high-strength square steel tube columns

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Transverse impact behavior of high-strength concrete filled normal- / high- strength square steel tube columns Xiaoqiang Yang , Hua Yang , Sumei Zhang PII: DOI: Reference:

S0734-743X(19)30749-3 https://doi.org/10.1016/j.ijimpeng.2020.103512 IE 103512

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International Journal of Impact Engineering

Received date: Revised date: Accepted date:

5 July 2019 16 January 2020 21 January 2020

Please cite this article as: Xiaoqiang Yang , Hua Yang , Sumei Zhang , Transverse impact behavior of high-strength concrete filled normal- / high- strength square steel tube columns, International Journal of Impact Engineering (2020), doi: https://doi.org/10.1016/j.ijimpeng.2020.103512

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Highlights    

Tests of 18 HSCFST members subjected to transverse impact loadings were conducted Impact force, deformation and energy absorption of HSCFST were investigated Influence of material strengths on the impact resistance was discussed in detail A FE model was developed, performing good prediction for impact responses of HSCFST

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Transverse impact behavior of high-strength concrete filled normal- / high- strength square steel tube columns Xiaoqiang Yanga,b, Hua Yanga,c,*, Sumei Zhangd a

Key Lab of Structures Dynamic Behavior and Control of the Ministry of Education,

Harbin Institute of Technology, Harbin, 150090, China b

School of Civil Engineering, Harbin Institute of Technology, Harbin, 150090, China

c

Key Lab of Smart Prevention and Mitigation of Civil Engineering Disasters of the

Ministry of Industry and Information Technology, Harbin Institute of Technology, Harbin, 150090, China d

School of Civil and Environment Engineering, Harbin Institute of Technology

(Shenzhen), Shenzhen, 518055, China

*Corresponding author. Postal address: Room 207, School of Civil Engineering, Harbin Institute of Technology, Huanghe Road #73, Nangang District, Harbin 150090, Heilongjiang Province, China; Tel: +86 451 86282079. E-mail addresses: [email protected] (Xiaoqiang Yang), [email protected] (Hua Yang*), [email protected] (Sumei Zhang).

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Abstract High strength materials and dynamic loadings are two of the hottest research topics for building structures. In this study, high-strength concrete filled square steel tube columns (HSCFST) using S690 along with Q355 structural steel subjected to transverse impact loadings were experimentally tested by a drop hammer tester, and the impact force, deformation, and energy absorption of such specimens were obtained. Results showed that HSCFST has great impact resistance, showing high impact force plateau value and small deflection. The influence of steel strength, concrete strength, steel ratio, and impact energy on impact resistance was deeply analyzed. Compared with normal-strength materials, using high strength steel can improve its impact resistance but the contribution is not as great as the increment of yield stress, whilst using the high-strength concrete has limited effect. Besides, a refined finite element (FE) method by ABAQUS/Explicit employing the rate-dependent constitutive model for S690 was adopted to simulate the dynamic responses of HSCFST under impact loadings, performing a good agreement with tests. This study provides the basic experimental data of HSCFST subjected to transverse impact and develops a reasonable FE model to predict its impact resistant performance, contributing to the related studies of HSCFST components.

Keywords: Concrete filled steel tube, impact resistance, high strength, drop hammer test, strain rate.

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1. Introduction Concrete filled steel tube (CFST) was suggested in the last century by researchers to improve the ductility and brittleness of concrete by confinement from outer steel tube, and meanwhile, the steel is also restricted by core concrete to improve its buckling behavior, which increases its bearing capacity [1-2]. Due to its excellent mechanical properties, CFST has been widely used as main structural member in modern grand structures and constructions, especially in China, e.g., super high-rise buildings (China Zun, 528m in height, 2018), TV towers (Canton Tower, 600m in height, 2009), transmission towers (Zhoushan Electricity Pylon, 380m in height, 2018), large-span arch bridges (Zhijing River Bridge, 430m in span, 2009), and bridge piers (Labajin Bridge, 183m in height, 2014). With the development of material and construction technology, the tendency to apply high-strength structural materials in civil engineering has also attracted the interests of engineers and researchers. Pouring the high-strength concrete (fc ≥ 60MPa) into high-strength steel (fy ≥ 460MPa) tubes can mitigate their negative performances of low ductility and high brittleness, and take advantages of the high-strength behavior for reduction of the cross-sectional area of structural members, material usages, and environmental pollutions, bringing benefits in both economic and environmental aspects. Therefore, high-strength concrete filled high-strength steel tubes (HSCFST) also have been used in practice recently – Abeno Harukas (Japan; fy = 590MPa, fc = 150MPa) and Obayashi Technical Research Institute (Japan; fy = 780MPa, fc = 160MPa) [3]. 4

The behavior of HSCFST members has been investigated by various researchers in the past decades. From 2002, Varma et al. [4, 5] conducted the studies on the seismic behavior of square HSCFST columns. From then on, researchers from Singapore, Australia, China, Japan, and other countries began to conduct experimental studies and finite element analyses for HSCFST members regarding axial compressive [6-10], eccentric compressive [11-13], bending resistant [14], torsion resistant [15], fire resistant [3, 16, 17] and seismic [18, 19] behaviors. In 2015, Liew [20] summarized a series of studies on HSCFST members and provided a design guide based on Eurocode 4 for HSCFST. Those experimental and theoretical studies of HSCFST members in terms of static and seismic behaviors have been widely investigated, and these research results showed that HSCFST has great performance for static and seismic conditions.

