Experimental studies on the chloride ion permeability of concrete considering the effect of freeze–thaw damage

Construction and Building Materials 236 (2020) 117556

Contents lists available at ScienceDirect

Construction and Building Materials journal homepage: www.elsevier.com/locate/conbuildmat

Experimental studies on the chloride ion permeability of concrete considering the effect of freeze–thaw damage Yuanzhan Wang a, Zhen Liu a,⇑, Kun Fu b, Qingmei Li a, Yuchi Wang c a State Key Laboratory of Hydraulic Engineering Simulation and Safety and Collaborative Innovation Center for Advanced Ship and Deep-Sea Exploration, Tianjin University, Jinnan District, 135 Yaguan Road, Tianjin 300072, People’s Republic of China b Power China Huadong Engineering Corporation Limited, Hangzhou 310014, People’s Republic of China c Tianjin Research Institute for Water Transport Engineering, Ministry of Transport, 2618 Xingang 2nd Road, Binhai New District, Tianjin 300000, People’s Republic of China

h i g h l i g h t s
The effect of minimum freezing temperature on chloride diffusion is investigated.
In tidal zone, frost damage significantly affects chloride diffusion in concrete.
A model is proposed to predict the chloride profiles of frost-damaged concrete.

a r t i c l e

i n f o

Article history: Received 21 June 2019 Received in revised form 28 October 2019 Accepted 8 November 2019

Keywords: Chloride ion permeability Freeze-thaw damage Compressive strength Normalized dynamic elastic modulus Time-dependent model Minimum freezing temperature

a b s t r a c t A series of laboratory experiments were conducted to investigate the effect of freeze–thaw (FT) cycles on mechanical properties and chloride permeability of concrete with the strength class of C30 in this paper. The mechanical property tests were performed on the specimens subjected to 0, 5, 15, 30, 50, 75 and 100 standard FT cycles to determine the compressive strength, normalized dynamic elastic modulus (NDEM) and mass loss. Also, the chloride natural diffusion tests of concrete specimens, after 0, 5, 15, 30 and 50 FT cycles, were conducted in a self-design tidal cycling simulation device to investigate the chloride ingress into concrete in marine environment of tidal zone. Based on the Fick’s second law, a time-dependent model was developed to predict the chloride profiles of frost-damaged concrete. Additional FT cycling tests at different minimum temperatures of freezing process (4 and 11 °C) and the chloride natural diffusion tests were conducted to study the effect of minimum temperature on the chloride diffusion in concrete. The experimental results revealed the significant influence of FT action on the mass loss, compressive strength and NDEM. After 50 FT cycles, the mean of NDEM dropped to 60%, indicating that the specimen was destroyed by FT action. With the decrease of minimum freezing temperature, the loss of NDEM increased apparently, performing a good linear relation. A substantial effect of FT damage on the chloride profiles was observed: compared to the counterparts of the control groups without being exposed to FT attack, chloride concentration was larger for the frost-damaged concrete, and this difference increased with the extension of exposure duration. A linear relation between the apparent chloride diffusion coefficient and the loss of NDEM was also obtained regardless of the exposure time. Moreover, it was found that the proposed model was applicable for the prediction of chloride profiles in concrete subjected to FT cycles at different minimum freezing temperatures. In other words, the proposed model can be used to describe the chloride ingress into frost-damaged concrete, regardless of the cause of the damage (due to the number of FT cycles or the minimum freezing temperature). Ó 2019 Elsevier Ltd. All rights reserved.

1. Introduction

⇑ Corresponding author. E-mail addresses: [email protected] (Y. Wang), [email protected] (Z. Liu), [email protected] (K. Fu), [email protected] (Q. Li), [email protected] (Y. Wang). https://doi.org/10.1016/j.conbuildmat.2019.117556 0950-0618/Ó 2019 Elsevier Ltd. All rights reserved.

In practical engineering, chloride contamination and freeze– thaw (FT) attack are the main causes which lead to the durability deterioration of marine concrete structures in cold regions [1–4]. Based on the Fick’s second law, many investigations have been carried out to quantify the chloride profiles in concrete without con-

2

Y. Wang et al. / Construction and Building Materials 236 (2020) 117556

sidering the effect of FT cycles [5–7]. To some extent, a welldesigned concrete, in which the pores and micro-cracks are discontinuous, is originally watertight, performing a relatively well chloride resistance [8]. When subjected to FT cycles, the frost damage normally leads to the expansion of micro-cracks and the reduction of cover-layer [9,10]. It is obvious that this damage will accelerate the chloride ingress into concrete [11]. The rapid chloride migration (RCM) method, in previous experimental studies, was used by many scholars to investigate the chloride resistance performance of frost-damaged concrete. For instance, Kuosa et al. and Zhao et al. [12,13] indicated that the chloride ingress increased significantly due to frost cracks, and a great linear relation was found between the chloride diffusion coefficient DRCM and normalized dynamic elastic modulus (NDEM). Wang et al. [14] studied the FT influence on the chloride diffusion resistance of concrete containing fly ash and silica fume by the RCM test. Above researches were based on the standard FT cycling tests, in which the minimum freezing temperature was a constant and the FT damage was caused by different number of FT cycles. As a matter of fact, the temperature was also one of the main crucial factors for the frost damage [15]. Ferreira et al. [16] pointed out that the minimum freezing temperature took a considerable effect on the chloride resistance of concrete and a positive linear relationship existed between DRCM and frost damage, in which the damage resulted from different minimum temperatures of freezing process (5, 10 and 20 °C). However, the investigations mentioned above were based on the results of RCM test, which was an electrically accelerated test method [17,18]. Besides, the voltage used in the RCM test generates heat, which in turn affects the diffusion rate of chloride [19,20]. In a word, there exists a difference between the RCM test and chloride natural diffusion test. Many scholars investigated the chloride natural penetration into frost-damaged concrete by immersion tests, in which the specimens were soaked in NaCl solution. Wang et al. [21] studied the chloride natural ingress into concrete by the FT cycling experiments in NaCl solution with the cycle time of 3 h. The results indicated that the chloride diffusion coefficient decreased at first and then increased with FT cycles. Li et al. [22] also conducted FT cycling tests using NaCl solution and developed a mesoscopic chloride diffusion model of concrete exposed to FT cycles. The above investigations simulated the chloride diffusion into concrete in winter when the temperature was relatively low. However, due to this relatively low temperature, chloride diffusion may not be particularly severe in this case. Boddy et al. [23] indicated that the rate of chloride diffusion increased with the increase of temperature. Consequently, after a FT attack in winter, the chloride penetration of concrete structures will be more serious in the summer with a higher temperature. Wang et al. [24] carried out 7-days and 28-days immersion tests to investigate the coupling effect of compressive load and FT cycles on the water absorption and chloride permeability of concrete. As revealed by this literature, the FT cycles took a significant effect on capillary absorption and chloride permeability of concrete regardless of the compressive load. To quantify the chloride profiles in frost-damaged concrete, Zhang et al. [25] conducted 60-days immersion tests to investigate the FT effect on mechanical behavior and chloride permeability of high strength concrete. Then, a predictive model based on Fick’s second law was proposed to predict the chloride contents in high strength concrete subjected to FT cycles. Ma et al. [26] investigated the FT effect on chloride permeability of recycled powder concrete by 60-days soaking tests. The results in this literature indicated that the apparent chloride diffusion coefficient Da increased linearly with the decrease of NDEM. Nevertheless, the exposure duration of the immersion test was one time point (60 days) in these researches, without considering the time dependence of Da. Zhang et al. [27,28] carried out 3, 10

