed to accelerate the corrosion process. A construction steel bar (6 mm in diameter) was centrally located (embedded length of 80 mm) in all the mortar specimens. Group N were cured in fog room conditions (95% RH, 20°C) for 14 days and afterwards partially (1/3 of the height) submerged in 7% NaCl solution

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in a lab environment for 7 months. Group P is in the same technical conditions as the group N, but a cathodic protection current was applied to the specimens since 60 days of cement hydration. The third group (R) is partially immersed in demineralised water and used as reference specimens. By means of cathodic protection, the corrosion protection is achieved by supplying impressed direct current (DC) to the steel reinforcement. The repulsion of anions (e.g. chloride) is a beneficial one as far as the corrosion risk of the steel is concerned. This chloride removal technique will reduce chloride level near the steel surface, and thereby delay the corrosion initiation of steel. The quantitative characterization of material microstructure by means of SEM and energy dispersive X-ray analysis (EDXA) is expected to provide insight into the structural alterations induced by cathodic protection, and therefore, helps to explain the mechanisms underlying the positive effects of cathodic protection techniques. In general, different electrical circuits with equally good fit results can be used to model the response of a system to applied signals [2]. In the present study, mainly depressed semicircles appeared in the impedance response of the freely-corroding samples (N), while the reference (R) group showed capacitive behaviour at low frequencies. The effect of depressed semi-circles is attributed to non-ideal capacitive behaviour which response instead with ideal capacitance has to be modelled with constant phase element (CPE) [ 1, 21. As investigated here and also available in literature [2, 31, this response is due to the system inhomogenity. When the electrode is rough, or in the case of irreversible redox reactions taking place on the electrode surface like the corroding samples in this study, the impedance plot is distorted with respect to the ideal semicircle. Thus, the CPE is employed in modelling the response in the investigated systems. The model for EIS is composed of the various components. Resistances R1 and R2, corresponding to the high frequency arcs of the reference samples. R1 is related to the electrolyte resistance, R2 to the concrete cover resistance. The corresponding CPE in the high frequency domain (Ql and Q2) indicate non-ideal capacitance of double layers in the bulk mortar in terms of pore network and interfacial transition zones (ITZ)between cement paste and aggregate. The resistance R3, corresponding to the intermediate arc, is employed to represent the pore solution resistance, having a significant contribution to the overall impedance response in the reference samples and negligible effect in the corroding samples. The corresponding Q2 represents the non-ideal capacitance of double layers in the pores of the mortar structure. The intermediate arc is not indicative of a charge transfer but is due to surface film on the pore walls and R2 deals with its ionic resistance. The low frequency domain is normally used to evaluate charge transfer process in combination with mass transport process [4]. The resistance Rq, corresponding to the low frequency arcs, was used for estimation of the polarization resistance of the embedded steel. The corresponding Q3 is used to determine the non-ideal interfacial capacitance of the steel surface. The equivalent electrical circuit for the corroding specimens contains the same elements as the reference, the difference appearing only in the pore resistance, which as mentioned above, is believed to be negligible in respect to the bulk mortar resistance due to the higher pore interconnectivity in the corroding samples. The reference mortar shows an almost straight line inclined to the vertical axis in the low frequency zone throughout the investigation period, indicating passive state of the steel reinforcement. In contrast, the N mortar presents a (depressed) semi-circle, typical for active corrosion of the steel reinforcement.

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INTERFACE MICROSTRUCTUREALTERATIONS

XRD analysis of the steel reinforcement surface clearly revealed the presence of iron oxides, iron oxyhydroxides (FeOOH), mainly deposited on the steel surface. Hematite, magnetite and maghemite are categorized in the group of iron oxides. The group of iron oxyhydroxides (FeOOH) varies in chemical compositions and presents different morphologies, according to which they can be categorized as goethite (a-FeOOH), lepidocrocite (y-FeOOH), akaganeite [Fe3+(0,0H,C1)].Goethite and lepidocrocite are prominent as corrosion products in rust. The fiequent occurrence of Green Rust, magnetite and lepidocrocite is associated with an ample supply of Fez' ions in the steel-mortar interface. It is expected that goethite, akaganeite, lepidocrocite are present in both N and P specimens, however, the dimension and morphology of the corrosion products are significantly different. The element mapping reveals that cathodic protection current can efficiently prevent chloride ions to penetrate through the concrete cover to the steel surface; in contrast, chloride ions are concentrated at the immediate surface of the steel reinforcement in N mortars.

relatively large an2 intact protective Ca(dH)z-layer.Further magnification ofthe corrosion products (right, 2000x) indicates main morphology of goethite is cotton-ball and tiny flowerlike structures; the latter is typical for goethite and lepidocrocite.

Fig. 2 Freidel's salt in cathodic protected mortar (left) and monosulphate (of plate morphology) in reference mortar (right). The needle-like crystals in both images are ettringite.

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SEM images in this study reveal tiny flower-like goethite growing on a relatively dense layer which is mainly composed of non-stoichiometric magnetite bartially substituted by calcium) and maghemite. In addition to the different proportions of corrosion products between the N and the P mortars, another important aspect is the crystallinity of iron oxychloride and goethite. In the P mortar, iron oxychlorides are whiskery and forms a relatively flat and compact layer. In contrast, iron oxychlorides of lamellar and globular types are found in the N mortar. In the P mortar, calcium rich layer is relatively intact (Fig. 1 left); goethite and lepidocrocite, is of semi-crystalline characteristics (Fig. 1 right). The favourable morphology (smaller dimension and lower crystallinity) of goethite can be attributed to the relatively high pH value and lower salinity (i.e., lower concentration of chloride ions) in the protected sample. SEM observations also show the presence of hexagonal crystals of Friedel's salt (Fig. 2, left) on the steel surface and in bulk mortar in both P and N mortars, surrounded by hydration products. As mentioned in the introduction section, the C3A phase of Portland cement has the ability to complex with the dissolvable chloride, resulting in formation of insoluble Friedel's salt 3CaO.Al203.CaC12.1OH20 or 3CaO.Al203.NaCl.lOH20. This combining of C3A phase with free chlorides in hydrated cement, results in the reduction of the corrosion-inducing dissolvable chlorides in the pore solution and also retards further ingress of chloride ion in the concrete. In the reference (R) mortar, well crystallized ettringite (Fig. 2, right) is the primary sulphoaluminate phase detected at early ages. Monosulphate was not detected by XRD at early ages in the reference mortar, though this does not necessarily indicate its absence, as much of the monosulphate at early ages is likely poorly crystalline and not detectable by XRD. The outer hydration products in R mortar are hardly of single phases and likely contain small amounts of other phases, such as calcium hydroxide, ettringite, monosulphate (AFm), etc., which are intermixed with the C-S-H gel (Fig. 2, right). Interpretation of these data is not straightforward. Substantial presence of ettringite and monosulphate are evidenced in the R mortar. Some of the monosulphate was deposited in small cavities or in the hollow shells remaining from the complete dissolution of very small cement particles, but much of the monosulphate observed were small platy lumps occluded by the fibrillar C-S-H in the outer product (see Fig. 2, right). These deposits of monosulphate were relatively large (often several micrometers in size). Two possible mechanisms have been suggested for the reactions between ions in pore solutions and cement paste components to form ettringite [S], i.e., topochemical (i.e. replacement of preexisting mineral phases) and through solution (i.e. direct precipitation from solution). The conversion of the AF, phase, monosulphate to Friedel's salt can be discussed in terms of the structural similarities between the phases. Friedel's salt and monosulphate are both members of the AF, family of structurally related phases, which have layered structures ~ ] ~net +. (Fig. 2), the basic building unit of which has the constitution [ C Q A ~ ~ ( O H ) ~The positive charge of this layer is balanced by anions: either monovalent, e.g., OH-, C1-, or divalent, e.g., CO?-, SO?-, etc. To maintain charge balance, the number of interlayer anions may be variable; for example, 2Cl--SO:-. The different numbers of anions required for charge balance, as well as their different sizes and polarisabilities, control the variable interlayer water contents and control the exact layer stacking sequence. Hence, Friedel's salt (in which the balancing ion is mainly C l 3 is readily distinguished from monosulphate, the main AF, type phase occurring in CT-free cement pastes.

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CORRELATION BETWEEN EIS AND MICROSTRUCTURE

In this section, the different elements of the equivalent circuits will be correlated to specific interfaces in the composite microstructure. The resistance of bulk mortar (represented by R1 and R2) is mainly dependant on properties of the pore solution and of the interfacial transition zone (ITZ)between cement paste and aggregate. The evolution of polarisation resistance (R4) reflects the electrochemical phenomena at the steel-paste interface. Song [6] proposed a model of equivalent circuit for interpreting the experimental Nyquist plot. According to the author, there are basically three types of paths in concrete structure: (1) continuous conductive paths (CCPs), discontinuous conductive paths (DCPs), and "insulator" conductive paths (ICPs). The CCPs are the continuously connected micro-pores, which could be a series of capillary cavities connected through pore necks. The discontinuous micro-pores in the concrete form the DCPs, whose continuity is blocked by the cement paste layers, which is also denoted as "discontinuouspoints" (DPs) in the paper. The discontinuous pores can also be connected to continuous pores as dead ends. Apart from the DCPs and CCPs, the continuous concrete matrix, which is composed of cement paste particles, acts as "insulator" paths (ICPs) in the concrete. Theoretically, current conduction through CCP occurs by ions migration in the pore solution, i.e., Ohm's law operates in this case. Therefore, the total impedance of all the CCPs in the concrete could statistically be described as a resistance RCCP.It can be expected that RCCP is inversely proportional to porosity and pore connectivity, whereas positively proportional to the resistivity of pore solution, and to the tortuosity (mainly dependent on volume fraction and shape of aggregate grains) of the transport paths (CCPs). It is also related to geometry of the concrete material (e.g. thickness of the concrete specimen). Specimen geometry and concrete mixing proportions are constant in this study, so only porosity, pore connectivity and resistivity of pore solution are relevant to RCCPin this case. The resistance of CCPs (denoted as RCCP)can be associated with the experimental values of R1 and R2 by RCCP = Rl+Rz. Compared with CCP, DCP has more complicated impedance expression because of the discontinuous points (DPs). The impedance of DCP can be considered consisting of two parts: the continuous portion of DCP and the discontinuous point (i.e. cement paste layers). At the DP point, current has to "penetrate" througb the cement paste layer. However, the cement paste has high resistivity and is usually regarded as an "insulator," so a DC current is difficult to "penetrate" through such a cement paste layer. However, the "discontinuouspoint" can also be treated as a double parallel plate capacitance (CDP)with the cement paste as its dielectric. The continuous portion of DCPs would have an impedance similar to that of CCP, and can be described as a pure resistance RCP= ( R ~ + R ~ ) R ~ For / R z theoretical . details of this equivalent circuit model, see Song [6]. During the process of cement hydration, a large amount of hydration products (C-S-H gel) are generated, which would block CCP paths and narrow the DCP paths or thicken the DP layers (cement paste). All these changes in CCP and DCP paths would increase RCCPand RCP,resulting in gradual increment of R1 and Rz with hydration time. Xie et al. [7] attributed R1 to the resistance of the pore solution, and Rz to an interface resistance between the solid (unhydrated cement and hydration products) and liquid (i.e. pore solution) phases in the cement paste. It was found out that R1 increase marginally with hydration time, whereas the influences of cement hydration on RZare much more significant. In the reference mortar, substantial existence of monosulphate was evidenced. In the freely corroding (N group) and protected (P) mortars, monosulphate would bind with chloride ions and transform into Friedel's salt (Fig. 2). The mechanism of Friedel's salt formation in cement pastes has not been unambiguously identified. Literature suggested two mechanisms

