Experimental and numerical studies on impact behaviors of recycled aggregate concrete-filled steel tube after exposure to elevated temperature

Accepted Manuscript Experimental and numerical studies on impact behaviors of recycled aggregate concrete-filled steel tube after exposure to elevated temperature

Wengui Li, Zhiyu Luo, Chengqing Wu, Vivian W.Y. Tam, Wen Hui Duan, Surendra P. Shah PII: DOI: Reference:

S0264-1275(17)30911-5 doi:10.1016/j.matdes.2017.09.057 JMADE 3390

To appear in:

Materials & Design

Received date: Revised date: Accepted date:

1 July 2017 21 September 2017 26 September 2017

Please cite this article as: Wengui Li, Zhiyu Luo, Chengqing Wu, Vivian W.Y. Tam, Wen Hui Duan, Surendra P. Shah , Experimental and numerical studies on impact behaviors of recycled aggregate concrete-filled steel tube after exposure to elevated temperature. The address for the corresponding author was captured as affiliation for all authors. Please check if appropriate. Jmade(2017), doi:10.1016/j.matdes.2017.09.057

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ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

Experimental and numerical studies on impact behaviors of recycled aggregate concrete-filled steel tube after exposure to elevated temperature

Wengui Lia, Zhiyu Luoa, Chengqing Wua, Vivian W.Y. Tamb, Wen Hui Duanc, Surendra P. Shahd Centre for Built Infrastructure Research, School of Civil and Environmental Engineering, University of

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a

School of Computing, Engineering and Mathematics, Western Sydney University, Penrith, NSW 2751, Australia c

Department of Civil Engineering, Monash University, Clayton VIC 3800, Australia

Center for Advanced Cement-Based Materials, Department of Civil and Environmental Engineering Northwestern

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b

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Technology Sydney, Sydney, NSW 2007, Australia

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University, Evanston, IL 60208, USA

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Corresponding authors: Email addresses: [email protected] (Wengui Li) 1

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Abstract: This study investigated the impact behaviors of recycled aggregate concrete-filled steel tube (RACFST) after exposed to elevated temperatures by experimental and numerical studies. The impact test on RACFST was conducted by a split Hopkinson pressure bar (SHPB) with 100 mm-diameter. After the validation of finite element method simulation by the experimental results,

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parametric analysis were applied to analyze the effects of RAC strength, steel strength and steel ratio

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on the impact behaviors and deformation properties of the RACFSTs exposed to elevated

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temperatures ranging from 20 oC, 200 oC to 500 oC and to 700 oC. The results show that both the increases in RAC strength, steel strength and steel ratio significantly enhance the impact resistance

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of RACFST. However, when RAC strength deteriorates quickly after exposure to high temperature,

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the increase in RAC strength just slightly improves the impact properties of RACFST, especially after exposure to elevated temperatures higher than 500 oC. In conclusion, increasing steel ratio is an

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effective way for improving the impact behaviors of RACFST, but resulting in higher cost for

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practical application. For the impact design of RACFST, steel strength and steel ratio are two essential factors in terms of impact behavior enhancement and cost efficiency.

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Keywords: Recycled aggregate concrete-filled steel tube (RACFST); Split Hopkinson pressure bar

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(SHPB); Impact behavior; Numerical simulation; Elevated temperature

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1. Introduction In many countries, there are more and more concrete wastes in the process of urbanization and building demolition, which cause serious problems. Recycled aggregate concrete (RAC) is treated as an ideal way to reuse concrete wastes for new concrete production. However, considering the lower

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performances of RAC than that of natural aggregate concrete (NAC), related methods are necessary

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to ensure safety and reliability in promoting the RAC application [1,2]. NAC-filled steel tube

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(NACFST) becomes increasingly popular for its attractive properties such as high strength, excellent ductility and ease of construction [3,4]. Inspired by the successful applications of NACFST,

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RAC-filled steel tube (RACFST) has been also developed in the recent years [5,6]. Steel tube can

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improve mechanical properties of RAC by providing confinement, and performance differences between RAC and NAC become less influential because of steel tube playing an important role in the

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loading process. This sort of composite member (RACFST) is a potential way for the application of

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RAC. Therefore, the mechanical performances of RACFST need comprehensive studies for producing RACFST for structural applications.

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Nowadays, the increasing extreme events such as terrorist attacks, chemical explosions,

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high-speed crashes and explosions caused by vehicles have aroused widespread concerns for the safety of building structures. In the process of blast and explosion, tremendous energy released during a very short period of time causes great damage for building structures, resulting in huge loss of life and property. At the same time, fire is also a major disaster for building structures. Exposure to fire or high temperature significantly reduces material properties and decreases the strength and stiffness of building structure. To an extreme case, building structures are subjected to fire exposure and then to blast loading, which largely increase the risk of structure failure. For example, the tragic 3

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events such as the September 11 attacks (New York, 2001) and the Tianjin explosion (Tianjin China, 2015) have raised the awareness in ensuring the structural resistance to combined action of fire and blast loading, particularly for some critical infrastructures. Therefore, to apply the RACFST members in the structural engineering, it is necessary to investigate the impact behaviors of RACFST

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suffering from both high temperature and impact loading. However, relevant researches on RACFST

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are very limited. Yang et al. [7] and Shakir et al. [8] studied the lateral impact behavior of square

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RACFST and circular RACFST members respectively, and found that there are no obvious differences between the impact resistances of RACFST and NACFST. Yang et al. [9] also

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investigated the compressive behavior of the RACFST after exposure to elevated temperatures, and

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reported that the RACFST stub columns exhibit slightly poorer mechanical performances compared with the corresponding NACFST specimens. Li et al. [10] found that the mechanical behavior

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differences between RACFST and NACFST become further obvious with the increase of recycled

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coarse aggregate (RCA) replacement ratio and exposure temperature. However, there are plenty of available of researches conducted on NACFST. Yousuf et al. [11,12]

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carried out experiments in studying the transverse impact resistance of concrete-filled mild steel

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columns and concrete-filled stainless steel columns. The results indicated that the stainless steel specimens exhibit better energy-dissipating characteristics compared with their mild steel counterparts, especially when concrete is used to fill the hollow tubes. Xiao et al. [13] investigated the effect of carbon fiber reinforced polymer (CFRP) on impact responses of CFRP confined NAC stub columns. The results indicated that the impact failure patterns are related to the impact energy. Increasing the thickness of steel tube and providing additional transverse confinement by CFRP can enhance the impact-resistant behaviors. Mirmomeni et al. [14] studied the influence of the specimen 4

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size on the impact responses of NACFST. The size-dependent behaviors of the NACFST were found to be a function of the level of confinement, which the circumferential steel tube imposes on the concrete. Wang et al. [15] experimentally studied the resistance of NACFST under close-range blast loads, and found that the NACFST columns are still able to retain a large portion of their axial load

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capacities even after exposure to close-range blast events. The finite element analysis (FEA) models

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have also been wildly used for studying the impact or blast behavior of NACFST [3,16-19]. The

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effects of various structural and load parameters on the impact responses of the NACFST column were fully evaluated, and the CFRP wrapping was found to be a promising strengthening technique

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in controlling the global failure of full scale NACFST columns subjected to vehicular impact. In

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terms of high temperature or fire resistance of the NACFST, valuable results have been obtained by experimental studies [20-22] and numerical analyses [23-25]. However, the researches on combined

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action of fire and impact on NACFST are very scarce. Only few researches [26,27] have been done

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on the impact behaviors of NACFST at elevated temperatures. Mirmomeni et al. [28] investigated the effect of high temperature on the mechanical performances of NACFST which had partial

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damage due to high strain rate loading. Results indicated that for NACFST, variation of residual

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properties is dependent on the level of pre-induced damage as well as the exposed temperature. Although the researches on NACFST can provide references for the estimation of impact resistances of the RACFST after exposure to fire, due to the properties difference between RAC and NAC, it is unsafe to apply RACFST for practical application without verifications based on experimental and numerical analyses. Until now, there is still no research conducted for investigating performance of the RACFST subjected to combined action of elevated temperature exposure and impact loading, such as impact loading before and after fire exposure. In this study, experimental and 5

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numerical investigations were conducted to study the impact behaviors of RACFST. A SHPB impact experiment was carried out on the RACFST with different RCA replacement ratios (0, 50% and 100%) and different exposure temperatures (20 oC, 200 oC, 500 oC and 700 oC). Based on validation by experimental results, numerical simulation was applied to investigate the effects of different

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parameters such as RAC strength, steel strength and steel ratio on the impact properties of RACFST

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after exposure to high temperature. The related results can provide valuable insights into the impact

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design of RACFST after exposure to fire.

