Residual mechanical properties of compound concrete containing demolished concrete lumps after exposure to high temperatures

Fire Safety Journal 105 (2019) 62–78

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Residual mechanical properties of compound concrete containing demolished concrete lumps after exposure to high temperatures

T

Bo Wu, Yong Yu, Xin-Yu Zhao∗ State Key Laboratory of Subtropical Building Science, South China University of Technology, Guangzhou 510640, PR China

ARTICLE INFO

ABSTRACT

Keywords: Compound concrete Demolished concrete lumps High temperature Residual mechanical properties Shape effects Volume dilation

The structural use of compound concrete (CC) made of demolished concrete lumps (DCLs) and fresh concrete (FC) has been demonstrated to be a feasible alternative for waste concrete recycling. Previous studies have put focus on the properties of CC at ambient temperature, but little is known about the fire related performance of such material. To address this gap, a total of 159 CC specimens were tested after exposure to high temperatures up to 600 °C. The samples' failure modes, residual mechanical properties, and deformation characteristics were presented. The findings suggest that despite the thermal exposure, the CC specimens possessed similar relative compressive strength and elastic modulus to their FC counterparts; however, above 200 °C both CC and FC showed a slightly decreased capacity of strength retention with temperature compared to recycled aggregate concrete. The CC specimens' shape effect was prominent, since as the exposed temperature and the replacement ratio increased, the cube-to-cylinder strength ratio rose dramatically from 1.23 to 1.72. Moreover, it was found that incorporating DCLs had a non-negligible impact on the specimens’ ambient and post-fire dilation behaviors. The ratio of the critical load (i.e. the load at which the dilations initiated) to the peak load exhibited a decreasing trend with the DCLs content; but such effect almost diminished at 600 °C. Lastly, a stress–strain relationship of CC after exposure to high temperatures was proposed.

1. Introduction The construction industry is developing rapidly in tandem with the global industrialization and urbanization. It is a vital sector that contributes greatly to a flourishing society; yet it also has high environmental impact [1–5]. It consumes lots of natural resources such as sand and gravel, and quarrying them degrades the environment. At the same time, a substantial number of old buildings which have reached the end of the service life are being demolished. The increasing demolition waste, however, is normally disposed of in landfills, rendering the landfill capacity a scarce resource in many countries [3,5]. To mitigate the side effects of construction and demolition, much effort has been devoted to finding techniques for recycling construction and demolition waste (CDW). The hope is to limit the burdens on natural resources and waste landfills to within reasonable bounds [2,6]. Recycling waste concrete is one effective technique. In producing recycled aggregate concrete (RAC), crushed waste concrete is used to replace, partly or totally, virgin aggregate [7,8]. Up to now, a large number of studies have been conducted aimed at quantifying the physical, mechanical (short-term) and durability proprieties of RAC and at evaluating its economic feasibility and life-cycle impact [9–19]. ∗

Compared with concrete made with natural aggregates (NAs), RAC has three key distinctions: (i) a higher volume of micro-cracks and pores produced in the crushing process; (ii) a certain amount of old mortar attached on the surface of the recycled aggregate (RA); (iii) more interfacial transition zones distributed in RAC, in particular the two aggregate–mortar interfaces [20,21]. Those features are the primary reasons RAC is generally considered inferior to concrete made with NAs [5]. However, after special treatment [22–24] or simply by adjusting mix proportions [25,26], the mechanical performance of RAC can be effectively improved. Considering that currently only less than 1% of RAs is used in structural concrete applications [27], more viable options for old concrete recycling would be welcome. With this thought, the author's group has been perusing a new solution that advocates directly mixing demolished concrete lumps (DCLs, also referred to as recycled concrete lumps (RCLs) in literature [28]) with fresh concrete (FC) in casting structural members. The resulting concrete comprising DCLs and FC can be termed compound concrete (CC). Chinese standard (GB/T 25177-2010 [29]) requires pieces of RA to have an equivalent diameter smaller than 31.5 mm. In contrast, the DCLs we have been dealing with are coarsely crushed and can have a

Corresponding author. E-mail address: [email protected] (X.-Y. Zhao).

https://doi.org/10.1016/j.firesaf.2019.02.008 Received 4 December 2018; Received in revised form 30 January 2019; Accepted 17 February 2019 Available online 19 February 2019 0379-7112/ © 2019 Elsevier Ltd. All rights reserved.

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Fig. 1. Example applications of various kinds of members made with CC containing DCLs.

size ranging from 60 to 300 mm. The size of DCLs can be chosen depending primarily on the cross-sectional dimension of the member which will be cast containing those DCLs. For example, DCLs 60–80 mm in equivalent diameter are preferred for floor slabs while 250–300 mm pieces can be adopted for bridge piers. Due to the large size of DCLs, this recycling method has been recognized as having several advantages [28]: (i) the manufacturing process is substantially simplified because the second-stage crushing and screening are eliminated; (ii) an increased recycling ratio can be achieved since a large amount of cement mortar in the DCLs can be retained and hence be recycled. Following this recycling method, the authors’ group has conducted a series of experimental studies addressing the mechanical properties of CC [e.g., 30–32]. Moreover, the practical problems of casting CC in real projects have been examined. Fig. 1 shows some pilot applications implemented in the past few years. The research and application results have indicated that CC delivers static performance qualitatively similar to that of new concrete. CC is thus applicable to many kinds of concrete members including building slabs, concrete-filled steel tubes, and massive concrete elements like column foundations and retaining walls as long as the replacement ratio is controlled to less than 33%. Beyond that the casting on site may become difficult. The previous studies of CC have been confined to the static tests conducted at room temperature. From the life-span risk management perspective, the fire-related behavior of CC definitely needs to be investigated. To that end, the present study was designed to characterize the residual mechanical behavior of CC after high-temperature exposure. In parallel with this research, a number of experimental studies have reported the post-fire response of RAC. In the earlier studies of Teranishi et al. [33] and Yang and Hou [34], the residual compressive strength of RAC samples after exposure to high temperatures was observed to be lower than that of natural aggregate concrete (NAC). Eguchi et al. [35] investigated the fire-induced spalling of RAC with a replacement percentage up to 100%. They observed no explosive spalling from either RAC or NAC during heating, thus the fire