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underlying the conversion from monosulphate to Friedel's salt [8]. The conversion of hydroxy-AF, to Friedel's salt by ion exchange or by the absorption of C1- as Friedel's salt forms by precipitation. Friedel's salt consists of two principal [Ca2Al(OH)62H20]+ layers that require balancing negative charge for stability. The availability of Cl- (from NaCl ingress) would satisfy charge neutrality in the solid but this would disturb the ionic charge balance in the pore solution. To compensate for this, an equivalent amount of Na' ions would be required to leave the pore solution, i.e. be absorbed into solid phases. In a cement paste system, the main ions in the pore solution are OH-, K', Na+, and Ca". Ions migrate to the interface from the pore solution, and pass through the interface, then continue to travel in the solid phase. In this case, the solid is usually an ion conductor, otherwise the ions would be accumulated in the vicinity of the interface in the solid side and would finally stop the current conduction. At the interface between the pore solution and cement gel, then the cations K', Na', and Ca2+would be "piled up" in the cement solid, which would dramatically change the properties of the cement solid. Fig. 3 reveals that concrete resistance of the R mortar is gradually increasing, in correspondence to the continuous hydration process. Porosity and pore connectivity are further reduced at more matured hydration stage. The much higher value of RCCPfor the R mortar (see Fig. 3c) is attributed to the fact that the R mortars were submerged in demineralised water. In contrast, the N and P mortars (both submerged in 7% NaCl solution) present much lower resistance, reflecting the significant effect of chloride ingress on electrical property of concrete. 400

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The N2 mortar witnessed a sharp drop in both RCCP(Fig. 3a) and in RCP(Fig. 3b) around 150 days of cement hydration. Petre-Lazar and Ghard proposed a simplified model to correlate the valence of iron ions in corrosion products, volume expansion of the corrosion products, as well as porosity and thickness of the steel-paste interfacial zone with mechanical aspects of reinforced cementitious materials [9]. According to this model, the estimated crack initiation period for the N2 mortar is about 130 days. This is in line with the time of the sharp drop in mortar resistance for the N2 specimen. This implies that the corrosion-induced cracks yield formation of channels and networks of CCPs. Fig. 3 reveals that RCCPof the P specimen maintains relatively stable throughout the investigation period and that the value of RCPindicates an obvious trend of increase since 160 days of cement hydration. The estimated crack initiation period for the P mortar is about 4 years. This can be mainly attributed to the favourable morphologies of corrosion products (very small dimensions and semi-crystallined) in the P mortars. The formation of small amounts of these tiny oxides can diffuse through the porous structure of the bulk mortar and locally reduce the porosity. This explains the estimated crack initiation period of 4 years for the P mortar. The polarisation resistance (Fig. 4, associated with R4, usually expressed in ohm.cm2) indicates the intensity of corrosion phenomena (charge transfer) at the steel-mortar interface. Cross-section images (SEM) of the reinforced mortar specimens are used to investigate the pore structure characteristics and the corrosion products at the steel-mortar interface. The quantitative image analysis techniques involve morphological opening operation on the pore space (to measure total porosity and critical pore size) and skeletonization of the pore network (to calculate pore connectivity). Details of the methodology for pore structure characterization can be referred to [lo]. In general, a higher porosity is corresponding to a higher pore connectivity, however, this is not necessary the case. For example, the porosity in the steelmortar interfacial zone is higher in the R (12.68%) than in the N (10.87%) mortar, that latter is resulting from the accelerated hydration in the presence of chloride ions. However, the pore connectivity (a dimensionless parameter, for definition, see [ 101) in the R mortar (at 120 days of hydration) is 0.18, significantly lower than in the N specimen (0.27). 2O M 5

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Fig. 5 SEM images (SOOX, cross-section of specimens immersed in 7% NaCl solution for 7 months) of the steel-mortar interface microstructure in P (left) and in N (right) mortars. This can be partially attributed to the presence of large amount of Hedley grains in the R mortar, which contribute to the total porosity but are isolated from the pore network. The formation of Hedley grains is assumed to be associated with CsA phase. The average porosity at the steel-mortar interfacial zone is 8.53% for the P mortar (Fig. 5 left) at 210 days of hydration, accompanied by a pore connectivity of 0.06. At this stage, volume expansion of corrosion products already initiated cracks in the N mortar; hence, the connectivity of void space is about 0.37, significantly higher than the P mortar. The polarisation resistance for the reference mortar (R) exceeds that of N mortar by one order of magnitude (around 1x106 Ohm.cm2) at 35 days of cement hydration. The Rp value dropped to 5x105 Ohm.cm2 at 50 days of cement hydration and remains stable at 2 . 5 ~ 1 0 ~ Ohm.cm2 afterwards. This sharp drop at 50 days can be correlated partially with structural morphology aspect of the material. It was observed that monosulphate were accumulated to a significant extent on the steel surface in the R mortar. At the hydration stage of around 2 months, monosulphate is well crystallised into plate structures, with relatively large size of a few micrometers. This morphology is responsible for the relatively coarser pore structure (e.g. critical pore size is 2.85 pm at the steel-mortar interfacial zone, at 120 days of cement hydration) in the R mortar. At the same hydration stage, the critical pore size is 1.23 pm for the N mortar. Accumulation of expansive corrosion products at the steel surface will lead to radial compressive stresses. De Wind and Stroeven [ 111 made an approximation of the induced stress and strain state by means of an elastic analysis. The corrosion-induced internal displacement of the interface layer leads to tensile stresses that will be going to exceed considerably the ultimate tensile strength of the material during the service life of constructions in spite of creep and relaxation effects. In an early stage of the of service life a construction, corrosion-induced microcracking will be inevitable as indicated in Fig. 5 . Concrete quality dominantly governs the capacity of accumulating a multitude of small isolated cracks. For decreasing quality the tendency to form macrocracks due to material heterogeneities will enhance and spalling off will be the structural effect. Hence, the freely corroding concrete structures will experience crack propagation and thus incurred damage in a relatively short time (the calculation result is about 2-3 years). In contrast, the service life of the protected structures will be significantly prolonged by one order of magnitude.

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SUMMARY AND CONCLUSIONS This contribution aims at exploring the microstructure alterations in concretes subjected to chloride-induced corrosion and cathodic protection techniques, as well as the correlation between the microstructure and the electrochemical process. For the freely corroding mortar specimens, the resistance of continuous conductive paths (CCPs) is increasing in the first 4 months as a result of cement hydration. However at about 130 days of cement hydration, the excessive volume expansion of corrosion products (of lamellar and flower-like morphologies) leads to cracks in the steel-mortar interfacial zone. The cracks further propagate into the bulk mortar and yield a sharp drop in the resistance of CCPs. The cathodic protection induces favourable microstructure changes in the specimen, in terms of dimension and morphological aspects of the corrosion products. The tiny corrosion products difhse through the material structure and fill in voids and open space, thus block a part of the conductive paths. As a consequence, the resistance of CCPs are maintained at a stable level and displays a trend of increase after 160 days of cement hydration. The polarisation resistance reflects the electrochemical phenomena at the steel-paste interface. Microstructure investigations and morphological studies of pore structure characteristics (by means of quantitative image analysis) supporting the experimental results of EIS. Combination of electrochemical measurements and microstructure characterisations allows breaking down the EIS measurements to different interface structures in concrete composites, and sheds light on the fundamental mechanisms underlying the efficiency of cathodic protection techniques. REFERENCES

1. M.F. Montemor, A.M.P. Simoes, M.G.S. Ferreira, Chloride-induced corrosion: fundamental and techniques, Cement Concrete Composites, 25,2003, pp 491-502 2. Impedance Spectroscopy Emphasizing Solid Materials and Systems, John Wiley 8z Sons, 1987 3. A.A. Sagues, S.C. Kranc and E.I. Moreno, Corrosion Science, 37, 1995, pp 1097-1113 4. J. Flis, L. Dawson, J. Gill and G.C. Wood, Corrosion Science, 32, 1991, pp 877-892 5. H. Lee, R.D. Cody, A.M. Cody, P.G. Spry, The formation and role of ettringite in Iowa highway concrete deterioration, Cement and Concrete Research, 35,2005, pp 332-343 6. G. Song, Equivalent circuit model for AC electrochemical impedance spectroscopy of concrete, Cement Concrete Research, 304 1,2000, pp 1723-1730 7. P. Xie, P. Gu, Z. Xu, J.J. Beaudoin, A rationalized AC impedance model for microstructural characterization of hydrating cement systems, Cement Concrete Research, 23, 1993, pp 359-367 8. M. Deng, M. Tang, Formation and expansion of ettringite crystals, Cement and Concrete Research, 24, 1994, pp 119-126 9. I. Petre-Lazar, B. GBrard, Mechanical behaviour of corrosion products formed at the steel concrete interface. Testing and Modelling. American Society of Civil Engineers, Austin 2000 10. J. Hu, Porosity of Concrete - Morphological Study of Model Concrete, PhD thesis, Delft University of Technology, Delft 2004 11. G. de Wind, P. Stroeven, Chloride penetration into offshore concrete and corrosion risks. Proceedings of Katharine and Bryant Mather International Conference on Concrete Durability, ACI Publications, Detroit 1987, pp 1679-1690

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

EFFECT OF POROSITY AND FRACTURE TOUGHNESS ON EXPLOSIVE SPALLING OF CONCRETE Dita MATESOVA, Zbyngk KERSNER Institute of Structural Mechanics Brno University of Technology, Faculty of Civil Engineering Vevefi 33 1/95,602 00 Bmo, Czech Republic, e-mail: [email protected],[email protected]

ABSTRACT Quasibrittle materials like concrete and mortar when subjected to high temperatures disintegrate (gradually or explosively)due to physical and chemical changes. The goal of this paper is evaluation of the combined effect of transport properties and resistance against crack growth on the explosive spalling occurrence. The proportions of these two features are dominant for explosive spalling occurrence apart from structure dimensions, heating rate and water content. The investigated material is concrete on the basis of Portland cement, natural mined sand (0/4 nun) and aggregate (4/8 mm). Four w/c ratios are studied 0.3,0.4,0.5 and 0.6. The total volume of opened pores is measured by using the high pressure intrusive mercury porosimetxy method to evaluate the transport properties. The resistance against crack growth is quantified by the effective fracture toughness. The combination of these properties is compared to the susceptibility of particular mixtures to the explosive spalling failure. This susceptibility is evaluated on a number of specimens subjected to the high temperature up to 1000°C in the medium heating rate of 2OWmin. Each specimen is supplied with a thermocouple, so the point of explosion can be recorded.

