Creep model of concrete with recycled coarse and fine aggregates that accounts for creep development trend difference between recycled and natural aggregate concrete

Cement and Concrete Composites 103 (2019) 303–317

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Creep model of concrete with recycled coarse and fine aggregates that accounts for creep development trend difference between recycled and natural aggregate concrete

T

Yue Genga,b, Muzi Zhaoa,c, Hua Yanga,b,∗, Yuyin Wanga,b a

Key Lab of Structures Dynamic Behaviour and Control of the Ministry of Education, Harbin Institute of Technology, Harbin, 150090, China Key Lab of Smart Prevention and Mitigation of Civil Engineering Disasters of the Ministry of Industry and Information Technology, Harbin Institute of Technology, Harbin, 150090, China c School of Civil Engineering, Harbin Institute of Technology, Heilongjang, Harbin, 150090, China b

A R T I C LE I N FO

A B S T R A C T

Keywords: Creep model Recycled aggregate Creep development

A new creep model of recycled aggregate concrete (RAC) incorporated with coarse recycled aggregates (CRA) and fine recycled aggregates (FRA) is proposed in this paper. The model accounts for the difference in the creep development trend between RAC and natural aggregate concrete (NAC). Experiments are performed to investigate how the incorporation of FRA affects the creep behavior of RAC with the variation in the CRA replacement ratio, and how the incorporation of RA influences the creep development trend. The results show that 1) the incorporation of FRA could reduce the negative effect of CRA on the creep of concrete, and 2) the FRA and CRA influence creep development in concrete. Finally, based on experimental results, current creep models for RAC are modified to account for the aforementioned influences.

1. Introduction The use of recycled aggregates obtained by crushing the construction and demolished concrete waste (CDW) can preserve natural resources and reduce the necessity of CDW disposal [1–5]. Over the past decades, the structural use of recycled concrete with coarse recycled aggregates (CRA) has been gradually assimilated in several codes [6–8]. However, the structural use of fine recycled aggregates (FRA; less than 5 mm in size) that accounts for 20%–50% of the recycled aggregate production is considerably limited in several codes [8,9] because of its high water absorption, high porosity, low apparent density, etc. The structural application of recycled aggregate concrete (RAC) incorporated with the FRA is limited. Accordingly, researchers have mainly focused on its short-term mechanical properties (e.g., compressive strength, elastic modulus, and splitting strength) [10–12], durability (e.g. carbonation depth and chloride migration coefficient) [13–15] and shrinkage behavior [10,12,16], whereas its creep behavior has not drawn considerable attention. Compared to the RAC with the FRA, the creep behavior in the recycled concrete that only contains CRA has been well-studied. In particular, several experiments have been conducted to investigate the influence of the apparent density and water absorption of CRA [17], the water-to-cement ratio (w/c) of the

∗

parent concrete [18,19], the w/c of the resulting concrete [18,20], and the residual mortar content [21] on creep behavior of the recycled concrete solely incorporated with the CRA. To account for the effect of these parameters on the creep behavior of RAC, creep models have been proposed [17,18,21,22]. The increase in the potential of using the FRA in structural concrete can be attributed to the investigations on the mechanical behavior of concrete with the FRA in the last 20 years. It has been found that by employing reasonable methods, the mechanical properties of RAC with the FRA can be improved. For instance, by incorporating fly ash [23] or by replacing only the recycled fine aggregates with a size of 2–5 mm in the RAC [19,24], the RAC compressive strength, elastic modulus, and slump could be similar to those of natural aggregate concrete (NAC) with differences of 9.7%–13.6%, 2.6%–10.0%, and 2.6%–10.1%, respectively. Since 2001, Brito et al. [11,25–27] have conducted investigations on improving the mechanical behavior of RAC with the FRA by adjusting the weight of extra water in the concrete mixture. In these investigations, the compressive strength, elastic modulus, and slump of the RAC are also similar to those of NAC with maximum differences of 2.2%–5.4%, 11.8%–13.5%, and 6.5%–8.9%, respectively. The structural application of concrete with the FRA requires a suitable creep model. However, the currently available RAC creep models

Corresponding author. School of Civil Engineering, Harbin Institute of Technology, Harbin, 150090, China E-mail addresses: [email protected] (Y. Geng), [email protected] (M. Zhao), [email protected] (H. Yang), [email protected] (Y. Wang).

https://doi.org/10.1016/j.cemconcomp.2019.05.013 Received 16 October 2018; Received in revised form 1 April 2019; Accepted 15 May 2019 Available online 16 May 2019 0958-9465/ © 2019 Elsevier Ltd. All rights reserved.

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only account for the influence of recycled aggregates on the final creep value by applying an amplification factor to the NAC creep model. None of the models account for the difference in the creep development trend between the RAC and NAC. The differences in creep development trends are found in studies reported in literature [18,20] that pertain to concrete solely incorporated with the CRA. It is found that because of the considerable amount of extra water introduced into recycled aggregates, the RAC specimens require a longer period for the creep to stabilize than in the NAC samples [18]. Because of this, the creep deformation in the RAC may be significantly higher than that in the NAC in the latter period of its service life (e.g., 50 years). Such an effect can be more pronounced when both CRA and FRA are adopted because more extra water is introduced. Existing creep tests [27–32] on RAC with the FRA either had a loading duration of less than 100 d [27–30] or a FRA replacement ratio of less than 20% [31,32]. With such a short duration or low replacement ratio, the difference in creep development between the RAC and NAC cannot be observed. Moreover, most studies have mainly focused on the influence of FRA on the creep of concrete with a nil replacement of CRA [27–29] or at particular substitution level of CRA [30]. There is no experiment that has investigated whether the influence of FRA on the creep would vary at different substitution levels of CRA. Therefore, a new creep experiment with a longer loading duration and different incorporation ratios of both FRA and CRA is performed to develop a new creep model of the RAC that considers the influence of FRA and/or CRA on the creep development in concrete. Forty specimens are tested for 14 months. The incorporation ratio of FRA (rFRA) is varied between 0% and 100%. At each rFRA ratio, the incorporation ratio of CRA (rCRA) is also made to vary between 0% and 100%. Based on the test results, a new creep model of the RAC is developed by modifying the available creep model by 1) introducing the “creep improvement factor,” kr-imp, to account for the different influences of FRA at different rCRA ratios; 2) introducing the “creep development factor” to consider the effect of the extra water in the premoistened RA on the creep development trend of RAC [18,20]. Finally, the proposed model is benchmarked against experimental data in literature to evaluate the model's adequacy in predicting the RAC creep behavior.