performance of RAC could be considered comparable to that of NAC. Similar conclusions were drawn by Vieira et al. [36], by Sarhat and Sherwood [37], and by Zhao et al. [38]. Those authors agree that there is no clear correlation between the residual mechanical properties after heating and the coarse RA replacement ratio. After being exposed to 400, 600 and 800 °C, the compressive strength of conventional concrete and concrete made with 100% RA decreased similarly. In summary, most prior studies have reported that both NAC and RAC suffered significant deterioration due to the thermal exposure; yet the residual mechanical properties of RAC were similar to, or at least only slightly worse than those of NAC. However, there is still significant evidence that leads to the contradictory conclusions. A group led by Kou and Poon [39] investigated the residual mechanical properties and the durability after high-temperature exposure of RAC with replacement ratios of 0, 50 and 100%. They found that NAC suffered more deterioration than RAC after heating to 500 or 800 °C, and that the mortar in NAC exhibited a larger increase in porosity than that in RAC. They attributed this finding to the fact that the thermal expansion coefficients of RA and new mortar are similar, reducing the amount of thermal micro-cracking. Similarly, Xiao and Zhang [40] confirmed that the beneficial effects of incorporating RA were more evident at higher replacement percentages. Xuan et al. [41] also noted that after exposure to 600 °C the thermal deterioration of NAC was more serious than that of RAC. More recently, Khaliq [42] has pointed out that recycled highstrength concrete shows better fire performance in terms of its mechanical properties and physical stability at elevated temperatures than concrete containing natural aggregates. Another relevant yet unclear aspect is that, in assessing the residual strength of any fire-damaged concrete, the rules for comparing and translating residual compressive strength relating samples with different shapes are still not established, even for NAC. In fact only very few investigations are available on this issue. Thus it is not surprising that inconsistent results have been reported. Li and Guo [43] showed that the residual compressive strength of cubes and prisms declines at the same rate after exposure to high temperature. However, Xiao et al. [44] observed a different trend. They found that the cylinder-to-cube 63

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strength ratio decreased with the temperature of exposure. Bamonte and Gambarova [45] also suggest a reduction of the conversion factor for translating the residual strength of cubes to that of cylinders. For CC under current consideration, incorporating large DCLs has intrinsic size and shape effects on the compressive strength even at ambient temperatures [30–32]. After fire exposure, the strength-conversion rules for CC may differ from those for NAC, or for CC itself at room temperature. Moreover, the post-fire deformation feature of CC is equally important, since it directly relates to safety and repair decisions of structures after a fire. Thus the objective of this study is: (i) to clarify the effect of elevated temperature and replacement ratio on the residual mechanical properties of CC containing DCLs; (ii) to investigate any shape effect inherent in CC after thermal exposure. This effect is reflected primarily through the strength-conversion rules; (iii) to identify the deformation characteristics and to suggest a stress–strain relationship of CC after exposure to high temperatures. The unique findings of CC in the above aspects in comparison with NAC or RAC also constitute the novelty of this work.

DCLs was thus determined to be 31.0 MPa derived from the average result of the core tests. The water absorption ratio of the DCLs was 4.10% measured in accordance with ASTM standard C127-12 (2012) [48]. The FC was from one batch of ready-mix concrete provided by a local supplier. The mix proportions are provided in Table 2. River sand and crushed granite (with a maximum particle size of 25 mm) were used as fine and coarse aggregate in the FC. Chemical analysis showed that the fine and coarse aggregate in the DCLs was also siliceous as that in the FC. Table 3 illustrates the primary composition of the FC and DCLs. As mentioned before, to capture the specimens’ temperature rise in the fire tests, additional FC specimens (termed the companion specimens) were fabricated. Chromel-alumel (Type-K) thermocouples were embedded in those specimens before casting. During the casting of the CC specimens (Fig. 2), the DCLs were first pre-wetted by sprinkling tap water on the lumps intermittently. After that, DCLs and FC were alternately placed into the molds while vibrating the mix continuously to ensure compactness. All of the specimens were cured in air for 120 days before the heating tests.

2. Experimental procedure

2.2. Heating procedure and temperature measurements

2.1. Specimens

Except for the unheated specimens, the rest of specimens were heated in batches to the target exposure temperatures (200, 400, and 600 °C). A large, circular electric furnace (internal size: 1100 mm in diameter and 1500 mm deep; see Fig. 3) was used which had a heating capacity up to 1200 °C. The furnace temperature was controlled in the fire tests by a programmable controller reading temperatures from two shielded thermocouples located at the mid-height of the furnace. According to the current design codes (e.g., the European standard EN1992-1-2 (2004) [49]), concrete after exposure to 500 °C will be heavily damaged and its contribution to structural resistance can be neglected. In view of that, the maximum exposure temperature for all samples was taken as 600 °C. This was also adopted by some other investigators (e.g., Zega and Maio [50]) and would not impair the current research conclusions. The heating rate reported in the literature [e.g., 6, 37, 50, 51] varies but commonly lies in the range of 1–5 °C/min. In this study a moderate rate of 3.5 °C/min was used for the initial heating until the furnace temperature reached the desired target value. Subsequently the furnace temperature was kept constant. It was observed that in the later stage of heating the specimen center temperature increased extremely slowly (i.e. entering into the steady thermal state [38,43]). Therefore the heating was terminated once the recorded center temperature reached 95% of the furnace temperature. After that the specimens were cooled naturally to room temperature. Fig. 4 gives the specimen center temperature as well as the furnace temperature varying with time. A plateau (or a slight drop) in the center temperature curves appeared at about 100–150 °C. This might indicate the release and evaporation of moisture in the concrete [52]. Once the free moisture had been evaporated completely the center temperature recovered and rose again. Noticeably, the specimen's shape had an effect on the center temperature variations. As shown in Fig. 4, the cubes and cylinders appeared to be heated more quickly than the prisms, especially after the temperature plateau (or drop). This was because the prism was longer than the cube and larger than the cylinder, hence retarding the heat absorption to some extent.

The mechanical test program consumed 72 CC specimens made of DCLs and FC, accompanied by 54 reference specimens cast with FC alone. The CC specimens were manufactured by substituting DCLs for some of FC at proportions of 20% and 30%. Table 1 summarizes the main test parameters. Each individual specimen was designated with a label “##D-η-T”, where “##” refers to the shape of the specimen (CU: cube; CY: cylinder; or PR: prism); “D” indicates the lateral dimension of the specimen (the edge length for a cube or prism, and the diameter for a cylinder); “η” indicates the replacement ratio of DCLs (i.e. the weight of DCLs as a fraction of the total weight of a CC specimen); and “T” represents the exposure temperature. For the cylindrical and prismatic samples, the height-to-lateral dimension ratio was uniformly set to 2.0. Thus, taking “CY200-0.3-400” for example, it indicates a ϕ200 × 400 mm cylindrical specimen with a replacement ratio of 30% and exposed to a maximum temperature of 400 °C. All of the specimens were grouped as shown in Table 1. Group 1 comprised only FC specimens which were not heated and thus served as controls. Group 2 included the FC specimens with the same lateral dimensions but experiencing different thermal exposures (20, 200, 400 or 600 °C). That group was further divided into three shape sub-groups: the cubes, the cylinders and the prisms. For convenience of comparison, the FC specimens listed in the last three rows of Group 1 are relisted in each sub-group of Group 2. Groups 3 and 4 were designated as in Group 2 except that the samples included DCLs (a 20% replacement ratio for Group 3; 30% for Group 4). Each sample was cast and tested in triplicate to check the reproducibility of results. That generated a total of 126 specimens as is shown in Table 1. Additional replicate specimens were also prepared for two purposes: (i) to monitor the specimens’ center temperature (which added 24 FC specimens); and (ii) to replace specimens which possibly spalled too severely to conduct subsequent compressive loading. The DCLs were sourced from a concrete road demolition site. Waste slabs of concrete were collected and crushed into smaller pieces 65–90 mm in equivalent diameter. The average size of the DCLs was approximately 80 mm, suitable for many applications. The compressive strength of the DCLs was determined by compressing 15 ϕ100 × 100 mm core samples drilled at the site. As specified in the Chinese standards CECS 03 (2007) [46] and BS EN 13791 (2007) [47], the strength of 150 mm standard cubes can be considered equivalent to that of 100 mm cylinder cores. The cube compressive strength of the