Keywords Porosity, fracture toughness, explosive spalling, high temperature, water-cement ratio.

INTRODUCTION Quasibrittle materials like concrete and mortar when subjected to high temperatures disintegrate due to physical and chemical changes. The attention should be focused generally on two types of concrete deterioration: (1) the gradual degradation of material quantified by e.g. the concrete fracture and mechanical parameters and (2) the explosive spalling, a brittle failure that occurs suddenly and violently. The latter phenomenon is very dangerous and not exactly predictable so far. It is believed that the primary reason for the explosive spalling in cementitious materials is the pressure build-up of volatiles and the secondary role is played by the build up of strain energies caused by thermal stresses [l-31. The latter is supported by the observation that the temperature range in which the explosive spalling occurs coincides with the time of high thermal gradient between the surface and the center of specimen [4]. Explosive spalling can take place at temperatures around 200°C and even higher [3-61. The pressure build-up is related to the strength, composition and, more specifically, to the content of vapors that on one hand are formed at a specific temperature and on the other hand to the pressure relief due to permeability. Sullivan [7] used the fractional factorial method and showed that the

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deterioration at elevated temperatures cannot be attributed to a single factor; it is rather affected by several parameters including the heating rate, wlc, the curing, the type of aggregate and fibers, and the load. Similar results were reported by Khoury et al. [8]. An effective prevention of explosive spalling is e.g. addition of polymeric (PP) fibers [2,9-131 to the cement matrix as they reduce a pore pressure build-up. According to Kalifa et al. [9] PP fibers are partially absorbed by the cement matrix when melted at the temperature around 17OOC and leave a pathway for gas. So they contribute to the creation of a more permeable network than the matrix, which allows the outward migration of gas and results in the reduction of pore pressures. Other opinions say that such a reduction is attributed to the formation of weak and porous interface zone between the fibers and the matrix that is readily transformed into continuous cracks upon the heating [6] or to the incompatible movements of the fibers with the matrix that produce microcracks long before melting [7]. Thus the transport properties are substantially increased. Hertz and Sorensen [14] developed the method which predicts the spalling behavior of specific concrete with specific moisture content and which takes into account the effect of restrained thermal expansion. This method also showed that PP fibers may hinder explosive spalling of dense concrete as well as if the thermal expansion is restrained. The aim of this paper is the experimental evaluation of the corporate effect on the opened porosity that quantifies the transport properties and the effective fracture toughness that reflects the resistance against crack development susceptibility of concrete to fail in the manner of explosive spalling. These two features are believed to be the very dominant parameters affecting the explosive spalling occurrence aside from the structure dimensions, the heating rate and the water content. Also the measurements of compressive strength are presented as stresses generated during heating which result often in explosive spalling failure and are not necessarily tensile. They may be compressive or a combination of both [7]. MATERIAL Four concrete mixes with wlc ratios 0.3, 0.4, 0.5 and 0.6 were designed. In all cases, the amount of aggregate was constant and the content of water and plasticizer was adjusted in order to obtain mixtures with the target wlc ratios and similar consistency (co!e slump S2: 50-100 mm). As a binder and filler, cement CEM I1 B 32.5 R (according to CSN EN 197-1 [ 151) from Hranice and aggregate of two fractions 014 and 418 mm from Tovaeov locality were used, respectively. Superplasticizer FM 794 was used for wlc = 0.3 and FM 350 for wlc = 0.4,0.5 and 0.6. The designed mixtures compositions are given in Table 1.

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Table 1 Designed compositions of concrete mixtures.

CEM II B 32,5 R [kg/m3] Water mg/m3] Sand DTK 014 TovaEov [kg/m3] Gravel HTK 418 TovaEov [kg/m3] Plastificizer [Urn3]

Waterlcement ratio [-I 0.3 0.4 0.5 532 468 414 159.6 187.2 207 1040 1040 1040 645 645 645 10.714 2.857 1.429

0.6 369.5 221.5 1040 645 0.714

Three different types of laboratory specimens were casted. Cubes 150 mm, cylinders 300 mm high and 150 mm in diameter and beams 80~80x480mm for compressive strength,

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porosity and fracture toughness evaluation, respectively. Laboratory specimens were unmolded afier 24 hours of setting and placed to humid environment of conditioning chamber.

EXPERIMENT The following parameters of four different concrete mixtures were evaluated (1) compressive strength, (2) porosity, (3) fracture toughness and (4) potential to explosive spalling occurrence. (1) Compressive strength was measured on cubes 150 mm in three duplicates per each wlc ratio after 140 days of setting.

(2) Porous structure was investigated by using the high pressure intrusive mercury porosimetry method. This method is based on the phenomenon that hydrophobic liquids do not infiltrate to the pores of material without applying of outer pressure. This method is utilized in material engineering for the assessment of the pores size and distribution in the total volume of material. For the purpose of this paper the total volume of pores is presented. Porosimetr PoreSizer 93 10 f j Micromeritics was used for experimental measurements in this case. Four duplicates from two different cylindrical specimens for each w/c ratio were evaluated. (3) Fracture toughness was calculated on the basis of load-deflection curves obtained from three point bending tests (3PBT) performed on centrally notched beams 8 0 ~ 8 0 ~ 4 8mm 0 after 140 days of setting. The effective crack model by Nallathambi and Karihaloo (see e.g. [ 161) was used for calculation of effective fracture toughness of three duplicates for each wlc ratio.

(4) Fractured beams 80~80x240mm from 3PBT were used for the evaluation of explosive spalling occurrence. The specimens were heated in a programmable muffle furnace at the nominal heating rate of 20°C/min, which is in the range of medium heating rates, up to 1000°C. The specimens were equipped with type K (alumel-chromel) thermocouple that was embedded at the center of the specimens immediately after the cast. The thermocouple’s wires were isolated by ceramic tubes and were connected to a data logger that recorded the temperature at the center of the specimens every 10seconds. The explosive spalling occurrence was quantified by means of (i) the relative amount of specimens that spalled regarding to the total amount of specimens heated for a certain w/c ratio and (ii) average percentage weight of spalled material with regard to the original specimen volume. For wlc ratios 0.3, 0.4,0.5 and 0.6 the number of heated specimens was 9, 8, 10 and 6 , respectively. RESULTS AND DISCUSSION

The figures 1-4 and Table 2 show the experimentally evaluated characteristics vs. the w/c ratios of tested mixtures. The compressive strength in Figure 1 typically decreases from wlc ratio 0.3 to 0.5. The value of wlc 0.6 is very similar to wlc 0.5. The decrease of strength with increasing wlc ratio is the consequence of the increasing porosity (total intrusion volume), see Figure 2. Proportionally, very similar results on porosity of cementitious mortars with various WIC ratios were obtained by Lafhaj et al. [17]. The effective fracture toughness in Figure 3 represents the resistance of material against a crack development. It has mostly the same trend as the compressive strength in Figure 1.

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Figure 2 Total intrusion volume (mean values with standard deviations and coefficient of variation) vs. wlc ratio. Table 2 gives the percentage weights of spalled parts of the heated specimens with regard to their total weight. Zero value indicates that no spalling occurred. Data from Table 2 are graphically expressed in Figure 4. The points connected to the three different lines represent three different statistics of the results. The points connected with the thick line represent the percentage amount of spalled specimens with regard to the total amount of specimens heated for given wlc ratio. The two thin lines connect points that reflect the average relative weight of spalled material. The continuous one connects values that involve all specimens and the dashed one includes only specimens that spalled. The specimens most susceptible to explosive spalling and with regard to the total number of spalled specimens are wlc ratios 0.3 and 0.6. The most resistant to the violent failure is wlc ratio 0.5. According to

585

Effect ofporosity and fracture toughness on explosive spalling of concrete

the values connected by the dashed line the amount of spalled parts (only spalled specimens included) decreases with increasing wlc ratio except wlc=O.5 (only one spalled specimen - not statistically significant). The trend of the relative weight of spalled parts (all specimens included) is similar except of ~ 1 ~ 0again. . 5 According to Sullivan [7] the concrete with wlc ratio 0.35 (0.35 and 0.25) appeared to be most susceptible for reinforced (plain) concrete to explode from the range of wlc ratios 0.25,0.35 and 0.5.

-Meanvalue

COV 12

2.0

0.0

'

I

0.3

0.4

i2 ' 0

I

I

0.5

0.6

WatedCement Ratio 1-1

Figure 3 Effective fracture toughness (mean values with standard deviations and coefficient of variation) vs. wlc ratio. Table 2 Quantification of explosive spalling by relative weight of spalled parts of the specimen.

\To. of specimen 1 2 3 4

5 6 7 8 9 10

wlc = 0.3

wlc = 0.4

12.61 11.25 12.68 70.96 7.87 11.64 6.40 0 32.89

8.70 0 9.16 4.12 34.18 9.06 0 36.36

-

-

-

wlc = 0.5 0 0 0

0 0 0 0 0 18.37 0

wlc = 0.6 10.44 5.58 0 5.36 6.07 9.43 -

-

The modes of explosives failure were various independently on wlc ratios, see Figure 5. The most frequent type of damage was (a) - one side damage, the other modes (bd)were rather unique and (e) is the specimen after heating to 1000°C which did not spall. The temperature in the center of specimens at the moment of explosion was in the range of 1802 10°C (the higher wlc ratio the lower temperature at violent failure).

Dita MTESOVA, Zbynik KERSNER

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Spalling (all spec. included)

- - Spalling (onlyspalled specimens included) -(3.

+hunt

of Spalled Specimem

1

0.3

0.4 0.5 WatedCement Ratio 1-1

100

0.6

Figure 4 Quantification of spalling tendency of concrete (mean values) vs. wlc ratio.

(b)

(c)

(4

Figure 5 Modes of explosive spalling failure. Figure 6 shows the combined effect of fracture toughness and porosity on the spalling behavior. It is depicted the dependency of fracture toughness on the total intrusion volume ratio vs. the quantified spalling behavior (three different statistics, as described above). The total intrusion volume represents the transport properties of the material and the fracture toughness represents the resistance against a crack development. So these two parameters can quantify the key properties of concrete as the susceptibility of material to explosive spalling. If we consider the continuous line in Figure 6 (left) that connects the points representing the relative weight of spalled material with respect to the total weight of heated specimens, we may say that the tendency of the explosive failure decreases with the decreasing ratio of the fracture toughness and the porosity. One more question arises. How to quantify the explosive spalling correctly? The three types of quantification presented in this study are: (I) percentage amount of failed specimen

Effect of porosity and fracture toughness on explosive spalling of concrete

587

and (2) and (3) percentage relative weight of material that spalled with respect to the total weight of specimen (all specimens included or only spalled specimens included). The following features of the damage are not considered by this quantification: the size of the pieces of damage part and the type of damage - surface (Figure 5c), one side (Figure 5a) or L o sides spalling (Figure 5b, d). ~

25

I

--

-o-

10

Spallrng (an specimens inchled) Spallrng (only spaued specimens inchded)

20 30 40 50 60 Fracture toughness I Total Intrusion Volume

10

20 30 40 50 61 Fracture toughness I Total Intrusion Volume

Figure 6 Combined effects of fracture toughness and porosity on spalling behavior.