Fig. 1. Grading curves of fine and coarse aggregates.

2.2. Mix proportions Eight mix designs are used to investigate the impact of both CRA and FRA replacement ratios on creep. In all cases with the incorporation of both CRA and FRA, the CRA replacement ratio equals or exceeds the FRA replacement ratio. For concrete mix designs with a CRA replacement ratio of 100%, three levels of FRA replacement ratios are investigated, namely 0%, 50% and 100%. For the concrete mix designs with the CRA replacement ratio of 50%, the FRA replacement ratios of 0% and 50% are included. For the concrete mix designs with a CRA replacement ratio of 0%, three levels of FRA replacement ratios are investigated, namely 0%, 50% and 100%. It is worth highlighting that the RAC with a CRA and FRA replacement ratios of 50% and 100%, respectively, is excluded in this long-term test because of the limitation in the number of loading rigs. This concrete mix is not tested because it is improbable to be adopted in real applications. More specifically, designers prefer the use of CRA rather than the FRA in structural concrete because the latter has considerably more severe side effects on concrete behavior than the former. In case the designers are considering the use of both FRA and CRA, it is unlikely that a FRA replacement level higher than that of CRA would be selected. To achieve acceptable slump values and mechanical properties, the CRA is used under the saturated-surface-dry (SSD) condition according to the guidelines of the Chinese standard [9] and suggestions in literature [36–38]. On the other hand, in the use of FRA, the water compensation method is employed according to the suggestions of Brito et al. [11,25–27]. Based on a series of experimental investigations, Brito et al. found that the FRA can absorb approximately 70% of its water absorption during a 10-min mixing period [11,25–27]. In this context, the amount of extra water that is used to compensate for the water absorption of the FRA is equivalent to the difference between the 70% of the FRA water absorption and the water content of air-dried FRA. It should be noted that the incorporation of CRA under the SSD condition and the use of water compensation method for the FRA may affect the slump values and creep behavior of the RAC in this test; this is discussed in a later section of this paper. The mix proportion details are summarized in Table 2. The target w/ c ratio listed in Table 2 represents the ratio of the assumed free water content to the cement content; the assumed free water content is obtained by subtracting 70% of the FRA water absorption from the total amount of water (i.e., the sum of the amount of water added to concrete mixes and the water content in fine aggregates). This ratio is constant in the experiment. The amount of superplasticizer to be added is determined by trials before mixing to ensure that all fresh concrete slumps are greater than 90 mm.

2. Experimental program 2.1. Materials In all concrete specimens in the experiment, ASTM Type Ι Portland cement is used; the 28-d compressive strength of concrete is 42.5 MPa. The natural fine aggregate (FNA) is river sand (class Ι conforming to GB/T14684-2011 [33]), and the natural coarse aggregate (CNA) is natural limestone gravel. The recycled aggregates are produced using waste concrete from a building in Harbin, China. In the production of recycled aggregates, the source concrete is crushed with a jaw crusher; thereafter, it is sieved with a square mesh. The sieved RA (with a nominal size larger than 25 mm) is jaw crushed for a second time to obtain a fraction size ranging from 0.15 to 25 mm. The sieved RA with a fraction size between 0.15 and 5 mm is employed as FRA, whereas those between 5 and 25 mm are used as CRA. For the purpose of comparison, all recycled aggregates that are used in this experiment have the same size distribution as the natural aggregates. The grading curves of coarse and fine aggregates in this experiment are presented in Fig. 1. The density, water absorption, and moisture content of aggregates are measured based on the Chinese code (JGJ 52–2006) [34], and the residual mortar content of the CRA is determined according to the method suggested by Abbas et al. [35]. All of the physical properties of aggregates are summarized in Table 1.

2.3. Preparation of specimens Concrete casting is conducted under laboratory conditions. In accordance with JCJ/T 240–2011 [9], the SSD condition is achieved by saturating the aggregates in water for 24 h and thereafter leaving them 304

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Table 1 Physical properties of aggregates. Aggregate

Grading (mm)

Oven-dried density (kg/m3)

DAc (kg/m3)

Water absorption (%)

Moisture content (%)

CRMd(%)

CNAa FNAb CRA FRA

5–25 0.15–5 5–25 0.15–5

2801 2597 2660 2222

2821 2626 2795 2472

0.70 1.11 5.07 11.23

– – – 5.63

– – 40 –

Note. a CNA, coarse natural aggregate. b FNA, fine natural aggregate. c DA, density of aggregates in the saturated-surface-day condition. d CRM, residual mortar content.

the conversion factor with the increase in the RA replacement ratio are also reported in literature. The differences among the values of conversion factors that are obtained from different tests are considerably large in both the RAC and NAC. None of the available tests report a difference among the conversion factors induced by the incorporation of recycled aggregates that is larger than existing differences obtained from previous NAC tests. In this context, the same conversion factor is used for the NAC and RAC according to the Chinese standard “Technical specification for application of recycled aggregate JGJ/T 240–2011” [9].