2.3. Compression testing The compression tests for the unheated and heated specimens were conducted on a 4000 kN capacity MATEST loading machine with Class 1 precision according to the European standard EN ISO 7500-1 (2004) [53]. Two TML linear variable differential transducers (LVDTs) with a 64

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Table 1 Details of specimens. Group

Specimen

Lateral dimension D (mm)

Replacement ratio η (%)

Exposure temperature T (°C)

Number of specimens

Group 1

CU150-0-20 CY150-0-20 PR150-0-20 CU200-0-20 CY200-0-20 PR200-0-20 CU200-0-20 CU200-0-200 CU200-0-400 CU200-0-600 CY200-0-20 CY200-0-200 CY200-0-400 CY200-0-600 PR200-0-20 PR200-0-200 PR200-0-400 PR200-0-600 CU200-0.2-20 CU200-0.2-200 CU200-0.2-400 CU200-0.2-600 CY200-0.2-20 CY200-0.2-200 CY200-0.2-400 CY200-0.2-600 PR200-0.2-20 PR200-0.2-200 PR200-0.2-400 PR200-0.2-600 CU200-0.3-20 CU200-0.3-200 CU200-0.3-400 CU200-0.3-600 CY200-0.3-20 CY200-0.3-200 CY200-0.3-400 CY200-0.3-600 PR200-0.3-20 PR200-0.3-200 PR200-0.3-400 PR200-0.3-600

150 150 150 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200 200

– – – – – – – – – – – – – – – – – – 20 20 20 20 20 20 20 20 20 20 20 20 30 30 30 30 30 30 30 30 30 30 30 30

20 20 20 20 20 20 20 200 400 600 20 200 400 600 20 200 400 600 20 200 400 600 20 200 400 600 20 200 400 600 20 200 400 600 20 200 400 600 20 200 400 600

6 6 6 3 3 3 3 3 3 4a 3 3 3 1a 3 3 3 1a 3 3 3 2a 3 3 3 3 3 3 3 3 3 3 3 1a 3 3 3 3 3 3 3 3

Group 2

Group 3

Group 4

a

+3

+3

+3

+3

+3

These specimens spalled heavily during the exposure, thus were replaced by the duplicates for mechanical test.

Table 2 Mix proportions of the fresh concrete. Water (kg/m3)

Cement (kg/m3)

Coarse aggregate (kg/m3)

Sand (kg/m3)

Fly ash (kg/m3)

Water reducer (kg/m3)

163

272

998

823

83

7.0

Note: The slump of the fresh concrete was about 165 mm. Table 3 Chemical composition of the coarse aggregate and mortar of the FC and DCLs. Concrete type

FC DCL

Component

coarse aggregate mortar coarse aggregate mortar

Composition (%) SiO2

CaO

Al2O3

Fe2O3

MgO

K2O

Na2O

SO3

Others

74.36 60.63 74.27 62.34

1.48 16.42 0.92 15.84

13.58 8.88 14.00 8.11

1.17 1.75 1.62 1.99

0.26 0.31 0.37 0.60

5.75 3.16 5.44 2.33

2.85 1.48 2.80 0.68

0.09 0.35 0.04 0.30

0.46 7.02 0.54 7.81

calibrated range of ± 5 mm and an accuracy of 5 × 10−4 mm were used to measure the axial deformations within the middle two-fifths of the cylindrical and prismatic specimens. Four pairs of 100 mm long strain gauges were glued, at 90° apart, on the surface of each specimen at mid-height to register the longitudinal and transverse strains. The upper and lower ends of the specimens were capped with high-

strength gypsum before loading. Then each specimen was preloaded to 10% of its predicted ultimate capacity to check the test device and the loading alignment [38,44]. After that the concentric compression load was applied under displacement control with a strain rate of 10 × 10−6/s in line with the ASTM C39 (2001) [54]. All of the measurements were automatically recorded by a data acquisition system. 65

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Fig. 2. Casting of specimens.

3. Test results and discussion

cylinders or prisms. This was due to the cubes’ higher surface-to-volume ratio, with thermal energy being absorbed more quickly. The water vapor pressure in the cubes thus increased presumably more quickly, resulting in more spalling. From Fig. 5(b) it can also be found that the CC specimens spalled less seriously than the FC specimens. Table 1 also indicates the cases of heavily spalled specimens. This phenomenon could be probably ascribed to the more porous microstructure of the DCLs. The DCLs generally had a higher volume of pores and micro-cracks than the new concrete, which was confirmed by the mercury intrusion porosimetry. The associated experimental results are summarized in Table 4. Samples for the porosimetry testing were obtained from the cement mortar of the unheated FC and DCLs. As can be seen, the total porosity of the mortar in the DCLs was measured as 16.4%, more than twice that of the mortar in the FC (7.6%). The average pore diameter of the two mortars was close, but for the mortar of the DCLs the pores larger than 50 nm were significantly more. As a matter of fact, spalling in fire is caused mainly by the build-up of pore pressure in concrete [55,56]. Higher porosity promotes water vapor movement, which lowers the pore pressure during heating. That explains why the spalling of the CC specimens was less severe compared to the FC specimens.

3.1. Visual inspection after heating Damage to concrete after fire exposure can be roughly detected by inspecting the concrete's surface. Fig. 5(a) compares the surface characters of selected pairs of FC and CC cylinders exposed to different temperatures. The appearance and color change of the CC specimens were generally similar to those of the reference FC specimens. The residual color of both after exposure to 200 °C was light grey, similar to that of the unheated specimens, and no noticeable cracks on the surface were observed. After heating to 400 °C the surface color changed to grey-white. Hairline cracks were noticed, but no explosive spalling. After 600 °C exposure both types of specimens were pink and showed more extensive thermal cracking. The color changes observed for the current specimens are overall consistent with literature reports [52,55]. Temperature-induced spalling was found only in some of the specimens heated to 600 °C. All of the severe spalling cases are shown in Fig. 5(b) and Table 1. Those specimens spalled generally when the center temperature reached 300–350 °C, as evidenced by the loud noises heard during the heating tests. Referring to Fig. 4, the spalling time corresponded to the moment when all the free water had escaped and the center temperature was increasing rapidly. Fig. 5(b) shows that the spalling was concentrated on the corners or ends of samples, at which heats could invade from multiple directions. Moreover, most of the spalled specimens were cubes rather than

3.2. Failure modes in compression Crack propagation in three typical CC specimens (CU200-0.2-400, CY200-0.2-400 and PR200-0.2-400) in the course of compressive loading is shown in Fig. 6, which also shows the final failure pattern of

Fig. 3. Furnace used in this study. 66

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Fig. 4. Temperature curves of companion specimens and furnace.

