Experimental study and numerical simulation of replaceable corrugated steel plate-concrete composite shear walls

Soil Dynamics and Earthquake Engineering 127 (2019) 105827

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Experimental study and numerical simulation of replaceable corrugated steel plate-concrete composite shear walls

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Wei Wang∗, Jiangliang Song, Mingyue Hou, Gewei Liu, Wanzhi Wang School of Civil Engineering, Xi'an University of Architecture and Technology, Xi'an, 710055, China

ARTICLE INFO

ABSTRACT

Keywords: Corrugated steel plate-concrete (CSPC) composite shear walls Tension-and-compression damper Replaceable corner damper Cyclic loading tests Seismic performance

In this paper, a new structural form of installing a replaceable corner tension-and-compression damper (RCTCD) at the bottom of corrugated steel plate-concrete (CSPC) shear walls was proposed. This new structural wall is called a replaceable corrugated steel plate-concrete (RCSPC) composite shear wall. After a strong earthquake, the damage is mainly concentrated in the RCTCDs, and the function of the structure is quickly restored by replacing the RCTCDs. To study the seismic performance of RCSPC shear walls, cyclic loading tests on one replaceable horizontally corrugated steel plate-concrete (RHCSPC) composite shear wall and one replaceable vertically corrugated steel plate-concrete (RVCSPC) composite shear wall were conducted. To verify the feasibility of replacement, both the RHCSPC and the RVCSPC structural walls experienced replacement and reloading processes. The test results indicate that the structural behaviour of the two structural walls were flexure dominant. The damage in the RCSPC structural walls were mainly concentrated in the RCTCDs. The replacement of the RCTCDs can be implemented conveniently after residual deformation occurs in the structure. Compared with the RHCSPC structural wall, the RVCSPC structural wall had better seismic performance. Subsequently, the experimental and numerical simulation results were compared, the numerical simulation could predict the shear strength of the RVCSPC and RHCSPC shear walls with good accuracy. In addition, simplified formulas were proposed to evaluate the shear buckling strength of the RVCSPC and RHCSPC shear walls. The calculated results agree well with the test results, verifying the accuracy of the proposed formula.

1. Introduction Shear walls are the main lateral force-resistance component of highrise building structures and have been widely used in seismic areas around the world. However, with the increasing height of high-rise and super high-rise buildings, higher performance of shear walls is also needed. The conventional reinforced concrete (RC) shear wall has been found from past earthquake experience to have poor deformability, low energy dissipation and low ductility [1,2]. Moreover, after an earthquake, to absorb the earthquake energy, significant residual drift and damage accumulates in the wall base plastic hinge region [3]. To improve the seismic performance of building structures, in recent years, researchers have performed a series of innovative studies on shear wall structure. Zona et al. [4] proposed an innovative hybrid coupled wall (HCW) system made of a single reinforced concrete (RC) wall coupled to two steel side columns by means of steel links and proposed a specific design procedure for studying the building structure. Asta et al. [5] analysed innovative steel frames with reinforced concrete infill wall (SRCW) systems in which energy dissipation is only carried out in the vertical elements of the steel frame, which are subjected mainly to axial

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forces, and the damage characteristics are verified by experiments and finite element software simulations. Hashemi et al. [6] proposed a hybrid damage avoidant steel-timber wall system using the innovative resilient slip friction (RSF) joint, and the efficiency of the system was studied through experimental joint component experiments. Pavi et al. [7] analysed the mechanical performance of a coupled steel plate shear wall that consists of two or several steel plate shear walls connected to each other at story levels under cyclic loading using the nonlinear finite element method. According to the results, the shear resistance and energy dissipation of the model increase with increasing capacity and length of the coupling beams. Similarly, researchers have also studied the seismic behaviour of flat steel plate concrete (FSPC) shear walls analytically [8,9] and experimentally. The FSPC shear wall could provide high initial elastic stiffness, great shearing capacity, and superior and excellent energy dissipation. However, in a recent study, it was found that the bond slip between the concrete and the flat steel plate of the composite shear wall and the buckling deformation were easily occurred. Therefore, as an improvement of FSPC shear walls, Wang Y. et al. [10] studied the seismic behaviour of the composite shear wall with concretefilled steel tubular frames and corrugated steel plates. The research showed

Corresponding author. E-mail address: [email protected] (W. Wang).

https://doi.org/10.1016/j.soildyn.2019.105827 Received 23 April 2019; Received in revised form 4 July 2019; Accepted 21 August 2019 0267-7261/ © 2019 Elsevier Ltd. All rights reserved.

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that the stiffness degradation of the proposed wall was slower than that of other walls, and the use of corrugated steel plate significantly improved the seismic performance while simultaneously increasing the ductility and reducing the damage. Author (Wang) et al. [11–13] completed a study addressing the seismic performances (including failure phenomena, failure mechanism, load carrying capacity, ductility and energy dissipation characteristics) of the corrugated steel plate-concrete (CSPC) composite shear wall and found that the CSPC shear wall had good bonding behaviour and seismic performance. However, the damage to the CSPC structural walls showed that the severe damage and residual drifts were still mainly concentrated at the bottom of the wall, and the residual drifts that occurred in the structures were difficult to repair after earthquakes, leading to a high repair cost. To reduce the aforementioned reasons, the resilient city was first proposed as the cooperative direction in the field of earthquake engineering in January 2009 [14,15]. The earthquake-resilient structure is the structure that can restore the structural function immediately after a severe earthquake without significant repair. To date, researchers have developed rocking structures [16,17], self-centring structures [18,19], and structures with replaceable structural components [20,21], which are the three main types of earthquake-resilient structures. Lu et al. [22] first proposed the concept of installing replaceable energy-dissipation components at the two bottom corners of the structural wall (the passive energy-dissipation component is a device on which damage is first concentrated). However, the test results showed that the lateral stiffness of the new structural wall installed with combined rubber bearings was much lower than that of the conventional RC structural wall, and the hysteresis behaviour of the replaceable component was also not very stable. Focusing on the residual deformation that will usually occur in the structural wall and the offset of the connection, which may make it difficult to replace the bearings properly. Q.Z. Liu et al. [23] improved the test scheme, and they optimized the performance of the replaceable components. Compared with the previous replaceable components, the stiffness and energy-dissipation capacity of the new replaceable component are much higher, and they also verified the feasibility of the replacement operation after a strong earthquake. In this paper, a new kind of replaceable corner tension-compression damper (RCTCD) installed at the two corners of CSPC shear walls was first studied. The cyclic loading tests on 1:2 scale horizontally corrugated steel plate-concrete (HCSPC) and vertically corrugated steel plate-concrete (VCSPC) shear walls installed with the RCTCDs were conducted to study the deformation capacity and failure modes, and both of the shear walls experienced replacement of the RCTCDs after unloading to verify the feasibility of the RCTCDs after a strong earthquake. In addition, the force-displacement hysteresis curves, skeleton curves, and seismic performance indexes (including the bearing capacity and displacement in each stage) were investigated. The failure characteristics, energy-dissipation capacity, deformation, stiffness and strength degradation characteristics of the structural walls were also analysed.