and 100-days immersion tests to determine the chloride ingress into frost-damaged concrete. It can be seen from the experimental results that Da increased significantly in approximate linearity with FT cycles and decreased considerably with the exposure duration. Above investigations were based on the immersion tests, which simulated the environment in underwater area. As a matter of fact, the deterioration of the marine concrete structures in tidal zone is severer than that in the underwater area [29]. Therefore, further investigation needs to be conducted to determine the chloride ingress into frost-damaged concrete in the tidal zone, in which the time dependence of Da also needs to be taken into account. In this paper, the laboratory FT cycling tests with a minimum freezing temperature of 18 °C were carried out to investigate the degradation indexes of mechanical performance after exposure to FT attack, i.e., the compressive strength, NDEM and mass loss of concrete specimens. The degradation levels of frost-damaged concrete were then quantified by these indexes. Subsequently, the chloride natural diffusion (CND) tests were conducted after several certain number of FT cycles to evaluate the effect of frost damage on the chloride diffusion of concrete in the tidal zone. A timedependent predictive model was then developed to predict the chloride profiles of concrete attacked by FT cycles in the tidal zone. Furthermore, additional FT cycling tests with minimum freezing temperatures of 11 and 4 °C, and then CND tests were carried out to study the effect of minimum freezing temperature on the chloride diffusion into concrete. 2. Test specimens and applied test methods 2.1. Materials and mix proportion of concrete All tested concrete specimens were prepared using the fine aggregates of the fineness modulus 2.7 with an apparent density of 2610 kg/m3 and coarse aggregates of the maximum grain size 20 mm with an apparent density of 2690 kg/m3. In order to eliminate the influence of chloride in tap water, the mixing water as well as the water used for curing later was the distilled water. And the cement was ordinary Portland cement (P.O. 42.5) with a density of 3100 kg/m3 produced by Tianjin Cement Plant (Beichen, Tianjin). The designed strength class of concrete specimens in the investigation was C30. The details of mix proportion and the 28days cubic compressive strength of concrete specimens are summarized in Table 1. 2.2. Test specimens From the mixes listed in Table 1, two kinds of specimens were designed and cast from the same batch of concrete, with a dimension of 100  100  100 mm3 and 100  100  400 mm3 respectively [21]. To eliminate the experimental material and measurement errors, three concrete specimens of each group (T1S-N and TE) were produced to measure the compressive strength, mass and dynamic elastic modulus. The details of the specimens used in this paper are given in Table 2. Note that the five specimens in the group T1C-N (N = 0, 5, 15, 30 or 50 FT cycles), T2C or T3C were used to determine the chloride profiles at five exposure durations, i.e., 30, 60, 80, 100 and 120 days, separately. Herein, the group without exposure to FT cycles (N = 0) was designated as the control group, and the rest were defined as the test groups. In this experiment, the preparation of specimens was based on the JTJ 270-98 [30]. Concrete was mixed in a forced action mixer, and all specimens were cast in the corresponding plastic molds from a same batch of concrete. These specimens were then cured in the standard curing box (T = 20 ± 3 °C, RH > 90%) for 24 h before demolding. After that, all of them were moved into a saturated cal-

3

Y. Wang et al. / Construction and Building Materials 236 (2020) 117556 Table 1 Mix proportion of experimental concrete. Water/cement ratio

Cement (kg/m3)

Water (kg/m3)

Fine aggregate (kg/m3)

Coarse aggregate (kg/m3)

Compressive strength at 28 days (MPa)

0.50

390

195

633

1180

40.26

Table 2 Specimen details used in test seriesa.

a

Group

Specimen dimensions

Ncs

FT cycles N

Tmin ( )

Usage

C0 T1S-N T1E T2E T3E T1C-N T2C T3C

100  100  100 mm 100  100  100 mm 100  100  400 mm 100  100  400 mm 100  100  400 mm 100  100  100 mm 100  100  100 mm 100  100  100 mm

1 37 3 3 3 55 5 5

/ 0, 5, 15, 30, 50, 75, 100 0, 5, 15, 30, 50, 75, 100 0, 30 0, 30 0, 5, 15, 30, 50 30 30

/ 18 18 11 4 18 11 4

The initial chloride content Compressive strength Mass and DEM Mass and DEM Mass and DEM Chloride profiles Chloride profiles Chloride profiles

Abbreviations: Ncs - Number of concrete specimens, Tmin - The minimum temperature reached of freezing process.

cium hydroxide solution for further curing until reached 28 days. Subsequently, the specimens were taken out, the surface slurry was wiped up and the surface of specimens was cleaned with alcohol. These samples were then ready for the FT cycling test. 2.3. Exposure to FT cycles The previous investigations showed that the deterioration of concrete caused by FT cycles was proposed to be linked to the internal moisture content of concrete [31,32]. The levels of FT damage were found to increase with the increases of the saturation [33]. In this study, after 27-days curing in the saturated calcium hydroxide solution, the concrete specimens had nearly the same saturation, ensuring the same initial conditions of samples. These samples were then subjected to FT cycles complying with a standard method [34], namely the FT cycling test. Prior to the FT cycling test, the specimens in groups T1S-0 and TE (i.e., T1E, T2E and T3E) listed in Table 2 were used to determine the compressive strength and mass, dynamic elastic modulus, respectively. Then, all specimens were put into the prismatic molds in batch. Subsequently, the distilled water was poured into these molds until the specimen was submerged. In this experiment, the exposure duration of one FT cycle was about three hours and twelve minutes. As revealed by Fig. 1, the core temperature of the specimen for the standard FT cycling test ranged from 18 ± 2 °C to 5 ± 2 °C. The minimum temperature reached in this

standard test was designated as T1. In this paper, T1 = 18 °C. To investigate the effect of minimum freezing temperature on the chloride diffusion in concrete, additional FT cycling tests were conducted at different lowest freezing temperatures (4 and 11 °C, denoted as T2 and T3, respectively). As listed in Table 2, the lowest temperatures reached of FT cycling test T1, T2 and T3 are corresponding with the group number T1, T2 and T3 separately. Immediately after a certain number of FT cycles, the prismatic samples in group TE (i.e., T1E, T2E and T3E) were taken out of the test box. The water on the surface of these specimens was wiped off, and then the mass and the dynamic elastic modulus of specimens were measured successively. Then, the following formulas Eqs. (1) and (2) were used to calculate the loss of relative mass and the normalized dynamic elastic modulus (NDEM), respectively:

Dmn ¼

m0  mn  100 m0

ð1Þ

where m0 refers to the mass of the specimens before the FT cycles (g). mn denotes the residual mass of the specimens after N FT cycles (g). Dmn is the normalized mass loss of the concrete specimens after N FT cycles (%).

NDEM ¼

Edn Ed0

ð2Þ

where Ed0, Edn are the dynamic elastic modulus of specimens in group TE before and after FT cycles (GPa), respectively. Meanwhile, three cubic concrete specimens in group T1S-N (N = 0, 5, 15, 30, 50, 75 and 100) after being subjected to the corresponding number of FT cycles were taken out of the test box to measure the compressive strength. This measurement was conducted in an automatic pressure testing machine according to GB/T 50081 [35]. Subsequently, the compressive strength fcu was calculated by the following equation:

f cu ¼ a

Fig. 1. The variation of core temperature of the test samples (Three FT cycles).

F A

ð3Þ

where fcu is the compressive strength (MPa) of a standard cubic specimen (150  150  150 mm3), accurate to 0.1 MPa. a denotes the size conversion factor; For the tested specimen with a dimension of 100  100  100 mm3, a = 0.95. F refers to the applied load by the testing machine when the specimen is destroyed (N). A is the area of the loading surface (mm2), which is taken as 10000 mm2 in the present experiment.