in air for 2 h until they are surface-dry. The specimens are prepared according to the following procedures. First, the aggregates are mixed in a laboratory mixer for 2 min with 1/3 of the total amount of water. Thereafter, cement is added to the mixer and agitated for another 3 min, after which the remaining amount of water, followed by the superplasticizer, is added and mixed further for 3 min. The fresh concrete is then cast in plastic molds and compacted in the vibration table. The following day, the specimens are removed from the mold and thereafter cured at a temperature of 20.0 ± 2.0 °C and humidity of 95% until the day before these are subjected to loading. 2.4. Test setup and method

2.4.3. Test on creep For each group of mixture, five 100 × 400 mm prisms are prepared for the long-term test, among which two are subjected to the design load at 28 d; the other three are prepared for shrinkage deformation measurements. The creep deformations are obtained by subtracting the measured shrinkage deformations from those measured in the corresponding loaded specimens. All 100 × 400 mm prisms are tested under the same conditions. The test is performed according to GB/T 50082–2009 [48]. To perform the long-term test, self-resisting loading frames are designed and introduced in a chamber in which creep and shrinkage tests are performed at a controlled temperature of 14.5 ± 3 °C and relative humidity (RH) of 64 ± 5%. The specimens that are included in the long-term tests are first loaded using a screw jack; thereafter, they are sustained with the nuts of prestressing rods (Fig. 2(a)). The loads are initially applied on the 28th day with an initial stress level of approximately 0.25. The loads are kept constant over long periods of time with a maximum deviation of 2% by tightening the nuts. The details pertaining to specimens subjected to long-term loads are summarized in Table 4. Electronic devices are employed to measure and monitor the deformation, loads, temperature, and RH values. The load values are

2.4.1. Slump test A slump test according to GB/T 50080–2002 is conducted after the mixing process [39]. The results are summarized in Table 3. 2.4.2. Test on strength and elastic modulus For each group of mixture, three 100 × 100 mm cubes and three 150 × 300 mm prisms are prepared to evaluate the 28-d compressive strength (fcu,100) and the elastic modulus (Ec28). These material tests are performed according to the standard of GB/ T 50081–2002 [40]; the details of the average test values are summarized in Table 3. The compressive strengths of 150 × 150 × 150 mm cubes (fcu) and the equivalent mean cylinder strengths (fcm28) are also listed in Table 3. The fcu is determined based on GB 50010–2010 [41] by multiplying a coefficient (0.95) to fcu,100; fcm28 is obtained using conversion factors provided by “CEB-FIP Model Code 1990” [42]. It is noteworthy that the conversion factors between the cubic strength and cylinder strength for RAC may differ from those that are used for the NAC. However, by investigating the test results of 79 groups of specimens that are reported in seven references [19,20,43–47], it is found that both descending and increasing trends in Table 2 Mix proportion of RAC and NAC. Concrete type

w/c

(w/c)target

Materials (kg/m3) Water

NAC RAC-rC0-rF50 RAC-rC0-rF100 RAC-rC50-rF0 RAC-rC50-rF50 RAC-rC100-rF0 RAC-rC100-rF50a RAC-rC100-rF100

0.45 0.47 0.48 0.45 0.47 0.45 0.47 0.48

0.45 0.45 0.45 0.45 0.45 0.45 0.45 0.45

180 186 193 180 186 180 186 193

Cement

400 400 400 400 400 400 400 400

Coarse aggregate

SPb

Fine aggregate

Natural

Recycled

Natural

Recycled

1191 1191 1191 595 595 0 0 0

0 0 0 594 594 1188 1188 1188

670 335 0 670 335 670 335 0

0 303 605 0 303 0 303 605

2 2 2 2 2 2 2 2

Note. a For the nomenclature, RAC-rC100-rF50 represents recycled aggregate concrete with CRA replacement ratio of 100% and with FRA replacement ratio of 50%. b SP, superplasticizer. 305

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Table 3 Material properties of concrete specimens. Concrete type

rCRA (%)

rFRA (%)

Concrete Strength (MPa) fcu,

NAC RAC-rC0-rF50 RAC-rC0-rF100 RAC-rC50-rF0 RAC-rC50-rF50 RAC-rC100-rF0 RAC-rC100-rF50 RAC-rC100-rF100

0 0 0 50 50 100 100 100

0 50 100 0 50 0 50 100

100

56.0 52.1 47.4 47.5 50.4 45.3 46.8 52.1

fcu

fcm28

53.2 49.5 45.0 45.1 47.9 43.0 44.5 49.5

42.4 40.0 36.7 36.8 38.9 35.2 36.4 40.8

Ec28 ( × 104MPa)

Slump (mm)

3.38 3.31 2.92 3.05 2.96 2.63 2.53 2.37

138 150 90 163 139 168 157 110

2.4.4. SEM test After the long-term test, the NAC specimen and creep test specimens with FRA substitution levels of 0% and 100% in the concrete with CRA substitution levels of 100% are sliced to investigate their microstructure properties in the interfacial transition zone (ITZ); each slice is 50 × 50 × 10 mm. The tests are conducted using VEGA3 TESCAN scanning electron microscope (SEM). Before the SEM examination, the surfaces of slices

recorded by load cells during the entire test duration. The creep and shrinkage deformations are measured by a 200-mm long digital mechanical strain gauge (DEMEC), as shown in Fig. 2(b). Four pairs of stainless-steel discs are glued to the four surfaces of each specimen, as shown in Fig. 2(c). The specimens for shrinkage tests are placed close to the specimens for creep tests to ensure that they are exposed to the same temperature and humidity conditions.