plate is open, and the lower end is welded to the lower connecting plate. The working mechanism of the RCTCD is illustrated in Fig. 3. When the replaceable corner tension-and-compression damper is subjected to tension, the outer corrugated steel plate is stretched alone, and the inner cross-core steel plate does not come into play. When it is compressed, the inner cross-core steel plate starts working, and the outer corrugated steel plate is subjected to compression together with a cross steel plate. The RCTCDs utilize the plastic deformation of the outer corrugated steel plate to dissipate energy. The working mechanism of RCTCDs can be divided into the following two stages. In stage I, both the outer corrugated steel plate and the inner cross steel plate remain elastic, the upper connection plate and the lower connection plate are relatively displaced under tension and compression, and the corrugated web plates undergo plastic deformation for energy dissipation. In stage II, the buckling deformation of the outer corrugated steel plate occurs, while the inner cross steel plate remains elastic. At this time, the inner cross steel plate begins to provide a pressure function to avoid transient damage to the shear wall. 3. Test program The replaceable vertically corrugated steel plate-concrete (RVCSPC) composite shear wall (named NEW_V) and the replaceable horizontally corrugated steel plate-concrete (RHCSPC) composite shear wall (named NEW_H) installed with RCTCDs were designed and constructed. The dimensions of all specimens are identical. The design parameters of the two new structural walls are identical except for the encased corrugated steel plates. The horizontally and vertically corrugated steel plates are considered as the variable for the two wall specimens. It is critical for the design of the RCTCD. If its stiffness is too large, the rest of the wall rather than the RCTCD might be damaged first during earthquakes. In contrast, if its stiffness is too small, the carrying capacity of the wall might be reduced too much. Therefore, the design bearing capacity of the RCTCD for NEW_V and NEW_H are both taken as 1.0. The main parameters of the replaceable corner tension-and-compression dampers are listed in Table 1. First, the NEW_V and NEW_H specimens were loaded. Then, they were unloaded after reaching the predetermined lateral displacement, and the RCTCDs were replaced. They were loaded again to study the feasibility of the replacement operation and the influence of the replacement of the RCTCDs on the seismic performance of the new structural wall. 3.1. Specimens In the new structural wall specimen, two RCTCDs were installed at the bottom of each side. To ensure that the steel plate and the concrete work together, shear studs with a diameter of 8 mm and a length of 60 mm were welded on both sides of the steel plate. The spacing of the shear stud is 200 mm. To facilitate replacement of the RCTCDs, the upper and lower steel plates are set at the two ends of the RCTCDs, as shown in Fig. 4. The highstrength bolts were used to connect the RCTCDs with embedded plates. The residual deformation often occurs in the shear wall after a strong earthquake, which may alter the position of the RCTCDs, as shown in Fig. 5. To realize the replacement of the RCTCDs after a strong earthquake, the RCTCD connectors must be able to accommodate a certain amount of misalignment. As shown in Fig. 6, the outer corrugated steel plate and the inner cross-core steel plate of the RCTCD for replacement are 5 mm shorter than that of the RCTCD first installed. The slot holes are set on both the lower connector plates and the upper connector plates. The total length of the slot hole is 42 mm to tolerate the ± 13 mm horizontal and vertical offset of the bolts, as shown in Fig. 7. After installation, the shim is stuffed into the gap between the lower connector and lower embedded plate to prevent looseness.

2. Description of corrugated steel plate-concrete composite shear wall with replaceable corner tension-compression dampers The dimensions of replaceable vertically (RVCSPC) and horizontally (RHCSPC) corrugated steel plate concrete composite shear walls are shown in Fig. 1 and Fig. 2, respectively. During a strong earthquake, the damage is expected to concentrate in the RCTCDs, and the other parts of the CSPC structural walls remain intact. The function of the CSPC structural walls can be restored quickly by replacing the RCTCDs after an earthquake. The configuration of the new replaceable corner tension-and-compression damper (RCTCD) is shown in Fig. 3. The replaceable part is located between the top and the bottom connection plate and is comprised of the outer corrugated steel plate and the inner cross-core steel plate. To realize the compression bearing capacity of the inner cross-core during the test, there is a certain height difference between the cross-core steel plate and the corrugated steel plate, the upper end of the inner cross-core steel

3.2. Material properties The parameters of the specimens are shown in Table 1. The mechanical properties of the steel plate and reinforcement are listed in 2

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Fig. 1. Dimensions and steel reinforcement of the RVCSPC. (a) Front view, (b) Vertically corrugated steel plate, (c) Section 1-1, (d) Section 2-2.