CONCLUSIONS The combined effect of the effective fracture toughness and the opened porosity on the susceptibility to explosive spalling failure in a medium heating rate of 2O0C/min in Portland cement based concrete with wlc ratios 0.3, 0.4, 0.5 and 0.6 was analyzed. The fracture toughness (porosity) decreases (increases) with increasing w/c ratio. The difference among average values of particular parameters for wlc ratios 0.5 and 0.6 is in the range of standard deviations. The concrete specimens were also subjected to compressive strength measurements; it has a very similar trend to effective fracture toughness. Three different criterions were used for quantification of explosive spalling behavior. According to the amount of spalled specimens, the most susceptible specimens for the explosive failure were wlc ratios 0.3 and 0.6, the most resistant is wlc ratio 0.5. If we consider that the transport properties (here represented by the total intrusion volume) and the strength (here represented by the compressive strength and fracture toughness) are the very dominant parameters affecting the explosive spalling failure occurrence then it is necessary to evaluate not only the single characteristics but also their proportions.

ACKNOWLEDGEMENTS This outcome has been achieved with the financial support of the Ministry of Education, Youth and Sports, project No. 1K04 111, Clutch.

Dita MATESOVA,Zbynik k X a N E R

588

REFERENCES

1. Ahmad, G. N., Hurst, J. P., An analytical approach for investigating the causes of spalling of high strength concrete at elevated temperatures. International Workshop of Fire Performance of High Performance Concrete, National Institute of Standards and Technology, Gaithersburg, 1997, pp 95-108 2. Persson, B., Fire resistance of self-compacting concrete, SCC. Materials and Structures, Vol. 37,2004, pp 575-584 3. Phan, L. T., Lawson, J.R., Davis, F.L., Effects of elevated temperature exposure on heating characteristics, spalling, and residual properties of high performance concrete. Materials and Structures, vol. 34,2001, pp 83-91 4. Phan, L. T., Carino, N. J., Effect of test conditions and mixture proportions on behavior of high strength concrete exposed to high temperatures. ACI Material Journal 99 (l), 2002, pp 54-66 5. Kalifa, P., Menneteau, F. D., Quenard, D., Spalling and pore pressure in HPC at high temperature. Cement and Concrete Research 30,2000, pp 1-13 6. Matesovk, D., Bonen, D., Shah, S. P., Factors affecting the resistance of cementitious materials at high temperatures at a high heating rate. Materials and Structures, in press 7. Sullivan, P.J.E., A probabilistic method of testing for the assessment of deterioration and explosive spalling of high strength concrete beams in flexure at high temperature, Cem. Concr. Composites 26,2004, pp 155-162 8. Khoury, G.A., Grainger, B.N., Sullivan, P.J.E., Transient thermal strain of concrete: literature review, conditions within specimen and behavior of individual constituents, Materials and Structures 37 (132), 1985, pp 131-144 9. Kalifa, P., Chene, G., Galle, Ch., High-temperature behavior of HPC with polypropylene fibres: From spalling to microstructure. Cement and Concrete Research 31, 2001, pp 14871499 10. Meda, A., Gambarova, P. G., Bonomi, M., High-performance concrete in fire-exposed reinforced concrete sections. ACI StructuralJournal 99 (3), 2002, pp 277-287 11. Bilodeau, A., Kodur, V.K.R. and Hoff, G.C., Optimization of the type and amount of polypropylene fibers for preventing the spalling of lightweight concrete subjected to hydrocarbon fire, Cem. Concr. Composite 26,2004, pp 163-174 12. Sullivan and Associates, Deterioration and spalling of high strength concrete under fire, Offshore Technology Report 2001/074,2001 13. Bostrom, L., Self-compacting concrete exposed to fire, In proc. of 3rd International Symposium on Self-compacting Concrete, Reykjavik, Iceland, 2003 14. Hertz, K.D.,Sorensen, L.S., Test methods for spalling of fire exposed concrete. Fire Safety Journal 40,2005, pp 466-476 15. CSN P EN 197-1 Cement - SloBeni, jakostni poEadavky a kritCria pro stanoveni shody. Part 1: Cementy pro obecnC pouiiti, 1993, in Czech 16. Karihaloo, B. L., Fracture mechanics of concrete. Longman Scientific & Technical, New York, 1995 17. Lafhaj, Z., Goueygou, M., Djerbi, A., Kaczmarek, M., Correlation between porosity, permeability and ultrasonic parameters of mortar with variable waterkement ratio and water content. Cement and Concrete Research 36,2006, pp 625-633

Proc. Int. Symp. "BrittleMatrix Composites 8" A.M. Brandt, KC. Li and I. H. Marshall, eds. Warsaw, October 23-25, 2006 ZTUREK RSI and Woodhead Publ., Warsaw 2006

HARDENING OF CALCIUM HYDROXIDE ANDCALCIUM SILICATE BINDERS DUE TO CARBONATION AND HYDRATION Ozlem CIZER', Jan CAMPFORTS', Koen van BALEN', Jan ELSEN' and Dionys van GEMERT' 'K.U.Leuven, Department of Civil Engineering, Building Materials and Technology Division, Kasteelpark Arenberg 40, B-300 1 Heverlee, Belgium, e-mail: [email protected] 2K.U.Leuven,Department of Geography and Geology, Physico-chemical Geology Section Celestijnenlaan200C, B-300 1 Heverlee, Belgium

ABSTRACT Hardening of calcium hydroxide and calcium silicate binders composed of cement, rice husk ash (RHA) and lime in different compositions were studied with mechanical strength, mercury intrusion porosimetry, thermal analysis and SEM. When cement is partially replaced with RHA and lime, hardening occurs as a result of combined hydration, pozzolanic reaction and carbonation reaction. While hydration of cement contributes to the early strength development of the mortars, carbonation is much more pronounced at later stage with the decrease in the cement content and the increase in the porosity of the mortars. RHA-cement mortars indicated a long-term strength development, which is lower than that of the reference cement mortar. This was attributed to the high water demand of the blended mortars due to the porous FWA grains, which resulted in an increase in their porosity. Strength reduction was recorded at the very early stage for RHA-cement-lime mortars containing 10%wt cement. This has been explained with the insufficient cement content, destructive effect of the calcium carbonate phases on the initially hydrated cement phases and partial carbonation of the initially hydrated phases.

Keywords RHA, cement, lime, hydration, pozzolanic reaction, carbonation. INTRODUCTION Use of pozzolana as a partial replacement with cement in mortar can improve its ultimate strength and modify its microstructure due to the formation of additional calcium silicate hydrate (C-S-H) phases through pozzolanic reaction between the pozzolana and calcium hydroxide formed during the hydration of cement. This leads to a decrease in the content of the calcium hydroxide in comparison to that of the hydrated cement. Therefore, presence of the pozzolana influences the progress of the cement hydration and its compounds. In fact, hydration mechanism becomes much more complicated since cement hydration and pozzolanic reaction follow different reaction processes at different rates. Hydration of cement proceeds much faster than the pozzolanic reaction that becomes usually effective between 3-14 days of hydration, which is after most of the alite in the cement has hydrated [l]. This period depends relatively on the reactivity and specific surface area of the pozzolana [l-2-31. Hydration of alite and belite has been reported to be accelerated with the presence of pozzolana [l-3-4-51. This can be explained by the fact that fine pozzolana grains act as a precipitation site for additional C-S-H hases and dissolution of alite and belite is accelerated as a result of the decrease in the Ca+Pions concentration due to their consumption through

590

Ozlem CIZER. Jan CAMPFORTS, Koen Van BALEN, Jan ELSEN, Dionys Van GEMERT

pozzolanic reaction [3]. Literature concerning the hydration of cement in the presence of additional lime is rather limited. Hydration of tricalcium aluminate has been reported to be slightly retarded in the presence of lime [6]. It has been found that hexagonal hydro-aluminates are formed with the hydration of tricalcium aluminate. In addition, formation of additional C-S-H by pozzolanic reaction has been reported to result in a pore size refinement effect leading to a decrease in the large pores and increase in the fine pores [7-8-91. Strength increase of the cement mortars blended with pozzolana is, therefore, correlated with the additional C-S-H formation and pore size refinement effect. Carbonation of the cement blended with pozzolana and lime is generally neglected in the literature. In this study, hardening of cement mortars in the presence of pozzolana and lime has been studied. Rice husk ash (RHA) has been used as a highly reactive artificial pozzolana. Cement was partially replaced with RHA and lime at different ratios. Results are presented for the hardening mechanism, strength development, porosity, and microstructural modification of the blended mortars with respect to the reference cement mortar.

RESEARCH PROGRAMME Materials RHA and commercial hydrated lime were used as a partial replacement with Portland cement (CEM I 52.5 N) in mortars. RHA provided fiom Tanzania is composed of high contents of silica (90.86%) by weight. XRD analysis indicated its amorphous phase due to the broad band between 15-30 28" (Figure 1). It contained certain amounts of cristobalite and tridymite as crystalline silica. This indicated that RHA was obtained by burning at relatively high temperatures (800-10OO0C), which led to the crystallization of the amorphous silica [lo]. In order to be used as a pozzolanic material, the ash was ground by means of a laboratory batch ball mill until certain fineness and surface area (1 3047 cm2/g) was reached. The ground RHA grains are mostly angular edged grains with varying particle sizes less than 50pm (Figure 1).