Fig. 2. Long-term test set-up. 306

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Table 4 Detailed information for long-term specimens. Concrete type

rCRA (%)

rFRA (%)

fcm28 (MPa)

a × a × h (mm)

t0 (days)

NL(kN)

nc

NAC RAC-rC0-rF50 RAC-rC0-rF100 RAC-rC50-rF0 RAC-rC50-rF50 RAC-rC100-rF0 RAC-rC100-rF50 RAC-rC100-rF100

0 0 0 50 50 100 100 100

0 50 100 0 50 0 50 100

41.1 41.8 42.2 38.9 40.3 35.9 37.5 38.8

100 × 100 × 400 100 × 100 × 400 100 × 100 × 400 100 × 100 × 400 100 × 100 × 400 100 × 100 × 400 100 × 100 × 400 100 × 100 × 400

28 28 28 28 28 28 28 28

100 104 102 93 100 89 101 103

0.24 0.25 0.24 0.24 0.25 0.25 0.27 0.27

by other researchers; they found that the percentage of increase vary between 23% and 56% in the RAC with a 50% rCRA and between 34% and 87% in the RAC with a 100% rCRA (e.g. [20,23,46]). The increase of creep deformation is because when pre-saturated CRA was incorporated, the water will bleed from the pores of aggregates to the cement paste, leading to the increase of the effective w/c ratios in the RAC mix. The increased w/c ratio would increase the creep deformation. This bleeding effect of water is confirmed by the SEM test in this study (Fig. 5(a) and (c)). Fig. 5(a) and (c) compare the microstructures at the ITZs of NAC and RAC-rC100-rF0 specimens. In the figures, it can be noted that because of the bleeding effect, the ITZ in the RAC solely incorporated with the CRA (i.e., specimen RAC- rC100-rF0) is more porous (Fig. 5(a)) than that in the NAC specimen (Fig. 5(c)). This high porosity of the microstructure of the RAC that is solely incorporated with the CRA could increase the creep deformation. When recycled fine aggregates are also included in the mix, the increase in the specific creep that is induced by the inclusion of CRA is less significant. In particular, in the RAC specimens with rFRA = 50% and rCRA = 100% (i.e., specimen RAC-rC100-rF50), the specific creep is only 6.2% higher than that in the RAC specimen with rCRA = 0% (i.e., RAC-rC0-rF50). In concrete mixes with rFRA = 100% and rCRA = 100% (i.e., RAC-rC100-rF100), the specific creep is only 7.8% higher than that in the specimen with rCRA = 0% (i.e., RAC-rC0-rF100). It can be observed in Fig. 4 that all of the RAC mixes that are incorporated with the FRA have a specific creep that is higher than that in the NAC specimens. In particular, in the RAC specimens with rFRA = 50% (i.e., RAC-rC0-rF50, RAC-rC50-rF50 and RAC-rC100-rF50), the specific creep is 33.8%–34.4% higher than that in the NAC specimens. On the other hand, in the RAC specimens with rFRA = 100% (i.e., RAC-rC0-rF100 and RAC-rC100-rF100), the specific creep is 22.8%–32.4% higher than that in the accompanying NAC specimens. It should be noted that in comparing the influence on creep deformation that is induced only by the incorporation of CRA, the effect that is induced only by the replacement of fine aggregates is less pronounced. For example, the specimens that use natural gravel and FRA have a specific creep that is only 23%–34% higher than that in specimens containing only natural aggregates, whereas in specimens that are provided only with the CRA, the specific creep can be 49%–53% higher than that in the NAC specimens. When recycled coarse aggregates are included in the mixes, the incorporation of FRA can even decrease the long-term deformation of RAC. In particular, in the RAC with a 50% CRA substitution level, the specific creep in specimens with a 50% rFRA ratio (i.e., RAC-rC50-rF50) is 9.2% lower than that in specimens with rFRA = 0% (i.e., RAC-rC50rF0). In the RAC with a CRA substitution level of 100%, the specific creep in the specimens with rFRA of 50% and 100% are 7.8% and 13.7% lower than that of the accompanying specimens with rFRA = 0%, respectively. The decrease in the creep deformation of the RAC mixes that are incorporated with the FRA as described above may be attributed to the continuous water absorption ability of the FRA. In particular, the FRA is mixed using the water compensation method that considers only 70% of the water absorption. With this method, the FRA still has the

are polished and coated by sputtering a thin gold layer. 3. Experimental results The range of slump values in all concrete mixes is 90–168 mm (Table 3); this range is acceptable in concrete construction in real applications. It is noteworthy that the incorporation of CRA can generally increase the RAC slump (Table 3), whereas the slump decreases with the replacement ratio of recycled fine aggregates. This variation in slump values in the test could have been induced by the water compensating methods employed for the recycled fine and coarse aggregates in this test. In particular, because of water bleeding inside the pores of pre-saturated CRA during mixing, the incorporation of CRA under the SSD condition would increase the effective w/c ratio of the RAC and consequently increase the slump. On the other hand, when the FRA is used in the experiment, the w/c ratio decreases with the replacement ratio. This is because only 70% of the FRA water absorption is considered in the water compensation, and the FRA may still have the capacity of absorbing more water during concrete mixing. This reduction in the effective w/c ratio could lead to the decrease in the slump value of the RAC. Fig. 3 depicts the specific creep (i.e. the creep strain produced by a sustained unit stress) of specimens with different FRA and CRA substitution percentages, along with certain predicted results; these are discussed in a later section. The creep deformation is calculated by deducting the shrinkage deformation (using the average value that is measured from the three accompanying shrinkage specimens) from the total deformation. As expected, the creep values in all specimens increase continuously under the sustained loads during the test. At the beginning of the test, the creep deformation substantially increases; after the first month, it starts to decrease. According to available longterm test results (loading duration longer than 10 years) [49–51], in the first year, the creep deformation would reach more than 80% of the value measured after 10 years under sustained loading. This indicates that although the creep strain grows over the entire service life of concrete, it mainly develops over the first year; hence, a one-year creep test can provide valuable information. Several creep experiments have been terminated at the end of the first year (e.g. [19,21,52–54]); accordingly, the creep test in this study is terminated at 420 d. 3.1. Influence of recycled aggregates on final creep deformation A detailed comparison of the specific creep between the NAC and RAC specimens is depicted in Fig. 4 to further explore the effect of rCRA (replacement ratio of CRA) and rFRA (replacement ratio of FRA) on the creep behavior of RAC specimens. As expected, the specific creep in RAC specimens with the CRA is higher than that of the other specimens that contain natural coarse aggregates regardless of the FRA replacement percentage. Considering the FRA substitution level of 0%, the RAC specimens with CRA substitution levels that vary between 50% and 100% have creep deformations that are 49%–53% higher than those of the NAC samples. Such an increase in creep deformation that is induced by the inclusion of CRA agreed well with those that have been observed 307