Tables 2 and 3. The average values of material tensile yield strength of steel plate was computed as 317 Mpa. The corresponding values of ultimate strength was found to be 477 Mpa. The average values of material tensile yield strength of reinforcements with diameters of 8 mm and 16 mm were computed as 545 Mpa and 445 Mpa, respectively. The corresponding values of ultimate strength was found to be 573 Mpa and 615 Mpa. The concrete of the specimens was cast in five stages, casting of the foundation beam with a height of 0.5 m at the first stage, and then the height of the concrete wall was divided into three stages to cast, and last, the casting of the remaining part. The compression strength of the concrete is 18.4 MPa and 20.4 MPa. The elastic modulus of the concrete is 3.23E4 MPa and 3.32E4 MPa.

displacement with three cycles for each amplitude. The displacement amplitude increment was 5 mm. Throughout the test process, the lateral displacement, force, and the steel plate strains were recorded electronically. To verify the feasibility of replacing the RCTCDs after a strong earthquake, the specimens NEW_V and NEW_H were continuously loaded first and then unloaded after the top drift ratio was close to 1/ 100, which is the allowable maximum inter-story drift ratio for RC shear wall structures specified in Chinese code [24]. Then, the RCTCDs were removed and replaced with new ones, and the specimens NEW_V and NEW_H were loaded by the force and displacement-controlled loading conditions once again until the final failure. 3.4. Instrumentation

3.3. Test setup and procedures

The specimens were extensively instrumented to monitor local responses (e.g., strains) and global reactions (e.g., applied lateral load and displacement). The strains of the corrugated steel plates of the structural walls at critical regions and the corrugated steel plates of the RCTCDs were measured during testing. The arrangement of the strain gauges in the specimens is shown in Fig. 10. Linear variable-displacement transducers (LVDTs) were installed to monitor rotational and translational displacement. The lateral displacements at the top and at the level with a height of 825 mm above the pedestal were measured by horizontal LVDTs. An LVDT was mounted on the pedestal to monitor the slip of the pedestal. The vertical and horizontal displacements of the RCTCDs were also measured. The arrangement of the LVTDs is shown in Fig. 11.

The test setup is illustrated in Fig. 8. The side near the actuator was the west side, the side away from the actuator was the east side, and the front side of the wall was the south side. The force- and displacement-controlled loading history was adopted in this test, as illustrated in Fig. 9. At first, a vertical load was exerted on the top of the specimens by hydraulic jacks and kept constant throughout the entire testing process. The axial compression load ratio for all the specimens was 0.15. The estimated yielding displacement of the specimens by initial numerical analysis before the tests was approximately 10 mm. At the first stage of loading, prior to reaching a displacement level of 10 mm, the horizontal cyclic loading was controlled by the force with one cycle for individual amplitude. The force amplitude increment was 50 kN. In the second stage, the loading was controlled by the 3

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Fig. 2. Dimensions and steel reinforcement of the RHCSPC. (a) Front view, (b) Horizontally corrugated steel plate, (c) Section 1-1, (d) Section 2-2, (e) Section 3-3.

Fig. 3. Design and working mechanism of the RCTCD. (a) Appearance, (b) Working mechanism.

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Table 1 Parameters of the specimens. Specimen

Steel plate

Wave angle

Thickness of steel plate

VSPCSW HSPCSW

Vertically corrugated steel plate Horizontally corrugated steel plate

45° 45°

3 mm 3 mm

4. Test results 4.1. Failure mode To describe the test phenomena better, the push direction was defined as the positive direction, and the pull direction was defined as the negative direction. The stage before the replacement of the RCTCD was defined as the first loading process, and the stage after the replacement of the RCTCD was defined as the second loading process. According to the analysis of the test phenomena, the stress process of the specimen in the first loading process was divided into the following three stages. It was assumed that the specimen yields when the top drift ratio reached 1% so that the damage could be limited to the replaceable area, and no significant damage occurred in the non-replacement area of the shear wall. (1) Elastic stage: Before reaching the cracking load, the lateral forcetop displacement hysteresis curves of each specimen basically changed linearly, and no obvious test phenomenon occurred. (2) Concrete cracking and crack development stage: After the elastic stage, the concrete began to crack, and the cracking loads of the specimens NEW_V and NEW_H were 100 and 150 kN, respectively (and the corresponding top drift ratios were 0.12% and 0.24%, respectively). When the loads were −100 and −150 kN, respectively, the two specimens made noises similar to friction, which originated from the local damage of the interface between the corrugated steel plate and concrete.

Fig. 5. Residual deformation.

oblique crack width at the bottom of the east side was widened, and a large number of staggered oblique cracks gradually appeared in the wall. Fig. 12 shows the damage in the specimens and the RCTCDs when the top drift ratios are 0.92% and 1.00%, respectively. When loaded to 18.01 mm (the corresponding top drift ratio was 0.92%), there were no new cracks at the bottom of the shear wall, but there was obvious outward bending deformation on the outer corrugated steel plate of the RCTCDs and signs of tearing in the weld seam, while the inner crosscore steel plate of the RCTCDs did not show significant deformation. The phenomena above which showed that the RCTCD had sustained local buckling damage. When the specimen NEW_H was loaded to 19.5 mm (the corresponding top drift ratio was 1.00%), a new vertical crack appeared at the bottom of the east side, and bending deformation occurred on the outer corrugated steel plate of the RCTCDs, which proved that the RCTCDs partially reached the buckling state. Then, the RCTCDs were removed from the specimens NEW_V and NEW_H. According to the analysis of the test phenomena, the stress process

The cracks of the two specimens were first horizontal micro-cracks in the corner of the shear wall beside the RCTCD, and then due to the effect of the restrained square steel tube columns, the micro-cracks of the two specimens were first horizontal and then oblique. When the lateral force-top displacement hysteresis curves of the two specimens showed a gentle transition, the critical state of the elastic phase was determined. At this time, the corresponding loads of the two specimens were 250 and 300 kN, respectively (and the corresponding top drift ratios were 0.32% and 0.67%, respectively). (3). Buckling stage: At this stage, the lateral force-top displacement hysteresis curves from the obvious turning point to the peak load point. When the specimen NEW_V was loaded to −14.1 mm, the original

Fig. 4. Embedded plate and connector. RCTCD, replaceable corner tension and compression damper. 5

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Fig. 6. Adjustment of the replaceable corner tension-compression damper (RCTCD). (a) Before replacement, (b) After replacement.

Fig. 7. Adjustment range. (a) Lower connector, (b) Upper connector. Table 2 Mechanical properties for steel of the replaceable tension-compression damper. Steel type

Sample no.

Thickness (mm)

Yield strength (MPa)

Ultimate strength (MPa)

Elastic modulus ( × 105 MPa)

Inner cross-core and outer corrugated steel plate

1 2 3 Average

3 3 3 3

315 315 320 317

480 470 480 477

2.43 2.01 1.83 2.09

Table 3 Mechanical properties for the steel reinforcement. Diameter

8 mm (Stirrup and longitudinal rebars) 16 mm (longitudinal rebar)

Sample no.