As known, building structures and infrastructures may suffer from accidental impact actions over the serving duration, such as vehicle impact, terrorist attack, gas explosion, and progressive collapse. Over the past decades, enormous loss of life and economic damage were reported because of those impact loadings on civil structures. Therefore, it is a priority to clarify the impact resistance of structural components, especially for HSCFST members mainly served as major structural members in those grand constructions. The impact resistance of normal-strength CSFT columns has been attracting extensive attention by researchers. In 2000, Prichard and Perry [21] began to conduct exploratory experiments for the study on concrete filled in different sleeve columns 5

under free-falling impact to investigate the dynamic responses. After the ―9.11‖ attack, scholars paid more attention to the impact of industrial and civil building structures. Researchers investigated the behaviors of normal CFST members under axial [22-26] and transverse impact loads [27-34] at normal temperature and high temperature, studying the effect of different strength of materials, steel ratio, boundary conditions, impact positions, axial compression ratio, size effect, etc. Meanwhile, they have put forward analytical models and simplified calculating methods of dynamic plastic bending moment, energy absorption, dynamic increase factor, etc. Recently, the impact resistance of new-type CFST members has been studied by scholars, such as cement composite filled steel tubular columns [35, 36], recycled aggregate concrete filled square steel tubular members [37], rigid polyurethane foam-filled steel tube [38, 39], concrete-filled stainless steel tubular columns [39-42], FRP strengthened CFST members [43, 44], concrete-filled double steel tubular columns [42, 45-48], and concrete-encased CFST members [49,50]. However, these researches have not involved the study on the impact resistance of HSCFST members. It should be noticed that steel and concrete are rate-dependent materials. As the strain rate increased under fast load, the strength of steel and concrete would increase. In the past decades, researchers have investigated the rate-dependent behavior of conventional strength structural steels (yield stress less than 460 MPa) at intermediate and high strain rates [51-55]. For high-strength structural steels (HSSS), the studies regarding its dynamical strain-rate behavior have been conducted [56-58], but it is still

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insufficient compared with those researches for its static performances [59-62]. Previously, the authors had investigated the dynamic behavior of S690 high-strength steel [58], showing less strain-rate sensitivity when compared with normal-strength steel. Besides, the plastic properties also have remarkable difference between the high-strength steel and normal steel, i.e., lower strain-hardening behavior and lower ductility. These different material properties may lead to different behavior at the member level, especially for those components subjected to large plastic strain and high strain rate when against impact loadings. Hence, the dynamic behavior of CFST members using high-strength structural materials need to be systematically studied, but not directly derived from the mechanical properties of those normal-strength CFST members. In this study, experimental tests of 18 square C60/C100 high-strength concrete filled S690 high-strength steel along with Q355 steel tube columns subjected to transverse impact loading were conducted by a drop hammer tester. The impact-resistant indicators - impact force and mid-span deformation, were experimentally obtained. The influences on the impact force, mid-span deflection, and energy absorption ability of such HSCFST members were investigated, and the effect of material strengths on the impact resistance was discussed in detail. Based on the test data, an FEA model considering dynamic rate-dependent constitutive models of high-strength materials was developed for predicting the dynamical responses of such components against transverse impact loadings.

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2. Experimental program 2.1 Specimen design A total of 18 square HSCFST specimens were tested in this paper at Harbin Institute of Technology (HIT), being classified as two series: (I) 6 CFST members using high-strength steel S690 and (II) 12 CFST members using normal-strength steel Q355. Four parameters were mainly considered in this study to evaluate the impact resistance of specimens: i.e., grade of steel, grade of concrete, tube thickness (steel ratio), and impact energy. The specimens have the same cross-section dimension (180 ⅹ180mm). The length of the clear span (L0) was 1500mm between the two simply supporting rollers. For Series I, all members were welded by two U-shape steel components fabricated using a 6mm commercial S690QL steel plate. Two different grades of high-strength concrete - C60 and C100, and three different impact energies varied by different impact masses and heights - 424kg with 3m, 424kg with 7m, and 544kg with 7m were considered. For Series II, the impact mass was 424kg with different impact heights of 3m, 5m, and 7m, respectively. Three nominal thicknesses - 4mm, 5mm, and 6mm of Q355 steel pipes were adopted, which is used to evaluate the influence of tube thickness (steel ratio). The detailed description for each testing specimen is listed in Table 1. To identify each specimen, the following labeling system was adopted: (1) the characters ―HS‖

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and ―NS‖ stand for high-strength steel (S690) and normal-strength steel (Q355); (2) the numbers ―3‖, ―5‖ and ―7‖ refer to 3m, 5m and 7m with impact mass 424kg, while ―7*‖ refer to 7m with impact mass 544kg; (3) the numbers ―100‖ and ―60‖ stand for the grade of concrete; and (4) the last numbers ―4‖, ―5‖ and ―6‖ refer to the nominal tube thickness (in unit: mm), while the actual steel pipe thickness ts is slightly varied. The schematic diagram of the numbering system is shown in Fig. 1.

2.2 Materials 2.2.1 Concrete Two grades of high-strength concrete: C60 and C100 were adopted in the study, which were produced in the laboratory at Harbin Institute of Technology. The mix properties of concrete are shown in Table 2. The concrete cubes (150ⅹ150ⅹ150mm) and prisms (150ⅹ150ⅹ300mm) being used to evaluate the compressive strength and elastic modulus, respectively, were cast at the same time as the testing specimens were fabricated, and then cured in the same laboratory condition. To simulate the curing condition of the concrete which was filled in the steel tube, the cube and prism specimens were carefully sealed with two layers of aluminum foil, two layers of plastic wraps, and two layers of scotch tape immediately after demolding (24 hours after casting). The material tests were conducted at 28 days and at the period of impact tests for member specimens (315 days later), of which mean values of compressive strength at 28 day (fcu,m_28), compressive strength during tests (fcu,m_test) and Young’s modulus (Ec) for material test results are shown in Table 2. 9

2.2.2 Steel The grade of normal-strength steel Q355 with three different nominal thicknesses (4mm, 5mm, and 6mm) and one grade of high-strength steel S690QL were used in this paper. Three tensile coupons for each type of steels were machined from the same batch of material with column members to conduct the standard tension tests by a universal electromechanical testing machine (Shimadzu AG-X Plus 250kN) at the speed of 0.9mm/min. Two pairs of strain gauges stuck in the middle of each coupon and an extensometer, together with the force sensor, were utilized to determine the tension properties of steels, as shown in Table 3. The tensile behavior of high-strength steel S690 also meets the demands of EN 1993-1-12.

