Effects of Redispersible Polymer Powder on Mechanical and Durability Properties of Preplaced Aggregate Concrete with Recycled Railway Ballast - International Journal of Concrete Structures and Materia

4.1

Compressive Strength and Modulus of Elasticity

The test results of compressive strength and elastic modulus for all mortar and concrete samples are summarized in Table 5; at least three specimens were tested for each sample.

The compressive strength tests at 2, 3 or 4 h of curing were conducted right after the surfaces of the specimens solidified. The initial setting time was delayed as the amount of polymer increased under the condition that the same W/C ratio and retarders were used in all mix samples. Thus, the specimens containing 0% or 2% polymer were tested at about 2 h of curing, the specimens with 6% polymer were tested at about 3 h of curing, and the specimens with 10% polymer were tested at about 4 h of curing. This delay in the setting time was likely attributed to the films the polymer particles produced around cement grains, as these films would inhibit the reaction between cement and water (i.e., cement hydration) (Su 1995; Wang et al. 2006, 2016).

The compressive stress–strain relationships of all twelve concrete samples (Table 4) acquired at 28 days of curing are presented in Figs. 13, 14 and 15. Note that there are only two plots in Fig. 15a, since one specimen of the sample (C-L-0) improperly failed during the test. The compressive strengths of all mortar and concrete samples at 28 days of curing are compared in Fig. 16. For a given polymer ratio, the mortar samples had a 50–175% higher compressive strength on average than the concrete samples. This implies that the compressive strength of PAC with recycled ballast aggregates was greatly affected by the ITZs between the mortar and aggregates (Najjar et al. 2014).

Compressive stress–strain curves with H-level aggregates: a C-H-0, b C-H-2, c C-H-6, and d C-H-10.

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Compressive stress–strain curves with M-level aggregates: a C-M-0, b C-M-2, c C-M-6, and d C-M-10.

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Compressive stress–strain curves with L-level aggregates: a C-L-0, b C-L-2, c C-L-6, and d C-L-10.

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Compressive strengths of all mortar and concrete samples at 28 days of curing.

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In both mortar and concrete cases, the compressive strength generally decreased as the polymer ratio increased (Fig. 16). With regard to the mortar samples, C-2, C-6, and C-10 had approximately 5.4, 8.2, and 24% lower compressive strengths than C-0, respectively. The reduction in strength was moderate up to a polymer ratio of 6%, but it was substantial at a ratio of 10%. For the concrete samples with H-level aggregates, C-H-2, C-H-6, and C-H-10 showed a compressive strength approximately 8.8, 16, and 23% lower than C-H-0. With a similar trend, C-M-10 and C-L-10 presented compressive strengths approximately 23 and 33% lower than C-M-0 and C-L-0, respectively. A lower compressive strength from a higher polymer ratio was likely attributed to the following: (1) the polymer particles hindered the cement from reacting with water, so that the strength development was delayed (Su et al. 1991; Su 1995; Jun et al. 2003). (2) Due to the low strength of the polymer itself, the polymer films had a similar effect to pores in the hardened concrete (Sakai and Sugita 1995).

Figure 17 plots the moduli of elasticity of all concrete samples, determined by ACI Committee 318 (2014) (see Table 5). The modulus of elasticity presented a tendency to decrease with increasing polymer ratio; the moduli of C-H-2, C-H-6, and C-H-10 were roughly 4.4, 8.5, and 12.2% smaller than that of C-H-0. Similarly, the moduli of C-M-10 and C-L-10 were 12.5 and 32.5% smaller than those of C-M-0 and C-L-0, respectively. This suggests that an increase in the polymer ratio generally caused a reduction in both the compressive strength and elastic modulus of PAC made from recycled ballast aggregates.

Moduli of elasticity of all concrete samples at 28 days of curing.

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Regarding the cleanness level of ballast aggregates, the specimens with H-level aggregates generally showed higher compressive strengths than those with L- and M-level aggregates for a given polymer ratio (Fig. 16). The trend of a higher strength with H-level aggregates was particularly distinct at 10% polymer ratio; the compressive strength of C-H-10 was roughly 18% and 26% higher than those of C-M-10 and C-L-10, respectively. A possible reason for this result is that denser ITZs between the mortar and aggregates were created using H-level aggregates with smaller amounts of dusty particles. However, the effect of the aggregate cleanness level was not apparent for polymer ratios of 2 and 6%.

4.2

Modulus of Rupture

Figure 18 compares the moduli of rupture (i.e., tensile strengths) of all concrete samples listed in Table 4. With regard to the effect of the polymer, the specimens with a polymer ratio of 10% generally displayed the largest tensile strengths; the moduli of rupture of C-H-10 and C-L-10 were about 13.1 and 5.4% larger than those of C-H-0 and C-L-0, respectively.

Tensile strengths and tensile-to-compressive strength (M/C) ratios.

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In addition, the tensile-to-compressive strength (M/C) ratios are presented in Fig. 16; the ratios ranged from approximately 0.75–0.14. Although the samples containing 2 and 6% polymer had slightly smaller moduli of rupture than those containing no polymer, their M/C ratios were similar to or larger than those with no polymer. Furthermore, the M/C ratio increased sharply at 10% polymer ratio (except for M-level aggregate concrete). Therefore, it can be concluded that the use of polymer enhanced the ITZs between the ballast aggregates and mortar, and consequently improved the tensile strength.