Yield strength (MPa)

Ultimate strength (MPa)

Elastic modulus ( × 105 MPa)

1 2 3 Average 1 2 3 Average

545 540 550 545 445 450 440 445

565 580 575 573 610 620 615 615

1.99 2.20 2.13 2.10 1.93 1.95 1.99 1.96

Fig. 8. Test setup.

phenomena of lateral force-top displacement hysteresis curves of each specimen were the same as the curves in the first loading stage.

of the specimen in the second loading process was divided into the following four stages. The purpose was to destroy the specimens thoroughly to measure the ultimate bearing capacity and seismic performance of the specimens.

(2). Concrete cracking and crack development stage: The cracking loads of the specimens NEW_V and NEW_H were 100 and 50 kN, respectively (and the corresponding top drift ratios were 0.20% and

(1) Elastic stage: No obvious cracks occurred in the shear wall, and the 6

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edge of the specimen NEW_H. Subsequently, the lateral force-top displacement hysteresis curves of the two specimens showed a gentle transition, and then, the critical state of the elastic phase was determined. At this time, the corresponding loads of the two specimens were 250 and 300 kN, respectively (and the corresponding top drift ratios were 0.53% and 0.89%, respectively). (3). Buckling stage: When the specimen NEW_V was loaded to 25 and -25 mm, respectively (the drift ratios were 1.28% and −1.28%, respectively), the corrugated steel plate of the RCTCDs at the bottom of the east and west of the wall had a slight bulging deformation. When the specimens of NEW_V and NEW_H were loaded to 31.4 and 26.4 mm, respectively, the positive directions were to the peak loads 563.28 and 439.44 kN, respectively (the drift ratios were 1.61% and 1.36%, respectively). New horizontal and diagonal cracks occurred, and the crack width also increased. The corrugated steel plate of the RCTCDs in specimen NEW_V exhibited serious yielding deformation, and then the cross-core steel plate in the RCTCD began to provide vertical bearing capacity. However, the deformation of the RCTCDs in specimen NEW_H was slight. (4). Failure stage: Fig. 13 shows the damage in the specimen at final failure. When the top drift ratio of specimen NEW_V and NEW_H were 2.60% and 1.81%, respectively, the outer corrugated steel plates of the RCTCDs in specimens NEW_V and NEW_H experienced serious buckling failures. The damage in the outer corrugated steel plates of the RCTCDs is shown in Fig. 14. Both the east and the west bottoms of the specimens

Fig. 9. Diagram for load history.

0.10%, respectively). The development direction of the diagonal cracks of specimen NEW_V occurred in the west bottom. In addition, a small piece of concrete spalled off from the east edge of the specimen NEW_H. When the specimens NEW_V and NEW_H were loaded to −200 and 200 kN, respectively, new cracks extended to the foundation beam of specimen NEW_V. A small piece of concrete spalled off from the west

Fig. 10. Arrangement of strain gauges. (a) Corrugated steel plate of specimen NEW_V, (b) Corrugated steel plate of specimen NEW_H, (c) Square steel tube column of new structural wall specimen, (d) RCTCD of new structural wall specimen. 7

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curves do not decrease significantly, which is due to the lack of the plastic failure stage. (1) At the initial stage of loading, all the specimens are in the elastic stage, and the skeleton curve is close to a straight line. (2) After the yield load, before the replacement of the RCTCDs, the peak load of specimen NEW_V is slightly higher than that of NEW_H, but the ductility of specimen NEW_H is better than that of specimen NEW_V. After the replacement of the RCTCDs, the peak load of specimen NEW_V is much higher than that of specimen NEW_V, and the ductility of specimen NEW_V is 1.28 times that of specimen NEW_H. (3) After the peak load, the bearing capacities of the specimens after the replacement of the RCTCDs begin to decrease. During the loading process, the RCTCDs in the specimen NEW_H buckle prematurely, and concentrated plastic deformation occurs at the toe of the wall, which results in serious damage to the specimens. The force drop rate of specimen NEW_H is much faster than that of NEW_V, which proves that the bearing capacity and ductility of specimen NEW_V are much better than those of specimen NEW_H.

Fig. 11. Arrangement of the linear variable-displacement transducers.

4.4. Characteristic loads and displacements

were seriously damaged, and the longitudinal rebars at the toes of the walls were buckled. The damage in longitudinal rebars at the toes of the specimens is shown in Fig. 15. The bottom of the specimens NEW_V and NEW_H formed a horizontal through-crack with a maximum width of 4 and 10 mm, respectively. It was confirmed that the RCTCD yielded before the yielding of the bottom edge of the wall. Comparing the first and the second loading processes of the specimens NEW_V and NEW_H, the damage in the second loading process was much more severe than in the first loading process, which was due to the accumulation of damage that occurred during the first loading process.

The yield point in the skeleton curve is defined to evaluate the ductility coefficients of the specimens by using the equivalent elastoplastic method suggested by Park [25]. The ultimate limit point is defined as the point on the descending section of the skeleton curve with a 15% force degradation. The lateral loads and displacements at individual characteristic points and the ductility coefficients for all the specimens are listed in Tables 4 and 5. Compared with specimen NEW_H, the load-carrying capacity of specimen NEW_V increased and the deformation capacity and ductility coefficient also increased. The secant stiffness at the cracking point regarded as the initial elastic stiffness is 54.57 and 48.78 kN/mm for the specimens NEW_V and NEW_H, respectively, before the replacement of the RCTCDs. For the specimens NEW_V and NEW_H after the replacement of the RCTCDs, the initial elastic stiffness is 35.15 and 31.18 kN/mm, respectively. Table 4 shows that the ultimate bearing capacity of specimen NEW_V is higher than that of specimen NEW_H. Table 5 shows that the yield displacement of specimen NEW_V is similar to that of specimen NEW_H, which indicates that there is good bonding performance between the corrugated steel plate and the concrete, and it is more difficult to produce out-of-plane buckling deformation and bond slip between the corrugated steel plate and the concrete.