2.3 Equipment and procedure To examine the transverse impact behaviors of CFST column specimens, a series of experimental impact tests were performed at HIT using a drop hammer equipment (see Fig. 2). The drop hammer setup consists of two 20.6m guide rails with supporting frames (maximum drop height is 19-20m which depends on the test conditions), a drop hammer with different indenter and weight blocks (from 424kg to 1090kg), a rigid support platform to apply boundary constraints, an electromagnet release mechanism and a data recording system [63]. Two supporting setups were designed to simply support the specimen at both ends, and two steel plates with a long screw were mounted for each supporting setup to ensure safety if the specimen rebounded and rotated, as shown in Fig. 2. 10

In this study, the mass of drop hammer was 424kg / 544kg with a wedge-shaped indenter, and the impact energy varied with different impact heights (3m, 5m, and 7m). For each test, the drop hammer was lifted up to the specified height, and it fell down alongside the guide rails and impacted the specimen as a free-falling body once the electromagnet release mechanism was manually triggered. The impact force would be recorded by a load cell mounted between the indenter and drop hammer frame, and the displacement of the specimen at the middle span was recorded by the high-speed linear variable differential transformer (LVDT) with a 100,000 Hz high-speed data acquisition system. In addition, a high-speed video camera with 2000 frames per second was adopted to monitor the whole test process.

3. Results & Discussion 3.1 Experimental results Eighteen square HSCFST members were systematically tested by the drop hammer tester, and key parameters showing dynamic responses for such specimens under transverse impact loadings were obtained. Table 4 shows the summary of information and key parameters of test results for those specimens under transverse impact loading. The detailed discussion would be presented in the following sections. 3.1.1 Failure mode All the specimens were simply supported in the tests, so that the main failure location was in the middle area that is the impact point. Fig. 3 shows the typical

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failure mode of a square HSCFST member subjected to transverse impact loading. Note that obvious overall flexural deformation was observed, and a plastic hinge appeared in the middle span. In the top area of the mid-span section, where the steel tube directly against impact indenter, the local buckling (see Fig. 3(a)) was observed, where the top surface of the steel tube was bulging on both sides of impact position. For higher impact height, the side surfaces of the square steel tube were also bulging due to the high impact energy. When the outer steel tube was cut and removed after the experiments, the typical flexural failure mode of core concrete could be seen clearly, shown in Fig. 3(b). There are numerous cracks in the tensile zone opposite the impact position and the obvious concrete crushing in the compressive zone of the mid-span section. A few cracks could be noticed at the locations of both pinned supports. Except for the positions of the impact area and support points, there is no obvious damage phenomenon for the test specimens. Fig. 4 shows the overall residual bending shape comparisons between the typical HSCFST specimens with different impact energy, steel ratio and strength, and concrete strength. As the impact energy increased, the phenomenon of overall bending would be further exacerbated, and the maximum local indentation δmax at the impact point and the mid-span residual deformation ur increased from 0.6mm and 9.0mm to 2.5mm and 38.8mm, respectively, as shown in Fig. 4(a). Those specimens with thinner tube thickness exhibited more damage, i.e., greater local and overall

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deformations (Fig. 4(b)), owing to the less dynamic bending resistance. Compared with the specimen using Q355, as seen in Fig. 4(c), the ur of S690 HSCFST member under the same impact energy was about 20% smaller, while the δmax was almost the same. However, the influence of concrete strength on flexural failure mode was not very obvious. The reasons for the different influences on those related performance indicators would be discussed in detail in sections 3.1.2 and 3.1.3. 3.1.2 Impact force As one of the key parameters for impact analysis, the impact force, which is the contact force between impact indenter and specimen, would be first discussed. The experimental impact force time histories for each specimen obtained from the tests were plotted in Fig. 5. The experimentally measured curves were somewhat oscillating on account of the use of the strain-type load cell, but basically could be divided into three stages, i.e., vibration stage, stable stage, and descending stage. In order to clearly show these three stages of the experimental curve and the associated physical quantities, Fig. 6 shows the typical impact force and mid-span displacement time histories for HSCFST members against impact loadings. Once the drop hammer indenter reached the top surface of the specimen, a localized depression at the area of impact position was developed, and at the same time, the impact force increased sharply to the peak value Fm and then decreased soon with the following vibration stage. At the same time, the specimen accelerated by the impact loading from zero velocity and deformed gradually. During the impact force decreased from peak value, the drop hammer

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continued to decelerate due to the resisting force of the specimen, and then the velocity of the specimen is larger than that of drop hammer. Hence, they temporarily separated and the impact force dropped to zero until they touched again. When the velocity of specimen deformation is almost the same as that of impact indenter, the specimen and drop hammer moved together, and the impact force kept relatively stable for a period of time, which was defined as a ―stable stage.‖ Finally, the impact energy was mostly dissipated and the drop hammer began to rebound, and at the same time, the impact force started to descend gradually. When it decreased to zero, the drop hammer and specimen separated completely. A completed impact process was finished. Note that due to the gravity effect, the drop hammer would continue to impact the specimen and rebound several times until the kinetic energy was completely dissipated. Compared to the original impact energy, the remaining kinetic energy of drop hammer was small after the first impact, so that this study only investigated the first impact process. It should be noticed that if the impact energy was so slight that the specimen did not undergo a certain plastic strain, the impact force time history will only have vibration stage, while if the impact energy was large enough to destroy the specimen directly, there is no smooth descending stage due to the unloading process, but instead a sudden drop during the plateau stage. Therefore, the impact energies set in these tests were aimed at obtaining large plastic strain for the specimen to evaluate its dynamical response, but not destroying the specimens directly. As reported above, at the start, the impact force would reach peak value sharply.

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The value of peak impact force Fm influenced by initial stiffness of specimen, contact surface between specimen and indenter, the frequency of data sampling and filtering, etc. was somewhat uncertain. During the stable stage, however, the impact force remains basically constant. The fact is, therefore, that the plateau value of impact force is a suitable performance index to be used to estimate the impact response of structures subjected to impact loadings when undergoing overall deformation. In the actual test, however, the clear starting point t1 of the stable stage for impact force time history was difficult to be defined. A method for the post-peak mean force suggested by Wang et al. [35] was adopted as the plateau value Fp to depict the value of impact force during the stable stage, shown as follows.