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Fig. 3. Specific creep (C(t,t0)) of RAC and NAC specimens during the long-term test.

ratio in both the new cement paste and the ITZ. The reduction in the effective w/c ratio would decrease the creep deformation in the new cement paste and increase the stiffness of the ITZ around the CRA; this would further decrease the creep deformation in the resulting concrete by providing a better restraint to the deformation of the hydrated cement paste. The enhancement of the ITZ around the CRA that is induced by the incorporation of FRA can be confirmed with the SEM images of the new ITZs between the aggregates and new cement paste in the RACs (Fig. 5). In comparing Fig. 5(a) and (b), it can be observed that the ITZ around the CRA in the RAC that is incorporated with the FRA is less porous than that in the RAC that is solely incorporated with the CRA. The increased density of the ITZ indicates higher stiffness; hence, it would more effectively restrain the creep deformation in the new cement paste. 3.2. Influence of recycled aggregates on creep development Fig. 4. Specific creep (C(t,t0)) of RAC with different CRA and FRA replacement ratio at the end of the long-term test.

Comparing Fig. 3(a) and (f), it can be observed that the RAC specimens require longer periods for the creep to stabilize compared with those of the NAC samples. In particular, the creep deformation in the NAC specimens practically stops developing after 200 d under sustained loading; in the next 220 d, it only increases by 18.5% (Fig. 3(a)). On the other hand, the creep increase in the RAC specimens remain

capability of absorbing more free water from both the new cement paste and the ITZ around the CRA during the concrete mixing and hardening processes. This leads to the reduction in the effective w/c 308

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Fig. 5. SEM images for ITZ of RAC and NAC specimens after long-term test.

considerable after 200 d under sustained loading; it is 26.8% of the final value at the end of the test. The difference between the creep development trend in the RAC and NAC specimens can be well illustrated in Fig. 6 by normalizing the creep deformations with respect to the final creep deformations at the end of the test (i.e., using the ratio of ε(t) to ε(tend)). In Fig. 6, the continuous growing trend in the creep of RAC is reflected by the lower values of ε(t)/ε(tend) at the start of the creep experiment when compared with that of the control NAC. For example, considering the time of 145 d, ε(t)/ε(tend) of the RAC with a 100% CRA is 8.2% lower than that of the NAC specimen. The percentage increases to 16.0% when both coarse and fine aggregates are fully replaced by recycled materials. By observing the creep development trend in the RAC specimens that are incorporated only with the CRA (Fig. 6(b)) or only with the FRA (Fig. 6(c)), it can be noted that the RAC specimens have a lower ε(t)/ε(tend) than the NAC specimens although the difference is relatively small from a quantitative point of view. Geng et al. [18] also found that the creep development rate in concrete mixes incorporated with recycled coarse aggregates was lower than that in mixes with natural aggregates only; their difference ranged 11.9%–15.8%. The continuous development of creep deformation in the RAC specimens may be attributed to the water supplemented to the new cement paste by the recycled aggregates, which contain a considerable amount of water. More specifically, drying creep depends on water loss. To achieve a satisfying slump value (higher than 90 mm in this paper), the CRA is used under the SSD condition, whereas in the FRA, a certain amount of extra water is included in the concrete mixture, and 1/3 of the total water is agitated with the combination of fine and coarse aggregates for 2 min before the addition of cement to allow time for the FRA to absorb the water. In this context, both the CRA and FRA in the hardened concrete contain relatively large amounts of water in the porous structure. The water loss in the new cement paste triggers the

concrete creep and causes a moisture gradient between the old and new cement pastes. Because of this gradient, the water within the porous structures of recycled aggregates is diffused to the new cement paste; this compensates for the water loss in the RAC specimens during drying. Because the water loss is compensated, the RAC requires a longer time to reach humidity equilibrium between its specimens and the atmosphere; hence, the drying creep deformation continues to develop over a longer period. The lower values of the relative creep strain of recycled concrete can also be attributed to various time-dependent developments of recycled concrete properties. In particular, because of the internal curing effect of recycled concrete, the time-dependent development of the RAC compressive strength could be greater than that of the NAC specimen. This leads to lower values of the relative creep strain that is consistent with conclusions reported in references [23,55]. To date, most of the RAC creep models have been obtained by multiplying an amplification factor to the creep model of normal concrete. In particular, the expressions of the amplification factor are regressed based on the measured results, which are terminated at 60–1000 d [20,21,27]. However, the equations that are adopted to describe the development of creep with time use the same expressions as those used in normal concrete (e.g., the models of Brito [17], Fathifazl [21], and Geng [18]). Such type of modification to build the RAC creep model cannot describe the difference in the creep development trend between the RAC and NAC, as presented in Fig. 6; hence, they will probably underestimate the creep deformation in the RAC in the latter part of its 50-year service life (Fig. 7). In Fig. 7, the prediction that uses Geng's model is obtained by introducing the amplification factor to the EC2 model [56] as suggested by Geng [18]. The EC2 model is one of the commonly recommended models [27,57–60] because of its capability of capturing both the creep development and final creep value in the NAC well [58,61]. The 309

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Fig. 7. Difference in the predicted long-term deformation between the modified EC2 model and regression based on short-time test for NAC and RAC specimens.

regression based on the short-term test data is obtained using the hyperbolic expression suggested by Neville [62]; it is proven capable of describing the increasing trend of long-term deformation with time in concrete. As shown in Fig. 7, the deformation that is calculated by the EC2 model [56] moderately approximates the experimental result of the NAC. On the other hand, for the RAC specimens, by modifying the final creep coefficient only, the creep development trend that is calculated by Geng's model [18] considerably differs from that obtained using the regressed test data; this results in a 48.7% underestimation of the creep value at 50 years. Therefore, it is necessary to modify the coefficient that is used to describe the development of creep with time in the recycled concrete.