Fig. 5. Surface character and temperature-induced spalling.

67

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at 80–90% of the peak stress. With further loading the cracks continued to grow and new fine cracks formed. The resistance decreased rapidly after the peak, but not so acutely as that in the cylinders. Ultimately the lateral sides of the specimens spalled, leading to an hour-glass failure mode. The temperature exposure seems to have had no significant effect on the sequence of failure progression, except that more surface cracks appeared on the specimens that had experienced higher temperature (600 °C), as shown in the rightmost photographs in Fig. 6. The interior morphologies of some typical fire-exposed CC specimens are shown in Fig. 7. With exposure to higher temperature, microcracks concentrated increasingly along the aggregate–mortar interfaces. This might arise from the thermal incompatibility between the aggregates and the matrix. The above damage localization was more prominent at temperatures higher than 400 °C. After that temperature of exposure, disintegration of the aggregate–mortar assemblage at some regions was more prone to occur under mechanical stresses. Nevertheless, the cohesion between DCLs and FC was still sufficiently strong, corroborated by the observation that the DCLs were not

Table 4 Pore parameters of the cement mortar of the FC and DCLs at room temperature. Concrete type

FC DCL

Porosity (%)

7.6 16.4

Average pore diameter (nm)

Pore volume (ml/g) ≤50 nm

50–100 nm

> 100 nm

8.5 8.8

0.0340 0.0334

0.0022 0.0033

0.0041 0.0097

the corresponding reference FC specimens (CU200-0-400, CY200-0-400 and PR200-0-400). As can be seen, the damage progression and failure mode of the CC specimens were quite similar to those of the corresponding FC specimens. In the case of cylindrical specimens, when the applied load reached 75–90% of the ultimate capacity, cracks began to develop near the midheight. They then propagated towards the specimens’ ends, producing a vertical columnar crack map. At the final stage the cracks bridged together and several small pieces of concrete spalled off. As for the cubic and prismatic specimens, initial cracking took place

Fig. 6. Damage progression and failure modes in compression tests. 68

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Fig. 7. Interior damages of CC specimens.

delaminating from the FC skeleton during the mechanical loading after fire exposure to any test temperature. As Fig. 7 shows, the microcracking was spread over the CC specimens, rather than lumped solely at the DCLs–FC boundary faces. This is possibly because of the thermal compatibility between the new and old cement mortar, both in essence being siliceous (see Table 3).

3.3.1. Compressive strength Table 5 summarizes the measured averaged compressive strengths of the unheated and heated specimens in all groups. In the table the symbols “f”, “E” and “ ˜ ” indicate the compressive strength, the elastic modulus and the peak strain, respectively; the subscripts “cu”, “cy”, and “pr” denote, respectively, the cubic, cylindrical and prismatic specimens; the subscripts “CC” and “D” denote a CC specimen and its lateral dimension, respectively. Table 5 clearly suggests that the compressive strength of the CC and FC specimens decreased with exposure to higher temperature. Moreover, the loss in strength was not proportional to the exposure temperature. From 20 °C to 200 °C the compressive strength decreased only slightly (1.5% for cubes, 3.4% for cylinders, and 9.1% for prisms,

3.3. Residual mechanical properties In this subsection, a range of residual mechanical properties of CC after thermal exposure are discussed and quantified. Also, comparison of the behavioral difference, if any, between the fire-exposed CC, RAC and NAC is attempted. 69

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Table 5 Measured and calculated residual mechanical properties of the specimens in Groups 1 to 4. Cubes

fcu,

CU150-0-20 CU200-0-20 CU200-0.2-20 CU200-0.3-20 CU200-0-200 CU200-0.2-200 CU200-0.3-200 CU200-0-400 CU200-0.2-400 CU200-0.3-400 CU200-0-600 CU200-0.2-600 CU200-0.3-600

Cylinders

D, cc

Avg. ± Std.

Cal.

41.7 39.3 39.4 39.3 39.6 38.4 38.2 32.4 29.6 30.1 20.8 19.9 20.1

– 39.3 37.3 36.3 37.9 36.0 35.0 31.5 29.9 29.1 20.1 19.0 18.5

± ± ± ± ± ± ± ± ± ± ± ± ±

1.6 0.6 0.3 0.1 0.9 0.9 0.2 1.6 1.3 0.8 2.5 1.0 1.5

CY150-0-20 CY200-0-20 CY200-0.2-20 CY200-0.3-20 CY200-0-200 CY200-0.2-200 CY200-0.3-200 CY200-0-400 CY200-0.2-400 CY200-0.3-400 CY200-0-600 CY200-0.2-600 CY200-0.3-600

fcy,

Ecy,

D, cc

˜ cy,

D, cc

Prisms

D, cc

Avg. ± Std.

Cal.

Avg. ± Std.

Avg. ± Std.

35.0 32.1 31.1 29.9 31.5 29.6 28.9 24.5 22.1 22.0 12.8 12.0 11.7

– 32.1 28.5 26.8 29.6 26.3 24.7 23.2 20.6 19.4 13.9 12.3 11.6

28.5 ± 0.5 27.8 ± 0.7 25.5 ± 2.8 21.3 ± 1.0 16.8 ± 0.6 16.4 ± 0.3 12.5 ± 2.2 6.3 ± 1.1 5.1 ± 1.3 5.0 ± 0.9 1.3 ± 0.1 1.2 ± 0.1 1.1 ± 0.1

2268 ± 63 2437 ± 80 2890 ± 756 3291 ± 179 2820 ± 226 3300 ± 186 3855 ± 563 5216 ± 416 5389 ± 541 5960 ± 362 10187 ± 765 11040 ± 336 11443 ± 452

± ± ± ± ± ± ± ± ± ± ± ± ±

0.8 1.4 0.5 1.6 0.4 1.2 1.1 0.4 1.2 1.0 0.4 0.9 1.1

PR150-0-20 PR200-0-20 PR200-0.2-20 PR200-0.3-20 PR200-0-200 PR200-0.2-200 PR200-0.3-200 PR200-0-400 PR200-0.2-400 PR200-0.3-400 PR200-0-600 PR200-0.2-600 PR200-0.3-600

fpr,

Epr,

D, cc

D, cc

˜ pr,

D, cc

Avg. ± Std.

Cal.

Avg. ± Std.

Avg. ± Std.