4.2. Lateral force-top displacement relationship The lateral force-top displacement hysteresis curves of each specimen are shown in Fig. 16. It can be seen from Fig. 16 that the initial hysteresis curve of each specimen is basically symmetrical in both the first and the second loading stages. The hysteresis curve at the initial stage of loading is linearly developed with no residual deformation, and each specimen is in elastic working status. As the specimen enters the yielding stage, the hysteresis loops before the replacement of the RCTCDs roughly coincide with those after replacement. With the cracking of the concrete, the stiffness of the specimens begins to degenerate, each specimen is in the yielding stage, and the residual deformation before the replacement of the RCTCDs increases slowly after unloading, which is highly beneficial to replacing the RCTCDs. After reaching the peak load, the bearing capacity and stiffness of the specimens NEW_V and NEW_H after the replacement of the RCTCDs gradually degenerate, and the shuttle-shaped hysteresis curve of the specimen NEW_V tends to be plumper, which indicates better ductility and energy dissipation capacity. However, when the specimen NEW_H reaches the peak load, the RCTCDS and the toe of the wall are seriously damaged, which leads to the instability of the specimen, and the hysteresis curve shows a rapid decline, which indicates poor ductility and energy-dissipation capacity.

4.5. Displacement ductility coefficient The displacement ductility coefficient is an index used to measure the deformation capacity of the structure after yield. As shown in Table 5, the displacement ductility coefficient of specimen NEW_V is 1.28 times that of specimen NEW_H, which indicates that the ductility of specimen NEW_V is better. However, the reason that the ductility coefficients of the specimens NEW_V and NEW_H are not high is that both specimens undergo a secondary reloading process after the replacement of the RCTCDs, which leads to the degradation of stiffness and bearing capacity of the specimens and ultimately reduces the ductility coefficients.

4.3. Skeleton curve

4.6. Stiffness degradation

The lateral force-top displacement skeleton curves of the specimens before and after the replacement of the RCTCDs are shown in Fig. 17. For the specimens NEW_V and NEW_H, the skeleton curves after the replacement of the RCTCDS indicate an obvious S-shape, which indicates that the specimens undergo three stages of elastic, elastoplastic and plastic failure under low-frequency cyclic loads. For the specimens NEW_V and NEW_H before the replacement of the RCTCDs, the skeleton

Under the condition of constant displacement amplitude, the stiffness of structural members decreases with the increase in repeated loading times, which is called stiffness degradation and can be represented by loop stiffness, K1, where K1 refers to the ratio of the average load to the average displacement under multiple cycles of the same displacement amplitude. As shown in Fig. 18, at the early loading stage, for the 8

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Fig. 12. Damage in the specimens and RCTCDs at the top drift ratio of 0.92% and 1.00%. (a) NEW_V, (b) NEW_H, (c) NEW_V, (d) NEW_H.

specimens NEW_V and NEW_H before and after the replacement of the RCTCDs, the stiffness of the specimen NEW_V is larger than that of the specimen NEW_H. At the later loading stage, for the specimens NEW_V and NEW_H after the replacement of the RCTCDs, the stiffness of the specimen NEW_H is smaller than that of the specimen NEW_V, which is due to the premature buckling failure of the RCTCD.

NEW_H are relatively gentle. At the later loading stages, after the replacement of the RCTCDs, the bearing capacity reduction coefficient of specimen NEW_V decreases moderately, whereas the bearing capacity of specimen NEW_H decreases sharply. The sudden decline in the bearing capacity of the specimen NEW_H in the failure stage is due to the premature buckling failure of the RCTCDs.

4.7. Degradation of bearing capacity

4.8. Energy dissipation capacity

Bearing capacity degradation refers to the characteristic that the bearing capacity decreases with increasing loading cycles, which can be expressed by the bearing capacity reduction coefficient η. The bearing capacity reduction coefficient η refers to the ratio of the peak load value of the last cycle to the peak load value of the first cycle at the same displacement amplitude. The bearing capacity reduction coefficient displacement curve for each specimen is shown in Fig. 19. At the early loading stages, before and after the replacement of the RCTCDs, the bearing capacity reduction coefficients of the specimens NEW_V and

The energy dissipation capacity of each specimen can be comprehensively evaluated by the energy dissipation per half cycle, which is calculated according to the lateral force-top displacement hysteresis curves. The energy consumption-half cycle curves, cumulative energy consumption-half cycle curves, and equivalent viscous damping coefficient-cycle curves are shown in Figs. 20–22. From Figs. 20 and 21, it can be seen that as the half-cycle number increases, the energy dissipation of each specimen continuously increases. However, the energy consumption of the specimen NEW_V 9

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Fig. 13. Damage in the specimens at final failure. (a) The front of NEW_V (drift ratio 2.60%), (b) The back of NEW_V (drift ratio 2.60%), (c) The front of NEW_H (drift ratio 1.81%), (d) The back of NEW_H (drift ratio 1.81%).

Fig. 14. Damage in the outer corrugated steel plates of the replaceable corner tension-compression dampers. (a) NEW_V, (b) NEW_H.

Fig. 15. Damage in the longitudinal rebars at the toes of the specimens. (a) NEW_V, (b) NEW_H. 10

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Fig. 16. Lateral force-top displacement hysteresis curves. (a) Before replacement of specimen NEW_V, (b) After replacement of specimen NEW_V, (c) Before replacement of specimen NEW_H, (d) After replacement of specimen NEW_H, (e) Before replacement of specimens NEW_V and NEW_H, (f) After replacement of specimens NEW_V and NEW_H, (g) Before and after the replacement of specimen NEW_V, (h) Before and after the replacement of specimen NEW_H.