4. Modification of existing creep model for RAC By modifying the EC2 model [56], a new creep model for calculating the creep deformation in the RAC is proposed in this paper. Based on all the investigations described above, it can be concluded that the incorporation of RA affected both the creep development trend and final creep value. The effect of RA on the creep development can be considered by introducing a coefficient, kr-Δw/c, into the creep development model of EC2, in which the additional water, Δw, that is induced by the incorporation of RA is considered as the main parameter. The incorporation of recycled aggregates has both detrimental and beneficial effects on the final creep value of the RAC. In particular, these side effects include 1) an increase in the total water content that is induced by the inclusion of RA and 2) creep of the residual cement paste; both of these are accounted for by applying an amplification factor, kr-RP, to the model. In factor kr-RP, the volume content of the residual cement paste, VRP, is the key parameter. The beneficial effect is the water content reduction in the fresh cement paste and in the ITZ around the CRA because of the continuous water absorption ability of FRA. This effect is accounted for by the introduction of a modification factor, kr-imp, which contains two key parameters: the water absorption of FRA, wFRA, and the volume of CRA, VCRARAC. Moreover, if the original concrete from which RA is obtained has been previously subjected to sustained loading, only the recoverable creep of the residual cement paste would be present [22]. For this case, the recoverable creep coefficient, kr-RC, in which the volume of residual cement paste (VRP) is the main parameter, has to be introduced to the creep model.

Fig. 6. Creep development (εc(t)/εc(tend) of RAC and NAC specimens during the long-term test.

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development in all RAC specimens is presented by the modified model, whereas the EC2 model tends to underestimate the creep development trend in the RAC in the latter period of the test. In particular, the deviation between the experimental results and those of the modified model is 5.5%–10.7%, whereas the deviation between the experimental results and the underestimated results of the EC2 model is 12%–35.7%. It is noteworthy that the difference between the experimental results [18,21,32,63] and the values predicted by the traditional model for specimens that are reported in literature is not remarkable, as shown in Fig. 10. This is because those reported in literature only belong to one group of specimens that are produced by using FRA with a replacement percentage of 19%; all the other specimens are only incorporated with the CRA. With the low incorporation ratio of FRA, the total additional water, Δw, induced by the inclusion of RA is small. As a result, the deviation between the creep development in the RAC and that in the NAC is also small. Among the existing 23 groups of RAC specimens that are incorporated with the FRA, only one group of specimens is included in this comparison [29–32] because it is the only group [32] that is subjected to sustained loading for a relatively long time; hence, it is the only group that provides sufficient information for prediction. Most of the reported creep tests on the RAC with the FRA are terminated within 90 d; during this period, the water within the recycled aggregates does not have ample time to diffuse to the new cement paste. This indicates that further experimental work on the RAC with a wide range of parameters and subjected to test durations longer than 200 d is necessary to verify the reliability of the proposed model.

4.1. Creep development factor (kr-Δw/c) The coefficient of creep development in the EC2 model [56] is presented in Eq. (1). 0.3

(t − t0) ⎤ βcN (t , t0) = ⎡ ⎢ (β + t − t0) ⎥ ⎣ H ⎦

(1)

where t denotes the age of concrete (in days) when the calculation of creep deformation is necessary, and t0 is the age of concrete at the first loading (in days); βH represents the coefficient related to the relative humidity of the environment. The relative humidity is the main factor that triggers the movement of water within the recycled aggregates into the new cement paste. Therefore, the creep development coefficient, βcR(t,t0), of the RAC can be predicted by multiplying kr-Δw/c to βH, as given by Eq. (2). 0.3

(t − t0) ⎤ βcR (t , t0) = ⎡ ⎢ (k r− Δw/c β + t − t0) ⎥ H ⎦ ⎣

(2)

The additional water, Δw, is equal to the water content of pre-saturated CRA plus the water compensation for the FRA. In our test, the amount of water compensation for the FRA is 70% of the water absorption capacity of FRA. It is noteworthy that the water added to compensate for the water absorption of FRA is considered to perform the same function as the water resulting from the pre-saturation of CRA. This is because the water added to compensate for water absorption is assumed to be inside the FRA after the mixing process. More specifically, according to the conclusion drawn by Brito et al. [11,25–27], the FRA can absorb approximately 70% of its water absorption during the 10-min mixing process. Based on this result, the amount of the extra water used to compensate for the water absorption of the FRA in this test was the difference between 70% of the FRA water absorption and the water content of air-dried FRA. Ideally, the extra water could be exactly absorbed by the FRA during the mixing process; hence, the extra water could be inside the FRA after mixing process. The relationship between kr-Δw/c and Δw/c is established through a regression analysis that is based on the experimental results shown in Fig. 6. The variation of kr-Δw/c as a function of Δw/c is illustrated in Fig. 8, where it can be observed that kr-Δw/c is relatively proportional to Δw/c, and that the relationship given by Eq. (3) exists.

k r− Δw/c = 87

Δw +1 c

4.2. Residual cement paste coefficient (kr-RP) According to Neville [62], the creep of the NAC is triggered by the creep of the cement paste, whereas the coarse and fine aggregates in the NAC restrain it. The creep in the NAC is related to the total volume fraction of natural aggregates in the NAC by Eq. (4). NAC c NAC = (1 − VNA )