 F=

t2

t0

p

F  t  dt

t 2  t0

(1)

Where t0 and t2 refer to the time moment of reaching peak impact force and reaching maximum displacement, respectively, as shown in Fig. 6. F(t) denotes the impact force time history. As shown in Fig. 5, the Fp calculated in this paper was relatively close to impact force when ignoring the slight oscillation of impact force versus time curve in the stable stage, which means that the method of obtaining Fp was suitable for the tests. All plateau values which were obtained by post-peak mean forces for all impact experimental tests of HSCFST members in this paper are listed in Table 4. It can be seen clearly in Fig. 7 that the plateau values were almost the same with the increasing impact energy Ei, which confirmed that this key parameter - plateau 15

value of impact force Fp was associated with inherent properties of the specimen but not impact energy for the impact conditions in this study. As the impact became more serious, a very slight growth of plateau value can be noticed owing to the strain-rate effect of materials. However, as shown in Fig. 5, the time duration of the impact process was obviously further extended, on account of the increasing energy. For those members with C60 concrete, the impact force was marginally lower (less than 8%) than that with C100 high-strength concrete in Fig. 7(a) and (b). Meanwhile, it might be inferred that the strength of concrete had more influence on impact resistance for those specimens with lower steel strength and lower impact energy, owing to the more contribution to the dynamic bending resistance from concrete for such members. However, it still can be concluded the strength of concrete had limited influence on the impact resistance of HSCFST. Note that the thickness of steel tubes (steel ratio) had a significant influence on the Fp for CFST members under transverse impact action, as shown in Fig. 7(c). When the steel ratio increased from 9.5% to 15.9%, the average Fp grew by 71% from 248kN to 425kN. In Fig. 7(d), it was noticed that the average Fp climbed when adopting S690 to replace Q355. The conclusion could be considered - the strength of the steel pipe and steel ratio of CFST members can improve the impact resistance of CFST subjected to impact loading, and the HSCFST structural components can bear impact loading well. However, it is a fact that although the yield stress of S690 steel is 722MPa which is almost 1.8 times compared to Q355 with 406MPa yield strength, the average Fp of

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HSCFST using S690 is only 1.24 times of that for Q355 members. It is a reason that compared with normal mild steel, the strain-rate effect of high-strength steel is less obvious, referring to the previous study on rate-dependence for S690 by Yang et al. [48], which may cause more than 10% reduction of Fp for S690 specimens. Another reason is that the strain-hardening of high-strength structural steel is much less than that of normal-strength steel, with ultimate stress 758MPa of S690 which is only 1.39 times that of Q355 with 546MPa.That means as the plastic strain develops, the gap of strength will be gradually narrowed. The impact of these two influencing factors will further increase as the impact energy increases, on account of the larger plastic strain and higher strain rate. Besides, the fact that steel ratio of S690 members is lower than the specimen using Q355 by 1.1% will also contribute to the reduction of Fp for S690 members to some extent, which is an initial defect of the tests, but it is not the main factor led to the above situation. Due to the lower plastic hardening behavior and dynamic strain-rate properties, the dynamical bending resistant capacity of high strength steel members would not increase as obviously as ordinary steel, especially for such members against impact loadings with large plastic strain. Therefore, the analyses for HSCFST members against impact loading should be based on the rate-dependent model for high-strength steel, on account of the fact that using the dynamic constitutive model of normal-strength steel satisfying would overestimate its impact resistance. 3.1.3 Deformation Due to the heavy indenter directly impact to the top surface of specimens, the

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local indentation was observed after the test. Unlike the hollow steel tube without the support of core concrete, the maximum local indentations of CFST members in this paper were relatively slight, which were less than 3mm listed in Table 4. The cross-section of HSCFST members where directly facing the impact loads still remained intact, so that it had sufficient bending resistance to dissipate the impact energy by global deformation. Another important deformation parameter, mid-span displacement, was experimentally obtained by the LVDT with a high-speed data acquisition system for each specimen. As seen in Fig. 6, although the impact force time history severely oscillated at the beginning of the test, the displacement increased almost linearly. As the velocity of motion decreased, the displacement tended to be stable reaching the peak point. After that, due to the impact energy was mostly absorbed and dissipated, the drop hammer and specimen were rebounded and the elastic deformation was gradually recovering. The mid-span displacement time histories of all tests could be seen in Fig. 8, and the maximum mid-span displacement um of each tested member was much larger than the maximum local indentation at the impact position, as listed in Table 4. Unlike the plateau value of impact force, the um was basically linearly related to the impact energy, as shown in Fig. 7. This is also indicated that the impact force had limited dependence of impact energy, because the integral of force and displacement reflects the energy absorption of energy to some extent. The phenomenon that the test specimens with C60 and C100 high-strength concrete

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respectively had almost the same peak displacement also confirmed the limited effect of grades of concrete on transverse impact-resistant performance. The increasing steel ratio would improve the bending stiffness and bending resistance, so that under the same impact energy, the deflection became smaller as the thickness of the steel tube continued to increase, as seen from Fig. 7(c). Likewise, the higher strength of steel pipe also could enhance the dynamic bending resistance (i.e., plastic moment after the plastic hinge formed) of such specimens against impact. The elastic deformation is greater for higher strength steel, leading to less total deflection, and under the same impact energy, the rotation will be reduced since the plastic hinge moment increases, leading to the reduction of peak displacement um. As seen in Fig. 7(d), the peak deformation of test specimens under impact energy of 12.5kJ and 29.1kJ decreased by about 13% when using S690 to substitute normal-strength steel Q355. Therefore, using high strength steel would effectively reduce the deformation, and increase the impact force as discussed before, indicating that HSCFST members had good impact resistance. It is the fact that, however, the yield strength of steel did not completely determine the impact resistance (dynamical bending resistance) of HSCFST members subjected to impact transverse loading undergoing large plastic strain, and the strain-hardening behavior, ultimate stress, and strain-rate effect took an important role in it, as discussed in 3.1.2. The improvement of impact behaviors - impact force and deformation for HSCFST members using S690 is not as good as the increase in yield

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stress of used steel, owing to the lower strain-strengthening behavior and relatively less obvious strain-rate effect of S690 high strength steel. 3.1.4 Energy absorption The total energy of impact process for each specimen shown in Table 4 was calculated by the gravitational potential energy of the drop hammer (Ei=mgH), neglecting the energy loss during the falling of the drop hammer. The total energy would be dissipated by three sections during the first impact process, i.e., local indentation, overall deformation, and remaining kinetic energy. However, the aforementioned local indentation was insignificant and the remaining kinetic energy was also slight, so that the global deformation may absorb most of the impact energy. To quantify the energy absorptions of overall deformation, for the HSCFST specimen under transverse impact loading, the typical impact force - mid-span displacement curve could be seen in Fig. 9. The filled area in Fig. 9 enclosed by the force-displacement curve represented the energy absorbed by global bending deformation. The energy absorptions of global bending deformation (Eg) for all the tests were calculated and summarized in Table 4, and the corresponding ratios to total impact energy (Eg/Ei) were also calculated and shown in Table 4. It can be seen that the energy absorption ratio (Eg/Ei) for overall deformation was about 0.74 on average, confirming that global bending deformation consumed most of the impact energy, while the local indentation did not play a major role in energy dissipation. The fact

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that the localized depression of steel pipe was small so as to not significantly weaken the bending capacity of the cross-section is more conducive to the bearing capacity and energy consumption of overall deformation for the structural component. It can be noticed that, as the total impact energy increased, the energy absorption ratio was slightly increased, and it could be considered basically stable. There is also no obvious correlation between the energy absorption ratio and the inherent properties of the specimen in this paper - strength of steel, strength of concrete and steel ratio. To investigate this aspect, more studies are needed in the future.