34.9 33.3 34.0 33.7 31.7 30.3 29.8 24.5 22.9 23.1 13.9 13.7 13.9

– 33.3 30.8 29.5 30.7 28.4 27.3 24.1 22.3 21.4 14.4 13.3 12.8

28.7 ± 0.6 29.1 ± 1.5 28.0 ± 2.3 27.3 ± 0.3 17.2 ± 1.9 16.7 ± 0.4 15.9 ± 0.7 6.5 ± 0.3 5.9 ± 0.5 5.9 ± 0.5 1.4 ± 0.2 1.5 ± 0.0 1.5 ± 0.1

2218 2357 2413 2536 2877 2907 2946 4981 5132 5049 9271 9684 9429

± ± ± ± ± ± ± ± ± ± ± ± ±

1.0 0.7 0.3 1.8 1.1 2.1 0.3 0.6 0.7 1.3 0.7 1.7 0.4

± ± ± ± ± ± ± ± ± ± ± ± ±

60 120 236 318 432 194 328 295 289 460 387 304 379

Note: (1) Avg. = average value; Std. = standard deviation; Cal. = calculated value; (2) Units: compressive strength in MPa; elastic modulus in GPa; peak strain in με.

averaged for both the CC and FC specimens). However, once exceeding 400 °C a marked strength loss of at least 20–30% was witnessed. After heating at 600 °C, the deterioration percentages for cubes, cylinders, and prisms were, respectively, 48.4%, 60.7%, and 58.9% on average. The strength loss is closely related to the physical and chemical changes occurring in concrete at elevated temperatures [57–62]. As has been reported, before 200 °C only a minor loss in strength takes place. Starting approximately at 400 °C the dehydration action commences to weaken the bonds in the calcium-silicate-hydrate (C-S-H) gel. Up to 600 °C, decomposition of the calcium hydroxide has been noticeable, together with the thermal expansion of aggregates. The microstructure of concrete is thus significantly damaged and the loss of macroscopic strength becomes appreciable. The changes in the relative compressive strength for CC samples are related to the exposure temperature, replacement ratio and specimen shape, as plotted in Fig. 8(a)–(c). The relative strength was calculated as the average percent retained strength of the heated specimens with respect to the average strength of the unheated reference specimens. As clearly shown, the relative strength decreased significantly with temperature, which was definitely the most affecting factor among the three test variables. From Table 5 and Fig. 8(a)–(c) it is also found that at ambient temperature the absolute strengths of the CC specimens were comparable to, or slightly lower than those of the control FC specimens. After different thermal exposures the CC specimens’ compressive strengths were always lower than those of their FC counterparts. However the relative strength was only moderately influenced by the incorporation A

,f

=f

T , 200, CC / f , 200, CC

=

1.00 + 0.15 × T /1000 1.00 0.07 × T /1000

1.61 × (T /1000)2 1.46 × (T /1000)2

temperature-induced deterioration of properties, the end effects contributed more to the strength than they would at room temperature, especially for cubes and at higher temperatures. As a result, the CC cubes were capable of retaining half of their strength even after being exposed to 600 °C. To be more clear, Fig. 8(d) compares the relative strengths of the three shapes of CC samples after exposure to high temperatures. The predictions using the European standard EN1992-1-2 (2004) [49] that is specific only for the natural siliceous or calcareous aggregate concrete cylinders are also plotted. Apparently, the code's prediction is still well correlated with the test results of the CC cylinders and prisms; yet for the CC cubes neither of the code's two predictions (siliceous and calcareous aggregates) can agree with the test data. As a supplement, an equation was derived empirically from the current test results to evaluate the relative strength of CC cubes (with 200 mm edge-length) after fire exposure:

Acu,f = fcu,

T 200, CC / fcu, 200, CC

1.61 ×

(T /1000)2

= 1.00 + 0.15 × T /1000 (1a)

where the variable T is the exposure temperature. A good fit of this equation to the data of the CC cubes is obtained (with a correlation coefficient R2 = 0.969), as clearly shown in Fig. 8(d). Thus, in combination with the Eurocode equation that has achieved a good correlation for the CC cylinders and prisms, the relative strength of the CC samples with different shapes and with a lateral dimension of 200 mm after fire exposure up to 600 °C can be summarized as:

for cubes for cylinders and prisms

of DCLs for all exposures and all shapes. Comparatively, the shape effect on the relative strength was nonetheless remarkable. At 400 °C, for the CC cylinders and prisms the average relative strengths were, respectively, 73.6% (Fig. 8(b)) and 69.8% (Fig. 8(c)), both lower than that for the CC cubes (78.1%, Fig. 8(a)). After heating at 600 °C the discrepancy became more evident—39.3% for cylinders, a similar magnitude of 41.1% for prisms, and a notably higher value of 51.6% for cubes. The CC cubes' better retention of their compressive strength might be attributed to the end restraint effect from the friction between the machine platens and the specimens. The friction forces restricted the samples’ lateral expansion and provided confinement to the local concrete adjacent to the platens, thus enhancing the overall bearing capacity. For the short cubes the end restraint effect was more pronounced than for the cylinders and prisms [63,64]. Further, due to the

(1b)

where A*, f is the relative strength; the subscript “*” refers to the specimen's shape (“cu” for cubes; “cy” for cylinders, and “pr” for prisms). Consequently, fcy, 200, CCT, for example, indicates the residual strength of a 200 mm CC cylinder. To see more clearly any difference in the relative compressive strengths of NAC, RAC, and CC, some typical and well-documented test results are collected in this study [38,41,50,58,61,65], as presented in Fig. 8(e). To be comparable, only samples made with siliceous aggregates are included. From Fig. 8(e), it is clear that (especially from the fitting curves) the decreasing rate of the relative compressive strength of the RAC samples was slightly lower than that of the NAC and CC ones, especially above 200 °C; for the CC samples, its relative compressive strength was similar to the NAC ones. Note that the interfacial transition zones (ITZs) are commonly believed to have a thermal resistance effect to prohibit heat transfer [66]. This may 70

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Fig. 8. Relative compressive strength.

explain the positive effect of the relatively loose ITZs dispersed in RAC which led consequently to a better retention in strength as compared to NAC. However, as CC had less amount of, and more concentrated loose ITZs (at the DCLs–FC boundaries) compared to RAC, the former’ relative strength was different from that of the latter, but varied similarly to that of NAC.

Note that Eq. (1b) is only applicable to the 200 mm CC specimens. This size is deemed suitable considering the relatively large size of DCLs. However, in practice the 150 mm cylinder or cube strength is routinely used. Thus a bridge should be constructed. Our group has summarized the following formulae for predicting the compressive strength of 200 mm cubes, cylinders and prisms of CC 71

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at room temperature based on a large volume of test data [32].

f,

= R × [f*,

200, CC

R = 1.0

200, FC

× (1

) + f*,

200, DCLs

ratio increases from 0 to 30%. Such an increase is consistent with the results of a previous study [32]. It is presumably due to the greater lateral dilation of CC compared to FC which would result in a more impact of the end effect on the compressive strength for the cubic CC specimens. The following similar expression can be established to relate fpr, 200, T T CC to fcy, 200, CC .