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Fig. 17. Lateral force-top displacement skeleton curves. (a) Before replacement of specimens NEW_V and NEW_H, (b) After replacement of specimens NEW_V and NEW_H, (c) Before and after the replacement of specimen NEW_V, (d) Before and after the replacement of specimen NEW_H.

increases faster than that of the specimen NEW_H in the failure stage, which indicates that the replaceable vertically corrugated steel plateconcrete (RVCSPC) composite shear wall has a stronger energy-dissipation capacity, which is due to premature buckling failure of the RCTCDs and eventually leads the replaceable horizontally corrugated steel plate-concrete (RHCSPC) composite shear wall to fail to make full use of its energy-dissipation capacity. As shown in Fig. 22, the equivalent viscous damping coefficient of each specimen increases with increasing cycle number. For the specimens NEW_V and NEW_H before the replacement of the RCTCDs, the

equivalent viscous damping coefficient of the specimen NEW_H is larger than that of the specimen NEW_V. However, after the replacement of the RCTCDs, the equivalent viscous damping coefficient of the specimen NEW_H is smaller than that of the specimen NEW_V, which indicates that the replaceable vertically corrugated steel plate-concrete (RVCSPC) composite shear wall has a stronger energy-dissipation capacity. In other words, the reason that the energy-dissipation capacity of the RHCSPC shear wall is too small, on the one hand, is due to the premature instability and failure of the RCTCDs at the toes of the

Table 4 Comparison of load-carrying capacity. Specimen number

Loading direction

Cracking load (kN)

Cracking load ratio

Yield load (kN)

Yield load ratio

Peak load (kN)

Peak load ratio

Ultimate load (kN)

Ultimate ratio

NEW_V (before)

Positive Negative Average Positive Negative Average Positive Negative Average Positive Negative Average

100.00 100.00 100.00 150.00 150.00 150.00 – – – – – –

0.67

437.58 411.25 424.42 370.47 334.04 352.26 515.71 442.97 479.34 385.56 371.22 378.39

1.20

503.96 460.37 482.17 425.96 388.34 407.15 563.28 481.41 522.35 439.44 416.00 427.72

1.18

– – – – – – 421.57 389.17 405.37 389.21 350.42 369.82

–

NEW_H (before) NEW_V (after) NEW_H (after)

–

1.27

12

1.22

1.10

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Table 5 Comparison of displacement and ductility. Specimen number

Loading direction

Cracking displacement (mm)

Yield displacement (mm)

Displacement at peak load (mm)

Ultimate displacement (mm)

Ultimate drift ratio

Ductility coefficient

Average

Ratio of ductility coefficient

NEW_V (before) NEW_H (before) NEW_V (after) NEW_H (after)

Positive Negative Positive Negative Positive Negative Positive Negative

2.30 2.76 4.67 5.41 3.90 3.60 1.90 2.02

12.49 14.33 14.03 14.02 14.45 14.09 14.19 14.04

18.02 18.03 19.52 19.53 31.42 25.02 26.43 26.42

– – – – 44.26 37.84 34.06 28.99

– – – – 2.3% 1.9% 1.7% 1.5%

– – – – 3.06 2.68 2.40 2.06

–

–

specimen RHCSPC shear wall, which leads to the formation of the plastic hinge at the toes of the wall. On the other hand, it is due to the error in the process of the RCTCDs in that the distance between the inner cross-core steel plate and the upper connector steel plate of the RCTCD in the specimen RHCSPC is too large. After the buckling of the outer corrugated steel plate of the RCTCD, the inner cross-core steel plate fails to provide vertical bearing capacity. In summary, if the stiffness of the RCTCDs could be reduced to match the stiffness of the rest of the wall, then the deformation ability of the specimen NEW_H might be improved to a certain extent.

– 2.87

1.28

2.23

to measure the strain of the cross-core steel plates. Actually, the plastic strain mainly concentrated on the yield segment of the outer corrugated steel plate, and the elastic strain of the other part was relatively small. The axial deformation-top displacement skeleton curves of the two RCTCDs on one side of the specimen are shown in Fig. 24. The compression deformation of the RCTCDs was much larger than the tensile deformation. The axial deformation of the outer RCTCD was larger than that of the inner RCTCD. The axial deformation was roughly proportional to the top displacement. Because of the presence of the residual deformation, the connectors of the RCTCDs should be specially designed so that they could tolerate a certain degree of misalignment. The total length of the slot hole was 42 mm to tolerate the ± 13 mm horizontal offset of the bolts. Therefore, the replacement of the RCTCD can be implemented after a strong earthquake when the construction details of the RCTCD proposed in this study are adopted.

4.9. Strains Fig. 23 shows the strain of the RCTCDs of specimen NEW_V and specimen NEW_H. Because of the symmetrical arrangement of the RCTCDs at the two bottom corners of the specimens NEW_V and NEW_H, only the strain gauge of C1 and C2 of the right side RCTCD is analysed. With the increase in the top displacement, both strain gauges C1 and C2 rose up. The strain gauge C2 yields before the strain gauge C1, which means that the RCTCD first breaks at the bottom part, and this is consistent with the phenomenon during the loading process. Table 6 shows the corresponding displacement of the RCTCD when yielding, and the partial incomplete data acquisition is due to the strain gauge failure, the main reason for the excessive deformation of the RCTCDs during the loading process.

5. Finite element simulation and verification Two finite element models of the RVCSPC and RHCSPC were developed using Abaqus to simulate the behaviours of the composite shear walls. Both of the element models had the same boundary frame and axial compression ratio for comparison purposes. The meshing of the finite element models is shown in Fig. 25. In the finite element, solid elements C3D8R was used to simulate the corrugated steel plates, H-shaped steel beams, square steel tube columns, concrete, and the RCTCDs. The truss element T3D2 was applied in the simulation of the rebars. The mesh dimensions of the steel plate, RCTCDs and concrete were 10 cm, 2.5 cm and 10 cm, respectively. There were 600 elements in total. Tie constraints were used to connect the

4.10. Deformation of the replaceable corner component Because there were no strain gauges set in the inner cross-core steel plate of the RCTCDs, the axial displacements of the RCTCDs were used

Fig. 18. Loop stiffness-displacement curves of the specimens. (a) Before replacement of specimens NEW_V and NEW_H. (b) After replacement of specimens NEW_V and NEW_H. 13

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Fig. 19. Bearing capacity reduction coefficient displacement curves. (a) Before replacement of specimens NEW_V and NEW_H. (b) After replacement of specimens NEW_V and NEW_H.