4

Y. Wang et al. / Construction and Building Materials 236 (2020) 117556

2.4. Chloride natural diffusion (CND) test In this investigation, the following three experiments were carried out to measure the chloride profiles of concrete: (1) The chloride natural diffusion (CND) test in the corrosion tank of a selfdesigned tidal cycling automatic device, which can simulate the marine environment in the tidal zone [36,37], (2) the grinding powder test and (3) the chloride concentration test using the CLE chloride rapid testing device. After a certain number of FT cycles, all other surfaces (except for one exposed surface) of the concrete specimens in group TC (i.e., T1C-N, T2C and T3C) provided in Table 2 were sealed with epoxy resin to insure the one-dimensional diffusion of chloride in concrete. Then, the CND test was carried out in the self-designed tidal cycling automatic device. As shown in Fig. 2, the self-designed tidal cycling automatic device was composed of the storage water tank, the test tank, the air-blast system, and the control system of flow rate, temperature and humidity. The flow rate control system can effectively adjust the flow rate of water and help the tidal cycling device to simulate the real-time rising and ebb tide process. In addition, this device can also control the marine environment factors such as temperature, humidity and wind. As known that the chloride content in seawater is about 3.5% (by weight), a 3.5% sodium chloride solution was applied for the simulation. In the present study, a diurnal tide (T = 20 ± 3 °C, RH = 60–80%) was simulated, i.e., the durations of rising tide and ebb tide were both 12 h. After the expected exposure durations (30, 60, 80, 100, or 120 days) under the chloride tidal cycling environment, the samples were taken out of the corrosion tank, being ready for the next experiment. Before the grinding powder test, the surface of concrete specimens was cleaned with distilled water and then dried. Subsequently, these specimens were put into a grinding machine and milled into powder starting from the diffusion surface, as illustrated in Fig. 3. A long-term field test exposed in marine tidal environment by Zhang et al. [38] discovered that the penetrating depth of chloride tended to be relatively stable at the depth of 20 mm when the exposure duration was less than 480 days. As a result, the powder was collected with a deepest depth of 20 mm in this paper. As can be seen in Fig. 3, the powder of Layer I was milled with 1 mm intervals, starting from the exposed surface, and powder of Layer II was collected with 2 mm intervals, ranging from 5 mm to 21 mm. According to the test code [30], prior to being kept in a stoving chest at 105 °C for 2 h, these powders were sieved with a 0.63 mm sieve.

Fig. 3. Locations of the grinding powder on the concrete specimen (The interval of Layer I and Layer II is 1 mm and 2 mm respectively).

After that, the powder was taken out and cooled to room temperature, preparing for testing the chloride concentrations. First, 2 g powder and 60 mL distilled water were successively added into an empty beaker. Second, the mix in this beaker was fully stirred to ensure that chloride was dissolved in the distilled water. Then, the supernatant of this well-stirred mix was taken to determine the free chloride ion content C (%, by weight of concrete) in the specimen using the CL-E chloride rapid testing device. This device, made in Shaoxing, was composed of a chloride ion selective electrode and a voltage test system, which can measure the chloride concentration of concrete.

3. Test results and discussions 3.1. Surface scaling and loss of relative mass The surface scaling of concrete specimens after 0, 5, 15, 30, 50, 75 and 100 FT cycles is illustrated in Fig. 4. As can be seen, the degree of surface scaling depends on the number of FT cycles. When this number is less than 15, there is no obvious erosion on the surface of the concrete specimen by visual inspection. At this time, the micropores on the surface of the concrete specimen attacked by FT cycles have not developed into large pores visible to the naked eye. When the number of FT cycles reaches 30, a

Fig. 2. The self-designed tidal cycling automatic device.

5

Y. Wang et al. / Construction and Building Materials 236 (2020) 117556

0 FT cycles

5 FT cycles

50 FT cycles

15 FT cycles

75 FT cycles

30 FT cycles

100 FT cycles

Fig. 4. Surface scaling of concrete samples after different FT cycles.

visible damage on the edge of the concrete surface begins to appear significantly. This phenomenon may be due to the fact that compared to the central region of the specimen, the edge is subjected to FT action in two directions, being more susceptible to FT damage [39]. The micropores in this area expand to macroscopic holes under the hydro-expansion action. With the continued increases of FT cycles, the surface erosion of the concrete specimen gradually develops from the edge to the center area until the surface of the whole test sample is covered with the holes left by this erosion. The damage degree of FT attack on the surface scaling of concrete specimens can be measured by the loss of relative mass [11]. As shown in Fig. 5, the normalized mass loss of concrete specimens varies with the number of FT cycles. It can be seen from Fig. 5 that the mass loss of the concrete specimen exhibits two distinct stages, being bilinearly correlated with the number of FT cycles [40]. Stage I: from 0 to 15 FT cycles, this loss caused by FT action is not obvious. Stage II: after 15 FT cycles, the loss of mass becomes significant, which is in accordance with the visual inspection of surface scaling (see Fig. 4). In general, the mass loss of frostdamaged concrete is not severe [41]. After 100 FT cycles, the mean value of this loss is approximately 0.9%. Through the regression analysis of the experimental data, a simpe formula is proposed to describe the relationship between the normalized mass loss Dmn of the concrete specimen and the number of FT cycles N, which is expressed as:

Fig. 5. Normalized mass loss of specimens exposed to FT attack.




Dmn ¼ 0:004N N 6 15 Dmn ¼ 0:091 þ 0:0105N N > 15

ð4Þ

The correlation coefficients R2 are 0.9 and 0.998 respectively. 3.2. Compressive strength Fig. 6 depicts the changes of residual compressive strength versus the number of FT cycles N. As shown in Fig. 6(a), the average compressive strength of concrete specimens decreases linearly with the increase of FT cycles, being in accordance with some other investigations [25,42]. To make a comparison with the results performed by Zhang et al. [25], the normalized compressive strength of concrete specimens is revealed by Fig. 6(b). As can be seen, even though the 28-d compressive strength has a difference, the slopes of both fitting lines are basically the same, manifesting that the test data in the proposed investigation is credible. An empirical formula through the regression analysis of the test data is illustrated in Fig. 6(a) to quantitively describe this linear change caused by FT action. An error analysis is carried out represented by error bars, in which the error bar represents the range of experimental results in parallel samples, i.e., the difference between the largest and smallest data. As can be seen from Fig. 6 (a), the relative error between the tested data and the values predicted by this empirical equation can be controlled within ±10%, performing a good prediction to the experimental results. Besides, the loss of compressive strength is significant. By 100th FT cycles, the mean of compressive strength is reduced to approximately 70% of the initial value. With the increase of FT cycles, especially after 50 FT cycles, there exists a considerable dispersion of the test values. It can be concluded that the extent of reduction caused by FT action in the compressive strength is severe. And this reduction is manifested as the damage degree and the dispersion degree of the compressive strength of concrete specimens. Moreover, an abnormality of the test results can be found in Fig. 6(a). The average compressive strength of specimen in group T1S-0 is slightly lower than that in group T1S-5. While considering the fact that tested results of specimens in group T1S-0 all lie below the values calculated by the empirical model, this abnormal data may result from the experimental errors or construction defects. To sum up, in the proposed investigation, the quantitative relationship between the compressive strength fcu of the frostdamaged concrete and the number of FT cycles N can be expressed as:

6

Y. Wang et al. / Construction and Building Materials 236 (2020) 117556

Fig. 6. Changes of the compressive strength versus FT cycles N: (a) the average compressive strength; (b) the normalized compressive strength.