Figure 1. XRD pattern and SEM micrograph of the RHA. Compositions of the mortars are given in Table 1. The mortars were prepared using standard quartz sand with maximum grain size of 2 mm [111. Binderkand ratio of the mortars was 1:3 by weight. Concerning the RHA-cement mortars, cement was replaced at 3o%-wt (RHA-C.3-7), 5o%-wt (RHA-C.5-5) and 70%-wt (RHA-C.7-3) ratio by RHA. Two types of ternary blended mortars composed of RHA, cement and lime were prepared. The ratio of the cement was kept 10%-wt for both compositions. While one (RHA-C-L.7-1-2) was composed of 70%-wt RHA and 20%-wt lime, the other (RHA-C-L.5-1-4) contained 50%-wt RHA and

591

Hardening of calcium hydroxide and calcium silicate binders due to carbonation and hydration

40%-wt lime. Due to the much more porous structure and higher specific surface area of the RHA and lime particles compared to the cement, they have higher water absorption capacity and water retentivity. This led to the use of more water for the blended mortars to reach similar workability, resulting in different waterhinder (WB) ratios. Table 1. Compositions of the cement mortars blended with rice husk ash and lime

I

Mortar

I

RHA

I

Cement

I

Lime

I

Sand

I

W/B

I

Flow

I

Methods Hardening of the mortars was studied using standard mortar beams (40x40~160mm) which were prepared in accordance with the European standards [ 111. Mortars were cured at the standard laboratory conditions (20°C, 60% R.H.) for 180 days. The progress of hardening was studied by means of mechanical strength, porosity, phenolphthalein staining, thermal analysis and Scanning Electron Microscope. Mechanical strength tests were carried out by compressive and three-point bending tests using the standard mortar beams at 7, 28, 60, 90, 120 and 180 days of hardening [ 111. Progress of carbonation of the mortars during curing was tested using phenolphthalein solution. Open porosity of the mortars was determined according to water saturation test by hydrostatical weighing [12]. Pore size distribution of the mortars was studied using Mercury Intrusion Porosimetry (Micromeritics AutoPore IV). Thermal analysis was performed on finely ground samples after they were dried by vacuum drylng at 0.025mbar. The analysis was carried out using a Netzsch STA 409 PC DSC-TGA system in static nitrogen atmosphere at a temperature range between 20-1000°C with a controlled heating rate 10"C/min. Microstructure of the fresh fractured surfaces coated with gold was analyzed using Philips XL 30s FG Scanning Electron Microscope (SEM) coupled with X-Ray Energy Dispersive System (EDS). RESULTS Mechanical strength Progress of compressive and flexural strength of the reference cement mortar and RHA-cement mortars are given in Figure 2 and Figure 3. While reference cement mortar yielded the highest compressive and flexural strength values at all stages, the values became lower as cement was partially replaced by RHA in 30%, 50% and 70%-wt ratio respectively. All the mortars achieved their ultimate compressive strength at 7 days of hardening and strength development continued until 28 days. After that, the compressive strength of only RHA-C.5-5 and RHA-C.7-3 mortars increased gradually until 180 days of hardening. RHA-cement mortars yielded an increasing flexural strength development until 180 days (Figure 3). Increase in the flexural strength of the RHA-C.3-7 mortar between 7 and 28 days was more than that of RHA-C.5-5 and RHA-C.7-3 mortars having lower cement content. Strength development of these two mortars between 28 days and 180 days was much more pronounced than that of RHA-(2.3-7.

Ozlem CIZER, Jan CAMPFORTS, Koen Van BALEN, Jan ELSEN, Dionys Van GEMERT

592

11,

0

IS

,

,

30

45

60

90

75

105

120

135

I50

165

180

195

Time ( d i p )

I+Cref

-+RHAC.3-7

+RHA-C.S-S

+RHA-C.7-3]

Figure3. Flexural strength of the reference cement mortar and RHA-cement mortars after 7,28,60, 90, 120 and 180 days of hardening. RHA-cement-lime mortars yielded relatively lower compressive and flexural strength values than those of the RHA-cement mortars. Reduction in compressive and flexural strength was recorded at the very early stage after 28 days of hardening (Figure 4 and Figure 5). Compressive strength of RHA-C-L.5-1-4 mortar did not change that much between 28 and 60 days while it decreased after 60 days. However, its flexural strength decreased between 28 and 60 days but after that it stayed almost at the same value until 120 days followed by a decrease at 180 days. Strength reduction was much more pronounced for RHA-C-L.7-1-2 mortar between 28 and 180 days even though its initial strength was higher than that of the other. Its compressive strength decreased from 6.23 N/mm2 to 2.67 N/mm2 and its flexural strength from 0.68 N/mm2 to 0.32 N / m 2 .

Hardening of calcium hydroxide and calcium silicate binders due to carbonation and hydration

00

593

8

0

30

60

90

120

150

I80

210

Figure 4. Reduction in the compressive strength of the RHA-C-L.7-1-2 mortar after 28 days and that of RHA-C-L 5- 1-4 mortar after 60 days.

"1

I

00 0

30

60

90

I*RHAc-L

I 20

IS0

I80

210

T h e (days) 5-14 4RHAC-L 7-1-21

Figure5. Reduction in the flexural strength of the RHA-cement-lime mortars after 28 days.

Open porosity and pore sue distribution Open porosity of the mortars determined at 60, 90 and 120 days of hardening are given in Table 2. While reference cement mortar had the lowest open porosity, RHA-cement mortars yielded an increase in the open porosity as the RHA content increased, leading to high WE3 ratio. RHA-C-L.7-1-2 and RHA-C-L.5-1-4 mortars had the highest porosity since they contained 20% and 40% lime respectively and only 10% cement by weight. Table 2. Open porosity of the mortars at 60,90 and 120 days of hardening.

Mortar Cref RHA-c.3-7 RHA-C.5-5 RHA-C.7-3 RHA-C-L.7-1-2 RHA-C-L.5- 1-4

60 days 13.6% 19.3% 21.8% 25.9% 30.7% 31.4%

Open porosity 90 days 120 days 14.1% 12.5% 19.5% 18.4% 22.2% 21.6% 25.7% 25.4% 30.8% 29.4% 31.1% 29.7%

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Ozlem CIZER, Jan CAMPFORTS, Koen Van BALE?( Jan ELSEN, Dionys Van GEMERT

Pore size distribution of the mortars was determined at 90 days of hardening (Figure 6). RHA-cement mortars had higher volume of smaller and larger pores when compared with the reference cement mortar. RHA-cement-lime mortars yielded the highest porosity with their higher volume of large and small pores than the rest. This is mostly related to their composition as they contain only 10%-wt cement and high W/B ratio.

0 001

001

01

10

I

im

IWO

Pore diameter Cm) S +RNA-C7-3

- 0 - M E - L 5 - 1 4 +RHAC-L

Figure 6. Pore size distribution of the mortars after 90 days. Carbonation depth Carbonation depth of the mortars increased as the content of the cement decreased in the mortar (Figure 7). RHA-C.7-3 containing the lowest cement content (30%-w&)reached a full carbonation depth at 60 days of hardening while other RHA-cement mortars still contained uncarbonated calcium hydroxide at the core at 180 days. RHA-cement-lime mortars reached a full carbonation depth at 180 days.

Cmf

RHA-C 3-7

W A S 5-5

W A C 7-3

M E - L 1-14

M<-L7-1-2

,el~y~,,lzs~~8aod,.90daysli20dayl.i~daysl

Figure7. Carbonation depth for the reference and blended cement mortars after 7,28,60,90,120 and 180 days of hardening. Thermal analysis (DTG) Progress of hardening of the mortars was followed by thermal analysis. Derivative of the weight loss (DTG) of the Cref, RHA-C.3-7 and RHA-C.5-5 mortars are given in Figure 8. Weight loss between 100°C and 200°C was attributed to the dehydration of hydrated phases and the loss at around 450°C was due to the decomposition of calcium hydroxide. Weight loss

595

Hardening of calcium hydroxide and calcium silicate binders due to carbonation and hydration

recorded between 500°C and 800°C was derived from the carbonated phases. Results revealed that partial replacement of cement by RHA led to a decrease in the calcium hydroxide content and an increase in the carbonated phases. Further decrease in the cement content resulted in a decrease in the hydrated phases. In the RHA-cement-lime mortars, carbonated phases were much more pronounced than the hydrated phases.

T2

! I

E

E

.-I 0

2W

4M)

600

Temperature pC)

800

1OW

0

200

400

600

Temperature(OC)

8W

1wO

0

2 0

400

6W

8 0

1000

Temperature(T)

Figure 8. Derivative of the weight loss (DTG) of the Cref, RHA-C.3-7 and RHA-C.5-5 mortars at 7,28,60,90 and 120 days of hardening.

Microstructure of the mortars Use of RHA and lime together with cement led to differences in the microstructure of the mortars. Reference cement mortar was characterized by needle-like and reticular C-S-H phases, and calcium hydroxide crystals formed inside the pores and in the matrix (Figure 9). In the RHA-cement mortars, flocs-like and fibrous-like C-S-H phases that were well connected to each other filled the pores of the mortar (Figure 10). Needle-like C-S-H phases were not observed that much in the matrix of these mortars.

Figure 10. SEM micrographs of flocs-like Figure 9. SEM micrographs of needle-like and reticular C-S-H phases formed inside the C-S-H phases formed in the matrix of pore of the Cref after 120 days of hardening. RHA-C.5-5 after 120 days of hardening.

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Ozlem CIZER, Jan CAMPFORTS, Koen Van BALEN Jan ELSEN, Dionys Van GEMERT

Plate-like, semi-circle and round RHA particles, which remained substantially unreacted or partially reacted after 120 days, were embedded within the matrix (Figure 1la-b). These RHA grains acted as a surface for the precipitation of additional C-S-H phases through pozzolanic reaction with calcium hydroxide formed during the cement hydration (Figure Ilc). When totally reacted, flocs-like C-S-H phases were formed (Figure 1Id). These hydrated phases played a bridging role between hydrated and unhydrated cement compounds by growing from the grain surface towards the surrounding matrix. Cluster of calcium hydroxide crystals embedded in the matrix were still detectable in the RHA-cement mortars after 120 days.

Figure 11. SEM micrographs of the matrix of RHA-C.5-5 mortar after 120 days of hardening, showing unreacted RHA grains (a-b), C-S-H phases started to form on the surface of a RHA grain (c), RHA grain that has reacted, forming amorphous C-S-H (d). RHA-cement-lime mortars exhibited a totally different microstructure when compared with the RHA-cement mortars (Figure 12). The matrix was rich in silica and carbonated phases and was composed of agglomerated dense phases with rather large pores in-between. Clusters of needle-like crystals and hexagonal plate-like crystals embedded locally within the carbonated matrix were observed (Figure 12).