corrugated steel plates to the square steel tube columns and the H-shaped steel beams. Additionally, the RCTCDs were connected with the upper and lower embedded plates of the shear walls by tie constraints. All the steels were embedded in concrete. For the boundary conditions, all six degrees of freedom of the ground beams were constrained since the foundation beams were fixed in the testing. Two analysis steps were established. The vertical load was controlled by Step 1, and the horizontal displacement was controlled by Step 2. The von Mises yield criterion was adopted in the analysis. A displacement-controlled lateral loading was applied to the centre of the beam-column joint, according to the loading protocols. To verify the accuracy of the numerical simulations, the predicted results are compared with the experimental results in terms of the hysteresis curves and failure mode, as shown in Fig. 26, Fig. 27 and Table 7. Fig. 26 demonstrates the hysteresis curves of specimens NEW_V and NEW_H after the replacement of the RCTCDs. Fig. 26 demonstrates that the finite element models could accurately predict the ultimate capacity of the specimens NEW_V and NEW_H. However, there are still slight differences between the numerical simulations and experimental results, which are the hysteresis curves obtained by Abaqus finite element simulations that are fuller, and the lateral stiffness of the finite element model is obviously higher than that of the experimental results. This is mainly due to the following two reasons: first, the finite element models cannot simulate the applied

forces, which are influenced by manufacturing error and lateral supports. Second, it is due to the bond-slip action between the corrugated steel plate and the concrete is neglected in the process of finite element analyses. Fig. 27 indicates the comparison of the failure modes of the specimens NEW_V and NEW_H, after the replacement of the RCTCDs, obtained from the Abaqus and test results, where “E” represents the experimental result, whereas “N” presents the numerical result. According to the finite element analysis, the stress distribution of the concrete of the finite element models of the RVCSPC and RHCSPC shear walls are mainly concentrated at the toes of the specimen and are basically consistent with the crack distributions in the tests. In addition, the stress and deformation of the RCTCDs at the toes of the two models are relatively larger, which meets the design requirements of the structural system with replaceable components: the failure is concentrated on the replaceable components to keep the main bearing components in good condition. Table 7 shows that the error values are controlled within 10% on average, which means that the predicted peak load and failure modes were in good agreement with the results obtained from the experiment, and also, the finite element models could simulate the performance of the RVCSPC and the RHCSPC shear walls well.

Fig. 20. Energy consumption-half cycle curves of specimens. (a) Before replacement of specimens NEW_V and NEW_H. (b) After replacement of specimens NEW_V and NEW_H. 14

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Fig. 21. Cumulative energy consumption-half cycle curves of specimens. (a) Before replacement of specimens NEW_V and NEW_H. (b) After replacement of specimens NEW_V and NEW_H.

6. Shear-bearing capacity

horizontal projection of the inclined panel width, c is the inclined panel width, and the size is shown in Fig. 28, and â is the global shear buckling coefficient. The calculation method of the elastic interactive shear buckling stress τEcr,I suggested by Yi is expressed as follows:

6.1. Shear buckling behaviour of the corrugated steel plate According to Yi et al. [26], the buckling strength of corrugated steel webs is controlled by interactive buckling which is attributed to the interaction between local and global shear buckling modes. The elastic local shear buckling stress equation is taken as follows: E cr , L

=

1 E cr , I

(1)

= CG

d t

1.5

t h

cr y

(2)

1

µ2) 4 ( )

3

4

(4)

= 1 0.614( 1/ s2

s s

< 0.6

0.6) 0.6 2 <

s s

2 (5)

6.2. Shearing capacity formula Based on the JGJ138-2016 Code for the design of composite structures [28], the shear-carrying capacity of the RVCSPC and RHCSPC composite shear walls are calculated through the superposition method, that is, adding the capacity of the reinforced concrete (Vc), the corrugated steel plate (Vp) and the RCTCDs (Vd). The shear forces Vc, Vp and Vd can be evaluated as follows:

5.045 E (1

1 E cr , G

where τcr is the critical shear buckling stress, τy is the shear yielding stress = fy / 3 , and λs is the shear buckling parameter, s = y / crE, I .

2

where CG is the coefficient unrelated with geometric parameters for the simplified global shear buckling stress, and is calculated as Equation (3), d is the corrugation depth, and h is the web height.

CG =

+

1

where E is the Young's modulus of elasticity, μ is the Poisson's ratio, t is the web thickness, and a is the flat panel width. The formula for the elastic global shear buckling stress is taken as follows: E cr , G

1 E cr , L

To account for the effects of inelasticity, residual stress, and initial deformations, Equation (5) was proposed in the design manual [27].

2

5.34 2E t 12(µ2 ) a

=

(3)

where η is the length reduction factor = (a + b)/(a + c), b is the

Fig. 22. Equivalent viscous damping coefficient-cycle curves of specimens. (a) Before replacement of specimens NEW_V and NEW_H, (c) After replacement of specimens NEW_V and NEW_H. 15

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Fig. 23. Strain of the corrugated steel plate in the replaceable corner tension-compression dampers of specimens NEW_V and NEW_H. (a) After replacement of specimen NEW_V, (b) After replacement of specimen NEW_H. Table 6 Displacement of the RCTCD at yield. Specimen

Loading direction

Yield displacement of point C1 (mm)

Yield displacement of point C2 (mm)

NEW_V

Positive Negative Average

22.14 – –

17.21 – –

Specimen

Loading direction Positive Negative Average

Yield displacement of point C1 (mm) – 14.65 –

Yield displacement of point C2 (mm) 12.93 14.93 13.93

NEW_H

1 0.5

0.5ft b w h w0 +

0.13NAw A

+

Vu =

fyh h w0 Ash Sv

cr Ap

(7)

where ξ is the reduction coefficient, τcr is the critical shear buckling stress, and Ap is the area of the steel plate. For the calculation of the shearing capacity of the corrugated steel webs in the RCTCD, it can be simplified as the steel plate connected with two sides. Referring to JGJ/T380-2015 Technical Code for Steel Plate Shear Wall [29], the formula for calculating the shear capacity of the RCTCD is taken as follows:

u

Vc =

0.6 0.5

Vp =

(6)

e

where λ is the shear span ratio, ft is the concrete tensile strength, bw is the width of the wall, hw0 is the effective depth of the wall, N is the axial compressive load, Aw is the area of the wall web, A is the gross crosssection area, fyh is the yield strength of horizontal distribution steel bar, Ash is the area of horizontal distribution steel bar, and Sv is the spacing of horizontal distribution steel bar.