f cu ¼ 41:47  0:1293N

ð5Þ 2

The correlation coefficient R = 0.9776. 3.3. Dynamic elastic modulus Fig. 7 and Table 3 indicates that the normalized dynamic elastic modulus (NDEM) of concrete specimens decreases significantly with the increasing FT cycles. After about 50 FT cycles, the average value of NDEM drops to 60%, denoting that the concrete specimens are already destroyed by the FT attack [40,43]. Therefore, in the present investigation, the CND test is focused on those undestroyed concrete specimens with the number of FT cycles not exceeding 50. Moreover, the residual NDEM of concrete specimens exposed to 100 FT cycles is approximately 20%, which is in accordance with the experimental results of the specimens with a similar water/cement ratio in literature [28]. Comparing with the normalized compressive strength shown in Fig. 6 (b), the loss of NDEM caused by FT action is more significant. In addition, it is obvious to notice that the NDEM exhibits a satisfactory linear relation with the number of FT cycles. On the basis of regression analysis, this linear relation can be expressed as follows:

NDEM ¼ 1  0:0078N

ð6Þ 2

The correlation coefficient R = 0.9877, showing a strong strength of this linear relation. Similar to the case of compressive strength, with the increase of FT cycles, the dispersion degree of

the test data increases obviously. As can be seen from a comparison of test results illustrated in Figs. 6(a) and 7, when the number of FT cycles is less than 50, the NDEM represents a better robustness than the compressive strength. Therefore, in this paper, the NDEM is recommended as the degradation index for the frost damage to quantify the damage degree of concrete [44]. In order to determine the bounds of the approximation error of the empirical formula, a forward error analysis between the experimental data and the empirical model results is shown in Fig. 7. As can be seen, the difference between the experimental results and the values calculated by Eq. (6) can be almost limited to ±15%. As revealed by Fig. 7, the test data of specimens outside the ±15% bounds is present at a higher number of FT cycles than 50. Consequently, Eq. (6) is credible to predict the NDEM of frost-damaged concrete, especially when the number of FT cycles is less than 50. Fig. 8 depicts the average NDEM of concrete specimens subjected to 30 FT cycles at different lowest freezing temperatures (4, 11 and 18 °C). As can be seen from this figure, the minimum temperature reached of FT cycles takes an apparent effect on the NDEM of specimens. Considering the fact that the FT cycling test is conducted using the distilled water, an assumption is adopted in this paper that there is no FT damage when the minimum freezing temperature is equal to 0 °C, which is the freezing point of water. That is when the minimum freezing temperature is 0 °C, the NDEM is equal to 1. Based on this assumption, an approximate linear relation is applied in Fig. 8 to determine the relationship between NDEM and the minimum freezing temperature. Besides, as have been demonstrated previously for the phenomena that the dispersion degree of compressive strength or NDEM usually increases with the increasing number of FT cycles, a similar case is present in the condition of different lowest freezing temperatures. In other word, the damage degree as well as the dispersion degree of the concrete specimen represented by NDEM increase with FT damage, regardless of the cause of this damage (due to the number of FT cycles or the minimum temperature reached of freezing process). 3.4. Chloride penetration

Fig. 7. The normalized dynamic elastic modulus (NDEM) of frost-damaged concrete samples in group T1E.

As mentioned in section 2.4, specimen in group C0 without being subjected to neither FT cycles nor chloride environment was used to determine the initial chloride concentration, and test results are listed in Table 4. With little difference between these test data, the mean of them is taken as the initial chloride concentration C0 (%) of concrete specimens. In this investigation, C0 = 0.02768. In order to eliminate the effect of the initial chloride

7

Y. Wang et al. / Construction and Building Materials 236 (2020) 117556 Table 3 The NDEM of concrete specimens exposed to different conditions of FT cycles. Group

Minimum freezing temperature ( )

Number of FT cycles

NDEM of parallel samples

T1E

18

T2E T3E

11 4

0 5 15 30 50 75 100 30 30

1.00 0.97 0.88 0.75 0.45 0.29 0.14 0.91 0.94

Fig. 8. The changes of average NDEM versus the minimum temperature reached of FT cycles.

concentration in specimens, the chloride content hereinafter described is the value which subtracts this initial value. The changes of chloride concentration profiles versus the exposure time of the CND test and the number of FT cycles are illustrated in Fig. 9(a) and (b), respectively, in which the following conclusion can be drawn: (1) The chloride concentrations in concrete decrease remarkably with the increase of diffusion depth, and gradually increase with the exposure time. (2) Under drywet cycling environment, the alternating action of water evaporation and capillary water absorption occurs on the concrete surface, which is defined as the convection. The mechanism of chloride ion ingression into unsaturated concrete is the coupling of diffusion and convection. Due to the seepage of pore solution toward the concrete surface, convection mainly happens in shallow concrete. According to test results, the convection zone refers to the layer ranging from the surface to the depth of 2.5 mm. This zone begins to appear in concrete specimens of the control group (group T1C-0) at 120-days exposure time, being consistent with the conclusion drawn by Zhang et al. [38], which is a significant characteristic of chloride diffusion in concrete under the wet-dry cycling environment [45]. However, when the concrete specimens are exposed to the FT cycles at the same time, the convection zone tends to appear earlier. For instance, this zone of frost-damaged concrete generally appears at 80-days exposure time, 40 days earlier than

Average values

1.00 0.94 0.90 0.81 0.67 0.34 0.26 0.88 0.94

1.00 0.95 0.91 0.84 0.79 0.58 0.21 0.88 0.96

1.00 0.95 0.90 0.80 0.64 0.40 0.20 0.89 0.95

the counterpart of specimen in the control group. As well, the effect of convection zone becomes more significant with the increase of exposure time and FT cycles. In deep concrete, diffusion is the control mechanism of chloride penetration. The concentration gradient is the main driving force to penetrate the chloride ion into concrete in the stable diffusion zone. Considering that the accurate determination of the depth of convection zone is difficult and the mechanism of chloride transport in deep concrete is what we are concerned about, many scholars use the Fick’s law to describe the diffusion characteristics of chloride in the stable diffusion region, and the effect of convection is attributed to the influence of diffusion boundary conditions [46]. (3) The chloride concentration at the same penetration depth and exposure time, except for the convection zone, increases with FT cycles, providing a potent evidence for the significant effect of the FT attack on the chloride diffusion in concrete. In addition, the differences between the chloride contents of the frost-damaged concrete and of the specimens in control group increase markedly with the extension of exposure duration under chloride environment. In other word, FT action takes a considerable effect on the long-term chloride diffusion in concrete. Besides, the test results show that the chloride ion penetration depth in concrete increases significantly with the number of FT cycles. To be clear, Fig. 10 illustrates the chloride profiles in concrete specimens at 80-days exposure time. It is obvious that the chloride penetration depth increases from 12 mm to 16 mm as the number of FT cycles increases from 0 to 50, denoted as x1 and x2, respectively. The test results of chloride profiles in concrete specimens exposed to 30 FT cycles with the minimum freezing temperature of 4 and 11 °C are illustrated in Fig. 11(a) and (b), respectively. As has been demonstrated previously for specimens after a certain number of FT cycles with a lowest temperature of 18 °C, a severer frozen damage usually results in a greater chloride concentration at the same depth and exposure time. This conclusion also can be confirmed by the test data shown in Fig. 11. In this condition, the convection zone also appears at about 80-days exposure time. Additionally, it is obvious to see from Fig. 11 that the FT effect on the convection zone is associated with the minimum freezing temperature and a lower temperature is accompanied by a more pronounced convection. In summary, the FT damage takes a significant influence on the chloride diffusion into concrete, including the chloride concentration, convection zone and penetration depth, and this influence will be amplified with the exposure time. As a matter of fact, the results of the ultrasonic pulse velocity test carried by Güneyisi