Hardening of calcium hydroxide and calcium silicate binders due to carbonation and hydration

597

Figure 12. SEM micrographs of the matrix of RHA-C-L.5-1-4 mortar (after 120 days) containing hexagonal and needle-like crystals embedded within the carbonated phases. DISCUSSION Replacement of the cement by RHA and lime influenced the hardening mechanism, mechanical strength, porosity and microstructure of the mortars. Thermal analysis results indicated that while hardening of reference cement mortar was governed by the cement hydration, that of RHA-cement mortars occurred as a result of early-stage hydration combined with pozzolanic reaction and carbonation at later stage (Figure 8). Decrease in the calcium hydroxide content in the RHA-cement mortars upon hardening gives evidence to its consumption by carbonation and pozzolanic reaction. Carbonation is much more pronounced with the decrease in the cement content and increase in the porosity of the mortars. The reaction started to appear much more after 60 days for MA-C.3-7 and after 28 days for RHA-C.5-5 and RHA-C.7-3. Pozzolanic reaction seems to start before 7 days and contributes to the formation of C-S-H phases upon hardening. The RHA-cement mortars achieved their ultimate strength at 7 days, confirming that early strength gain was mainly controlled by the rapid hydration of cement since pozzolanic reaction proceeds much slower [1-21. Contribution of the pozzolanic reaction and carbonation to the strength development of these mortars is also evident as the strength values increased gradually until 180 days while no considerable increase was recorded in the strength of the reference cement mortar after 28 days (Figure 2 and Figure 3). All the RHA-cement mortars indicated a long-term strength development even though their strength values were lower than that of the reference cement mortar. Strength values became lower as cement was partially replaced by RHA respectively at 30%-wt, 5o%-wt and 70%-wt ratio. This was ascribed to their higher W B ratio than the reference cement mortar, which derived from high water demand of porous RHA particles having high specific surface area. This resulted in an increase in the porosity with the formation of large pores (>0.1 pm) Figure 6). Increase in the volume of small pores (<0.1 pm) was also observed due to the additional C-S-H phases formed by the pozzolanic reaction. This can be verified with SEM micrographs as the matrix of the MA-cement mortars was mostly composed of well-connected flocs-like C-S-H phases with small pores in-between while that of the reference cement mortar was characterized by large pores containing needle-like and reticular C-S-H phases (Figure 9 and Figure 10). FW4-cement-lime mortars had relatively lower compressive and flexural strength values than those of the RHA-cement mortars. Strength reduction was recorded at the very early stage of 28 days (Figure 4 and Figure 5). This reduction was also reported in the literature for the same type of mortar composition after 28 weeks [13] and for ternary grouts containing

598

Ozlem CIZER, Jan CAMPFORTS, Koen Van BALEN Jan ELSEN, Dionys Van GEMERT

10%-wt cement after 60 days or more [14-151. As the cement content was relatively low (10%-wt), its hydration would be at the very early stage together with the pozzolanic reaction but carbonation would be expected to be much more effective at later stage. This could be verified with the thermal analysis results revealing that hydration and pozzolanic reaction were effective until 28 days when strength development was achieved. Carbonation started to be much more pronounced after 28 days when strength reduction was recorded. This could indicate that calcium carbonate phases formed afterwards could have a destructive effect on the initially hydrated cement phases. As pozzolanic reaction leading to additional C-S-H phases proceeds slowly, its contribution would be insufficient to the initially hydrated cement phases to overcome the effect of the calcium carbonate formation. In addition, partial carbonation of the C-S-H phases was also recorded by thermal analysis after 60 days with the slight decrease in the C-S-H phases and increase in the carbonated phases. The overaIl effect of this would be to weaken rather than to strengthen the existing C-S-H phases. This could also be the reason why RHA-cement-lime mortars exhibited a reduction in their early-stage strength. In agreement with the earlier research, it can be concluded that 10%-wt cement was insufficient for the long-term strength development of these ternary blended mortars [ 14-151. Carbonation depth of the blended cement mortars at any age was greater than that of the reference cement mortar. A correlation between carbonation depth and porosity of the mortars could be drawn since the carbonation depth increased with increasing porosity which contributes to the diffision of the carbon dioxide into the core where calcium hydroxide is still present to carbonate [16]. Carbonation depth of the mortars is in agreement with the thermal analysis results revealing the calcium hydroxide content and carbonated phases. Microstructural investigation by SEM revealed considerable differences among the reference cement mortar, RHA-cement and RHA-cement-lime mortars. While needle-like and reticular C-S-H phases were observed in the matrix of the reference cement mortar (Figure 9), flocs-like and fibrous-like C-S-H phases well connected to each other mostly composed the matrix of the RHA-cement mortars (Figure 10). Unreacted or partially reacted RHA grains could be detected in the matrix of the RHA-cement mortars even at the age of 120 days, indicating the ongoing pozzolanic reaction (Figure 1 la-b). The mechanism of the pozzolanic reaction could be observed as well. The reaction started on the surface of the RHA grains, leading to the precipitation of C-S-H phases. When totally reacted, they became flocs-like C-S-H phases playing a bridging role between hydrated and unhydrated cement particles by growing from the grain surface towards the surrounding matrix (Figure llc-d). This could help improving the microstructure of the interfacial transition zone between sand and the cement matrix. Different from the RHA-cement mortars, RHA-cement-lime mortars were characterized by rather poor microstructure composed of silica-rich carbonated phases that were agglomerated with rather large pores in-between. Clusters of hexagonal and needle-like crystals were embedded locally within the matrix. Lower strength values of these mortars could be related with their poor microstructure. CONCLUSIONS

Hardening of cement in combination with RHA and lime occurred as a result of combined cement hydration, pozzolanic reaction and carbonation. Initial strength development of the RHA-cement mortars was mostly governed by the cement hydration while pozzolanic reaction contributed to the long-term strength development. Carbonation was favoured at the later stage with the increase in the porosity of the mortars due to high WE3 ratio. High porosity led to relatively lower strength values when compared to the reference cement mortar. Compressive and flexural strength of each RHA-cement mortar increased gradually,

Hardening of calcium hydroxide and calcium silicate binders due to carbonation and hydration

5 99

indicating their long-term strength development. However, strength reduction was recorded at the very early stage of the mortars when cement in low content was used in combination with RHA and lime. This could be explained with the insufficient cement content, destructive effect of the calcium carbonate phases on the initially hydrated cement compounds and partial carbonation of the hydrated phases. The lowest cement content leading to long-term strength development for the ternary blended mortars will be further investigated. ACKNOWLEDGEMENT This study is the part of an ongoing Research Project (OT-project 3E030765) conducted at the Department and Master thesis research ‘Performance Evaluation of Pozzolanic Rice Husk Ash Binder in Tanzania’ by Jan Campforts at the K.U.Leuven. The research grant offered to Ozlem Cizer by the OT-project funded by K.U.Leuven is gratefully acknowledged. REFERENCES 1. Massazza, F., Pozzolana and pozzolanic cements. In: Lea‘s Chemistry of Cement and Concrete, Hewlett P.C. ed., 4th Edn., Arnold, London, 1998, pp 471-631. 2. Taylor, H.F.W., Cement Chemistry. Academic Press Limited, London, 1990. 3. Ogawa K., Uchikawa H., Takemoto K., The mechanism of the hydration in the system C3S-pozzolana. Cem. Concr. Res., 10, 1980, pp 683-696. 4. Wu Z.Q., Young J.F., The hydration of tricalcium silicate in the presence of colloidal silica. Journal of Materials Science, 19, 1984, pp 3477-3486. 5. Sharara, A., Didamony, H., Ebied, E., El-Aleem, A., Hydration characteristics of p-C2S in the presence of some pozzolanic materials. Cem. Concr. Res., 24, 1994, pp 966-974. 6. Collepardi M., Baldini G., Pauri M., Corradi M., Tricalcium aluminate hydration in the presence of lime, gypsum or sodium sulphate. Cem. Concr. Res., 8, 1978, pp 571-580. 7. Mehta P.K., Studies on blended Portland cements containing Santorin earth. Cem. Concr. Res., 11, 1981, pp 507-518. 8. Mehta, P.K. and Gjorv, O.E., Properties of Portland cement concrete containing fly ash and condensed silica fume. Cem. Concr. Res., 12, 1982, pp 587-595. 9. Chengzhi, Z., Aiqin, W., Mingshu, T., The filling role of pozzolanic material. Cem. Concr. Res., 26, 1996, pp 943-947. 10. Shinohara, Y., and Kohyama, N., Quantitative analysis of tridymite and cristobalite crystallized in rice husk ash by heating. Industrial Health, 42,2004, pp 277-285. 11. EN 196-1, Methods of testing cement-Part 1: Determination of strength. European Standard, 1987. 12. EN 1936, Natural stone test method - Determination of real density and apparent density and of total porosity and open porosity. European Standard, 1999. 13. Stroeven, P., Bui, D. D., Sabuni, E., Ash of vegetable waste used for economic production of low to high strength hydraulic binders. Fuel, 78, 1999, pp 153-159. 14. Toumbakari E., Lime-Pozzolan-Cement grouts and their structural effects on composite masonry walls. Ph.D Dissertation, K.U.Leuven, Belgium, 2002. 15. Van Rickstal, F., Toumbakari, E., Ignoul, S., Van Gemert, D. Development of Mineral Grouts for Consolidation. In Consolidation of Masonry, Ed. D. Van Gemert, In: Advances in Materials Science and Restoration, no. 1. Freiburg: Aedification Publishers, 2003, pp 61-76. 16. Van Balen, K. and Van Gemert, D., Modeling Lime Mortar Carbonation. Materials and Structures, 27, 1994, pp 393-98.

60 1

INDEX OF CONTRIBUTORS

AKITA Hiroshi, Japan 187, 195,205

FAIRBAIRN Eduardo de M.R., Brazil 25 1

ABRAMOWICZ Marian, Poland 285

FERN~DEZ-JIMENEZh a , Spain 55

ADOLPHS Jiirgen, Germany 35

FINN R, Germany 73

AICHER S., Germany 73

FORMAGINI Sidiclei, Brazil 25 1

ALTERMAN Dariusz, Japan 195

FULLEA J, Spain 123

ANDRADE Carmen, Spain 123 GAL16 Jure, Croatia 323 BANJAD-PECUR Ivana, Croatia 323

GARBACZ Andrzej, Poland 303

BASNAL Saurabh, USA 85

GNYP Olga P., Ukraine 517

BIER Thomas A,, Germany 149, 175

GOLEWSKI Grzegorz, Poland 537

BOHM Robert, Germany 487

GOLASZEWSKI Jacek, Poland 441

BROUKALOVA h a , Czech Republic 343

GONCHAR Olga, Ukraine 399

BUTNARIU Efrat, Israel 293

GOTS Vladimir, Ukraine 399

CAMPFORTS Jan, Belgium 589

HANZLOVA Hana, Czech Republic 2 13

CHEN Huisu, The Netherlands 25

HU Jing, France 42 1,57 1

CHUDOBA Rostislav, Czech Republic 36 1

HUFENBACH Werner, Germany 487

CIZER Ozlem, Belgium 589 COPUROCLU O w h a n , The Netherlands 387

JESSE Frank, Germany 275

C W P E R S Heidi, Belgium 99 KASPERKIEWICZ Janusz, Poland 195 DENIS Christophe, France 495