3.2 Finite element analysis 3.2.1 FEA model To further understand the dynamical responses of HSCFST subjected to transverse impact loadings, an FEA study on such specimens was conducted using ABAQUS/Explicit. Fig. 10 shows a general view of the representative FE model, in which a refined model of HSCFST specimen under transverse impact loading was created, and the details of the FE model regarding element type and size, contact interaction, loads and boundary conditions would be briefly presented as follows. The square steel tube and core concrete were all modeled as C3D8R solid element that is an 8-node linear brick, reduced integration, and hourglass control element, while the drop hammer indenter and support setups were modeled using 4-node 3-D bilinear rigid quadrilateral element (R3D4). The mesh size was set as 20mm, which is 1/9 of steel pipe width in this study.

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The specimen was simply supported by two rigid rolls that were fully fixed. The wedge-shaped indenter was also created as a rigid body with a reference point, and the reference point was constrained except for vertical direction. The mass and initial velocity of drop hammer were assigned for the reference point of drop hammer, where the initial velocity was defined as a predefined field by ABAQUS, calculated by

v  2 gH (H refers to impact height). The general contact model was employed for the interaction of FE model, where pressure-overclosure was defined as ―hard contact‖ in the normal direction for the whole model, while for tangential direction, ―frictionless‖ was employed to depict the contact behavior for FE model except for the surface pairs between core concrete and outer square steel tube. For the contact surface between concrete and steel tube, the ―penalty‖ model for friction formulation with friction coefficient 0.6 was adopted [64]. 3.2.2 Material properties 3.2.2.1 Steel For normal-strength steel, the five-stage stress-strain relationship suggested by Han et al. [64] was employed as the static material model, where the yield stress, Young’s modulus and Poisson’s ratio of steel were set according to the tension test results for Q355 in Table 3. To further consider the strain-rate effect, the Cowper-Symonds (C-S) model was adopted to depict the dynamic increase factor (DIF) as the increasing strain rate when against impact loadings. The C-S model was

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expressed as follows by Eq. (2). 1

d   p  1   s D

(2)

Where  s and  d refer to the quasi-static stress and the dynamic stress with the strain rate of  , respectively. D and p are material constants. Referring to [65], the values of coefficients D and p could be taken as 6844 and 3.91 respectively for normal-strength mild steel. As for S690 high-strength structural steel, it showed huge difference in tensile properties compared with normal strength steel, i.e., higher yield stress, shorter length of yield plateau, higher yield to tensile stress ratio, and lower elongation. Therefore, the five-stage stress-strain model was no longer valid for S690. Owing to the lower strain-hardening behavior and strain rate effect of S690, the previous models based on normal-strength steel would overrate the flow stress, especially for conditions at large plastic strain and high strain rate. The proposed Johnson-Cook (J-C) model and C-S model of S690 at [58] can perform acceptable predictions for S690 high-strength steel. To obtain more accurate FE results for HSCFST members using S690 subjected to impact loadings, a rate-dependent dynamic constitutive model for S690 [58] was employed in this paper. Hence, the static stress-strain relationship in the FE model was set as  =722+400 p 0.57 , and the C-S model with coefficients: D=18404 and p=2.91 fitted for S690 was also used to consider the strain-rate increase for S690. 3.2.2.2 Concrete

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The Concrete Damaged Plasticity (CDP) model embedded in ABAQUS/Explicit was adopted to model the compressive and tensile properties for both C60 and C100 high-strength concrete. Referring to [66], the dilation angle, eccentricity, fb0/fc0, K, and viscosity parameter of plasticity coefficients for the CDP model were set as 30°, 0.1, 1.16, 0.667, and 0.0005 in this analysis, respectively. The core concrete stress-strain model of square CFST members for analysis in ABAQUS reported by Han et al. [64] was employed in this paper, considering the steel confinement, which had been proved that it could predict well for the FE analyses of square CFST members. In addition, the strength of concrete also has obvious sensitivity to strain rate so that a rate-dependent model, depicting the dynamic increase behavior of core concrete as the strain rate increases, was required. It can be noticed that the rate-dependent models of concrete for considering compressive and tensile DIF was reported by CEB-FIP model code 1990 [67], while the latest version of CEB-FIP model code 2010 [68] ignored the influence of quasi-static concrete strength which was considered in the previous equation. Actually, as the strength of concrete increases, the sensitivity to the strain rate for high-strength concrete decreases [69]. Hence, the previous model suggested by [67] was employed to calculate the DIF of concrete in this analysis, as shown in Eq. (3) and Eq. (4). For

compression:

1.026s  ,  f cd / f cs    /  co   1/3   f cd / f cs   s   /  co  ,

(3)

24

  30s1   30s1

For tension:

1.016 s  ,  f td / f ts    /  to   1/3   f td / f ts  s   /  to  ,

  30s1   30s1

(4)