2. Experimental program

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2.1 Material properties

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Recycled coarse aggregate used in this test was obtained from a RCA manufacturing supplier in Shanghai, China. Gravel from an aggregate production plant in Changsha, China was used in this test

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as natural coarse aggregate (NCA). The physical properties of the NCA and RCA are shown in Table

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1. RAC samples of 150 mm ×150 mm ×150 mm with RCA replacement ratios of 0%, 50% and 100% were prepared to measure the static compressive strength. The mix proportions of the NAC and RAC

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samples are shown in Table 2. In the mix design, extra water was added for ensuring the same

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effective water to cement ratio of RACs based on the water absorption of RCAs from original condition to saturated surface dry condition [29]. The compressive strengths of RAC with RCA replacement ratios of 0, 50% and 100% were 42.8 MPa, 43.66 MPa and 35.27 MPa respectively. The elastic moduli of RACs were 3.22×104 MPa, 3.16×104 MPa and 2.54×104 MPa respectively. The steel tubes for RACFST preparation were Q235 seamless steel tubes obtained from the same batch of steel products. The average yield strength and elastic modulus of the steel were 320.1 MPa and 201.5 GPa respectively based on the three duplicated samples. 6

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2.2 Specimen preparation Seamless steel tubes with the outer diameter of 96 mm and wall thickness of 2 mm were used for specimen preparation. RAC was cast into each hollow steel tube layer by layer, and then was vibrated by a poker vibrator. The RACFST specimens were naturally cured in the laboratory. In order

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to ensure a uniform state of stress along the specimen longitudinally and avoid apparent strength

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increases because of remarkable end-friction effects, length to diameter ratio was set as 0.5 for this

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SHPB test [26,27,30]. After 28-day curing in the laboratory, specimens were cut to small specimens with height of 50 mm by a mechanical processor. A double sided grinding machine was used to

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further ensure that the two ends of each specimen were precisely parallel to each other. The typical

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RACFST specimen for SHPB impact test is shown as Fig. 1. The design of RACFST specimen is shown in Table 3. In the experiments, three duplicate specimens were prepared for collecting average

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2.3 Experimental program

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value. Thus, a total of 36 RACFST specimens were tested in the study.

RACFST specimens were heated using a program controlled electrical furnace at the Key Laboratory

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of Building Safety and Energy Efficiency (Hunan University), Ministry of Education, P.R China. The

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initial temperature (room temperature) for specimens was about 20 oC and the heating rate was set at 10 oC per minute [31]. After reaching each specified temperature, the temperature was held for 3 hours for ensuring that the specimens are uniformly heated [32]. The RACFST specimens were then naturally cooled down to room temperature. As shown in Fig. 2, a 100 mm cross-sectional SHPB device was applied to experimentally investigate the impact behaviors of RACFST specimens after exposure to different elevated temperatures. The SHPB system is mainly consisted of striking bar, incident bar and transmission bar. 7

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A soft cloth was used to cover one end of the incident bar (close to striking bar) to filter high frequency oscillation of the impact wave and reduce the fluctuation range. Vaseline was smeared on the contact surfaces of the RACFST specimen and pressure bar (incident bar and transmission bar) for reducing the frictions. In the SHPB test, a constant air pressure value of 0.9 MPa was set for each

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specimen, which corresponded to the impact velocity of 8.79 m/s of striking bar.

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In the impact test, after the striking bar impacting the incident bar, incident impulse  i , reflected

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impulse  r and transmission impulse  t are subsequently generated. These impulses can be acquired by the strain gauges on the surface of the incident and transmission bars. The typical

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impulse signals are shown in Fig. 3. Based on these impulses, the stress  , strain  , and strain rate 

EA t AS

 

2 c0 l0

  

2c0 r l0

t

  dt

(2)

r

(3)

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0

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(1)

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

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of the specimen during the impact process can be obtained by Eqs. (1), (2) and (3).

where E and A refer to the elastic modulus and cross-sectional area of SHPB bar; C0 denotes the

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elastic wave velocity of the pressure bars; As and l0 represent the cross-sectional area and length of the RACFST specimen respectively.

3. Experimental results The damage of RAC100FST specimens after exposure to elevated temperatures are shown in Fig. 4. It is obvious that there were no visible damages (microcracks) or large deformation. Only the color of the steel tube and RAC turned dark. Fig. 5 reveals the impact failure patterns of RACFST specimens which were exposed to different elevated temperatures. It shows that after suffering from high 8

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temperature exposure and impact loading, most of the RACFST specimens still kept their original shape but the impact damages became more and more severe with the increase of exposure temperature. For all of the specimens, the results show that the steel tube has yielded under impact loading. Besides, different degree of damage can be also observed from the inner RAC. For example,

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only a small amount of cracks in inner RAC could be found after impact for the specimens without

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high temperature exposure. However, when the elevated temperature reached 700 oC, the obvious

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cracks appeared on the surface of inner RAC and even propagated through the RCAs under impact loading. At the same time, it can be found that the middle of the specimen slightly swelled up after

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impact. Fig. 5 also indicates that the RCA replacement ratios seem not to obviously affect the failure

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pattern of RACFST after impact. Based on the previous study [33], RAC and nanoparticle modified RAC without confinement were crashed into small particles or powder by SHPB impact test.

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However, the RACFST specimens can still maintain their integrity after the combined effects of high

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temperature exposure and impact loading. It implies that the RACFST exhibits sufficient high temperature and impact resistances, which is suitable for the application in structural engineering.

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The impact stress-strain curves of RACFST specimens exposed to elevated temperatures are

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shown in Fig. 6. With the increase of elevated temperatures, the compressive stress of RACFST specimens decreases, while the compressive strain increases. For instance, after exposed to temperature of 200 oC, there was only slight decrease of stress, but the declining trend became more obvious at temperatures of 500 oC and 700 oC. Beyond a certain threshold of high temperature, the impact performance of RACFST was obviously reduced. To be more precise, compared with the RAC100FST specimen at 20 oC, specimen at 200 oC exhibited 1.4 % reduction in the peak stress, while the specimen at 700 oC presented 31.0% reduction in the peak stress. As for the effect of RCA 9

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replacement ratios, the RACFST specimens with higher RCA replacement ratios presented lower compressive stress. The decline became more obvious for higher temperature exposure, which may be due to the poor high temperature resistance of RCA [34]. At temperature of 200 oC, the RAC100SFT exhibited 1.7% lower peak stress than that of RAC0SFT, while at temperature of 700 o

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C, the RAC100SFT presented 9.3% lower peak stress. It can be concluded that increasing the

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content of RCA reduces the high temperature resistance of RACFST and further decreases the impact

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performance. But the variation in the stress-strain curves of RACFST specimens with the different RCA replacement ratios was in a slight range, which is less evident than the variation caused by high

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temperature exposure. It indicates that the RACFST is suitable for the actual structural application in

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terms of the impact or blast resistances after fire exposure.

4. Numerical analysis

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4.1. Numerical FEM model

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Based on the experimental studies, numerical models were established in the ABAQUS/Explicit module. The one-quarter numerical finite element model of RACFST is shown in Fig. 7. The

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pressure bars in the SHPB system and the RACFST specimens are all regular cylinders, therefore the

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one quarter model was adopted to save the computing time. A mesh convergence study was performed for suitable mesh sizes in achieving the balance between the accuracy and computing time. The cross-section size of RACFST specimens and pressure bars are the same, as shown in Figs. 1 and 2. The initial impact velocity of 8.79 m/s along axial was assigned to the striking bar for simulating the impact behaviors of RACFST in the experiments. The trial study shows that the buffer bar did not influence the incident impulse, reflected impulse and transmission impulse. Thus, the buffer bar was omitted in the model. Besides, because the small size of RACFST specimen and long 10

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duration of high temperature treatment, uniform strength was assumed for each section of the RACFST after exposure to elevated temperature. The impact duration was recorded by acquisition system by 0.0032 s, and the time step was set to 0.003 s in the numerical analysis accordingly. Surface-to-surface contact and hard contact in the normal direction were considered between steel

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and RAC in the FEM model. As Vaseline was smeared on the contact surfaces of RACFST specimen

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and the pressure bars (incident bar and transmission bar), the friction between these surfaces were

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ignored. For the contact between steel rube and RAC, coulomb friction model in the tangential direction was adopted, and the friction factor was set as 0.25 [35]. The 8-node brick element with

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reduced integration (C3D8R) was applied in this numerical model.

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4.2. Properties of RAC

The damage plasticity model was adopted for the RAC in this numerical analysis [36-39]. Based on

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the analysis and trial calculation of a large number of experimental data, the stress-strain constitutive

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model for core concrete of NACFST was put forward by Han et al. [40], as shown in Eqs. (4) and (5). This model has been widely used in the simulation of NACFST. Previous researches on this model

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for NACFST and RACFST under ranging from quasi static to impact loading showed that desired

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results [39,41,42]. Thus, the compressive stress-strain relationship for core RACs at room temperature was determined by this numerical model. It should be noted that the length to diameter ratio of the impact specimen in this study was far less than that of quasi-static compression and conventional impact specimens. The RACFST impact specimen has better stability and confinement by the steel tube better when the deformation is large. Han et al. [40] found that when the confinement factor  is greater than a certain value (around 1.0 for circular CFST) due to the sufficient constraint, the core RAC in steel tube wouldn't exhibit the descending segment, and the 11

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stress can be maintained and even slightly increase in a large strain range. As a result, when the confinement factor  is close to the critical value of 1.0, and the real constraint for impact specimens under large strain is obviously greater than that of the corresponding quasi-static ones. The RAC in impact specimens may not have obvious descending segment, as shown in Eq. (5). At

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the same time, there is not sufficient time for developing cracks under impact loading, which also

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reduces the declining trend of stress-strain curves for core RAC. In this study, base on the above

x ,x 1  ( x  1) 2  x

where x   /  o ,

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y

(4) (5)

y   /  o ,  0  f c', r ,  o  ( c  800 r0.2  106 )(1  r /  ) ,  r  f y As / f ck ,r Ac ,

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y  2x  x2 , x  1

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analysis, the stress of RAC was defined as maintaining its peak value after the peak strain.

 c  (1300  12.5 f c',r ) 106 , and  0.5) 7 ]

 ( f c', r )0.5  0.12 , θ = 65.715r 2 - 109.43r  48.989 ;

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r

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  0.5(2.36  105 )[0.25 (

where r refers to RCA replacement ratio; f c',r is compressive strength of the cylindrical RAC;

RACFST.