= [0.2 ln(Le / He ) =

(8)

u Le t

He tk

0.05 ln( e ) + 0.68]

cr

(9) (10)

where Vu is the design value of the shear capacity of the steel plate shear wall, τu is the design value of the ultimate shear strength of steel, Le is the span of the steel plate shear wall, He is the height of the steel plate shear wall, λe is the relative height-thickness ratio of the steel plate shear wall, and εk is the steel grade correction factor = 235 / fy . From the above Equations 8–10, the Vd formula for calculating the shear capacity of the RCTCD can be obtained as follows:

Fig. 24. Axial deformation-top displacement skeleton curves of the inner replacement corner tension-compression dampers (RCTCDs). (a) Before replacement of specimens NEW_V and NEW_H, (b) After replacement of specimens NEW_V and NEW_H. 16

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Fig. 25. Meshed finite element model. (a) Concrete, (b) Embedded horizontally corrugated steel, (c) Embedded vertically corrugated steel, (d) Replaceable corner tension-compression damper.

Vd =

0.6 u le t 0.5

the formulas are basically consistent with the test shear strength (Vexp), and the error values are controlled within 10% on average, which shows that the theoretical formulas fit well and have certain reliability.

(11)

Therefore, the shear-carrying capacity of the mixed section of the corrugated steel plate composite shear wall is as follows:

V = Vc + Vp + Vd

7. Conclusions

(12)

A new type of earthquake-resilient structural wall was proposed in this study. Cyclic loading tests were carried out on the new structural wall specimens. Based on the test results, the following conclusions can be drawn:

Table 8 shows the calculated shear strength of the RVCSPC and RHCSPC shear walls using the abovementioned formulas. In addition, the comparison of the test shear strength (Vexp) to the calculated values (Vcal) is also shown in Table 8. The calculated values (Vcal) obtained by

(1) The structural behaviour of all specimens was flexure dominant. As

Fig. 26. Comparison of the lateral applied force-displacement curves obtained from the numerical and experimental results. (a) After replacement of the specimen NEW_V, (b) After replacement of the specimen NEW_H.

17

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Fig. 27. The failure mode obtained from the finite element analysis. (a) Concrete of the RVCSPC, (b) Steel of the RVCSPC, (c) Concrete of the RHCSPC, (d) Steel of the RHCSPC, (e) RCTCDs of the RVCSPC, (f) RCTCDs of the RHCSPC. Table 7 Comparison of test and finite element results. Specimen

NEW_V (RVCSPC) NEW_H (RHCSPC)

Peak load (kN)

Table 8 Calculated and measured shear-carrying capacity of the specimens NEW_V and NEW_H.

Vnum/Vexp

Vnum

Vexp

579.16 462.58

522.35 427.72

1.10 1.08

Type

Vcal (kN)

Vexp (kN)

Vcal/Vexp

RVCSPC (NEW_V) RHCSPC (NEW_H)

509.30 439.36

522.35 427.72

0.975 1.027

specimens after the replacement of the RCTCDs remained better. (3) The replaceable corrugated steel plate-concrete (RCSPC) composite shear wall, especially the replaceable vertically corrugated steel plate-concrete (RVCSPC) composite shear wall, had higher energy dissipation capacity. Compared with the replaceable horizontally corrugated steel plate-concrete (RHCSPC) composite shear wall, the ductility coefficient of the RVCSPC increased by 28%. (4) The initial stiffness of the replaceable vertically corrugated steel plate-concrete (RVCSPC) was higher than that of the replaceable horizontally corrugated steel plate-concrete (RHCSPC), and the degradation of bearing capacity and stiffness was slower. The lateral bearing capacity and ultimate displacement of the RVCSPC increased by 22% and 30%, respectively, compared with those of

Fig. 28. Trapezoidal corrugated web.

expected, in the new structural wall specimens, the damage mainly concentrated on the replaceable corner tension-compression dampers (RCTCDs) and the main part was well protected. (2) The replacement of the RCTCDs could be implemented conveniently even after the residual deformation occurred in the structure. The seismic performance of the new structural wall 18

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the RHCSPC. (5) The results obtained by Abaqus were basically in agreement with the experimental results. The failure was concentrated on the replaceable components to keep the main bearing components in good condition. Therefore, the results of the finite element could be used in engineering analysis. (6) The simplified formulas were developed to evaluate the shearing capacity of the RVCSPC and RHCSPC shear walls with a difference of less than 10%. These results showed that the theoretical formulas fit well and are highly accurate.

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Acknowledgments The authors are grateful for the financial support received from the National Natural Science Foundation of China (grant nos.51578449 and 51878548) and the key project of the Nature Science Foundation programme of Shaanxi Province (grant no.2018JZ5013). Appendix A. Supplementary data Supplementary data to this article can be found online at https:// doi.org/10.1016/j.soildyn.2019.105827. References [1] Wood S, Wright J, Moehle J. The 1985 Chile earthquake: observations on earthquake-resistant construction in Vina del. Urbana: University of Illinois at Urbana -Champaign; 1987. [2] Liao FY, Han LH, Tao Z. Performance of reinforced concrete shear walls with steel reinforced concrete boundary columns. Eng Struct 2012;44:186–209. [3] Gu AQ, Zhou Y, Xiao Y, Li QW, Qu G. Experimental study and parameter analysis on the seismic performance of self-centering hybrid reinforced concrete shear walls. Journal of Soil Dynamics and Earthquake Engineering 2018;116:409–20. [4] Zona A, Degée H, Leoni G, Dall'Asta A. Ductile design of innovative steel and concrete hybrid coupled walls. J Constr Steel Res 2016;117:204–13. [5] Dall'Asta A, Leoni G, Morelli F, Salvatore W, Zona A. An innovative seismic-resistant steel frame with reinforced concrete infill walls. Eng Struct 2017;141:144–58. [6] Hashemi A, Zarnani P, Masoudnia R, Quenneville P. Seismic resistant rocking coupled walls with innovative Resilient Slip Friction (RSF) joints. J Constr Steel Res 2017;129:215–26. [7] Pavir A, Shekastehband B. Hysteretic behavior of coupled steel plate shear walls. J Constr Steel Res 2017;133:19–35. [8] Shafaei S, Ayazi A, Farahbod F. The effect of concrete panel thickness upon composite steel plate shear walls. J Constr Steel Res 2016;117:81–90. [9] Dey S, Bhowmick AK. Seismic performance of composite plate shear walls. Structure

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