(2a)

× ]

(2b)

×

where R* is a factor accounting for the shape effect. The subscript “*” again represents the specimen's shape; η refers to the replacement ratio; α* is a constant that can be taken as 0, 0.32 or 0.13 for CC cubes, cylinders and prisms, respectively (see Ref. [32] for more details); f*, 200, FC and f*, 200, DCLs are, respectively, the 200 mm sample strengths of the FC and the DCLs, respectively. The strength f*, 200, DCLs can be related to (in only a nominal sense) the 150 mm cube strength of the DCLs (i.e. fcu, 150, DCLs, normally obtained by core tests in practice) by employing the following relationship:

f*,

=

200, DCLs

f*,

200, FC

f*,

150, FC

× f*,

150, DCLs

fpr, 200, CC T fcy, 200, CC T

fpr, 150, DCLs =

fcy,

150, FC

fcu, 150, FC fpr,

150, FC

fcu, 150, FC

150, DCLs

(3b)

× fcu,

150, DCLs

(3c)

Equations. (3a)–(3c) have been established based on the assumption that both the scale- and shape-effect laws affecting the strengths of FC at room temperature are also applicable to DCLs. That has been validated in previous studies [30–32]. Inserting Eq. (2a) into Eq. (1) then yields the following expression:

f,

200, CC

(0

T

=A

, f

× R × [f*,

30%, 20 ° C

T

200, FC

× (1

) + f*,

200, DCLs

× ] (4)

600 °C)

With Eqs. (1)–(4) at hand, the residual strengths of 200 mm CC cubes, cylinders, and prisms can be expeditiously predicted using the strength of 150 mm samples of the DCLs and FC at room temperature. During the prediction, the ratio f*, 200, FC/f*, 150, FC in Eq. (3a) can be obtained according to the code specifications. In this research, the measured values of f*, 200, FC and f*, 150, FC (see Table 5) were instead used. The predicted results were then compared with the experimental outcomes, as shown in Table 5. Clearly the predictions correlate well with the measurements. The last unsolved issue on the compressive strength of fire-exposed CC is the shape effect. To reflect that, the strength conversion factors considering the effects of replacement ratio and exposure temperature are developed. Combining Eqs. (1), (2) and (4), the following expression can be obtained linking fcu, 200, CCT with fcy, 200, CCT: fcu, 200, CCT fcy, 200, CCT

=

fcu, 200, FC fcy, 200, FC

1

1.23 0.32 ×

×

Rcu Rcy

×

Acu, f A cy, f

=

1

1.23 0.32 ×

×

1.00 + 0.15 × T / 1000 1.00

0.07 × T / 1000

fcy, 200, FC

×

Rpr Rcy

×

Apr, f

= 1.03 ×

A cy, f

1.03 × (1 0.13 × ) 0.32 × + 0.0247 ×

2

=

1

1 1

0.13 × 0.32 × 1.03 0.19 ×

(6)

Note that because the ratio Apr, f/Acy, f in the above equation is unity (cf. Eq. (1)), the resulting prism-to-cylinder strength ratio is independent of the exposure temperature. Fig. 9(b) shows the predicted and experimental strength ratios varied with the replacement ratio and the exposure temperature. Evidently the predictions agree well with the test results. Moreover, the prismatic strength of CC is generally higher than the cylinder strength. Similar relationships have been observed in previous strength studies on NAC [67,68]. With greater replacement the prism-to-cylinder strength ratio increases, analogously to the former case of the cube-tocylinder strength ratio. The above observations may be explained by the end restraint effect. That effect is more pronounced for the prisms than for the cylinders due to the former’ larger end area. The additional outer concrete of the prisms may have a confining effect on the inner concrete, thus enhancing the prisms’ load-bearing capacity [69]. Besides, the direction of casting may have an effect. Referring back to Fig. 2, the prism specimens were cast horizontally but the cylinders vertically. The prisms were intended to simulate the casting of CC beams or precast CC members cast flat in plant. The quality of horizontally-cast specimens is often slightly better than that of vertically-cast ones [67,68]. With more DCLs being incorporated, the effect of casting direction would become more significant, eventually contributing to an increased prism-to-cylinder strength ratio.

(3a)

× fcu,

fpr, 200, FC

1

where the subscript “150” indicates a 150 mm sample. If the shape indicator “*” is “cu”, a link between fcu, 200, DCLs and fcu, 150, DCLs can immediately be created. When it is “cy” or “pr”, the following additional transition equations are needed.

fcy, 150, DCLs =

=

3.3.2. Elastic modulus The elastic modulus was quantified in the compression tests using the widely-adopted equation (σ2 – σ1)/(ε2 – 0.00005) (in MPa), where σ1 and σ2 are the stresses corresponding to the 0.00005 longitudinal strain and 40% of the predicted ultimate strength, respectively, and ε2 is the longitudinal strain at the stress σ2 [70]. The resulting average values are tabulated in Table 5. It is clear that the elastic modulus decreased with exposure to higher temperatures. Also, the elastic modulus was consistently negatively affected by the incorporation of DCLs. The higher the replacement level, the lower the elastic modulus. In comparison with the compressive strength, incorporating DCLs had a more significant effect on the elastic modulus below 200 °C. At higher temperatures, however, that adverse effect attenuated. At 600 °C the elastic modulus seemed no longer sensitive to the replacement ratio. The relative elastic moduli for the cylinders and prisms are presented in Fig. 10(a) and (b). That normalized modulus was defined as the ratio of the residual modulus after thermal exposure to that at room temperature. From the figure it can be concluded that the replacement ratio had no significant effect on the relative modulus, nor did the specimen's shape have much effect. To quantify the modulus deterioration of CC after high-temperature exposure, the following formula is developed through curve fitting:

1.61 × (T / 1000)2 1.46 × (T / 1000)2

× 1.00 + 0.08 × T /1000 + 0.35 × (T /1000)2]

(5) Fig. 9(a) presents a data grid comparing the predicted and measured values of the strength ratio fcu, 200, CCT/fcy, 200, CCT at various replacement ratios η and exposure temperatures T. The figure shows that the predictions are generally acceptable and that as the exposure temperature rises, the cube-to-cylinder strength ratio increases, only slightly below 400 °C, but markedly afterwards. This is due primarily to the end effect previously discussed. Interestingly, the strength ratio also shows a slight increasing tendency with the replacement ratio. An increase of 3.7–6.5% is observed for all exposures as the replacement

A*,

E

= E*,

200, CC

+ 1.66 ×

T/E

*, 200, CC

= 1.05

(T /1000)2 (20

°C

2.68 × T /1000 T

(7)

600 °C) E*, 200, CCT

where the subscript “*” indicates a prism or cylinder; is the residual elastic modulus of the CC after being heated to temperature T; 72

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Fig. 9. Shape effect on residual compressive strength.

decreased slightly and then recovered, presenting a non-linear variation. At higher replacement ratios the modulus ratio showed an increasing tendency on the whole. This may be caused by the casting direction effect which made the prisms stiffer than the cylinders, and that effect became more significant when more DCLs were incorporated.