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f ck , r equals to 0.67 f cu for normal strength of RAC;  r denotes the confinement factor for

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After exposure to elevated temperatures, the peak stress and peak strain of RAC are affected by high temperature damage. Besides, high strain rate may also influence the mechanical properties of core RAC. However, there are limitations for experimental studies to accurately obtain the real stress-strain relationship of core RACs under the combined effects of triaxial confinement, high temperature exposure, and impact loading in this study. Therefore, based on the previous researches [39,42,43], the stress-strain model of core RACs was determined by considering the high temperature exposure and strain rate effect. Eqs. (6) and (7) were proposed by Han et al. [43] and Li 12

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et al. [10] for estimating the residual mechanical properties of core concrete after exposure to elevated temperature, which were applied in this numerical analysis. The strain rate effect of concrete is usually deemed to the combined result of the Stefan effect, the cracking propagation effect and inertial effect [29,44]. When the concrete is not confined by steel tube, it is crashed into

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small pieces under the impact loading and numerous microcracks develop in the cement mortar and

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aggregates. However, from Fig. 5, there was no obvious damage of core RACs during the impact

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process. It implies that there are undoubtedly no obvious Stefan effect and cracking propagation effect for core RAC. Furthermore, due to the constraint provided by steel tube, the inertia effect of

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core RAC is also greatly reduced [30]. Thus, the CEB empirical equation [45] for unconstrained

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concrete probably overestimates the strain rate effects of core RAC. In this study, it is presumed that the strain rate effect was ignored for core RAC, and the strength increase was regard to be caused by

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the confining pressure provided by steel tube. The damage factor which is related to stiffness

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degradation is acquired based on the assumption of strain equivalence and experimental results from Birtel et al. [46]. At the same time, the tension constitutive relationships of core RAC was adopted

o

1  2.4(T  20) 6  10 17

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 o (T ) 

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using the fracture energy model [10].

 o (T )  [1  (1500T  5T 2 )  106 ] o

(6) (7)

4.3. Properties of steel tubes The steel tube used for RACFST was made of low carbon steel. The mass density of the low carbon steel is 7800 kg/m3. The experimental result indicates that steel tube is likely to reach the plastic yield under the impact loading. Elastic-plastic model was adopted for describing the constitutive behavior of the low carbon steel. According to Han et al. [43], a typical two-stage stress-strain curve 13

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was used as shown in Eqs. (8) and (9). In this model, the Poisson's ratio of steel after exposure to elevated temperatures was taken as a constant value of 0.3, and the decrease of elastic modulus under high temperature was estimated according to the research by Lu et al. [47].

  Es (T ) ,    y (T )

(8)

  f y (T )  E1 (T )[   y (T )],    y (T )

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(9)

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where E s (T ) denotes the elastic modulus of the steel after suffering from elevated temperature of

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T ;  y (T )  f y (T ) / Es (T ) refers to the yield strain of steel; E1 (T )  0.01Es (T ) is the elastic modulus in the strengthening phase; f y (T ) represents the yield strength of steel after exposure to

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elevated temperature of T .

f y (T )  f y , T  400C

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The conversion formula of f y (T ) from the yield strength at ambient temperature is: (10)

(11)

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f y (T )  f y [1  2.33  104 (T  20)  5.88  107 (T  20) 2 ], T  400C

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In this study, the strain rate effect of steel is considered based on the Cowper-Symonds model [48]. The dynamic yield function of this model is shown in Eq. (12). (12)

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 dy   y (1   / D 1/ p )

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where  dy denotes the dynamic yield strength of steel;  y represents the yield strength of steel under quasi-static load;  refers to the corresponding strain rate; D and P are material parameters, which are 40.4s 1 and 5 in this study. 4.4. Properties of SHPB pressure bar The pressure bars in the SHPB system were made of high strength steel, which is always considered in elastic state without any damage during the impact process. The elastic modulus and Poisson's ratio of the high strength steel are 210 GPa and 0.3, respectively. The mass density of steel is 7800 14

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kg/m3.

5. Numerical model validation 5.1 Stress wave and impact failure patterns Fig. 8 shows the typical stress wave propagation in specimen (core RAC and steel tube). Before

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0.0012 s, the stress wave transmitted from one end of incident bar to the other end rapidly. At 0.0012

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s, the main part of the stress wave was very close to the interface of incident bar and specimens. At

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this time, the axial stress value of RAC and steel tube were still very low. At 0.0015 s, part of the stress wave was transferred into the specimen, and the axial stress value for both steel tube and RAC

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achieved a large value as shown in Fig. 8. At 0.0018 s, the stress wave already returned and the stress

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wave on the incident bar became reflected wave. The axial stress value of RAC still increases, but the corresponding stress value of steel tube decreased. This indicates that the propagation of the

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stress wave seems slightly faster in steel tube than in RAC. When the time reached 0.0021 s, the

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majority of the stress wave began to depart from the RACFST specimen, and the axial stress of the steel tube and RAC returned to small values.

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Fig. 9 displays the comparison on experimental and numerical impact failure patterns of RACFSTs,

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which are exposed to different elevated temperatures. Generally, the numerical impact failure patters were consistent with the experimental impact failure patterns. Both the numerical and experimental results show that the steel tube yielded during impact loading. Similar with the experimental results, the deformation of specimens increased with the rise of the temperature. It means more severe impact damage in the RACFSTs when suffering from high temperatures. For example, in the numerical result, a marked increase in deformation was observed at temperature of 700 oC.

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5.2 Validation of pulse signal The typical strain data acquired from the incident bar and the transmission bar in FEA model are shown in Fig. 10. According to the different RCA replacement ratios and different exposure temperatures, comparisons were made by selecting parts of the specimens during impact. For ease of

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comparison, the simulated pulse signals were moved to the same start time point as the experimental

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signals. During the real impact experiments, a soft cloth was used to filter high frequency oscillation,

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which was not considered in the FEA model because of the unavailable mechanic parameters of the soft cloth during impact. Therefore, the transmitted pulse of simulated pulse signals presented

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relatively some fluctuations. Furthermore, the simplified end condition (omit the buffer bar) led to

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the pulse signals after the transmission pulse (  t ), which displays some differences between the numerical results and the experimental results, but does not influence the values of incident pulses,

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reflected pulses and transmission pulses. Overall, the simulated pulse signals are in good agreement

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with the experimental pulse signals in terms of the incident and reflected pulse as well as transmitted pulse.

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5.3 Impact stress-strain curves

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According to the two-wave calculation formula, the stress-strain curves can be obtained based on the pulse signals. The comparisons of simulated stress-strain curves and experimental stress-strain curves are shown in Fig. 11. For RACFSTs with different RCA replacement ratios exposed to temperatures of 20 oC and 200 oC, the numerical stress-strain curves exhibited slightly higher stress values compared with the experimental ones. When the suffered temperature reached 500 oC, the numerical stress values were in good agreement with the experimental results. However, for high temperature of 700 oC, the numerical stresses became lower compared to the corresponding 16

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experimental values. In general, the numerical stress-strain curves agreed with the experimental stress-strain curves very well. The differences between the numerical results and the experimental results mainly occurred when the strain was in a large range. But the representative impact strength value of NACFST is usually taken at the end of elastic-plastic stage [27,49]. In this case, the

PT

corresponding strain was still moderate, and the numerical stresses were very close to the

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experimental stresses. In terms of the strain, for NACFST specimen, the maximum numerical strain

SC

agreed well with the corresponding experimental values when the temperatures were less than 500 oC. When exposed to temperature of 700 oC, the maximum numerical strain was slightly greater than that

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of the experimental one (0.0528 and 0.0451 respectively). When the RACFST specimen exposed to

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elevated temperatures lower than 500 oC, compared with the NACFST, the differences between the maximum numerical strain and the maximum experimental strain slightly increased with the increase

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of the RCA replacement ratio. However, all these differences in all elevated temperatures were still

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within acceptable limits. Based on the above analysis, this numerical model on RACFST can be used to conduct parametric analysis.

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For the previous numerical studies on NACFST, most of the numerical models were established

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only based on empirical equations rather than the measured experimental parameters. For all the numerical parameters, it is difficult to obtain them by real experiment. Therefore, some parameters are often provided by empirical equations. However, it should be noted that plenty of parameters only depended by empirical calculation are likely to cause possibilities of error. In this study, the numerical results really are in good agreement with the experimental results. Thus, some of the mechanical parameters were not obtained from direct experimental studies. Although based on the previous studies, it is feasible to simulate the NACFST after high temperature exposure under impact 17

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

loading with the mentioned empirical equations, establishing numerical models through parameters obtained by experiments are still necessary for the future numerical research to build a numerical model, which is suitable for large variety of temperature exposures and impact loading. In conclusion, this numerical model established by empirical equations and experimental results may be appropriate

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for the current parameter rages of high temperature and impact velocity.