As regards the effect of the cleanness level of ballast aggregates, the concrete samples of H-level aggregates showed higher tensile strengths than those of M- or L-level aggregates (Fig. 18). The modulus of rupture of C-H-10 was 24.3 and 11.3% higher than that of C-M-10 and C-L-10, respectively. However, the M-level aggregate concrete samples had rather smaller moduli of rupture than the L-level aggregate concrete samples; i.e., the moduli of rupture of C-M-10 and C-M-0 were approximately 14.6 and 6.0% smaller than those of C-L-10 and C-L-0, respectively. This suggests that partial aggregate cleaning might cause a reduction in the tensile strength of PAC with recycled ballast aggregates.

4.3

Shrinkage Strain

The results of the shrinkage tests for the concrete samples are presented in Fig. 19. Overall, the shrinkage strain rapidly increased at the early curing stage, but after about 15–20 days, the change in shrinkage strain became insignificant.

Shrinkage strains of concrete samples with 2, 6, or 10% polymer ratio.

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The concrete samples with 2% polymer ratio exhibited a significant increase in shrinkage strain during a few hours immediately after the beginning of the strain measurement (Fig. 19). In contrast, early expansion, signified by positive strain values, was observed for the samples with higher polymer ratios. In particular, it was the greatest in the samples with 10% polymer ratio (Fig. 19). The difference in early expansion in relation to the polymer ratio was possibly due to the following reasons: (1) re-emulsified polymer particles inhibited the reaction between cement and water (as discussed for the early compressive strength results), and consequently delayed the shrinkage. (2) At the same time, re-emulsified polymer particles might enlarge in volume to cause the expansion of concrete under the unconstrained condition (Ohama 1995).

The ultimate shrinkage strain of concrete gradually decreased as the polymer ratio increased; for example, C-H-6 and C-L-6 underwent more than 30% smaller peak strains compared with C-H-2 and C-L-2, respectively (Fig. 19). The reduction in the ultimate shrinkage strain with a higher polymer ratio was likely because continuous polymer films occupying micro-pores in the concrete hindered water evaporation (Su 1995; Persson 2001; Zhang and Kong 2014). Furthermore, it was possibly in part because re-emulsified polymer particles might enlarge in volume to offset the shrinkage.

With regard to the effect of the cleanness level of ballast aggregates, C-L-2 and C-L-6 underwent a smaller shrinkage strain than C-H-2 and C-H-6, respectively. This could be possibly because the fine particles attached to ballast aggregates filled the micro-pores of concrete and consequently reduced the shrinkage strain (Li and Yao 2001). However, the effect of the aggregate cleanness on the shrinkage strain of concrete was not as significant as that of the polymer ratio.

4.4

Freezing–Thawing Resistance

The results of the freezing–thawing resistance tests for the concrete samples are presented in Fig. 20. The test results are generally in accordance with the flexure test results. The freezing–thawing resistance of PAC with recycled ballast aggregates was not notably enhanced by the addition of 2 or 6% polymer. However, the concretes with a polymer ratio of 10% (C-H-10, C-M-10, and C-L-10) exhibited substantially higher freezing–thawing resistance. Among the H-level aggregate concretes, C-H-2 firstly reached an RDM value below 60% near the 150th cycle, and C-H-0 and C-H-6 followed later in turn. However, the RDM of C-H-10 remained larger than 80% after 300 cycles (Fig. 20). Similarly, the durability factor of the samples with 10% polymer was roughly 2.7–3.0 times larger than those with no polymer, whereas the factor was not increased by the use of 2 or 6% polymer (Table 6).

Relative dynamic moduli of elasticity (RDMs) of concrete samples with a H-level, b M-level, and c L-level aggregates.

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At a given polymer ratio, the concretes with H-level aggregates showed higher freezing–thawing resistance than those with L- or M-level aggregates (Fig. 20). Moreover, the durability factor of concrete was the highest with H-level aggregates for every polymer ratio (Table 6). This was likely in part because polymer played a more pronounced role in the ITZs of H-level aggregates with fewer alien particles, which would have hindered the formation of polymer films.

The mass of each specimen did not change much in all samples throughout the freezing–thawing tests (Fig. 20); only small mortar parts peeled from the surfaces of the specimens. This implies that the degradation of the freezing–thawing resistance of PAC with recycled ballast aggregates originated mainly from interfacial failures between the aggregates and mortar. The freezing–thawing test results were clearly visualized in the specimen sections that were cut after the completion of testing (Figs. 21 and 22): (1) the concretes with 10% polymer displayed the least damage in the ITZs, regardless of the aggregate cleanness level (Figs. 21d and  22d). (2) The concretes with L-level aggregates (Fig. 22) experienced more extensive damage in the ITZs than those with H-level aggregates (Fig. 21).

Specimen sections with H-level aggregates after freeze–thaw testing: a C-H-0, b C-H-2, c C-H-6, and d C-H-10.

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Specimen sections with L-level aggregates after freeze–thaw testing: a C-L-0, b C-L-2, c C-L-6, and d C-L-10.

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Figures 23 and 24 show the magnified images of damaged ITZs in H- and L-level aggregate concretes, respectively, taken by an optical microscope with a magnification factor of 160 after the freeze–thaw tests. The concretes with low RDM and DM values clearly experienced severe crack damages in the ITZs; the crack width in C-L-2 and C-L-6 was approximately of the order of 0.3–1.4 mm. In contrast, considerable damages were not found in the ITZs of the concretes with the highest RDM and DF values (e.g., C-H-10).

Damaged ITZs with H-level aggregates after freeze–thaw testing: a C-H-0, b C-H-2, c C-H-6, and d C-H-10 [magnification factor: 160].

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Damaged ITZs with L-level aggregates after freeze–thaw testing: a C-L-0, b C-L-2, c C-L-6, and d C-L-10 [magnification factor: 160].

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