KERSNER Zbyngk, Czech Republic 527,581

DESHPANDE Yogini, USA 131

KIM Haejin, USA 161

DICK-NIELSEN Lars, Denmark 221

KLEMM Agnieszka J., UK 45

DOMBROWSKI Katja, Germany 149

KLEMM Piotr, Poland 45

DOTSENKO Juliya V., Ukraine 517

KLEPACZKO Janusz R., France 547

DOVGAN Alexandra, Ukraine 459

KNAPEN Elke, Belgium 66 KOHOUTKOVA Alena, Czech Republic 343

ELSEN Jan, Belgium 589

KOICHEV Aleksandr A., Ukraine 5 17 KOIDE Hideo, Japan 187

602 KOLEVA Dessi A., The Netherlands 571

PODAGELIS Igor, Lithuania 459

KONRAD Martin, Germany 361

PODROUAEK Jan, Czech Republic 205

KOTYNIA Renata, Poland 109

PONIKIEWSKI Tomasz, Poland 45 1

KOVALCHUK Georgiy, Ukraine 55

POSTEK Eligiusz, UK 495

KlUVENKO Pave1 V., Ukraine 55,467

POULSEN Peter Noe, Denmark 221

KUDER Katherine G., USA 43 1

PUSHKARYOVA Ekaterina, Ukraine 399

KUMAGAI Shinsuke, Japan 3 15 KWASNIEWSKI Leslaw, Poland 303

RADLI&KA Aleksandra, USA 33 1 RADLIfJSKI Mateusz, USA 161

LANGKAMF' Albert, Germany 487

RANGARAJU Prasada Rao, USA 373

LI Victor C, USA 263

REINHARDT Hans Wolf, Germany 73

LUTSKM Yevgen S., Ukraine 517

RIZWAN Syed Ali, Germany 149,175

LYASHENKO Tatiana, Ukraine 459 SADOWSKI Tomasz, Poland 495,507,537 MATESOVA Dita, Czech Republic 58 1

SAMBORSKI Sylwester, Poland 507

MINDESS Sidney, Canada 15,239

SANCHEZ Javier, Spain 123

MOBASHER Barzin, USA 85,293

SANJEEVAN Poologanathan, UK 45

MOKHORT Mikola A, Ukraine 467

SCHLANGEN Erik, The Netherlands 387

MOZGAWA Wlodzimierz,Poland 477

SHAH Surendra P., USA 43 1

MU Edward, USA 43 1

SHINKEVITCH Elena S., Ukraine 517 SILVA Flavio de A., Brazil 251

NANTUNG Tommy, USA 16 1

SEOWIK Marta, Poland 351

NIZAMI Muhammad Sharif, Pakistan 175

SOMPURA Ketan, USA 373

NOVAK Drahomir, Czech Republic 205

SORANAKOM Chote, USA 85 SPfJRA Dukan, Czech Republic 285

OCHI Mitsukazu, Japan 3 15

STANG Henrik, Denmark 221

OHAMA Yoshihiko, Japan 315

STROEVEN Piet, The Netherlands 25,421,571

OJIMA Mitsuo, Japan 195

STROEVEN Martijn, The Netherlands 25

OLEK Jan, USA 131, 161,373

SZERSZEh Maria M., USA 263

OLESEN John Forbes, Denmark 221

SZWED Aleksander, Poland 263

OTA Masahiro, Japan 3 15 OZAKA Yoshio, Japan 187

TOLEDO FILHO Romildo D., Brazil 25 1

OZYURT Nilufer. USA 43 1

TOUMI Ahmed, France 409 TRAN Quoc-Thanh, France 409

PALOMO Angel, Spain 55

TURATSINZE Anaclet, France 409

PELED Alva, Israel 293 PICH6R Waldemar, Poland 477

VAN BALEN Koen, Belgium 589

603

VAN GEMERT Dionys, Belgium 65,589 VAN ITTERBEECK Petra, Belgium 99 VEGT Ilse, The Netherlands 559 VESELY Vklav, Czech Republic 527 VODICKA Jan, Czech Republic 213,285 VOhCHOVSKY Miroslav, Czech Republic 361 VOZNESENSKY Vitaly, Ukraine 459

*BORNY

Jaroslav, Czech Republic 213

WASTIELS Jan, Belgium 99 WEERHEIJM Jaap, The Netherlands 559 WEISS Jason, USA 331 ZANDER Anthony, USA 161 ZHANG Lihe, Canada 239

BRITTLE MATRIX COMPOSITES

8

Proceedings of the Eighth International Symposium on Brittle Matrix Composites BMC8, held in Staszic Palace, Warsaw,Poland, 23-25 October 2006

Also published previously: Brittle Matrix Composites 1 Proceedings of the 1" International Symposium and EUROMECH Colloquium 204, Elsevier Science Publishers, November 12-15, 1985 Brittle Matrix Composites 2 Proceedings of the 2ndInternational Symposium, Elsevier Science Publishers, September 20-22, 1988 Brittle Matrix Composites 3 Proceedings of the 3rdInternational Symposium, Elsevier Science Publishers, September 17-19, 1991 Brittle Matrix Composites 4 Proceedings of the 4'hInternational Symposium, Woodhead Publ. Ltd. (Cambridge) and Inst. of Economic Education (Warsaw), September 13- 15, 1994 Brittle Matrix Composites 5 Proceedings of the 5* International Symposium, Woodhead Publ. Ltd. (Cambridge) and BIGRAF (Warsaw) October 13-15, 1997 Brittle Matrix Composites 6 Proceedings of the 6* International Symposium, Woodhead Publ. Ltd. (Cambridge) and ZTUREK Research.-Scientific Institute (Warsaw), October 9- 1 1,2000 Brittle Matrix Composites 7 Proceedings of the 7" International Symposium, Woodhead Publ. Ltd. (Cambridge) and ZTUREK Research.-Scientific Institute (Warsaw), October 13-15,2003

INSTITUTE OF FUNDAMENTAL TECHNOLOGICAL RESEARCH POLISH ACADEMY OF SCIENCES

BRITTLE MATRIX COMPOSITES 8 Edited by

A.M. BRANDT Institute of Fundamental Technological Research Polish Academy of Sciences, Warsaw,Poland

V.C. LI Department of Civil and Environmental Engineering University of Michigan, Ann Arbor, USA

I.H. MARSHALL Monash University, Melbourne, Australia

WOODHEAD PUBLISHING LIMITED ZTUREK RESEARCH-SCIENTIFIC INSTITUTE CAMBRIDGE and WARSAW 2006

0 2006 Institute of Fundamental Technological Research, Warsaw

WOODHEAD PUBLISHTNG LIMITED Abington Hall, Abington, Cambridge, England

ZTUREK RESEARCH SCIENTIFIC INSTITUTE Warsaw, Poland ISBN 1-84569-03 1-1 ISBN 83-89687-09-7 Conditions of sale: All rights reserved. No part of this publication may be reproduced or transmitted in any form or by any means, electronic or mechanical, including photocopying, recording, or any information storage and retrieval system, without permission in writing from the publisher.

*

British Library Cataloguing in Publication Data. A catalogue record for this book is available from the British Library.

The selection and presentation of material and the opinion expressed in this publication are the sole responsibility of the authors concerned. No responsibility is assumed by the Publishers for any injury andor damage to persons or property as a matter of products liability, negligence or otherwise, or fiom any use or operation of any method, products, instructions or ideas contained in the material herein.

Produced by:

ZTUREK Research-Scientific Institute 00-950 Warsaw 1, P.O. Box 674 http://www.zturek.pl, [email protected]

V

Preface The collection of papers presented in this volume of BMC is the eighth in a series of triennial international symposium focusing on Brittle Matrix Composites. It was held October 23-25, 2006 in Warsaw, Poland. The objective of the Brittle Matrix Composite Symposia is to bring together researchers in the broad field of composite materials with brittle cement, ceramic and polymer matrix, to share their latest research findings. While their application targets can be quite different, these composites with different matrices share many similar characteristics. Some examples illustrating the positive impact of one community on another include the mechanics of fibedmatrix debonding, fiber bridging behavior, application of non-conventional components, and multiple cracking processes. The BMC Symposia provide a forum aimed at encouraging and enhancing crossdisciplinary knowledge exchange. It is also the main conference in composites where researchers from Eastern and Western Europe, and many other countries outside of Europe, gather once every three years to update developments in composite research and development in various countries. The series of BMC Symposia was initiated by the highly successful EUROMECH 204 Colloquium on Brittle Matrix Composites held in Poland, in Jablonna in November 1985, renamed BMC 1. Three years later (September 1988) the Second Symposium BMC 2 was organized in Cedzyna, also in Poland. This was followed by the Third International Symposium on Brittle Matrix Composites BMC 3 in Warsaw, September 1991. BMC 4, 5, 6 and 7 were also held in Warsaw in 1994, 1997, 2000 and 2003, respectively. In each case, the Institute of Fundamental Technological Research (IFTR) of the Polish Academy of Sciences served as the host. From the early days of BMC 1 in 1985 to the present BMC 8, these Symposia have served both as effective forums for dissemination of knowledge and a catalyst for innovation. They have undoubtedly contributed to international collaboration both scientifically and personally. With regard to the latter, there can be few countries in the world without knowledge of the BMC Symposia! The subject of the Symposia is composite materials with matrices behaving as brittle in normal or special conditions. Brittle matrix composites are applied in various domains (civil engineering, mechanical equipment and machinery, vehicles, etc.). In the last decades their importance and diversity amongst engineered materials has continuously increased. Examples of the materials covered in the accepted papers principally include: aggregate-binder composites (concretes, fiber concretes, polymer concretes), sintered materials (ceramics), and other composites with brittle matrices. Various approaches to the material engineering problems are presented in the papers. Amongst others: mechanical properties, strength and toughness, rheology, analysis of materials structure and microstructure,

VI

various degradation effects, crack propagation and control, test methods and new test results, computation methods and manufacturing processes, durability assessment of materials and structures, applications of new materials and their behavior in structures. The present volume is the proceedings of BMCS. The papers have been selected on the basis of two-stage peer reviews. The volume contains 55 papers prepared by 129 authors from 19 countries. Collectively, this set of papers represent the latest advancements in the field of Brittle Matrix Composites. This volume should prove useful to experienced researchers and to students entering this field as a helpful reference. Particular thanks are due to the authors of the articles in this volume, for their readiness to present the results of their outstanding investigations in the form of original contributions. Grateful thanks are extended to members of the International Advisoty Board for their significant help in reviewing the papers and in numerous other problems encountered during the organizational stages. The tireless efforts and creative attitude of local Organizing Committee during preparations for this event is highly appreciated. Without their dedicated work it would not have been possible to publish this volume before the Symposium date. The support from the Institute of Fundamental Technological Research of the Polish Academy of Sciences was essential for organization of this Symposium as well as those in past years. The scientific sponsorship of RILEM is acknowledged with gratitude. We sincerely hope that this volume, along with the seven previously published in this series, will contribute to the development of science and technology in the field of composite materials.