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6. Parametric analysis and discussions

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In order to analyze the impact behaviors of RACFST exposed to elevated temperature, parametric analysis was conducted for investigating the effects of different factors such as the RAC strength,

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steel strength and steel ratio on the impact behaviors of RACFST under impact velocity of 8.79 m/s

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after suffering elevated temperatures. As the RCA replacement ratio seems not significantly influence the mechanical behaviors of the RACFST, only the specimens with 100% RCA replacement ratio

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(RAC100FST) was used in the numerical analysis. Related results will provide design references for

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impact resistances of RACFST after suffering high temperature. 5.1 RAC strength

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Parametric analysis was conducted on the impact behaviors of RACFST with different RAC

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strengths from 30 MPa, 40 MPa to 50 MPa. The impact stress-strain curves of RACFST with different RAC strength are shown in Fig. 12. It is found that with the increase of RAC strength, the impact stress of the RACFST increased, but the impact strain decreased. The impact strength improvement was not very obvious. Referring to in Huo et al. [27], the impact stress at the end of the elastic-plastic stage is taken as the impact strength. Fig. 13(a) reveals the impact strength improvement of RACFST increases with the increase of the core RAC strength. The impact strength enhancement factor ( K c ) was defined in Eq. (13). Increasing RAC strength seems not significantly 18

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

to improve the RACFST impact strength. For example, the maximum K c was only about 1.10 for RAC100FST20 with C50 RAC strength. Moreover, with the increase of exposure temperature, the K c of RACSFT tended to decrease. At the temperature of 200 oC, the impact strengths of RACFST

with RAC strength grade of C40 and C50 were 1.04 and 1.10 times of that of RACFST with RAC

PT

strength grade of C30 respectively. At high temperature of 500 oC, these ratios decreased to 1.02 and

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1.08 times respectively. Then, they further declined to 1.01 and 1.02 times respectively for high

SC

temperature of 700 oC. The enhancement of impact strength of RACFST caused by the increase of RAC strength grade can be ignored when at the temperature of 700 oC. Generally, RAC is more

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sensitive than steel tube to the high temperature exposure. Therefore, after exposure to high

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temperature, the contribution of core RAC to the overall mechanical properties of RACFST is actually reduced. Therefore, increasing the RAC strength is unlikely to be a effective way for

D

enhancing the impact properties of RACFST, especially after exposure to temperatures more than

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500 oC. To improve the impact resistances of RACFST in structure design, it seems not an effective

Kc =

 dm  d 30

(13)

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to increase the RAC strength.

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where  dm refers to impact strength of RACFST which has core RAC strength of Cm from 30 MPa, 40 MPa to 50 MPa;  d 30 denotes the impact strength of RACFST which has RAC strength of 30 MPa. The maximum strain in the impact stress-strain curve was used to evaluate the deformation capacity of RACFST under impact loading. The maximum strain for RACFST with core RAC strength of 30 MPa, 40 MPa and 50 MPa after exposed to high temperature of 700 oC were 0.0552, 0.0544 and 0.053 respectively. Fig. 13(b) reveals the impact deformation improvement of the 19

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

RACFST as the increase of the core RAC strength. The impact deformation decrease coefficient (  c ) was defined as shown in Eq. (14).

ηc =

 dm  d 30

(14)

where  dm refers to maximum impact strain of RACFST which has core RAC strength of Cm

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from 30 MPa, 40 MPa to 50 MPa;  d 30 denotes maximum impact strain of RACFST which has

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core RAC strength of 30 MPa.

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With the increase of RAC strength,  c of RACFST after high temperature exposure decreased

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accordingly. In other word, the RACFST with higher core RAC strength exhibited relatively less impact deformation. However,  c of RACFST increased with the increase of exposure temperature.

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For temperature of 700 oC,  c for all the RAC strength was close to 1.0. Thus, it indicates that when exposure temperature reaches a certain value, increasing RAC strength has no obvious effect

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on the impact deformation of RACFST. The lowest  c appeared at temperature of 20 oC. For instance, the maximum impact deformation of RACFST with core RAC strength of 40 MPa and 50 MPa were about 0.922 and 0.818 times of the maximum impact deformation of RACFST with core

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RAC strength of 30 MPa respectively. On the other hand,  c was 0.986 and 0.960 times

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respectively at high temperature of 700 oC. Thus, in impact resistance design, increasing the RAC strength can only slightly reduce the deformation of RACSFT structures. 5.2 Steel strength The effect of steel strength from 235 MPa, 345 MPa to 420 MPa on the impact behaviors of RACFST was also investigated. The impact stress-strain curves of RACFST with different steel strength exposed to elevated temperatures are shown in Fig. 14. It is obvious that the shape of impact stress-strain curves for RACFST with different steel strength was similar to each other. Fig. 15(a) 20

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

reveals the rate of enhancement for the impact strength of RACFST with the increase of steel strength, in which the definition of impact strength enhancement factor ( K s ) is shown in Eq. (15). It is apparent that the impact strength of RACFST increased with the increase of the steel strength. After exposed to elevated temperatures, the impact strength of RACFST exhibited obvious

PT

improvements by adopting higher strength steel, and the impact strength improvement was more

RI

obvious when RACFST exposed to higher temperatures. The maximum increase of impact strength

SC

of RACFST reached around 1.47 times. For example, when exposed to temperature of 20 oC, the impact strengths of RACFST with 345 MPa and 420 MPa steel tubes were about 1.19 and 1.29 times

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of the corresponding impact strength of RACFST with 235 MPa steel respectively. When exposed to

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temperature of 500 oC, these ratios increased to 1.28 and 1.46 times respectively. When exposed to elevated temperature of 700 oC, this impact strength enhancement factors further increased to 1.30

D

and 1.47 times respectively. Exposed to high temperature of 200 oC and 500 oC, due to the high

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temperature sensitivity of RAC, the core RAC exhibited obvious strength degradation but the strength degradation of steel was much mild. Thus, the steel tubes played a more important role in

o

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the loading capacity than RAC after high temperature exposure. When the temperature rose to 700

AC

C, although the core RAC strength was further reduced, the strength of steel also experienced a

great decline, which reduced the strength difference between RACFSTs with different steel strength. For example, the enhancement for impact strength of RACFST after exposed to temperature of 700 o

C by increasing the steel strength is only slightly higher than that of temperature of 500 oC. After

exposure to higher temperature, the steel tube contributed greater contribution for impact resistance of RACFST than that at lower temperature, and the steel strength difference is still significant after high temperature exposure. Thus, to protect fire heated RACFST from impact failure, adopting high 21

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strength steel is one of efficient methods for improving the impact resistances for RACSFT structures.

 dn  d 235

Ks =

(15)

Where  dn refers to impact strength of RACFST which has steel strength of Qn from 235 MPa,

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345 MPa to 420 MPa;  d 235 denotes impact strength of RACFST which has steel strength of 235

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MPa.

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Fig. 15(b) displays the impact deformation decrease coefficient (  s ) of RACFST after exposure to

NU

high temperatures. For instance, after exposed to temperature of 700 oC, the maximum impact strain of RACFST with Q235 and Q345 steel were 0.0616 and 0.0529 respectively. The definition of  s is

MA

shown in Eq. (16). It is obvious that as the enhancement of the steel strength,  s decreased accordingly, showing better deformation resistance. However, with the increase in the exposure

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D

temperature, the improvement of impact deformation of RACFST by increasing the steel strength became obviously less. Moreover, when beyond a certain temperature, adopting higher strength steel

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has no obvious effect for improving the impact deformation of RACFST. It shows that the lowest  s appeared at temperature of 20 oC. The impact deformation of RACFST with steel strengths of 345

AC

MPa and 420 MPa were 0.697 and 0.553 times of the impact deformation of RACFST with steel strength of 235 MPa. When the exposure temperature reached 700 oC, these ratios increased to 0.859 and 0.771 times respectively. However, compared with enhancing the RAC strength, increasing steel strength is more effective for decreasing the impact deformation of RACFST, which can be considered in the impact or blast resistance design of RACFST structures.

ηs =

 dn  d 235

(16)

22

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where  dn refers to the maximum impact strain of RACFST which has steel strength of Qn from 235 MPa, 345 MPa to 420 MPa;  d 235 denotes the maximum impact strain of RACFST which has steel strength grade of 235 MPa. 5.3 Steel ratio

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The effect of different steel ratio on the impact behaviors of the RACFST was studied by changing

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the thickness of the steel tube. The steel tube thicknesses of RACFST were 2.0 mm (steel ratio of

SC

8.9%), 3.0 mm (steel ratio of 13.8%) and 4.0 mm (steel ratio of 19.0%) respectively. After exposure to high temperatures, the impact stress-strain curves of RACFST with different steel thickness are

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displayed in Fig. 16. It is clear that the impact stress of RACFST was apparently enhanced, and the

MA

impact strain was reduced with the increase of steel tube thickness. For the RACFST with steel thickness of 4 mm, the impact strength reached around 150 MPa at temperature of 200 oC. After

D

exposure to temperature of 700 oC, the impact strength reached around 110 MPa. The improvements

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by increasing steel ratio were much greater compared with those by increasing the strengths of RAC and steel tube.

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Fig. 17(a) displays the improvement of impact strength of RACFST after exposure to elevated

AC

temperatures by increasing steel tube thicknesses. The impact strength enhancement factor ( K ) is defined in Eq. (17). K of RACFST was greatly enhanced by applying higher thicknesses of steel tubes, which was especially obvious for the RACFST suffered from higher temperatures. The maximum improvement achieved about 1.81 times. For example, at temperature of 20 oC, the impact strength of RACFST with steel tube thickness of 3 mm and 4 mm were 1.21 and 1.37 times of the impact strength of RACFST with steel tube thickness of 2 mm respectively. After exposure to temperature of 500 oC, these ratios increased to around 1.28 and 1.52 times. When the exposure 23

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temperature reached 700 oC, the impact strength improvement further increased to 1.44 times and 1.81 times respectively. On the other hand, the confinement factor of RACFST means constraint of core RAC by steel tube was also enhanced. Increasing of the steel tube wall thickness effectively improved the impact strength of RACFST exposed to high temperatures. Particularly, with the

PT

increase of elevated temperature, the core RAC experienced greater decline than that of steel tube.