and E*, 200, CC is the elastic modulus of unheated CC. The correlation coefficient R2 of this proposed formula is 0.987 for the test data in this study. Some previously published experimental results [6,71,72] are also presented in Fig. 10(c) for comparison. It can be seen that the proposed formula performs better than the Eurocode in predicting the relative elastic modulus. The modulus ratio Epr, 200, CCT/Ecy, 200, CCT is plotted in Fig. 11(a) against the replacement ratio and the exposure temperature. It is clear that as the exposure temperature increased the modulus ratio first

3.3.3. Peak strain The peak strains of the cylindrical and prismatic CC samples

Fig. 10. Relative elastic modulus. 73

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Fig. 11. Shape effect on residual elastic modulus and peak strain.

registered in the compression tests are given in Table 5. As expected, the exposure temperature increased the peak strains, especially after the critical temperature 400 °C where significant physical and chemical changes in concrete begin [52,59]. The average peak strains at 400 °C and 600 °C were, respectively, 2.00 and 3.86 times the peak strains at room temperature. The peak strains increased with the DCL content, in general. This was mainly due to the greater deformation of DCLs associated with their lower strength than that of FC. Nevertheless, as the exposure temperature increased the effect of the replacement ratio was smaller overall, similarly to the previous modulus observations. All in all, the fire-induced acute deterioration diminished the effect on deformity caused by the incorporation of DCLs. Fig. 12(a) compares the influences of the replacement ratio and the specimen's shape on the relative peak strain (the ratio of the peak strain recorded after thermal exposure to that at room temperature) for the CC cylinders and prisms. It is evident that, in general, the relative peak strain was not significantly affected by the specimen's shape, while the role of the replacement ratio was dependent on the temperature. At 200 °C the effect of the replacement ratio on the relative peak strain was minimal. At higher temperatures, however, the effect of incorporating DCLs became progressively more significant. At 400 and 600 °C the relative peak strain generally decreased with the replacement ratio. Fig. 11(b) exhibits how the peak strain ratio ˜ pr, 200, CCT/ ˜ cy, 200, CCT changed with the replacement ratio and the exposure temperature. It shows that the peak strain of the CC prisms was generally smaller than

that of their cylindrical counterparts. This could be explained by the casting direction effect. As the replacement level increased the disparity became larger. And the peak strain ratio first increased and then decreased as the temperature increased, another non-monotonic variation as with the former case shown in Fig. 11(a). Curve fitting suggests the following equation for predicting the peak strain of CC after thermal exposure:

A*,

= ˜*,

200, CC

+ 10.6 ×

T /˜ *, 200, CC

= 1.05

(T /1000)2 (R2

1.74 × T /1000 (8)

= 0.974) T

where the subscript “*” refers to a prism or a cylinder; ˜ *, 200, CC and ˜ *, are the peak strains of CC after elevated temperature exposure and at room temperature, respectively. That expression is plotted in Fig. 12(b) together with the current and some previous relevant test data [6,51,73]. As seen, the proposed predictor agrees better with the test values than the code prediction. 200, CC

3.4. Deformation characteristics 3.4.1. vol dilation Concrete exhibits a notable volume change (dilation) in the inelastic range. For CC the volume dilation is of especial interest and significance, since confining devices such as steel tubes are often needed to suppress the potential weakness induced by the incorporation of DCLs [74]. The efficacy of such confinement is closely related to the volume dilation.

Fig. 12. Relative peak strain. 74

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Fig. 13. Volume dilation characteristics.

Fig. 13(a) and (b) are experimental stress–volumetric strain curves prepared to check the effects of the replacement ratio and exposure temperature on the volume changes occurring in CC after fire exposure. Each data point is the average result of three identical samples. The stress is normalized with respect to the peak stress. The volumetric strain is defined as ε1 + 2ε2, where ε1 and ε2 are, respectively, the longitudinal and transverse strains. The normalized critical stress, indicated with the crosses in the figure, is the normalized stress at which the volume change switched from contraction to expansion. It is clear from the figure that the replacement ratio had a significant effect on the dilation behavior of CC. As the replacement ratio increased, the normalized critical stress generally decreased; in other words, the onset of dilation became earlier. Also, the volumetric strain at a specific stress decreased with the replacement ratio. Those observations were understandable due to the greater deformability of the DCLs. As a result, the CC specimens expanded laterally more than the FC specimens. This is also why the restraining effect exerted by the loading platen on the CC specimens was more significant. The exposure temperature affected the CC's dilation greatly. At higher exposure temperatures the normalized critical stress dropped substantially—by about 50–60% from 20 to 600 °C (see Fig. 13(b)). At the same time, the stress–volumetric strain curves became more and more rounded and convex, indicating a greater increase in the volumetric strain and a smoother transition in the volume changes. The variation of the normalized critical stress with the replacement level was very limited at 600 °C, as Fig. 13(a) clearly shows. This is consistent

with the previous conclusion that exposure to 600 °C tends to generally reduce the effect of replacement ratio on CC's post-fire deformations. 3.4.2. Longitudinal stress–strain relationship The longitudinal stress–strain curves of the CC and FC prisms after various temperature exposures are shown in Fig. 14(a). The CC specimens show several features in common with the FC specimens. As the exposure temperature increased, the peak stress and the elastic modulus decreased, the peak strain increased, and the slopes of the descending branches became more gradual. To clearly examine the effects of the exposure temperature and replacement ratio on the stress–strain curves, Fig. 14(b) and (c) provide non-dimensionalized versions of those curves. For both the CC and FC prisms, as the exposure temperature increased, the slope of the ascending branches declined slightly while the descending branches were steeper on the whole. This accords with the observations published by Liu et al. [62] when investigating the behavior of recycled aggregate concrete. It represents a decrease in the relative ductility of concrete after fire exposure. The pre-peak portion of the curves was not much affected by the incorporation of DCLs, but the softening after the peak was generally greater at larger replacement ratios, for all exposures. Liu et al. [75] reported similar results for CC at room temperature. The inherent weakness of the DCLs became highlighted once the catastrophic post-peak failure was initiated. Fitting the test data suggests the following uniaxial stress–strain constitutive model for reproducing the full-range of CC behavior after 75

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Fig. 14. Longitudinal stress–strain curves and effects of their primary influential factors.

76

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exposure to high temperatures. This empirical model modifies Popovics’ well-accepted relationship for plain concrete [76]. It is expressed in a fractional form:

Y=

nX 1 + Xn

n

T Y = / fpr,

X= / n=

(9b)

T pr, 200, CC

(9c)

Epr,

=1

200, CC

fpr,

None. Acknowledgments The authors would like to acknowledge research grants from China's National Key R&D Program (grant 2017YFC0803300), China's National Natural Science Foundation (grant 51438007), Guangzhou's Key Project of Science (Technology) Research (grant 201607020005), and the State Key Laboratory of Subtropical Building Science (grants 2017KC17, 2017ZB30 and 2018ZC03).