Where fcd and ftd refer to the dynamic compressive and tensile strength, respectively; fcs and fts are the quasi-static compressive and tensile strength;  is strain rate,

 co =-30 10-6 s1 and  to =3 10-6 s1 ; s =1/ (5+9 fc / fco ) ,  s =1/ 10  6 f c / f co  , fco =10MPa ; log  s =6.156s -2 ; log s =7.112 s -2.33 . Detailed information could be found in [67]. Note that these models are suitable for the compressive strength from 12MPa to 80MPa for the concrete cylinder. For the C100 high-strength concrete in this study with characteristic cylinder compressive strength of about 87MPa, however, these models are still employed to consider the DIF on account of the absence of experimental data for C100 concrete at high strain rates. 3.2.3 FEA results By the refined FE model established in this paper, the FE predicted curves of impact force and displacement time histories were obtained. The comparisons of impact force time histories between experimental results and FEA predicted results for each specimen were plotted in Fig. 5, showing a good agreement with each other. The plateau values of impact force derived from FEA results are also listed in Table 4, and the corresponding ratio to test value was 1.071 on average with 0.068 for standard deviation. Besides, Fig. 8 shows the mid-span displacement versus time curves derived from both the experimental test and the FEA method. Compared with the test data, the FE results showed satisfying consistency, for both peak value and overall trend. For the comparison of peak displacement, the mean value was 0.992, with a 25

standard deviation of 0.044 for the ratio of FE results to experimental results, which could be observed in Table 4. Hence, the developed FE model could predict the transverse impact performance of HSCFST members with reasonable prediction accuracy. To further verify the validity and extended applicability of developed FE model, previous studies of square CFST subjected to transverse impact loading by drop hammer tester conducted by other researchers [37, 39, 41] were employed, of which different types of drop hammer, different boundaries, and different axial load ratios were considered. The developments of the FE model and material model are the same as Section 3.2.1 and 3.2.2. In particular, in order to apply the axial load to the specified axial compression load ratio and then keep the axial load stable during the entire impact process, a large length and low stiffness spring was modeled to achieve the purpose of applying axial load. Firstly, a specified displacement was set up using ABAQUS/Explicit smooth step method to compress the spring to get the specified axial load. Subsequently, keeping the spring in compression, the drop hammer was modeled to impact the specimen until the impact process ended. The axially compressive force versus time curve was plotted in Fig. 11, showing that it was relatively stable during the impact process. Finally, the impact force and mid-span displacement time histories of square CFST members with specified axial load ratio during the entire impact process were obtained. Fig. 12 shows the comparisons of impact force and global defection time-histories between the tests from previous

26

studies and the predicted FE model. The developed FE model gave relatively good estimations on the main impact performance indexes - impact force and mid-span deformation. In general, the developed FE model may reasonably simulate the CFST members subjected to transverse impact loadings, considering different drop hammers, boundaries, impact energy, steel ratio, strength of materials, and axial compressive load ratio. Particularly, the impact FE model established in this paper employed the dynamic constitutive models of high-strength structural steel and obtained a good prediction for those HSCFST using S690 high-strength structural steel. The verified model in this study could be used to investigate the transverse impact behaviors of square HSCFST members.

Conclusions This paper elaborated the experimental tests on 18 square HSCFST specimens subjected to transverse impact loading, to evaluate the influences of material strengths, steel ratio, and impact energy on impact resistance. Different important performance indexes - impact force, deformation, and energy absorption were deeply investigated. Moreover, a refined FE model that employed suitable rate-dependent constitutive models for high-strength materials was developed, performing excellent prediction accuracy with test results. Based on the test results and finite element analysis, several conclusions were obtained as follows.

27

HSCFST members have excellent transverse impact resistance. Under impact loadings, the overall failure mode of simply supported HSCFST is flexural bending with slight local indentation at the impact position. With the increasing impact energy, the plateau value of impact force is almost constant, while the deformation of HSCFST would increase, which is almost linearly related to the impact energy. The steel ratio has the most effective influence on impact force and global defection, because it improves the bending resistance by increasing the steel ratio. The strength of steel can affect the impact performance, but the improvement of impact force plateau value (or the reduction of defection) is not as good as the yield stress of steel, owing to the lower strain strengthening behavior and less obvious strain-rate effect of high-strength steel. The grade of concrete has a limited influence on the impact force and overall deformation. Although the energy absorption ratio was slightly increased as the total impact energy increased, it could be considered basically stable, and there is no obvious correlation between the energy absorption ratio and the inherent properties of the specimen. The mean energy absorption ratio is 0.74 for HSCFST members in this paper.

An FE model employing rate-dependent models for S690 was developed to predict the dynamic responses of HSCFST against transverse impact loadings. Compared with the test results, it shows that the FE model has good agreement with tests. Moreover, the FE model was further verified by other test results with a 28

different drop hammer, boundaries, and axial load ratios, performing reasonable prediction accuracy.

Acknowledgments This work was supported by the National Natural Science Foundation of China (Grant No.51678194), and the National Key Research and Development Program of China (Grant No.2018YFC0705700).

Author Statement Xiaoqiang Yang: Conceptualization, Investigation, Methodology, Data curation, Formal analysis, Validation, Writing- Original draft preparation, Writing- Reviewing and Editing. Hua Yang: Conceptualization, Resources, Visualization, Writing- Reviewing and Editing, Project administration, Funding acquisition. Sumei Zhang: Supervision, Resources, Writing- Reviewing and Editing.

29

Declaration of Interest statement

We wish to confirm that there are no known conflicts of interest associated with this publication, and there has been no significant financial support for this work that could have influenced its outcome. We confirm that the manuscript has been read and approved by all named authors and that there are no other persons who satisfied the criteria for authorship but are not listed. We further confirm that the order of authors listed in the manuscript has been approved by all of us. We confirm that we have given due consideration to the protection of intellectual property associated with this work and there are no impediments to publication, including the timing of publication, with respect to intellectual property. In so doing we confirm that we have followed the regulations of our institutions concerning intellectual property. We understand that the Corresponding Author is the sole contact for the Editorial process (including Editorial Manager and direct communications with the office). He/she is responsible for communicating with the other authors about progress, submissions of revisions, and final approval of proofs. We confirm that we have provided a current, correct email address which is accessible by the Corresponding Author and which has been configured to accept email from [email protected]

30

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36

Tables Table 1 Summary of testing specimens. Series I

II

Steel Concrete L0

No.