RI

Thus, the steel tube contributes more to the impact resistances of RACFST, which makes the

SC

advantage of RACFST with higher steel ratio becomes more obvious for the RACSFT structures

Kα =

 dt d2

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with consideration of impact or blast resistances. (17)

MA

where  dt refers to the impact strength of RACFST which has steel tube wall thickness of t from 2 mm, 3 mm to 4 mm;  2 denotes impact strength of RACFST which has steel tube wall thickness

D

of 2 mm.

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Fig. 17(b) shows the impact deformation decrease coefficient (  ) of RACFST after different high temperature and definition of  is shown in Eq. (18). With the increase in the steel tube thicknesses,

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 decreased accordingly. It means that the RACFST with higher thickness of steel tubes exhibit

AC

smaller deformation under impact loadings. Similar to the  by changing the RAC strength and steel strength,  increased with the increase of elevated temperature. Thus, for RACFST subjected to increasing elevated temperature, the reduction of impact deformation by raising the steel ratio becomes less effective. For example, for the RACFST exposed to temperature of 700 oC, the impact deformations of the RACFST with thicknesses of 3 mm and 4 mm steel tube were reduced to around 0.663 and 0.494 times of the corresponding impact deformation of the RACFST with steel tube thickness of 2.0 mm respectively. The  values were lower than those achieved by enhancing the 24

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RAC strength and steel strength. Thus, increasing the steel ratio obviously improved the impact deformation resistance of RACFST. The lowest  was found when exposed to temperature of 200 C. For example, the  values were about 0.567 and 0.373 for RACFST with wall thickness of 3

o

mm and 4 mm respectively. Therefore, in the practical structure design, increasing the thickness of

PT

steel tube effectively enhances the deformation resistance capacity of RACSFT under impact or blast

 dt d2

(18)

SC

ηα =

RI

after fire exposure.

NU

where  dt refers to the maximum impact strain of RACFST which has steel wall thickness of t from 2 mm, 3 mm to 4 mm;  d 2 denotes the maximum impact strain of RACFST which has steel

MA

tube thickness of 2 mm.

7. Conclusions

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D

The impact behaviors of RACFST after exposure to elevated temperatures were experimentally and numerically investigated in this study. The effects of RAC strength, steel strength and steel ratio on

CE

the impact mechanical behaviors and deformation properties of RACFST were conducted by parametric analysis on as well. The main conclusions were drawn as following:

AC

(1) During the SHPB test, the stress wave propagation process, pulse signals and stress-strain curves acquired by numerical simulation were consistent with the experimental results. The numerical model was validated by the experimental results for investigation on impact behaviors of RACFST exposed to elevated temperatures. (2) Increasing the RAC strength slightly enhanced the impact strength of RACFST, and the enhancement magnitude declined with the increase of elevated temperature. With the increase of core RAC strength, the impact deformation of RACFST decreased, but the reduction magnitude 25

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

tended to decline with the increase of elevated temperature. (3) Increasing steel strength exhibited enhanced impact strength of RACFST after exposure to elevated temperature, which was more obvious when suffered from increased temperature. When adopting higher steel strength, the impact deformation greatly decreased, but the reduction

PT

magnitude decreased with the increase of elevated temperature.

RI

(4) Compared with increasing the RAC strength and steel strength, increasing the steel ratio was a

SC

more effective way for improving the impact strength and decreasing impact deformation of RACFST. The improvement of impact strength was more obvious for RACFST when subjected

NU

to higher elevated temperature.

MA

(5) Increasing steel ratio is an effective way for improving the impact behaviors of RACFST, but it brings about an increase in cost. Thus, in the practical design, the impact performances of

D

RACSFT need more comprehensive analysis based on the basic design parameters, such as RAC

energy.

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Acknowledgement

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strength, RCA replacement ratio, steel strength, thickness of steel tube (steel ratio) and impact

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The authors gratefully acknowledge the research grants from the Australian Research Council (DE150101751 and IH150100006). The authors are also grateful for the financial supports from University of Technology Sydney Early Career Researcher and Blue Sky Research Scheme Grants, Australia, as well as the National Engineering Laboratory for High-speed Railway Construction, Central South University, P.R. China.

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References [1] W.G. Li, J. Xiao, C. Shi, C.S. Poon. Structural behaviour of composite members with recycled aggregate concrete - an overview. Advances in Structural Engineering, 18(6) (2015) 919-938. [2] W.G. Li, Z. Sun, Z. Luo, S.P. Shah. Influence of relative mechanical strengths between new and

PT

old cement mortars on the crack propagation of recycled aggregate concrete. Journal of

RI

Advanced Concrete Technology, 15(3) (2017) 110-125.

SC

[3] F. Zhang, C. Wu, H. Wang, Y. Zhou. Numerical simulation of concrete filled steel tube columns against BLAST loads. Thin-Walled Structures, 92 (2015) 82-92.

NU

[4] J. Xiao, W.G. Li, Y. Fan, X. Huang. An overview of study on recycled aggregate concrete in

MA

China (1996-2011). Construction and Building Materials, 31 (2012) 364-383. [5] G. Li, X. Zhao, P. Wang, X. Liu. Behaviour of concrete-filled steel tubular columns

D

incorporating fly ash. Cement and Concrete Composites, 28(2) (2006) 189-196.

PT E

[6] J. Xiao, W.G. Li, C.S. Poon. Recent studies on mechanical properties of recycled aggregate concrete in China-A review. Science China Technological Sciences, 55(6) (2012) 1463-1480.

CE

[7] Y.F. Yang, Z.C. Zhang, F. Fu. Experimental and numerical study on square RACFST members

AC

under lateral impact loading. Journal of Constructional Steel Research, 111 (2015) 43-56. [8] A.S. Shakir, Z.W. Guan, S.W. Jones. Lateral impact response of the concrete filled steel tube columns with and without CFRP strengthening. Engineering Structures, 116 (2016) 148-162. [9] Y.F. Yang, R. Hou. Experimental behaviour of RACFST stub columns after exposed to high temperatures. Thin-Walled Structures, 59(4) (2012) 1-10. [10] W.G. Li, Z. Luo, Z. Tao, W.H. Duan, S.P. Shah. Mechanical behavior of recycled aggregate concrete-filled steel tube stub columns after exposure to elevated temperatures. Construction and 27

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

Building Materials, 146 (2017) 571-581. [11] M. Yousuf, B. Uy, Z. Tao, A. Remennikov, R. Liew. Behaviour and resistance of hollow and concrete-filled mild steel columns due to transverse impact loading. Australian Journal of Structural Engineering, 13(1) (2012) 65-80.

PT

[12] M. Yousuf, B. Uy, Z. Tao, A. Remennikov, R. Liew. Transverse impact resistance of hollow and

RI

concrete filled stainless steel columns. Journal of Constructional Steel Research, 82 (2013)

SC

177-189.

[13] Y. Xiao, Y. Shen. Impact behaviors of CFT and CFRP confined CFT stub columns. Journal of

NU

Composites for Construction, 16(6) (2012) 662-670.

MA

[14] M. Mirmomeni, A. Heidarpour, X. Zhao, R. Al-Mahaidi, J. Packer. Size-dependency of concrete-filled steel tubes subject to impact loading. International Journal of Impact Engineering,

D

100 (2017) 90-101.

PT E

[15] H. Wang, C. Wu, F. Zhang, Q. Fang, H. Xiang, P. Li, Z. Li, Y. Zhou, Y. Zhang, J. Li. Experimental study of large-sized concrete filled steel tube columns under blast load.

CE

Construction and Building Materials, 134 (2017) 131-141.

AC

[16] L. Chen, Y. Xiao, G. Xiao, C. Liu, A. Agrawal. Test and numerical simulation of truck collision with anti-ram bollards. International Journal of Impact Engineering, 75(1) (2015) 30-39. [17] M.I. Alam, S. Fawzia, X. Liu. Effect of bond length on the behaviour of CFRP strengthened concrete-filled steel tubes under transverse impact. Composite Structures, 132 (2015) 898-914. [18] S. Aghdamy, D.P. Thambiratnam, M. Dhanasekar, S. Saiedi. Computer analysis of impact behavior of concrete filled steel tube columns. Advances in Engineering Software, 89 (2015) 52-63. 28

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

[19] M.I. Alam, S. Fawzia, X.L. Zhao. Numerical investigation of CFRP strengthened full scale CFST columns subjected to vehicular impact. Engineering Structures, 126 (2016) 292-310. [20] L.H. Han, W.H. Wang, H.X. Yu. Experimental behaviour of reinforced concrete (RC) beam to concrete-filled steel tubular (CFST) column frames subjected to ISO-834 standard fire.

PT

Engineering Structures, 32(10) (2010) 3130-3144.

RI

[21] M. Neuenschwander, M. Knobloch, M. Fontana. ISO standard fire tests of concrete-filled steel

SC

tube columns with solid steel core. Journal of Structural Engineering, 143(4) (2016) 04016211. [22] Y. Yao, X.X. Hu. Cooling behavior and residual strength of post-fire concrete filled steel tubular

NU

columns. Journal of Constructional Steel Research, 112 (2015) 282-292.