200, CC 200, CC / pr, 200, CC

0.11 × T /1000 + 0.77 × (T /1000)2

= 1 + 0.31 ×

Declarations of interest

(9a)

200, CC

Epr,

behavior under axial load. The ratio of the critical load to the peak load exhibited a decreasing trend with the DCL content. But such effect became less important after 600 °C exposure.

(9d) (9e) (9f)

where σ and ε are, respectively, the CC's compressive stress and strain; α and β are calculated parameters empirically derived and accounting for the influence of the exposure temperature and the replacement ratio, respectively. To validate that above model, the predicted stress–strain curves were compared with the test curves in Fig. 14(a). There was good agreement between the predictions and the measurements in terms of the peak stresses as well as the ascending and descending slopes.

Appendix A. Supplementary data Supplementary data to this article can be found online at https:// doi.org/10.1016/j.firesaf.2019.02.008. References [1] M. Wijayasundara, P. Mendis, R.H. Crawford, Integrated assessment of the use of recycled concrete aggregate replacing natural aggregate in structural concrete, J. Clean. Prod. 174 (2017) 591–604. [2] V.W.Y. Tam, M. Soomro, A.C.J. Evangelista, “A review of recycled aggregate in concrete applications (2000–2017), Constr. Build. Mater. 172 (2018) 272–292. [3] J. de Brito, N. Saikia, Recycled Aggregate in Concrete: Use of Industrial, Construction and Demolition Waste, Springer Science & Business Media, 2012. [4] N. Kisku, H. Joshi, M. Ansari, et al., A critical review and assessment for usage of recycled aggregate as sustainable construction material, Constr. Build. Mater. 131 (2017) 721–740. [5] J.Z. Xiao, Recycled Aggregate Concrete Structures, Springer, Berlin, 2018. [6] G.M. Chen, Y.H. He, H. Yang, et al., Compressive behavior of steel fiber reinforced recycled aggregate concrete after exposure to elevated temperatures, Constr. Build. Mater. 71 (2014) 1–15. [7] I.B. Topcu, S. Sengel, Properties of concretes produced with waste concrete aggregate, Cem. Concr. Res. 34 (8) (2004) 1307–1312. [8] W.G. Li, Z.Y. Luo, Z. Tao, et al., Mechanical behavior of recycled aggregate concrete-filled steel tube stub columns after exposure to elevated temperatures, Constr. Build. Mater. 146 (2017) 571–581. [9] C.S. Poon, Z.H. Shui, L. Lam, Effect of microstructure of ITZ on compressive strength of concrete prepared with recycled aggregates, Constr. Build. Mater. 18 (6) (2004) 461–468. [10] J.Z. Xiao, W.G. Li, Z. Sun, et al., Properties of interfacial transition zones in recycled aggregate concrete tested by nanoindentation, Cem. Concr. Compos. 37 (2013) 276–292. [11] C. Thomas, J. Setién, J.A. Polanco, et al., Fatigue limit of recycled aggregate concrete, Constr. Build. Mater. 52 (2014) 146–154. [12] S. Lotfi, P. Rem, J. Deja, et al., An experimental study on the relation between input variables and output quality of a new concrete recycling process, Constr. Build. Mater. 137 (2017) 128–140. [13] H.R. Zhang, Y.X. Zhao, Integrated interface parameters of recycled aggregate concrete, Constr. Build. Mater. 101 (2015) 861–877. [14] R.V. Silva, J. de Brito, R.K. Dhir, Establishing a relationship between modulus of elasticity and compressive strength of recycled aggregate concrete, J. Clean. Prod. 112 (2016) 2171–2186. [15] W.G. Li, Z.Y. Luo, Z.H. Sun, et al., “Numerical modelling of plastic–damage response and crack propagation in RAC under uniaxial loading, Mag. Concr. Res. (2017) 1–14. [16] M. Wijayasundara, P. Mendis, R.H. Crawford, Methodology for the integrated assessment on the use of recycled concrete aggregate replacing natural aggregate in structural concrete, J. Clean. Prod. 166 (2017) 321–324. [17] T.Y. Xie, A. Gholampour, T. Ozbakkaloglu, Toward the development of sustainable concretes with recycled concrete aggregates: comprehensive review of studies on mechanical properties, J. Mater. Civ. Eng. 30 (9) (2018) 04018211. [18] W.G. Li, Z.Y. Luo, C.Q. Wu, et al., Experimental and numerical studies on impact behaviors of recycled aggregate concrete-filled steel tube after exposure to elevated temperature, Mater. Des. (2017) S0264127517309115. [19] W.G. Li, Z.Y. Luo, C.Q. Wu, et al., Impact performances of steel tube-confined recycled aggregate concrete (STCRAC) after exposure to elevated temperatures, Cem. Concr. Com. 86 (2018) 87–97. [20] W.G. Li, J.Z. Xiao, Z. Sun, et al., Failure processes of modeled recycled aggregate concrete under uniaxial compression, Cem. Concr. Com. 34 (2012) 1149–1158. [21] Z.H. Duan, C.S. Poon, Properties of recycled aggregate concrete made with recycled aggregates with different amounts of old adhered mortars, Mater. Des. 58 (2014)

4. Conclusions To evaluate the residual mechanical properties of compound concrete containing DCLs after exposure to high temperatures, a comprehensive experimental program was designed and carried out. The observations from this study suggest the following useful conclusions: (1) Temperature-induced concrete spalling occurred only after 600 °C exposure. Generally, the cubes spalled more seriously than the cylinders and prisms. Moreover, the fresh concrete specimens were more susceptible to the spalling damage than the compound concrete. (2) After thermal exposure, extensive cracking produced along the aggregate–mortar interfaces dispersed in DCLs and FC during the compressive loading of CC specimens, but the DCLs and FC themselves generally remained integrated. (3) Overall, the relative compressive strength and elastic modulus of the CC specimens containing DCLs (up to 30% replacement ratio) were quantitatively similar to those of reference specimens made of FC alone after exposure to temperatures up to 600 °C. Moreover, a comparison indicates that above 200 °C both the CC and NAC specimens showed a slightly decreased capacity of strength retention with temperature compared to the RAC ones reported in the literature. (4) A set of equations are proposed which can predict the residual strength of CC cubes, cylinders and prisms based on the exposure temperature, the 150 mm cube strength of FC and DCLs at room temperature, and the replacement percentage. Equations to predict the elastic modulus and peak strain of CC after fire exposure are also proposed. (5) As the exposure temperature raised, the cube-to-cylinder strength ratio fcu/fcy increased monotonically—slightly below 400 °C, but markedly afterwards. The prism-to-cylinder strength ratio fpr/fcy did not have a similar, monotonic relationship with the temperature. In addition, with an increase in the replacement ratio, both fcu/fcy and fpr/fcy increased. (6) For all exposure temperatures, the ratio of the elastic moduli of prismatic and cylindrical CC specimens generally increased with the replacement ratio, while the peak strain ratio showed the opposite tendency. (7) The incorporation of DCLs had an impact on the CC's dilation 77

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