B

ts

Steel

Impact

Impact

Impact energy

grade

grade

/mm /mm /mm ratio(%) mass m /kg height H /m

HS3-100-6

S690

C100

6.0

14.8

424

3

12.5

HS7-100-6

S690

C100

6.0

14.8

424

7

29.1

HS7*-100-6 S690

C100

6.0

14.8

544

7

37.3

HS3-60-6

S690

C60

6.0

14.8

424

3

12.5

HS7-60-6

S690

C60

6.0

14.8

424

7

29.1

HS7*-60-6

S690

C60

6.0

14.8

544

7

37.3

NS3-100-4

Q355

C100

3.9

9.5

3

12.5

NS5-100-4

Q355

C100

3.9

9.5

5

20.8

NS7-100-4

Q355

C100

3.9

9.5

7

29.1

NS3-100-5

Q355

C100

5.0

12.1

3

12.5

NS5-100-5

Q355

C100

5.0

12.1

5

20.8

NS7-100-5

Q355

C100

5.0

12.1

7

29.1

NS3-100-6

Q355

C100

12.5

Q355

C100

5

20.8

NS7-100-6

Q355

C100

7

29.1

NS3-60-6

Q355

C60

3

12.5

NS5-60-6

Q355

C60

5

20.8

NS7-60-6

Q355

C60

15.9 15.9 15.9 15.9 15.9 15.9

3

NS5-100-6

6.4 6.4 6.4 6.4 6.4 6.4

7

29.1

1500 180

1500 180

424

Ei /kJ

Table 2 Mix proportions and basic material properties of concrete. Grade

Water 3

Cement

Fine

3

3

Coarse

Silica Plasticizer

3

3

3

fcu,m_28 fcu,m_test 2

2

Ec

kg/m

kg/m

kg/m

kg/m

kg/m

kg/m

N/mm

N/mm

N/mm2

C100

121

500

623

1156

50

10

102

110

50,800

C60

180

450

645

1156

-

1.5

63

80

42,400

Table 3 Tensile properties of steels. Yield stress

Tensile stress Young’s modulus Es Poisson's Elongation fu N/mm2 (× 105 N/mm2) ratio 758 1.96 0.30 15%

Grade

ts/mm

S690

6.0

722

Q355

3.9

402

545

2.03

0.28

27%

Q355

5.0

416

509

2.01

0.27

26%

Q355

6.4

406

546

2.05

0.28

26%

2

fy N/mm

37

Table 4 Summary of test results and FE simulation results. δmax

Fp,test

Fp,FE

/mm

/kN

/kN

0.684

0.6

520

501

19.5

0.669

1.6

525

37.3

27.8

0.745

2.5

HS3-60-6

12.5

8.2

0.661

HS7-60-6

29.1

20.2

HS7*-60-6

37.3

NS3-100-4

Series I

II

No.

Ei/kJ

Eg/kJ

Eg/Ei

HS3-100-6

12.5

8.5

HS7-100-6

29.1

HS7*-100-6

Fp,FE/Fp,test

um,test um,FE

um,FE/um,test

/mm

/mm

0.963

22.2

21.7

0.977

563

1.072

43.0

40.9

0.951

535

577

1.079

54.4

51.4

0.945

0.7

480

500

1.042

22.7

22.3

0.982

0.694

1.7

520

557

1.071

43.9

42.5

0.968

29.6

0.793

2.9

524

573

1.094

55.0

53.6

0.975

12.5

9.5

0.760

0.8

241

278

1.154

38.5

37.4

0.971

NS5-100-4

20.8

16.2

0.780

1.4

248

287

1.157

62.7

59.1

0.943

NS7-100-4

29.1

22.5

0.772

2.2

255

299

1.173

84.6

82.0

0.969

NS3-100-5

12.5

8.6

0.689

0.7

310

334

1.077

32.0

31.5

0.984

NS5-100-5

20.8

16.1

0.777

1.3

299

352

1.177

51.7

52.4

1.014

NS7-100-5

29.1

23.0

0.790

2.0

315

361

1.146

69.4

66.0

0.951

NS3-100-6

12.5

9.4

0.755

0.5

417

407

0.976

25.3

25.6

1.012

NS5-100-6

20.8

15.9

0.765

1.1

424

431

1.017

39.5

39.9

1.010

NS7-100-6

29.1

20.3

0.697

1.5

434

443

1.021

49.1

53.9

1.098

NS3-60-6

12.5

9.4

0.756

0.7

398

396

0.995

26.2

26.3

1.004

NS5-60-6

20.8

16.1

0.773

1.3

410

419

1.022

40.8

41.2

1.010

NS7-60-6

29.1

20.5

0.705

1.8

418

436

1.043

51.0

55.5

1.088

Average value

0.737

1.071

0.992

Std. dev.

0.045

0.068

0.044

38

Figures

Fig. 1 Schematic diagram of the labeling system.

(a) Drop hammer tester at Harbin Institutive of Technology

(b) Specimen and related setups

(c) Schematic view of specimen and supports

Fig. 2 Drop hammer system.

39

(a) Typical failure mode of outer steel pipe. (Impact along ―–Y‖ axis)

(b) Typical failure mode of core concrete. (Impact along ―–Y‖ axis) Fig. 3 Typical flexural failure mode of HSCFST against impact (HS7-100-6).

(a) Influence of impact energy

(b) Influence of steel ratio

(c) Influence of steel strength

(d) Influence of concrete strength

Fig. 4 Comparisons of overall bending shape for test specimens at various test parameters.