MA

[23] Y.F. Wang, Y.S. Ma, B. Han, S. Deng. Temperature effect on creep behavior of CFST arch bridges. Journal of Bridge Engineering, 18(12) (2013) 1397-1405.

D

[24] V. Albero, A. Espinos, M. Romero, A. Hospitaler, G. Bihina, C. Renaud. Proposal of a new

PT E

method in EN1994-1-2 for the fire design of concrete-filled steel tubular columns. Engineering Structures, 128 (2016) 237-255.

CE

[25] H. Guo, X. Long, Y. Yao. Fire resistance of concrete filled steel tube columns subjected to

AC

non-uniform heating. Journal of Constructional Steel Research, 128 (2017) 542-554. [26] J. Huo, Q. Zheng, B. Chen, Y. Xiao. Tests on impact behaviour of micro-concrete-filled steel tubes at elevated temperatures up to 400 °C. Materials and Structures, 42(10) (2009) 1325-1334. [27] J. Huo, Y. He, B. Chen. Experimental study on impact behaviour of concrete-filled steel tubes at elevated temperatures up to 800 °C. Materials and Structures, 47(1) (2014) 263-283. [28] M. Mirmomeni, A. Heidarpour, X.L. Zhao, J.A. Packer. Effect of elevated temperature on the mechanical properties of high-strain-rate-induced partially damaged concrete and CFSTs. 29

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

International Journal of Impact Engineering, (2017), dx.doi.org/10.1016/j.ijimpeng.2017.02.006. [29] J. Xiao, L. Li, L. Shen, C.S. Poon. Compressive behaviour of recycled aggregate concrete under impact loading. Cement and Concrete Research, 71 (2015) 46-55. [30] Y. Xiao, J. Shan, Q. Zheng, B. Chen. Experimental studies on concrete filled steel tubes under

PT

high strain rate loading. Journal of Materials in Civil Engineering, 21(10) (2009) 569-577.

RI

[31] J.S. Huo, Y.M. He, L.P. Xiao, B.S. Chen. Experimental study on dynamic behaviours of concrete

SC

after exposure to high temperatures up to 700 °C. Materials and Structures, 46(1-2) (2012) 255-265.

NU

[32] G.G. Mohamedbhai. Effect of exposure time and rates of heating and cooling on residual

MA

strength of heated concrete. Magazine of Concrete Research, 38(136) (1986) 151-158. [33] W.G. Li, Z.Y. Luo, C. Long, C.Q. Wu, W.H. Duan, S.P. Shah. Effects of nanoparticle on the

PT E

112 (2016) 58-66.

D

dynamic behaviors of recycled aggregate concrete under impact loading. Materials and Design,

[34] C. Laneyrie, A.L. Beaucour, M.F. Green, R.L. Hebert, B. Ledesert, A. Noumowe. Influence of

CE

recycled coarse aggregates on normal and high performance concrete subjected to elevated

AC

temperature. Construction and Building Materials, 2016, 111: 368-378. [35] M. Yousuf, B. Uy, Z. Tao, A. Remennikov, R. Richard Liew. Impact behaviour of pre-compressed hollow and concrete filled mild and stainless steel columns. Journal of Constructional Steel Research, 96(96) (2014: 54-68. [36] J. Xiao, W.G. Li, D.J. Corr, S.P. Shah. Simulation study on the stress distribution in modeled recycled aggregate concrete under uniaxial compression. Journal of Materials in Civil Engineering, 25(4) (2013) 504-518. 30

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

[37] J. Xiao, W.G. Li, D.J. Corr, S.P Shah.. Effects of interfacial transition zones on the stress-strain behavior of modeled recycled aggregate concrete. Cement and Concrete Research, 52(10) (2013) 82-99. [38] W.G. Li, Z.H. Sun, Z.Y. Luo, S.P. Shah. Influence of relative mechanical strengths between new

PT

and old cement mortars on the crack propagation of recycled aggregate concrete. Journal of

RI

Advanced Concrete Technology, 15(3) (2017) 110-125.

SC

[39] Y.F. Yang, Z.C. Zhang, F. Fu. Experimental and numerical study on square RACFST members under lateral impact loading. Journal of Constructional Steel Research, 111 (2015) 43-56.

MA

Press,Beijing China, 2007. (in Chinese).

NU

[40] L.H. Han. Concrete-filled steel tubular structures-theory and practice. China Science

[41] L.H. Han, G.H. Yao, Z. Tao. Performance of concrete-filled thin-walled steel tubes under pure

D

torsion. Steel Construction, 45(1) (2008) 24-36.

PT E

[42] R. Wang, L.H. Han, C.C. Hou. Behavior of concrete filled steel tubular (CFST) members under lateral impact: Experiment and FEA model. Journal of Constructional Steel Research. 80 (2013)

CE

188-201.

AC

[43] L.H. Han, Y.F. Yang, H. Yang, J.S. Huo. Residual strength of concrete-filled RHS columns after exposure to the ISO-834 standard fire. Thin-Walled Structures, 40(12) (2002) 991-1012. [44] J. Li, X. Ren. A review on the constitutive model for static and dynamic damage ofconcrete. Advances in Mechanics, 40 (3) (2010) 284-297. (in Chinese). [45] CEB. Concrete structures under impact and impulsive loading. Synthesis Report, Bulletin d’Information, No. 187, Lausanne: Comite Euro-International du Beton, (1988) 205-206. [46] V. Birtel, P. Mark. Parameterised finite element modelling of RC beam shear failure//ABAQUS 31

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

Users’ Conference. (2006) 95-108. [47] J. Lu, H. Liu, Z. Chen, X. Liao. Experimental investigation into the post-fire mechanical properties of hot-rolled and cold-formed steels. Journal of Constructional Steel Research, 121 (2016) 291-310.

PT

[48] Y. Deng, C.Y. Tuan. Design of concrete-filled circular steel tubes under lateral impact. ACI

RI

Structural Journal, 110(4) (2013) 691-701.

SC

[49] L.H. Han, X.L. Zhao, Z. Tao. Tests and mechanics model of concrete-filled SHS stub columns,

AC

CE

PT E

D

MA

NU

columns and beam-columns. Steel and Composite Structures, 1(1) (2001) 51-74.

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Table Table 1 Physical properties of NCA and RCA Table 2 Mix design proportions of NAC and RAC

RI

PT

Table 3 RACFST specimen design for SHPB impact test

Coarse aggregate

Bulk density

type

grading (mm)

(kg/m3)

NCA

5~20

RCA

5~20

Apparent density

NU

Aggregate

SC

Table 1 Physical properties of NCA and RCA

Water absorption

(%)

1420

2752

0.67

1235

2637

8.54

AC

CE

PT E

D

MA

(kg/m3)

33

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

Table 2 Mix design proportions of NAC and RAC Sand

NCA

RCA

Mixing water

Additional water

(kg/m3)

(kg/m3)

(kg/m3)

(kg/m3)

(kg/m3)

(kg/m3)

NAC

402

645

1148

0

RAC50

402

645

574

RAC100

402

645

0

PT

Cement

205

0

574

205

28

1148

205

56

AC

CE

PT E

D

MA

NU

SC

RI

Type

34

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

Table 3 Design of RACFST specimens for SHPB impact test

0

20

RAC0FST-500

96  48

0

500

RAC50FST-20

96  48

50

20

RAC50FST-500

96  48

50

500

RAC100FST-20

96  48

100

20

RAC100FST-500

96  48

100

500

RAC0FST-200

96  48

0

200

RAC0FST-700

96  48

0

700

RAC50FST-200

96  48

50

200

RAC50FST-700

96  48

50

700

RAC100FST-200

96  48

100

200

RAC100FST-700

96  48

100

700

RI

96  48

NU

RAC0FST-20

PT

Dimension (mm)

MA

Dimension (mm)

r (%) T (oC)

Specimens

SC

r (%) T (oC)

Specimens

Note: the number between 'RAC’ and ‘FST' refers to the RCA replacement ratio; the number after 'RACFST' refers to the exposed

AC

CE

PT E

D

elevated temperature; the 'RAC0FST' is equal to 'NACFST'.

35

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

Figures Fig. 1. RACFST specimen for SHPB test Fig. 2. SHPB apparatus for impact test: (a) Schematic diagram of SHPB system; (b) SHPB impact setup

PT

Fig. 3. Typical pulse signals during SHPB impact test

RI

Fig. 4. Typical appearance of RACFST after high temperatures exposure (RAC100FST)

SC

Fig. 5. Impact failure patterns of RACFST exposed to high temperature: (a) RAC0FST; (b) RAC50FST; (c) RAC100FST

NU

Fig. 6. Impact stress-strain curves of RACFST exposed to elevated temperature: (a) Exposed to

MA

temperature of 20 oC; (b) Exposed to temperature of 200 oC; (c) Exposed to temperature of 500 oC; (d) Exposed to temperature of 700 oC

D

Fig. 7. FEA model of SHPB impact experiment

PT E

Fig. 8. Stress wave propagation in core RAC and steel tube Fig. 9. Impact failure patters of RAC100SFT after exposure to high temperatures: (a) Experimental

CE

failure patterns of RAC100SFT; (b) Numerical failure patterns of RAC100SFT

AC

Fig. 10. Comparisons of experimental pulse signals and simulated pulse signals Fig. 11. Comparisons on simulated stress-strain curves and experimental imapct stress-strain curves: (a) RCA replacement ratio of 0%; (b) RCA replacement ratio of 50%; (c) RCA replacement ratio of 100% Fig. 12. Impact stress-strain curves of RACFST with different RAC strengths: (a) Exposed to temperature of 20 oC; (b) Exposed to temperature of 200 oC; (c) Exposed to temperature of 500 oC; (d) Exposed to temperature of 700 oC 36