Table 4 The initial chloride concentration C0 of specimens used in present study. Number

1

2

3

4

5

Average value

C0 (%)

0.026912

0.028565

0.030319

0.028565

0.024040

0.027680

8

Y. Wang et al. / Construction and Building Materials 236 (2020) 117556

Fig. 9. Changes of chloride profiles versus (a) the chloride exposure time and (b) the number of FT cycles.

possibility that it is this alteration of pore and cracks caused by FT cycles which gives rise to the extension of the chloride transmitting channel, finally leading to a considerable influence on the chloride diffusion into concrete. Consequently, not only the number of FT cycles, but also the minimum temperature reached of FT cycles lead to the damage of concrete specimens. And it is this damage instead of the number of FT cycles or the lowest temperature reached of freezing process that determines the chloride penetration and the convection zone of the frost-damaged concrete essentially. 4. Time-dependent predictive model of chloride profiles in frost-damaged concrete Fig. 10. The chloride profiles in specimens after different number of FT cycles at 80days exposure duration.

et al. and Ng et al. [47,48], the X-ray computed tomography and acoustic emission tests conducted by Shields et al. [33] indicated that, from the microscopic aspects, the pore structure and the FT cracks developed with frost induced damage. There is a strong

As has been demonstrated for the significant effect of FT damage on the chloride penetration in concrete specimens, a quantified description of the diffusion mechanism needs to be developed. In the present investigation, this quantified description is carried out by a modified model based on the Fick’s second law. Collepardi et al. [49] proposed the analytical solution of Fick’s second law to fit the chloride profiles in concrete without exposure to FT cycles, as given in Eq. (7):

Fig. 11. Changes of chloride profiles versus the exposure time after 30 FT cycles with different minimum freezing temperatures: (a) T = -4 °C and (b) T = -11 °C.

9

Y. Wang et al. / Construction and Building Materials 236 (2020) 117556





 x C ðx; tÞ ¼ C s 1  erf pffiffiffiffiffiffiffiffi 2 Da t

ð7Þ

where C (x, t) denotes the free chloride concentration (%) within the concrete at depth  (mm) and exposure time t (s). Cs is surface chloride concentration (%), and erf() is referred to as the error function of Gauss. The apparent chloride diffusion coefficient Da and the surface chloride concentration Cs vary with exposure duration t [50,51], according to Eqs. (8) and (9), respectively:

Da ðtÞ ¼ D28

 m t 28 t

ð8Þ

where Da(t) and D28 is the apparent chloride diffusion coefficient (m2/s) at the exposure time t and t28 (days), respectively, and usually t28 = 28 days. m refers to the age attenuation factor of concrete [50].

C s ðtÞ ¼ C s0 ð1  eat Þ

ð9Þ

where Cs(t) denotes the surface chloride concentration (%) of the concrete specimens corresponding to the exposure time t. Cs0 is the ultimate value of the surface chloride concentration (%). a is referred to as the time-dependent coefficient of Cs0 [51]. 4.1. Time-dependent model of apparent chloride diffusion coefficient Da(t, P)

4.1.1. Apparent chloride diffusion coefficient Da Assuming that the chloride ingress into frost-damaged concrete also follows the Fick’s second law, and then the error function solution Eq. (7) can be adopted to fit the chloride profiles illustrated in Fig. 9. Due to that the chloride ingress in the convection zone does not follow the Fick’s second law, the convection zone data is excluded [38]. In this way, considering that the penetration depth x , exposure time t and chloride concentration C(x, t) are known, the apparent chloride diffusion coefficients Da(t) and the surface chloride concentration Cs(t) after each certain number of FT cycles can be calculated through regression analysis. The regression analysis results of Da(t) and Cs(t) are shown in Figs. 12 and 15, separately. As revealed by Fig. 12, the apparent chloride diffusion coefficient Da(t) increases considerably in approximate linearity with the number of FT cycles, which is in good accordance with the results of concrete specimens at a 60-days immersion test obtained by Zhang et al. [28]. The larger the number of FT cycles is, the higher the apparent chloride diffusion coefficient becomes. When the number of FT cycles reaches 50, the apparent chloride diffusion

coefficient of the frost-damaged concrete is almost more than 1.75 times to the counterpart of the concrete in the control group. Even though there exists an anomalous point in the results of specimens at the 30-days exposure time, considering that the results at other exposure durations have a better linear relation, this anomaly may be caused by the construction defects of the specimen in group T1C-5. Besides, with the increase of exposure time, a dramatic decline of the apparent chloride diffusion coefficient of the concrete specimen after a certain number of FT cycles is observed in Fig. 12. This decline can be described by an age attenuation factor m in Eq. (8). Therefore, the results of apparent chloride diffusion coefficient D28 and corresponding age factor m are calculated through Eq. (8), which are subsequently given in Table 5. As presented in Table 5, with the increase of FT cycles, there is an obvious increase in the apparent chloride diffusion coefficient D28. By contrast, it can be seen that the age attenuation factor m is almost a constant equaling to the mean value 0.322 with a variation coefficient of 4.4%, which does not vary with FT cycles. Consequently, this constant is adopted as the representative value of the age factor m in the present study, which can be determined by a series of CND test on the specimens without exposure to FT cycles. 4.1.2. FT effect factor b of Da The loss of mass, compressive strength and NDEM are frequently applied as the degradation indexes to quantify the frozen damage in concrete [44]. As has been demonstrated previously, the NDEM of concrete specimens varies apparently with FT cycles and exhibits a better robustness than the compressive strength. Therefore, the loss of NDEM in the proposed investigation is adopted to represent the effect of FT cycles on chloride penetration in concrete. In this paper, an effect factor of the apparent chloride diffusion coefficient Da is proposed to consider this influence of FT damage on Da. This effect factor b(P) is defined as follows:

bðPÞ ¼

D28;P D28;0

ð10Þ

where D28,0, D28,P are the apparent chloride diffusion coefficients of concrete at time t28 before and after FT cycles, respectively, in which D28,0 needs to be obtained by CND tests of the concrete without being exposed to FT cycles. In the present research, D28,0 = 3.27  1012 m2/s. P refers to the loss of NDEM, as given in Eq. (11):

P¼

Ed0  Edn ¼ 1  NDEM Ed0

ð11Þ

As has been developed in the previous discussion, the data of the NDEM and the apparent chloride diffusion coefficients D28,0, D28,P is presented in Table 3 and Table 5, respectively. Based on Eqs. (10) and (11), the effect factor b and the loss of NDEM P can be obtained. As illustrated in Fig. 13, the effect factor b increases linearly with the NDEM loss P. And this relation in the proposed investigation can be stated by a simplified formulation, which is expressed as:

bðPÞ ¼ 1 þ 1:792P

ð12Þ

Table 5 D28 and corresponding m at time t28 of specimens exposed to different FT cycles.

Fig. 12. Changes of apparent chloride diffusion coefficient Da(t) versus FT cycles N.

FT cycles N

D28 (1012 m2/s)

m

R2

0 5 15 30 50

3.27 3.83 4.07 4.65 5.48

0.322 0.339 0.301 0.313 0.336

0.9538 0.9432 0.9417 0.9580 0.9850

10

Y. Wang et al. / Construction and Building Materials 236 (2020) 117556

Fig. 13. The change of effect factor b versus the loss of NDEM P.

Fig. 15. Changes of surface chloride concentration Cs(t) versus the number of FT cycles N.