A.M. Brandt V.C. Li I.H. Marshall Warsaw, Ann Arbor, Melbourne, July 2006

VII

SYMPOSIUM CO-CHAIRMEN: Prof. A.M. Brandt, Warsaw - Poland Prof. V.C. Li, Ann Arbor - USA Prof. I.H. Marshall, Melbourne - Australia INTERNATIONAL, ADVISORY BOARD: Prof. B. Barr, Cardiff - UK Prof. A. Bentur, Haifa - Israel Prof. J. Kasperkiewicz, Warsaw - Poland Prof. K. Kromp, Vienna - Austria Prof. S. Mindess, Vancouver - Canada Prof. Z. Mrbz, Warsaw - Poland Prof. Y. Ohama, Koriyama - Japan Prof. J. Olek, Lafayette - USA Prof. H.W. Reinhardt, Stuttgart - Germany Prof. K. Scrivener, Lausanne - Switzerland Prof. S.P. Shah, Evanston -USA Prof. H. Stang, Lyngby - Denmark Prof. P. Stroeven, Delft - The Netherlands Prof. R.N. Swamy, Sheffield - UK Prof. A. Vautrin, St. Etienne - France ORGANIZING COMMITTEE: Prof. J. Kasperkiewicz Prof. M.A. Glinicki Mrs. A. Gutweter Dr. D. J6iwiak-Niediwiedzka Prof. M. Marks Mr. M. Sobczak Mrs. J. Tymkiewicz

IX

Table of Contents Preface ...............................................................................................................................................

5

PLENARY INVITED PAPER High performance concrete: where do we go from here? Sidney MINDESS, Canada ..................................................................................................

15

CONTRIBUTIONS On connectivity of porosity in model cement paste Piet STROEVEN, Huisu CHEN, Martijn STROEVEN, The Netherlands ...........................

25

Moisture dependence of pore size and specific surface area of hardened cement paste determined with S A X S and inverse gas chromatography Jiirgen ADOLPHS, Germany ...............................................................................................

35

The effects of microstructural features of mortars on the laser cleaning process Poologanathan SANJEEVAN, Agnieszka J. KLEMM, Piotr KLEMM, UK, Poland

45

..........

Fly ash based geocements: genesis of microstructure and properties at hydration-dehydration process Pave1 KRIVENKO, Georgiy KOVALCHUK, Angel PALOMO, Ana FERNANDEZ-JIMENEZ, Ukraine, Spain ..................................................................

55

Influence of water-soluble polymers on the microstructure of cement mortars E k e KNAPEN, Dionys Van GEMERT, Belgium ...............................................................

65

Static and dynamic response of cellulose fiber gypsum-board wall elements Hans Wolf REINHARDT, R. FINN, S. AICHER, Germany ..............................................

73

Effect of material non-linearity on the flexural response of fiber reinforced concrete Chote SORANAKOM, Barzin MOBASHER, Saurabh BANSAL, USA .............................

85

The effect of durability on the design of self-bearing sandwich panels with cementitious composite faces Heidi CWPERS, Petra Van ITTERBEECK, Jan WASTIELS, Belgium ...........................

99

Debonding phenomena in FRP - strengthened concrete members Renata KOTYNIA. Poland ...................................................................................................

109

Fracture toughness variation of prestressing steels by bicarbonate solutions Javier S h C H E Z , Jose FULLEA, Carmen ANDRADE, Spain ..........................................

123

X

Sensitivity of rapid-setting self-consolidating concrete (RSSCC) to mixture production variables Yogini DESHPANDE, Jan OLEK, USA ...............................................................................

131

A discussion on the essential issues for successful production of self-compacting concrete (SCC) Syed Ali RIZWAN, Thomas A. BIER, Katja DOMBROWSKI, Germany ..........................

149

Preliminary optimization analysis of ternary mixtures for bridge decks Mateusz RADLINSKI, Jan OLEK, Haejin KIM, Tommy NANTUNG, Anthony ZANDER, USA ......................................................................................................

161

High performance self-compacting mortars containing pozzolanic powders Syed Ali FUZWAN, Thomas A. BIER, Muhammad Sharif NIZAMI, Germany,Pakistan ................................................................

175

Discussion on secondary flexure in uniaxial tension test of concrete Hiroshi AKITA, Hide0 KOIDE, Yoshio OZAKA, Japan ...................................................

187

Identification of uniaxial tension tests of concrete based on machine learning technique Dariusz ALTERMAN, Janusz KASPERKIEWICZ, Hiroshi AKITA, Mitsuo OJIMA, Japan, Poland ............................................................................................

195

Virtual 3D nonlinear simulation of uniaxial tension test of concrete Drahomir NOVAK, Jan PODROUZEK, Hiroshi AKITA, Czech Republic, Japan ..............................................................................

205

Comparison of basic mechanical-physical properties and frost resistance of common fine-grained concrete and brickconcryte with fibers and without fibers Hana HANZLOVA, Jaroslav V b O R N Y , Jan VODICKA, Czech Republic ......................

213

PLENARY INVITED PAPER Recent developments in the modeling of matrix crack propagation in Brittle Matrix Composites Henrik STANG, Lars DICK-NIELSEN, Peter NOE POULSEN, John FORBES OLESEN, Denmark ......................................................................................

221

CONTRIBUTIONS Compressive toughness of fibre reinforced concrete under impact loading Lihe ZHANG, Sidney MMDESS, Canada ..........................................................................

239

Behavior under compression and bending loads of multi-scale high performance steel fiber reinforced cement based composites Flhvio de A. SILVA, Sidiclei F O W G I N I , Romildo D. TOLEDO FILHO, Eduardo de M. R. FAIRBAIRN, Brazil ................................................................................

251

XI

Flexural response of reinforced beam with high ductility concrete material Maria M. SZERSZEN, Aleksander SZWED, Victor C.LI, USA, Poland .............................

263

Efficiency of multi filament reinforcement in cementitious composites Frank JESSE, Germany ........................................................................................................

275

Volume Changes of Fibre Concret: with Steel and Synthetic Fibres Jan VODICKA, DuSan SPURA, Marian ABRAMOWICZ, Czech Republic, Poland .............................................................

285

Impact behavior of fabric-cement based composites Efrat BUTNARIU, Alva PELED, Barzin MOBASHER, Israel, USA ..................................

293

Modeling of stress wave propagation in repair systems tested with impact-echo method Andrzej GARBACZ, Leslaw KWASNIEWSKI, Poland ....................................................

303

Strength development and epoxy resin-cement interaction in hardener-free epoxy-modified mortars Yoshihiko OHAMA, Mitsukazu OCHI, Shinsuke KUMAGAI, Masahiro OTA ................. 315 Non-destructive testing of polymer modified concrete Jure GALIC, Ivana BANJAD PECUR, Croatia ...................................................................

323

Quantifying variability in assessing the risk of early-age cracking in restrained concrete elements Aleksandra RADLdJSKA, Jason WEISS, USA ...................................................................

33 1

Structural performance and crack control of fibre concrete beams with conventional reinforcement Alena KOHOUTKOVA, Iva BROUKALOVA, Czech Republic .........................................

343

The analysis of crack formation in concrete and slightly reinforced concrete member in bending Marta SLOWIK, Poland ......................................................................................................

35 1

Multiple cracks bridg:d by multifilament yarns: impact of local scatter on ultimate load Miroslav VORECHOVSKY, Martin KONRAD, Rostislav CHUDOBA, Czech Republic, Germany ...............................................................

361

An investigation into deicer-induced ASR distress in concrete

Prasada Rao RANGARAJU, Ketan SOMPURA, Jan OLEK, USA ......................................

373

Experimental and numerical study of frost salt scaling of concrete Oguzhan COPUROGLU, Erik SCHLANGEN, The Netherlands ........................................

387

XI1

Stability of hydrosulfoaluminosilicate compounds and durability of concrete based on these compounds Ekaterina PUSHKAROVA, Vladimir GOTS, Olga GONCHAR, Ukraine ..........................

399

Durability of an overlay-old concrete interface: the role of a metal fibre reinforcement Quoc-Thanh TRAN, Ahmed T O W , Anaclet TURATSINZE, France .............................

409

PLENARY INVITED PAPER On science aspects of simulating practical manifestations of concrete Piet STROEVEN, Jing HU, The Netherlands .......................................................................

42 1

CONTRIBUTIONS Rheology of fiber-reinforced cement systems using a custom-built rheometer Katherine KUDER, Nilufer OZYURT, Edward Mu, Surendra SHAH, USA ......................

43 1

The influence of cement paste volume in mortar on the rheological effects of the addition of superplasticizer Jacek GOLASZEWSKI,Poland ...........................................................................................

441

The rheological properties of fiesh steel fibre reinforced self-compacting concrete Tomasz PONIKIEWSKI, Poland .........................................................................................

45 1

Correlation between bending resistance of epoxy composite specimens maintained in water and in petroleum Tatiana LYASHENKO, Vitaly VOZNESENSKY, Alexandra DOVGAN, Igor PODAGELIS, Ukraine, Lithuania ................................................................................

459

Inorganic geocements matrix in composite materials Pave1 V. KRIVENKO, Mykola A. MOKHORT, Ukraine ....................................................

467

Properties of clinoptilolite based autoclaved composites Wlodzimierz MOZGAWA, Waldemar PICHbR, Poland

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

477

Advanced models for the simulation of anisotropic damage in fibre and textile reinforced ceramics Werner HUFENBACH, Robert BbHM, Albert LANGKAMF', Germany ...........................

487

Modelling of metallic inter-granular layers in polycrystalline ceramics Eligiusz POSTEK, Tomasz SADOWSKI, Christophe DENIS, UK, Poland, France

495

.........

A new micromechanics based predictive method for porous ceramics behavior under compression Sylwester SAMBORSKI, Tomasz SADOWSKI, Poland ....................................................

507

The influence of modification of the structure of silicate materials on their properties after non-autoclaved hardening Elena S. SHINKEVITCH, Yevgen S. LUTSKIN, Olga P. GNYP, Aleksandr A. KOICHEV, Juliya V. DOTSENKO, Ukraine .................................................

517

R-curves from equivalent elastic crack approach: effect of structural geometry on fracture behavior Vaclav VESELY, ZbynEk KERSNER, Czech Republic .......................................................

527

Fracture toughness at shear (Mode 11) of concretes made of natural and broken aggregates Grzegorz GOLEWSKI, Tomasz SADOWSKI, Poland .......................................................

537

On fracture energy of concrete for short-time loading in tension Janusz R. KLEPACZKO, France .........................................................................................

547

Dynamic response of concrete at high loading rates. A new Hopkinson bar device. Ilse VEGT, Jaap WEERHEIJM, The Netherlands ...............................................................

559

Microstructure alterations underlying electrochemical process of chloride-induced corrosion Dessi A. KOLEVA, Jing HU, Piet STROEVEN, The Netherlands ......................................

571

Effect of porosity and fracture toughness on explosive spalling of concrete Dita MATESOVA, ZbynEk KERSNER, Czech Republic ....................................................

581

Hardening of calcium hydroxide and calcium silicate binders due to carbonation and hydration Ozlem CIZER, Jan CAMPFORTS, Koen Van BALEN, Jan ELSEN, Dionys Van GEMERT, Belgium ..........................................................................................

589

Index of Contributors ......................................................................................................................

601

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