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

Fig. 13. Effect of RAC strength on impact mechanical property and impact deformation of RACFST exposed to high temperature: (a) Impact strength enhancement factor; (b) Impact deformation decrease coefficient Fig. 14. Impact stress-strain curves of RACFST with different steel strengths: (a) Exposed to

PT

temperature of 20 oC; (b) Exposed to temperature of 200 oC; (c) Exposed to temperature of 500 oC;

RI

(d) Exposed to temperature of 700 oC

SC

Fig. 15. Effect of steel strength on impact mechanical property and impact deformation of RACFST exposed to high temperature: (a) Impact strength enhancement factor; (b) Impact deformation

NU

decrease coefficient

MA

Fig. 16. Impact stress-strain curves of RACFST with different steel ratios: (a) Exposed to temperature of 20 oC; (b) Exposed to temperature of 200 oC; (c) Exposed to temperature of 500 oC;

D

(d) Exposed to temperature of 700 oC

PT E

Fig. 17. Effect of steel ratio on impact mechanical property and impact deformation of RACFST exposed to high temperature: (a) Impact strength enhancement factor; (b) Impact deformation

AC

CE

decrease coefficient

37

ACCEPTED MANUSCRIPT

SC

RI

PT

Revised JMAD-D-17-04650R1

(b) Specimen dimension

NU

(a) RACFST specimen

AC

CE

PT E

D

MA

Fig. 1. RACFST specimen for SHPB test

38

ACCEPTED MANUSCRIPT

SC

RI

PT

Revised JMAD-D-17-04650R1

PT E

D

MA

NU

(a) Schematic diagram of SHPB system

(b) SHPB impact setup

AC

CE

Fig. 2. SHPB apparatus for impact test

39

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

4 3

PT

1 0 -1

RI

Voltage (V)

2

-2

-4

Incident and reflected pulse Transmitted pulse

0.0000

0.0004

0.0008

0.0012

NU

Time (s)

SC

-3

0.0016

AC

CE

PT E

D

MA

Fig. 3. Typical pulse signals during SHPB impact test

40

ACCEPTED MANUSCRIPT

(a) 20 oC

SC

RI

PT

Revised JMAD-D-17-04650R1

(b) 200 oC

(c) 500 oC

(d) 700 oC

AC

CE

PT E

D

MA

NU

Fig. 4. Typical appearance of RAC100FST after exposure to high temperatures

41

ACCEPTED MANUSCRIPT

PT

Revised JMAD-D-17-04650R1

NU

SC

RI

(a) RAC0FST

CE

PT E

D

MA

(b) RAC50FST

(c) RAC100FST

AC

Fig. 5. Impact failure patterns of RACFST exposed to high temperature

42

ACCEPTED MANUSCRIPT

120

120

100

100

80

80

Stress (MPa)

40

o

RAC0FST (20 C) o RAC50FST (20 C) o RAC100FST (20 C)

20 0 0.00

0.01

0.02

60 40

o

RAC0FST (200 C) o RAC50FST (200 C) o RAC100FST (200 C)

20

0.03

0.04

0 0.00

0.05

0.01

PT

60

0.02

RI

Stress (MPa)

Revised JMAD-D-17-04650R1

Strain 

0.04

0.05

(b) Exposed to temperature of 200 oC

SC

(a) Exposed to temperature of 20 oC

0.03

Strain 

120

120

NU

100 80

Stress (MPa)

80

40

0.02

0.03

Strain 

D

0.01

0.04

0.05

60 40 20 0 0.00

PT E

0 0.00

o

RAC0FST (500 C) o RAC50FST (500 C) o RAC100FST (500 C)

o

RAC0FST (700 C) o RAC50FST (700 C) o RAC100FST (700 C)

0.01

0.02

0.03

0.04

Strain 

(c) Exposed to temperature of 500 oC

(d) Exposed to temperature of 700 oC

CE

20

MA

60

Fig. 6. Impact stress-strain curves of RACFST exposed to elevated temperature

AC

Stress (MPa)

100

43

0.05

ACCEPTED MANUSCRIPT

SC

RI

PT

Revised JMAD-D-17-04650R1

AC

CE

PT E

D

MA

NU

Fig. 7. FEM numerical model of SHPB impact test on RACFST

44

ACCEPTED MANUSCRIPT

(b) t=0.0015

AC

CE

PT E

D

MA

NU

(a) t=0.0012

SC

RI

PT

Revised JMAD-D-17-04650R1

(c) t=0.0018

45

ACCEPTED MANUSCRIPT

PT

Revised JMAD-D-17-04650R1

RI

(d) t=0.0021

AC

CE

PT E

D

MA

NU

SC

Fig. 8. Stress wave propagation in the core RAC and steel tubes

46

ACCEPTED MANUSCRIPT

RI

PT

Revised JMAD-D-17-04650R1

MA

NU

SC

(a) Experimental failure patterns of RAC100SFT

D

(b) Numerical failure patterns of RAC100SFT

AC

CE

PT E

Fig. 9. Impact failure patters of RAC100SFT after exposure to high temperatures

47

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

0.0010

0.0002

0.0000 -0.0002 -0.0004

0.0000

SC

0.0002

Strain (

Strain (

0.0004

RI

0.0004

0.0006

-0.0002 -0.0004

Incident and reflected pulse

-0.0008 -0.0005 0.0000 0.0005 0.0010 0.0015 0.0020 0.0025 0.0030

-0.0006 0.0000

MA

Time (s)

NU

-0.0006

PT

0.0006 Experiment Simulation

0.0008

Transmitted pulse Experiment Simulation

0.0005

0.0010

0.0015

0.0020

0.0025

0.0030

0.0020

0.0025

0.0030

Time (s)

(a) RAC0FST20

0.0010 Experiment Simulation

D

0.0008 0.0006

CE

-0.0006

0.0002

Strain (

0.0000 -0.0002 -0.0004

0.0004

PT E

0.0002

0.0000 -0.0002 -0.0004

Incident and reflected pulse

-0.0008 -0.0005 0.0000 0.0005 0.0010 0.0015 0.0020 0.0025 0.0030

-0.0006 0.0000

Time (s)

AC

Strain (

0.0004

0.0006

Transmitted pulse Experiment Simulation

0.0005

0.0010

0.0015

Time (s)

(b) RAC50FST200

48

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

0.0010

0.0006 Experiment Simulation

0.0008

0.0004

0.0006

0.0002

Strain (

0.0002 0.0000 -0.0002 -0.0004 -0.0006

0.0000 -0.0002

Transmitted pulse Experiment Simulation

-0.0004

Incident and reflected pulse -0.0006 0.0000

-0.0008 -0.0005 0.0000 0.0005 0.0010 0.0015 0.0020 0.0025 0.0030

0.0005

0.0010

0.0006

0.0025

0.0030

0.0020

0.0025

0.0030

SC

Experiment Simulation

0.0004 0.0002

0.0002

Strain (

0.0004

NU

0.0006

0.0000 -0.0002

MA

-0.0004

Incident and reflected pulse

-0.0008 -0.0005 0.0000 0.0005 0.0010 0.0015 0.0020 0.0025 0.0030

0.0000

-0.0002 -0.0004

-0.0006 0.0000

Transmitted pulse Experiment Simulation

0.0005

0.0010

0.0015

Time (s)

D

Time (s)

PT E

(d) RAC100FST700

CE

Fig. 10. Comparisons of experimental pulse signals and simulated pulse signals

AC

Strain (

0.0020

RI

(c) RAC100FST500

-0.0006

0.0015

Time (s)

Time (s)

0.0008

0.0010

PT

Strain (

0.0004

49

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

160

160

o

Exposed to temperature of 20 C

o

120

120

100

100

60 40

60 40

RAC0FST20 experiment RAC0FST20 simulation

0.01

0.02

0.03

RAC0FST200 experiment RAC0FST200 simulation

20

0.04

0 0.00

0.05

160

Exposed to temperature of 500 C

0.05

o

NU

60 40 RAC0FST500 experiment RAC0FST500 simulation

0.02

0.03

80 60 40

RAC0FST700 experiment RAC0FST700 simulation

20 0 0.00

0.05

0.01

0.02

0.03

0.04

0.05

Strain (

PT E

Strain (

0.04

100

D

0.01

Stress (MPa)

80

MA

Stress (MPa)

0.04

120

100

0 0.00

0.03

Strain (

Exposed to temperature of 700 C

140

120

20

0.02

160

o

140

0.01

SC

Strain (

RI

0 0.00

80

PT

80

20

Exposed to temperature of 200 C

140

Stress (MPa)

Stress (MPa)

140

(a) RCA replacement ratio of 0%

160

160 o

o

60 40 20 0 0.00

Stress (MPa)

100

0.02

100 80 60 40

RAC50FST20 experiment RAC50FST20 simulation

0.01

Exposed to temperature of 200 C

120

AC

Stress (MPa)

120

80

140

CE

140

Exposed to temperature of 20 C

0.03

20

0.04

0 0.00

0.05

RAC50FST200 experiment RAC50FST200 simulation

0.01

0.02

0.03

Strain (

Strain (

50

0.04

0.05

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

160

160

120

Stress (MPa)

Stress (MPa)