The correlation coefficient R2 = 0.99861. Inserting Eq. (12) into Eq. (8) gives:

8 > < > :

Da ðt; PÞ ¼ bðPÞD28;0

t28 m t

bðPÞ ¼ 1 þ 1:792P

ð13Þ

D28;0 ¼ 3:27  1012 m2 =s; t 28 ¼ 28d; m ¼ 0:322

As described in the previous unit, there is no effect of FT action on the age attenuation factor m, which can be regarded to be a constant. As revealed by Fig. 14, this time-dependent model can make a convincing prediction about the apparent chloride diffusion coefficient Da(t, P) of concrete exposed to FT attack. And the error of this model can be controlled within ±10%.

4.2. Time-dependent model of surface chloride concentration Cs(t, P)

4.2.1. Surface chloride concentration Cs The research expanded by Zhang et al. [27] showed that after exposure to FT cycles, a damage gradient from surface to center of specimens was observed, with the severest damage nearby the surface. As a result, considering that the surface chloride concentration Cs relies on the properties of surface concrete, this damage nearby the surface may have a significant effect on the surface chloride concentration of concrete specimen.

Fig. 15 shows the surface chloride concentration of concrete specimens after different number of FT cycles and exposure time. It can be seen from Fig. 15 that the surface chloride concentration Cs increases with FT cycles. As expected, this increase of the surface chloride concentration Cs is rapid in the early stage of FT action. After that, this increase becomes slow, followed by an eventual leveling off regardless of the exposure time. Dai and Ng [52] demonstrated that it was the ice crystallization pressures within the pores and the micro-cracks that generate the internal damage of frost in cement pastes. With the development of cracks, the pressure of ice crystal can be relaxed during the process of FT cycles. Accordingly, the rate of cracking development at the concrete surface shall be slowed down with the increase of FT cycles. In addition, as revealed by Fig. 15, the surface chloride concentration at a certain number of FT cycles is time-dependent, increasing significantly with the exposure time. The difference in the surface chloride concentration Cs between the test group and control group also increases with the exposure time. That is to say, the FT attack takes a significant influence on the long-term durability of the concrete. This influence can be reflected in the surface chloride concentration. As described previously, Eq. (9) can be used to depict the time dependence of the surface chloride concentration. Based on this equation, the ultimate value of the surface chloride concentration Cs0 and its corresponding time-dependent coefficient a are calcu-

Fig. 14. (a) Comparison between empirical model and experimental results of apparent chloride diffusion coefficient Da(t); (b) Corresponding relative error analysis of this model.

11

Y. Wang et al. / Construction and Building Materials 236 (2020) 117556

lated through regression analyses using least square method, as illustrated in Table 6. As can be seen from Table 6, the maximum value of the surface chloride concentration Cs0 shows an apparent increasing tendency with the increase of FT cycles. In the present investigation, owing to the CND tests are conducted after every certain number of FT cycle, the FT damage of each specimen is certain during the chloride penetration process, regardless of the exposure time under the chloride environment. Thereby, there is a definite possibility that this damage makes a direct influence on the ultimate value of the surface chloride concentration. Besides, as shown in Table 6, the time-dependent coefficient a remains almost steady, indicating that the exposure time of the maximum value reached has little to do with the FT damage. Consequently, the mean value of these time-dependent factors is used as the typical value. In the present investigation, a = 0.0073. 4.2.2. FT effect factor c of Cs As has been demonstrated previously, the FT damage takes a significant influence on the maximum value of the surface chloride concentration. In order to reflect this influence, an effect factor c(P) is proposed based on the experimental results, which is defined as:

cðPÞ ¼

C s0;P C s0;0

ð14Þ

where Cs0,0, Cs0,P refer to the maximum value of the fitting curves of the surface chloride concentration of concrete before and after FT cycles (%), respectively. In this paper, Cs0,0 = 1.6427 (%). Similar to the form of Eq. (9), an empirical formula is suggested as follows to describe the effect factor c(P) in this paper:



cðPÞ ¼ 1 þ m 1  enP



ð15Þ

where m and n are the fitting parameters. As illustrated in Fig. 16, the parameters m and n in Eq. (15) are obtained through data regression, which fits the experimental results very well, with the correlation coefficient R2 of 0.9754. Inserting this formula into Eq. (9) gives:

8 ð1  eat Þ > < C s ðt; PÞ ¼ cðPÞC s0;0  cðPÞ ¼ 1 þ 0:352 1  e7:114P > : C s0;0 ¼ 1:6427 ð%Þ; a ¼ 0:0073

ð16Þ

where Cs (t, P) is the surface chloride concentration of concrete specimens at exposure time t and NDEM loss P. Eq. (16) is the time-dependent model of the surface chloride concentration considering the action of FT cycles. The maximum value of surface chloride concentration Cs0,0 and time-dependent coefficient a can be obtained from the CND experiments of specimens without exposure to FT action. To verify the correctness of this time-dependent model, a comparison is given between the empirical model and the experimental data of surface chloride concentration. As can be seen in Fig. 17, the prediction value of the surface chloride concentration is approximately equal to the experimental value, and the error is less than ±10%.

Table 6 The ultimate value of the surface chloride concentration of concrete specimens after different FT cycles. FT cycles N

Cs0 (%)

a

R2

0 5 15 30 50

1.6427 1.7764 1.9572 2.0987 2.1834

0.0072 0.0070 0.0074 0.0074 0.0075

0.9931 0.9437 0.9881 0.9967 0.9934

Fig. 16. Change of effect factor a versus the loss of NDEM P.

4.3. Time-dependent predictive model of chloride profiles C(x, t, P) As described in the previous sections, the time-dependent models of the apparent chloride diffusion coefficient Da(t, P) and of the surface chloride concentration Cs(t, P) in concrete exposed to FT action are proposed, expressed by Eqs. (13) and (16), respectively. Inserting these two formulas into Eq. (7) gives:

8 > > > > > > > > > > > > > > < > > > > > > > > > > > > > > :

   x ffi C ðx; t; PÞ ¼ C s ðt; PÞ 1  erf pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2

Da ðt;PÞ86400t

C s ðt; PÞ ¼ cðPÞC s0;0 ð1  eat Þ  cðP Þ ¼ 1 þ 0:352 1  e7:114P C s0;0 ¼ 1:6427 ð%Þ; a ¼ 0:0073  m Da ðt; PÞ ¼ bðPÞD28;0 t28 t

ð17Þ

bðPÞ ¼ 1 þ 1:792P D28;0 ¼ 3:27  1012 m2 =s; t 28 ¼ 28d; m ¼ 0:322

Up to the present, the time-dependent predictive model of chloride concentration C(x, t, P) in concrete exposed to FT attack is established based on the experimental data, which can be expressed by Eq. (17). The relative error between the experimental value and the empirical value is illustrated in Fig. 18. As can be seen, this predictive model fits well with the experimental data. Overall, the percent error between them is almost not exceed ±15%. Thereby, the proposed model can effectively reflect the chloride transmitting law in concrete specimens exposed to FT cycles. In other words, this model can be used to predict the chloride profiles in frost-damaged concrete. 4.4. Further validation of the proposed model As has been demonstrated previously, the time-dependent predictive model proposed in the present research is based on the test data of specimens exposed to the standard FT cycling test with the minimum freezing temperature of 18 °C. Therefore, this model is suitable for the prediction of the chloride concentration under this standard condition. Nevertheless, when it comes to the FT cycling test with a different minimum freezing temperature, the proposed model is not convincing. Since this model is based on the NDEM loss P instead of the number of FT cycles, this paper attempts to apply this model to a more general case. The tested data of specimens in group T2C and T3C, and the predicted chloride contents calculated by Eq. (17) are illustrated in Fig. 19. Overall, the predicted results agree well with the experimental data. As a matter of fact, the difference of the minimum temperature reached leads to different damage degree of concrete

12

Y. Wang et al. / Construction and Building Materials 236 (2020) 117556

Fig. 17. (a) Comparison between empirical model and experimental results of surface chloride concentration Cs(t); (b) Corresponding relative error analysis of this model.

be further proofed. Consequently, the proposed model (Eq. (17)) can make a good prediction for the chloride profiles in frostdamaged concrete, regardless of the cause of the frozen damage (due to the number of FT cycles or result from the minimum freezing temperature). 5. Conclusions

Fig. 18. Comparison between empirical model and experimental results of chloride concentration.