40

2400

2400 Test FEA

2000

2000 1600

Fp,test

800

400

0 -0.004 0.000 0.004 0.008 0.012 0.016 0.020

0 -0.004 0.000 0.004 0.008 0.012 0.016 0.020

0 -0.004 0.000 0.004 0.008 0.012 0.016 0.020

t/s

t/s

t/s

(b) HS7-100-6

2000

(c) HS7*-100-6

2000

Test FEA

1600 1200

2000

Test FEA

1600 1200

Fp,test

1200

Fp,test

800

Fp,test

800 400

400

400

Test FEA

1600

F / kN

800

F / kN

0 -0.004 0.000 0.004 0.008 0.012 0.016 0.020

0 -0.004 0.000 0.004 0.008 0.012 0.016 0.020

0 -0.004 0.000 0.004 0.008 0.012 0.016 0.020

t/s

t/s

t/s

(e) HS7-60-6 2000

2000 Test FEA

1600

2000

Test FEA

1600 1200

800

F / kN

1200

Fp,test

0.000

0.004

0.008

0.012

1200

800

0 -0.004

0.016

Fp,test

(g) NS3-100-6

0.000

0.004

0.008

0.012

0 -0.004

0.016

Test FEA

400

0.008

0.012

0.016

(i) NS7-100-6 Test FEA

1200

Fp,test

0.004

2000

1600

F / kN

1200

0.000

t/s

2000

1600

Fp,test

400

(h) NS5-100-6

2000

800

800

t/s

t/s

Test FEA

1600

400

400 0 -0.004

(f) HS7*-60-6

F / kN

(d) HS3-60-6

Test FEA

1600 1200

800

Fp,test

F / kN

F / kN

Fp,test

800

400

(a) HS3-100-6

F / kN

1200

Fp,test

800

400

Test FEA

2000

F / kN

1200

1200

2400

1600

F / kN

F / kN

1600

F / kN

Test FEA

800

Fp,test

400

400

0 -0.004 0.000 0.004 0.008 0.012 0.016 0.020

0 -0.004 0.000 0.004 0.008 0.012 0.016 0.020

0 -0.004 0.000 0.004 0.008 0.012 0.016 0.020

t/s

t/s

t/s

(j) NS3-60-6

(k) NS5-60-6

41

(l) NS7-60-6

1800 1500

Test FEA

1500

1200

1200

Fp,test

600

F / kN

900

900

Fp,test

600

Fp,test

300

0 -0.004 0.000 0.004 0.008 0.012 0.016 0.020

0 -0.004 0.000 0.004 0.008 0.012 0.016 0.020

0 -0.004 0.000 0.004 0.008 0.012 0.016 0.020

t/s

t/s

t/s

(m) NS3-100-5

(n) NS5-100-5

1800

(o) NS7-100-5

1800

Test FEA

1500

1800 Test FEA

1500

1200

600

1200

F / kN

900

900 600

Fp,test

900

Fp,test

600

Fp,test

300

300

Test FEA

1500

1200

F / kN

F / kN

900 600

300

300

Test FEA

1500

1200

F / kN

F / kN

1800

1800

Test FEA

300

0 -0.005 0.000 0.005 0.010 0.015 0.020 0.025

0 -0.005 0.000 0.005 0.010 0.015 0.020 0.025

0 -0.005 0.000 0.005 0.010 0.015 0.020 0.025

t/s

t/s

t/s

(p) NS3-100-4

(q) NS5-100-4

(r) NS7-100-4

Fig. 5 Comparisons of impact force versus time curves between tests and FEA results. u Vibration stage Stable stage Descending stage

um ur Mid-span displacement

t

F

Impact force

Fm Fp 0 t0

t1

t2

td

t

Fig. 6 Typical impact force and mid-span deformation time histories for HSCFST member against transverse impact loadings.

42

(a) Influence of concrete grade for HS

(b) Influence of concrete grade for NS

(c) Influence of steel ratio

(d) Influence of strength for steel and concrete

Fig. 7 Comparisons of plateau value of impact force and mid-span displacement versus impact energies for tests. 80

60 Ei=37.3kJ

40 Ei=29.1kJ

20

0.008

0.012

0.016

40 Ei=29.1kJ

0 0.000

0.020

0.004

0.012

(a) HS-100-6 Test FEA

Ei=12.5kJ

Ei=29.1kJ

u / mm

60

20

0.004

t/s

(d) NS-60-6

0.016

0.020

0.016

0.020

(c) NS-100-6 80 Ei=29.1kJ

40 Ei=20.8kJ

0 0.000

0.012

t/s

Test FEA

20

0.012

0.008

100

0.005

0.010

0.015

t/s

(e) NS-100-5

0.020

0.025

Test FEA Ei=29.1kJ

60

Ei=20.8kJ

40 20

Ei=12.5kJ

Ei=12.5kJ

0.008

0 0.000

0.020

(b) HS-60-6

Ei=20.8kJ

0.004

40

20

80

40

0 0.000

0.016

Ei=29.1kJ

t/s

80

u / mm

0.008

Test FEA

Ei=20.8kJ

Ei=12.5kJ

t/s

60

60 Ei=37.3kJ

20

Ei=12.5kJ

0.004

80

Test FEA

u / mm

0 0.000

u / mm

u / mm

60

Test FEA

u / mm

80

Ei=12.5kJ

0 0.000 0.005 0.010 0.015 0.020 0.025 0.030

t/s

(f) NS-100-4

Fig. 8 Comparisons of mid-span displacement versus time curves between tests and FEA results.

43

2400 2000

F / kN

1600 1200

Fp,test

800 400 0

0

10

20

30

40

50

60

u / mm

Fig. 9 Typical impact force versus mid-span displacement curve for HSCFST member under transverse impact.

Fig. 10 General view and mesh of FE model for HSCFST against transverse impact loading. 300 250

N / kN

200 150 100 50 0 0.00

0.01

0.02

0.03

0.04

0.05

t/s

Fig. 11 Time history of axial force obtained in the developed FE model for NC-0.3-6 [37].

44

90

400

60

200

30

150 Test (F) FEA (F)

800

Test (u) FEA (u)

1000

120

600

90

400

60

200

30

150 Test (F) FEA (F)

800

Test (u) FEA (u)

120

600

90

400

60

200

30

0 0 -0.006 0.000 0.006 0.012 0.018 0.024 0.030

0 0 -0.006 0.000 0.006 0.012 0.018 0.024 0.030

t/s

t/s

t/s

(a) Yang et al. (NC-0-6) [37]

(b) Yang et al. (NC-0.15-6) [37]

(c) Yang et al. (NC-0.3-6) [37]

120 Test (F) FEA (F)

160

F / kN

u / mm

60

30

40 0 -0.02

0.00

0.02

0.04

0.06

0.08

0.02

Test (u) FEA (u) 90

180 60 120 30

60

0 0.10

0.03

120 Test (F) FEA (F)

240

90

80

0.01

300

Test (u) FEA (u)

120

0.00

0 -0.02

u / mm

200

0 -0.01

0.00

0.02

0.04

0.06

0.08

t/s

t/s

(d) Remennikov et al. [39]

(e) Yousuf et al. [41]

0 0.10

Fig. 12 Comparisons of impact force and mid-span displacement between tests of previous studies and FEA results.

45

0 0.04

u / mm

1000

120

600

F / kN

F / kN

800

Test (u) FEA (u)

u / mm F / kN

150 Test (F) FEA (F)

u / mm F / kN

1000