120 100 80 60

0 0.00

100 80 60 40

40 20

o

Exposed to temperature of 700 C

140

RAC50FST500 experiment RAC50FST500 simulation

0.01

0.02

0.03

RAC50FST700 experiment RAC50FST700 simulation

20

0.04

0 0.00

0.05

0.01

160

140

Stress (MPa)

80 60 40 RAC100FST20 experiment RAC100FST20 simulation 0.01

0.02

0.03

0.04

160

PT E

140

0.02

0.03

0.02

0.03

0.04

0.05

o

Exposed to temperature of 700 C

120 100 80 60 40

RAC100FST500 experiment RAC100FST500 simulation

0.01

0.01

Strain (

Stress (MPa)

40

RAC100FST200 experiment RAC100FST200 simulation

160

CE

60

AC

Stress (MPa)

80

60

0 0.00

0.05

o

100

80

20

Exposed to temperature of 500 C

120

0 0.00

100

40

D

Strain (

20

o

NU

100

140

0.05

120

MA

Stress (MPa)

120

0 0.00

0.04

Exposed to temperature of 200 C

SC

160 o

Exposed to temperature of 20 C

RI

(b) RCA replacement ratio of 50%

20

0.03

Strain (

Strain (

140

0.02

PT

140

o

Exposed to temperature of 500 C

20

0.04

0 0.00

0.05

Strain (

RAC100FST700 experiment RAC100FST700 simulation

0.01

0.02

0.03

0.04

Strain (

(c) RCA replacement ratio of 100%

Fig. 11. Comparisons on numerical and experimental impact stress-strain curves

51

0.05

ACCEPTED MANUSCRIPT

140

140

120

120

100

100

80 60 40

0.02

0.03

0.04

0 0.00

0.05

Strain (

0.01

0.02

0.03

0.04

0.05

Strain (

(b) Exposed to temperature of 200oC

NU

(a) Exposed to temperature of 20oC 140

140

120 100 80

RAC100FST500-C30 RAC100FST500-C40 RAC100FST500-C50

20 0 0.00

0.01

0.02

PT E

40

D

60

0.03

0.04

Stress (MPa)

MA

120 100

80 60 40

RAC100FST700-C30 RAC100FST700-C40 RAC100FST700-C50

20 0 0.00

0.05

0.01

0.02

0.03

0.04

Strain (

CE

Strain (

(c) Exposed to temperature of 500 oC

(d) Exposed to temperature of 700 oC

Fig. 12. Impact stress-strain curves of RACFST with different RAC strengths

AC

Stress (MPa)

RI

0.01

RAC100FST200-C30 RAC100FST200-C40 RAC100FST200-C50

20

SC

0 0.00

60 40

RAC100FST20-C30 RAC100FST20-C40 RAC100FST20-C50

20

80

PT

Stress (MPa)

Stress (MPa)

Revised JMAD-D-17-04650R1

52

0.05

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

1.12

R2=0.95

0.95 R2=0.95

0.90

RAC100FST20 test RAC100FST20 fitting RAC100FST200 test RAC100FST200 fitting RAC100FST500 test RAC100FST500 fitting RAC100FST700 test RAC100FST700 fitting

2

R =0.83

1.04

PT

Kc

1.06

1.00

R =0.99

0.85

1.02 2

R =0.94

0.80

1.00 0.98

0.75 30 MPa

40 MPa

50 MPa

RI

1.08

2

R =0.87 2

c

1.10

1.05 RAC100FST20 test RAC100FST20 fitting RAC100FST200 test RAC100FST200 fitting RAC100FST500 test RAC100FST500 fitting RAC100FST700 test RAC100FST700 fitting

30 MPa

SC

RAC strength

40 MPa

50 MPa

RAC strength

(b) Impact deformation decrease coefficient

NU

(a) Impact strength enhancement factor

R2=0.95 R2=0.99

Fig. 13. Effect of RAC strength on impact mechanical property and impact deformation of RACFST

AC

CE

PT E

D

MA

exposed to high temperatures

53

ACCEPTED MANUSCRIPT

160

140

140

120

120

80 60 40

0.01

RAC100FST200-Q235 RAC100FST200-Q345 RAC100FST200-Q420

20

0.02

0.03

0.04

0 0.00

0.05

0.01

0.02

0.03

0.04

0.05

Strain (

NU

Strain (

(a) Exposed to temperature of 20 oC

(b) Exposed to temperature of 200 oC

160

140 120 100

D

80

40

PT E

60

0 0.00

0.01

0.02

0.03

0.04

120 100 80 60 40

RAC100FST500-Q235 RAC100FST500-Q345 RAC100FST500-Q420

20

140

Stress (MPa)

MA

160

RAC100FST700-Q235 RAC100FST700-Q345 RAC100FST700-Q420

20 0 0.00

0.05

0.01

0.02

0.03

0.04

Strain (

CE

Strain (

(c) Exposed to temperature of 500 oC

(d) Exposed to temperature of 700 oC

AC

Stress (MPa)

60

SC

0 0.00

80

40

RAC100FST20-Q235 RAC100FST20-Q345 RAC100FST20-Q420

20

100

RI

100

PT

160

Stress (MPa)

Stress (MPa)

Revised JMAD-D-17-04650R1

Fig. 14. Impact stress-strain curves of RACFST with different steel strengths

54

0.05

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

1.6

PT

0.8

1.2

2

R =0.95

RAC100FST20 test RAC100FST20 fitting RAC100FST200 test RAC100FST200 fitting RAC100FST500 test RAC100FST500 fitting RAC100FST700 test RAC100FST700 fitting

0.7

1.1

0.6

1.0

0.5

SC

Ks

1.3

0.9

2

R =0.97 2

R =0.93

RI

1.4

1.0

2

R =0.92

s

1.5

1.1 RAC100FST20 test RAC100FST20 fitting RAC100FST200 test RAC100FST200 fitting RAC100FST500 test RAC100FST500 fitting RAC100FST700 test RAC100FST700 fitting

0.4

0.9 235 MPa

345 MPa

235 MPa

420 MPa

NU

Steel strength

(a) Impact strength enhancement factor

2

R =0.96

2

R =0.95 2

2

R =0.92

345 MPa

R =0.92

420 MPa

Steel strength

(b) Impact deformation decrease coefficient

MA

Fig. 15. Effect of steel strength on impact mechanical property and impact deformation of RACFST

AC

CE

PT E

D

exposed to high temperature

55

ACCEPTED MANUSCRIPT

180

180

160

160

140

140

100 80 60 40

0.01

60

RAC100FST200-2mm RAC100FST200-3mm RAC100FST200-4mm

20

0.02

0.03

0.04

0 0.00

0.05

0.01

0.02

0.03

0.04

0.05

Strain (

NU

Strain (

(a) Exposed to temperature of 20 oC

(b) Exposed to temperature of 200 oC

180

160 140 120

D

100

Stress (MPa)

MA

180

60

PT E

80

20 0 0.00

0.01

0.02

0.03

0.04

160 140 120 100 80 60 40

RAC100FST500-2mm RAC100FST500-3mm RAC100FST500-4mm

40

RAC100FST700-2mm RAC100FST700-3mm RAC100FST700-4mm

20 0 0.00

0.05

Strain (

0.01

0.02

0.03

0.04

CE

Strain (

(c) Exposed to temperature of 500 oC

(d) Exposed to temperature of 700 oC

AC

Stress (MPa)

80

SC

0 0.00

100

40

RAC100FST20-2mm RAC100FST20-3mm RAC100FST20-4mm

20

120

PT

120

RI

Stress (MPa)

Stress (MPa)

Revised JMAD-D-17-04650R1

Fig. 16. Impact stress-strain curves of RACFST with different steel ratios

56

0.05

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

2.2

R =0.99

0.8

R2=0.99

0.7

R2=0.99

1.4

0.6 0.5

R2=0.99

1.2 1.0

0.3

0.8 3 mm

0.2

4 mm

2 mm

NU

Thickness of steel tube

RAC100FST20 test RAC100FST20 fitting RAC100FST200 test RAC100FST200 fitting RAC100FST500 test RAC100FST500 fitting RAC100FST700 test RAC100FST700 fitting

SC

0.4

2 mm

PT

K

1.6

0.9 2

RI

1.8

1.0



2.0

1.1 RAC100FST20 test RAC100FST20 fitting RAC100FST200 test RAC100FST200 fitting RAC100FST500 test RAC100FST500 fitting RAC100FST700 test RAC100FST700 fitting

2

R =0.96

3 mm

2

R =0.92 2

R =0.91

4 mm

Thickness of steel tube

(b)Impact deformation decrease coefficient

MA

(a)Impact strength enhancement factor

2

R =0.93

Fig. 17. Effect of steel ratio on impact mechanical property and impact deformation of RACFST

AC

CE

PT E

D

exposed to high temperature

57

ACCEPTED MANUSCRIPT

AC

Graphical abstract

CE

PT E

D

MA

NU

SC

RI

PT

Revised JMAD-D-17-04650R1

58

ACCEPTED MANUSCRIPT Revised JMAD-D-17-04650R1

Highlights: (1) Increasing RAC strength slightly enhances impact strength of RACFST, which declines with the

PT

increase of elevated temperatures.

RI

(2) Increasing steel strength enhances impact strength of RACFST, which is more obvious when

SC

suffered from higher temperature.

NU

(3) Increasing steel ratio is an effective to improve the impact strength and decreasing impact

MA

deformation of RACFST.

(4) Both the steel strength and steel ratio should be considered together for impact or blast design

AC

CE

PT E

D

for the RACFST structures.

59