Fig. 19. Comparison between empirical model and experimental results of chloride concentration at different minimum freezing temperature.

specimens subjected to FT cycles, which eventually affects the chloride ingress in concrete specimens. This influence mechanism indicates that the effect of minimum freezing temperature during FT cycles on the chloride transport performance in concrete is essentially the effect of FT damage on the chloride diffusion. With a percent error of less than ±15% (see Fig. 19), this point of view can

The effects of FT attack on the mechanical behavior and chloride permeability of concrete specimens with the designed strength class of C30 are clarified in the proposed investigation based on extensive laboratory FT cycling tests and chloride natural diffusion (CND) tests. Two coefficients c and b are proposed to consider the effects of FT damage on the surface chloride concentration and apparent chloride diffusion coefficient of the concrete specimens, respectively. Afterwards, a time-dependent model is proposed to predict the chloride profiles of frost-damaged concrete in the tidal zone. The following conclusions can be drawn: (1) Under the action of FT cycles, the damage degree of concrete, represented by the loss of compressive strength and NDEM, increases substantially with the increase of FT cycles or with the decrease of minimum freezing temperatures. At about 50 FT cycles, the NDEM decreases to 60% of the initial value. After 100 FT cycles, the NDEM is only approximately 20%. Besides, the FT effect on the mechanical performance of concrete is also reflected in the increase of the dispersion degree of the mechanical index. (2) Under the marine tidal environment, the damage caused by FT attack takes a considerable effect on chloride diffusion into concrete, including the chloride concentration, convection zone and penetration depth, and this effect will be amplified with the exposure duration under chloride environment. As the convection zone in concrete specimens without exposure to FT attack occurs at the exposure duration of 120 days, the FT cycles can shorten the occurrence time of the convection zone effectively. In addition, at the same exposure time, the apparent chloride diffusion coefficient and the surface chloride concentration increase apparently with the frozen damage. (3) Based on the experimental results of the standard FT cycling tests and the CND tests, a time-dependent predictive model is developed and verified to predict the chloride penetration in frost-damaged concrete. And further investigation indicates that the proposed model can be applicable for the con-

Y. Wang et al. / Construction and Building Materials 236 (2020) 117556

ditions with a different minimum freezing temperature. The relative error of chloride concentrations between the predicted values calculated by proposed model and the tested results in concrete specimens is less than ±15%, providing a strong evidence for the correctness of the predictive model. In this investigation, the tests were performed on the concrete specimens without air entrainment. Nevertheless, air-entrained concrete was widely used to improve the frost resistance. As a consequence, additional experiments should be taken to develop the time-dependent predictive model of chloride profiles in the frostdamaged concrete with air entrainment. Besides, the specimens used in this paper were first exposed to the rapid FT cycling test and then followed by the chloride diffusion. However, in fact, concrete structures in the marine environment are subjected to FT attack and chloride penetration simultaneously, which needs to be taken into consideration in the future. In addition, there exist many differences between the laboratory experiments and field exposed tests, such as the freezing rate, the lowest temperature reached, and the length of freezing-thawing period [53]. Thus, the question that how to use the laboratorial results to predict the chloride resistance of concrete in field needs further investigating.

Declaration of Competing Interest The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.

Acknowledgments We are grateful for the financial support from the National Key Research and Development Program of China (2016YFC0802204, 2016YFC0802201), the National Natural Science Foundation of China (51679166), and the National Natural Science Fund for Innovative Research Groups Science Foundation (51321065).

References [1] R.M. Ferreira, Optimization of RC structure performance in marine environment, Eng. Struct. 32 (5) (2010) 1489–1494. [2] M. Santhanam, M. Otieno, 5 – Deterioration of concrete in the marine environment, in: M.G. Alexander (Ed.), Marine Concrete Structures, Woodhead Publishing, 2016, pp. 137–149. [3] H. Wang, S.Q. He, C.H. Fei, Deterioration performance of damaged concrete beams under freezing-thawing cycle and chloride environment in coastal cities, J. Coast. Res. 295–299 (2018). [4] S.P. Yin, L. Jing, M.T. Yin, B. Wang, Mechanical properties of textile reinforced concrete under chloride wet-dry and freeze-thaw cycle environments, Cem. Concr. Compos. 96 (2019) 118–127. [5] T. Luping, J. Gulikers, On the mathematics of time-dependent apparent chloride diffusion coefficient in concrete, Cem. Concr. Res. 37 (4) (2007) 589–595. [6] Y.Z. Wang, L.J. Wu, Y.C. Wang, C.X. Liu, Q.M. Li, Effects of coarse aggregates on chloride diffusion coefficients of concrete and interfacial transition zone under experimental drying-wetting cycles, Construct. Build. Mater. 185 (2018) 230– 245. [7] Y.Z. Wang, L.J. Wu, Y.C. Wang, Q.M. Li, Z. Xiao, Prediction model of long-term chloride diffusion into plain concrete considering the effect of the heterogeneity of materials exposed to marine tidal zone, Construct. Build. Mater. 159 (2018) 297–315. [8] M. Soutsos, Concrete durability: A practical guide to the design of durable concrete structures, Thomas Telford Limited, London, 2010. [9] J.J. Valenza, G.W. Scherer, A review of salt scaling: I, Phenomenol. Cem. Concr. Res. 37 (7) (2007) 1007–1021. [10] J.J. Valenza, G.W. Scherer, A review of salt scaling: II, Mechan. Cem. Concr. Res. 37 (7) (2007) 1022–1034. [11] S. Kessler, C. Thiel, C.U. Grosse, C. Gehlen, Effect of freeze–thaw damage on chloride ingress into concrete, Mater. Struct. 50 (2) (2016) 121.