αNAC

× cp

(4)

where VNANAC denotes the total volume fraction of natural coarse and fine aggregates in the NAC; cp is the creep of the cement paste; αNAC represents the coefficient related to the stiffness of natural aggregates (it can be estimated as 1.33 according to Fathifazl et al.) [21]. Following Neville's speculation, this paper proposes equations for predicting the influence of residual cement paste on the creep value of the resulting concrete that is incorporated with CRA and/or FRA. In particular, the creep of concrete is assumed to be induced by the creep of the cement paste, whereas the creep of the cement paste is restrained by the natural coarse and fine aggregates in the RAC. The stiffness values of the original virgin fine and coarse aggregates are assumed to be the same as those of the new natural fine and coarse aggregates, respectively. Based on these assumptions, the amplification factor, φRAC/φNAC, is proposed, as given by Eq. (5):

(3)

Using Eq. (2), the predicted ratio, ε(t)/ε(tend), is compared to the value that is obtained from the RAC specimens (Fig. 9) of this study and those that are reported in literature (Fig. 10). It can be observed that a reasonable estimate of the creep

RAC RAC αRAC c pR φRAC (1 − VTNCA ) − VTNCA = × 1.33 NAC NAC φNAC cp (1 − VCNA − VFNA )

(5)

where αRAC can still be estimated as 1.33 when the ratio of ERAC/ENAC falls within the range 0.7–1.3 [21]; cpR is the creep of the cement paste in the RAC, including the creep of the residual cement paste and that of the new cement paste; cp is the creep of cement paste for NAC. Because of the potential water absorption ability of FRA, the value of cpR may be lower than that of cp. This difference between the two is considered in the creep improvement factor, kr-imp, which is presented in a later section. Therefore, φRAC/φNAC can be calculated as Eqs. (6) and (7); the residual cement paste coefficient, kr-RP, is given by Eq. (7):

φRAC = k r−RP k r−imp φNAC

Fig. 8. Variation of the coefficient kr-Δw/c as a function of the increment of water-to-cement ratio (Δw/c). 311

(6)

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Fig. 9. Comparison of measured and calculated ratio of ε(t)/ε(tend) for the RAC tested in this paper.

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Fig. 10. Comparison of measured and calculated ratio of ε(t)/ε(tend) for the RAC in the available literature.

(i.e., the sum of VOCACRA and VOFACRA) and that in the FRA (i.e., VOFAFRA) can be calculated using Eqs. (10) and (11):

1.33

RAC RAC 1 − VTNCA − VTNFA ⎞ k r−RP = ⎜⎛ NAC NAC ⎟ ⎝ 1 − VCNA − VFNA ⎠

(7)

VTNCARAC

represents the total volume fraction of natural where the coarse aggregates in the RAC; it is the sum of the volume fraction of the new natural coarse aggregates (VCNARAC) and that of the original virgin coarse aggregates in the CRA (VOCACRA), as given by Eq. (8). The total volume fraction of natural fine aggregates in the RAC is denoted by VTNFARAC; it is the sum of the volume fraction of the new natural fine aggregates (VFNARAC), the volume fraction of the original virgin fine aggregates in the CRA (VOFACRA), and that of the original virgin fine aggregates in the FRA (VOFAFRA), as given by Eq. (9). RAC RAC CRA VTNCA = VCNA + VOCA

(8)

RAC RAC CRA FRA VTNFA = VFNA + VOFA + VOFA

(9)

CRA CRA RAC FRA VOCA + VOFA = VCRA − VRP

(10)

FRA RAC FRA VOFA = VFRA − VRP

(11)

VCRARAC

where and VFRARAC represent the volume fractions FRA, respectively; VRPCRA and VRPFRA denote the volume

of CRA and fractions of

residual cement paste in the CRA and FRA. Substituting Eqs. 8–11 into Eq. (7), the residual cement paste coefficient, kr-RP, can be calculated as Eq. (12). 1.33

CRA FRA VRP + VRP ⎞ k r−RP = ⎜⎛1 + NAC NAC ⎟ 1 − VCNA − VFNA ⎝ ⎠

VRPCRA/VRMCRA

(12) CRA

The ratio (where VRM is the volume fraction of the residual mortar in the CRA) and that of VRPFRA/VFRARAC are

The volume fraction of the original virgin aggregates in the CRA 313

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Fig. 11. Comparison between measured and predicted ratio of creep for RAC and that for NAC reported in the available literature.

original concrete mix proportion, which may not be easily obtained. If this is the case, to predict kr-RP, choosing a particular value of VOP/VOM may be an easier approach. In particular, according to the Chinese specification of mix proportion design [65], as the w/c of concrete varies between 0.3 and 0.6, the ratio of the cement paste volume to that of the mortar varies from 0.48 to 0.76. In the prediction of residual cement paste coefficient, kr-RP, the use of 0.57 as the value of VOP/VOM results in a deviation of less than 12%. This deviation from the prediction results is obtained when the true value of VOP/VOM is used for the RAC with aggregates obtained from waste concrete with w/c of 0.3–0.6. Eqs (13) and (14) can then be simplified as Eqs. (15) and (16), respectively.

assumed to be equal to the ratio VOP/VOM (where VOP is the volume fraction of the cement paste in the original concrete that is used to produce recycled aggregates, and VOM is the volume fraction of the mortar in the original concrete). Based on these assumptions, VRPCRA and VRPFRA can be calculated by Eqs. (13) and (14): CRA VRP =

VOP RAC CRM VCRA VOM

(13)

FRA VRP =

VOP RAC VFRA VOM

(14)

where CRM represents the volume content of residual mortar; it can be measured according to Abbas et al. [35] or Domingo [64]. The proposed expressions (Eqs. (12)–(14)) for predicting the residual cement paste coefficient, kr-RP, require the knowledge of the

CRA RAC VRP = 0.57CRM VCRA

314

(15)

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Fig. 12. Comparison between the ratios of (φRAC/φNAC)ex measured from experiments and the predicted ratios of (φRAC/φNAC)pre. FRA RAC VRP = 0.57VFRA

kr-RC proposed by Fathifazl et al. [22] to consider the influence of the loading history for the residual cement paste on the creep of RAC (see Eqs. (20)–(22)).

