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?
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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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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
5
E
I
1 . 5 *
0
0
.? .
G
-Y
-- 20
4
- s - l o k d 0
" . . ; . " ' ; . " . ; . . . . ; ' " . ; ' . . ' r
I
2
3
4
5
5
sum r(%)
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 (=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.
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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 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,' 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
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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
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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.
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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.
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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
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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.
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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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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
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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.
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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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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
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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
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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
Syed Ali RIZWAN, Thomas A. BIER, Katja DOMBROWSKI
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
162
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 2.711 [Gf 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
204
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
206
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
216
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.
Comparison of basic mechanical-physical properties andpost resistance of commonfine-grained ...
2 19
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
223
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 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].
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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
24 1
Compressive toughness offibre reinforced concrete under impact loading
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m I
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Machine Columns
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-
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,
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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
,1-:::::__’r Uppe ( P l , P2, P3, P4, P5)
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-
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Fig. 2 Two sets of 5 tracking points, 100 mm apart in the vertical direction
0.9
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1
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g
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t
-
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Dx-Lowr Dx-Upper
v
z
co
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Fig. 3. Deformation vs. time history under impact
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3
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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0.5 0
0
0.2
0.6
0.4
0.8
1
Time (ms)
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
a
300 200 100
0 0
0.5
1
1.5 Time (ms)
2
2.5
3
Fig. 6 Load vs. time curve for FRC cylinder under impact Load vs. deformation By combining the results shown in Figs. 3 and 6 , a load vs. deformation curve can be obtained, as shown in Fig. 7. 700
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
0.2580
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.
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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
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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
c*l
Left over FRC cylinder Fig. 10 Sketch of FRC failure under impact load
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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
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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:
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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. ..
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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
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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
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3
4
5
6
7
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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
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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
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Fig. 1. Typical experimental stress-strain curves for ECC in uniaxial direct tension test. I
smin
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0000
0005
0010
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0020
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0030
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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