13

[12] H. Kuosa, R.M. Ferreira, E. Holt, M. Leivo, E. Vesikari, Effect of coupled deterioration by freeze–thaw, carbonation and chlorides on concrete service life, Cem. Concr. Compos. 47 (2014) 32–40. [13] J. Zhao, G. Cai, D. Gao, S. Zhao, Influences of freeze–thaw cycle and curing time on chloride ion penetration resistance of Sulphoaluminate cement concrete, Construct. Build. Mater. 53 (2014) 305–311. [14] L.C. Hao, Y.Z. Liu, W.J. Wang, J.G. Zhang, Y. Zhang, Effect of salty freeze-thaw cycles on durability of thermal insulation concrete with recycled aggregates, Construct. Build. Mater. 189 (2018) 478–486. [15] Z. Hu, H. Ding, J. Lai, H. Wang, X. Wang, S. He, The durability of shotcrete in cold region tunnel: a review, Construct. Build. Mater. 185 (2018) 670–683. [16] M. Ferreira, H. Kuosa, M. Leivo, E. Holt, Concrete performance subject to coupled deterioration in cold environments, Nucl. Eng. Design 323 (2017) 228–234. [17] T. Luping, L.-O. Nilsson, Rapid determination of the chloride diffusivity in concrete by applying an electrical field, Aci Mater. J. 89 (M6) (1992) 49–53. [18] T. Luping, Electrically accelerated methods for determining chloride diffusivity in concrete-current development, Magaz. Concr. Res. 48 (176) (1996) 173– 179. [19] J.G. Cabrera, P.A. Claisse, Measurement of chloride penetration into silica fume concrete, Cem. Concr. Compos. 12 (3) (1990) 157–161. [20] C. Andrade, Calculation of chloride diffusion coefficients in concrete from ionic migration measurements, Cem. Concr. Res. 23 (3) (1993) 724–742. [21] Y. Wang, M. An, Z. Yu, B. Han, W. Ji, Experimental and cellular-automata-based analysis of chloride ion diffusion in reactive powder concrete subjected to freeze–thaw cycling, Construct. Build. Mater. 172 (2018) 760–769. [22] B. Li, J. Mao, T. Nawa, Z. Liu, Mesoscopic chloride ion diffusion model of marine concrete subjected to freeze-thaw cycles, Construct. Build. Mater. 125 (2016) 337–351. [23] A. Boddy, E. Bentz, M.D.A. Thomas, R.D. Hooton, An overview and sensitivity study of a multimechanistic chloride transport model, Cem. Concr. Res. 29 (6) (1999) 827–837. [24] Y. Wang, Y. Cao, P. Zhang, Y. Ma, T. Zhao, H. Wang, Z. Zhang, Water absorption and chloride diffusivity of concrete under the coupling effect of uniaxial compressive load and freeze–thaw cycles, Construct. Build. Mater. 209 (2019) 566–576. [25] X. Zhang, L. Wang, J. Zhang, Mechanical behavior and chloride penetration of high strength concrete under freeze-thaw attack, Cold Reg. Sci. Technol. 142 (2017) 17–24. [26] Z. Ma, W. Li, H. Wu, C. Cao, Chloride permeability of concrete mixed with activity recycled powder obtained from C&D waste, Construct. Build. Mater. 199 (2019) 652–663. [27] P. Zhang, F.H. Wittmann, M. Vogel, H.S. Müller, T. Zhao, Influence of freezethaw cycles on capillary absorption and chloride penetration into concrete, Cem. Concr. Res. 100 (2017) 60–67. [28] P. Zhang, Y. Cong, M. Vogel, Z. Liu, H.S. Müller, Y. Zhu, T. Zhao, Steel reinforcement corrosion in concrete under combined actions: The role of freeze-thaw cycles, chloride ingress, and surface impregnation, Construct. Build. Mater. 148 (2017) 113–121. [29] M. Safehian, A.A. Ramezanianpour, Assessment of service life models for determination of chloride penetration into silica fume concrete in the severe marine environmental condition, Construct. Build. Mater. 48 (2013) 287–294. [30] JTJ 270-98, Testing code of concrete for port and watering engineering, Ministry of Communications of the People’s Republic of China, China, 1998. [31] G. Fagerlund, A service life model for internal frost damage in concrete, Univ. (2004). [32] M. Ferreira, M. Leivo, h. kuosa, D. Lange, The influence of freeze-thaw loading cycle on the ingress of chlorides in concrete, 2016. [33] Y. Shields, E. Garboczi, J. Weiss, Y. Farnam, Freeze-thaw crack determination in cementitious materials using 3D X-ray computed tomography and acoustic emission, Cem. Concr. Compos. 89 (2018) 120–129. [34] GB/T 50082-2009, Standard for test methods of long-term performance and durability of ordinary concrete, China Architecture and Building Press, Beijing (China), 2009. [35] GB/T 50081-2002, Standard for test method of mechanical properties on ordinary concrete, China Architecture and Building Press, Beijing (China), 2002. [36] Y.Z. Wang, Q.M. Li, C.a. Lin, Chloride diffusion analysis of concrete members considering depth-dependent diffusion coefficients and effect of reinforcement presence, J. Mater. Civ. Eng. 28 (5) (2016). 04015183. [37] Y.Z. Wang, C.X. Liu, Y.C. Wang, Q.M. Li, H. Liu, Time-and-depth-dependent model of chloride diffusion coefficient for concrete members considering the effect of coarse aggregate, J. Mater. Civil Eng. 30 (3) (2018) 12. [38] J. Zhang, J. Zhao, Y. Zhang, Y. Gao, Y. Zheng, Instantaneous chloride diffusion coefficient and its time dependency of concrete exposed to a marine tidal environment, Construct. Build. Mater. 167 (2018) 225–234. [39] W.A. Cordon, Freezing and thawing of concrete-mechanisms and control, American Concrete Institute and Iowa State University Press, US, 1966. [40] R. García-Giménez, M. Frías, I. Arribas, I. Vegas, R.V. de la Villa, V. Rubio, Freeze-thaw effect on the durability of binary cements containing activated coal-mining waste, Construct. Build. Mater. 190 (2018) 140–149. [41] J. Wu, X. Jing, Z. Wang, Uni-axial compressive stress-strain relation of recycled coarse aggregate concrete after freezing and thawing cycles, Construct. Build. Mater. 134 (2017) 210–219. [42] J. Wang, D. Niu, H. He, Frost durability and stress–strain relationship of lining shotcrete in cold environment, Construct. Build. Mater. 198 (2019) 58–69.

14

Y. Wang et al. / Construction and Building Materials 236 (2020) 117556

[43] C.W. Miao, R. Mu, Q. Tian, W. Sun, Effect of sulfate solution on the frost resistance of concrete with and without steel fiber reinforcement, Cem. Concr. Res. 32 (1) (2002) 31–34. [44] M.C. Rao, S.K. Bhattacharyya, S.V. Barai, Systematic Approach of Characterisation and Behaviour of Recycled Aggregate Concrete, Springer, 2019. [45] M. Khanzadeh-Moradllo, M.H. Meshkini, E. Eslamdoost, S. Sadati, M. Shekarchi, Effect of wet curing duration on long-term performance of concrete in tidal zone of marine environment, Int. J. Concr. Struct. Mater. 9 (4) (2015) 487–498. [46] C.H. Lu, Y. Gao, Z.W. Cui, R.G. Liu, Experimental analysis of chloride penetration into concrete subjected to drying-wetting cycles, J. Mater. Civil Eng. 27 (12) (2015) 10. [47] E. Güneyisi, M. Gesog˘lu, Z. Algın, H. Yazıcı, Effect of surface treatment methods on the properties of self-compacting concrete with recycled aggregates, Construct. Build. Mater. 64 (2014) 172–183.

[48] K. Ng, Y. Sun, Q. Dai, X. Yu, Investigation of internal frost damage in cementitious materials with micromechanics analysis, SEM imaging and ultrasonic wave scattering techniques, Construct. Build. Mater. 50 (2014) 478– 485. [49] M. Collepardi, A. Marcialis, R. Turriziani, The kinetics of penetration of chloride ions into the concrete, Il cemento 67 (4) (1970) 157–164. [50] M.D.A. Thomas, P.B. Bamforth, Modelling chloride diffusion in concrete: Effect of fly ash and slag, Cem. Concr. Res. 29 (4) (1999) 487–495. [51] M.K. Kassir, M. Ghosn, Chloride-induced corrosion of reinforced concrete bridge decks, Cem. Concr. Res. 32 (1) (2002) 139–143. [52] Q. Dai, K. Ng, Transmission X-ray microscope nanoscale characterization and 3D micromechanical modeling of internal frost damage in cement paste, J. Nanomechan. Micromechan. 4 (1) (2014) A4013005. [53] A. Duan, J. Qian, Effect of freeze–thaw cycles on the stress–strain curves of unconfined and confined concrete, Mater. Struct. 44 (7) (2011) 1309–1324.