(16)

If a higher degree of precision is required, the use of Eqs. (13) and (14) is still suggested. Substituting Eqs. (15) and (16) into Eq. (12), the resulting equation can then be re-arranged to yield Eq. (17) with the replacement ratios of CRA (rCRA) and FRA (rFRA).

1.33

k r−RC

1.33

RAC RAC rCRA CRM VCA + rFRA VFA ⎞ k r−RP = ⎜⎛1 + 0.57 ⎟ NAC NAC 1 − V − V CNA FNA ⎝ ⎠

(17)

Kt = 4.3. Creep improvement factor (kr-imp)

k r−imp = 1.3 + wre/c =

−

RAC 35VCRA (wre/c )

⎤ ⎥ ⎥ ⎥ ⎦

(20) (21)

VRPRAC

where the and represents the volume fraction for the new cement paste and that for the residual cement paste, respectively; the t is the duration of loading (days); the β donates the adjusted factor, which can be estimated to be 1 according to Fathifazl et al. [22]. RAC CRA FRA VRP = VRP + VRP CRA

(18)

and VRP where VRP respectively.

wabsorption − wnatural − wadd c

t 0.6 10 + t 0.6 VNPRAC

By applying the regression techniques to the experimental data in Fig. 1, this paper proposes equations to predict the creep improvement factor kr-imp (see Eqs. (18) and (19)) for the RAC with the incorporation of FRA. RAC 1.5VCRA

⎡ ⎞ ⎛ 1 ⎢ ⎟ = ⎢1 − βKt ⎜ RAC VNP ⎜ 1 + RAC ⎟ ⎢ VRP ⎠ ⎝ ⎣

(19)

(22) FRA

can be obtained with Eq. (15) and Eq. (16),

4.5. Model validation

where the wabsorption donates the mass of absorbed water in weight for FRA; the wnatural is the moisture content in weight for FRA before mixing; the wadd represents the weight of extra water inserted into concrete mix due to the incorporation of FRA.

The predicted results are compared with the creep data of 54 groups of RAC specimens used in this study and 14 other groups reported in literature [20,21,29–32,46,63,64,66–70]. The benchmarking specimens have a relatively wide range of material properties, 28-d mean cylinder strength (fcm) of 18–52 MPa, replacement ratios (rCRA and rFRA) that range from 0% to 100%. The comparative results of specimens tested in this paper are

4.4. Recoverable creep coefficient (kr-RC) This paper follows the equations of the recoverable creep coefficient 315

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Natural Science Foundation of China (No. 51678195), by the Postdoctoral Research Plan of Heilongjiang (LBH-Q16098) and by the Open-funds of Key Lab of HIT (HITCE 201701). The authors are very grateful to Prof. Dawn E. Lehman for her assistance to enhance the writing.

presented in Fig. 3. Some representative comparisons between experimental data reported in literature and predicted results obtained from the test are presented in Fig. 11. In this figure, the effect of the incorporation of RA on the creep behavior of RAC specimens is presented by the ratio of the creep coefficient of RAC (φRAC) to that of the NAC (φNAC) at the end of tests. It can be observed in Figs. 3 and 11 that the calculated values agree well with experimental data on the development trends of creep and creep deformation (the maximum difference is 16%). In order to evaluate the accuracy of the proposed model in predicting the creep values comprehensively, the measured φRAC/φNAC at the end of the tests are compared with those predicted for the 54 groups of specimens (Fig. 12). In this figure, the experimental data from the 23 groups of specimens with FRA are color highlighted, whereas the RAC specimens incorporated only with the CRA are in gray. It can be observed that the proposed model can only reasonably estimate the creep of the RAC—the maximum deviation is 21%, the mean is 0.993, and the COV is 0.115.

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5. Conclusion The experimental investigations in this study demonstrated how the incorporation of FRA affected the creep behavior of concrete with different replacement levels of the CRA. Forty specimens were included in the long-term experiment with the FRA substitution levels (rFRA) of 0%, 50%, and 100% for the concrete with CRA substitution levels (rCRA) of 0% and 100%; specimens with rFRA and rCRA of 0% and 50%, respectively, were also investigated. The long-term deformations were measured over 14 months. A new creep model was developed to predict the creep behavior of RAC incorporated with both CRA and FRA. The experimental results showed that the creep deformation in the RAC was 23–53% higher than that in the NAC specimens. In the RAC with 0% CRA, the creep deformation increased with the increase in the FRA incorporation ratio; whereas in the RAC with 100% CRA, the creep deformation decreased with the increase in the FRA incorporation ratio. In particular, when the natural fine aggregates were fully replaced by the FRA, the creep deformation increased by 34% in the RAC with 0% rCRA, whereas the creep deformation decreased by 13.7% in the RAC with 100% rCRA. Furthermore, the creep development trend was significantly influenced by the replacement of recycled aggregates because of the supplemented water to the new cement paste from the saturated recycled aggregates; this may have led to the 48.7% underestimation of the creep deformation in the RAC in the latter period of the 50-year service life of concrete using existing RAC creep models. In this paper, the proposed creep model of the RAC has equations to predict creep development considering the supplemented water to the new cement paste from the saturated recycled aggregates. The model also has equations for the final creep values that account for the influence of 1) the reduction in the water content of fresh cement paste because of the potential ability of FRA to absorb water during the mixing and hardening processes of the concrete (the FRA is not fully saturated) and 2) the enhancement of ITZ around the CRA induced by the incorporation of FRA. According to the comparative results of the current experimental data, the new model was found capable of predicting the creep development trend; it was found that the predicted creep values in the RAC specimens have a maximum deviation of 21%, mean value of 0.993, and COV of 0.115. It is recommended that further experiments can be conducted on the RAC specimens with wider parameter ranges using CRA and FRA from different sources and with loads sustained for longer durations to further validate the reliability of the proposed model. Acknowledgements The research work in this paper was supported by the National 316

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