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    Experimental Investigation of Steel Plate Shear Walls under Shear-Compression Interaction

    This paper describes the derivation of the equation for evaluating the strength of bridge steel plate  reinforced concrete structure (SC) and the experimental results of SC panels subjected to in-plane shear.

    Two experimental research programs were carried out. One was the experimental study in which the influence of the axial force and the partitioning web were investigated, another was that in which the influence of the opening was investigated.

    In the former program, nine specimens were loaded in cyclic in-plane shear. The test parameters were the thickness of the surface high building steel plate, the effects of the partitioning web and the axial force. The experimental results were compared with the calculated results, and good agreement between the calculated results and the experimental results was shown.

    In the later programs, six specimens having an opening were loaded in cyclic in-plane shear, and were compared with the results of the specimen without opening. FEM analysis was used to supplement experimental data. Finally, we proposed the equation to calculate the reduction ratio from the opening for design.

    Four scaled one-storey single-bay steel plate shear wall  specimens with unstiffened panels were tested to determine their behaviour under cyclic loadings. The shear walls had moment-resisting beam-to-column connections. Four different vertical loads, i.e., 300 kN, 600 kN, 900 kN, and 1200 kN, representing the gravity load of the upper storeys were applied at the top of the boundary columns through a force distribution beam. A horizontal cyclic load was then applied at the top of the specimens. The specimen behaviour, envelope curves, axial stress distribution of the infill steel plate, and shear capacity were analyzed. The axial stress distribution and envelope curves were compared with the values predicted using an analytical model available in the literature.

    To investigate the shear resistance of single-bay low alloy plate shear walls (SPSWs), a large number of experiments have been conducted using low cyclic loading. Driver et al. [1] carried out a cyclic test of a four-storey SPSW. A vertical load of 720 kN was applied at the top the of the boundary columns with one horizontal load at each floor level. Their results indicated that the final deflection at the top floor is nine times larger than the yield deflection. The test specimen proved to be initially very stiff and had an excellent ductility and energy dissipation. Later, because of damage to the 1st storey, only the upper three storeys of this SPSW specimen was further tested by Behbahanifard [2] to verify their finite element model. Moghimi and Driver [3] carried out a test of a large-scale two-storey SPSW specimen. A vertical load was also considered in their test. The results indicated an excellent performance. In addition, high ductility and energy dissipation were observed. Qu et al. [4] performed a two-phase experimental program on a full-scale two-storey SPSW with reduced beam section connections and composite floors. Their first-phase test was pseudodynamical tests using three ground motions of decreasing intensity. The buckled infill steel plate was replaced by new panels in the subsequent test. Their results showed that the repaired specimen could survive and dissipate significant energy without severe damage to the boundary frame. The final storey drifts reached 5.2% and 5.0% at the first and second storey. Other experimental research included works on reduced beam section anchor [5], low-yield point SPSW [6], unstiffened perforated SPSWs [7], partially connected SPSWs [8], SPSWs with semirigid connected frame [9], shake table test of buckling restrained SPSWs [10], self-centering SPSW [11], and the use of light-gauge SPSWs [12].

    However, until today, physical experiments on the cyclic behaviour of SPSWs under concurrent gravity and horizontal load have not been reported. Most investigations were performed numerically, e.g., Elgaaly and Liu [13] compared the shear carrying-capacity of a SPSW with and without gravity load by the finite element method. The authors concluded that the gravity load has little effect on the shear-carrying capacity. This might be caused by the low magnitude of compression and thin infill wall considered in their analyses. Zhang and Guo [14] performed finite element analyses on the behaviour of SPSWs with precompression from the adjacent columns. Their research showed that the shear capacity of SPSWs was significantly impaired by the precompression. Their previous research [15] also showed that the gravity load acting at the top of the boundary columns has significant effects on the shear load-carrying capacity.

    To evaluate the adequacy of an analytical model available in the literature in predicting the influences of gravity loads on the cyclic performance of the SPSW, physical experiments were performed in this paper. Four scaled SPSWs under compression-shear interaction were designed and tested. The envelope curves, the axial stress distribution and the maximum shear capacity obtained from the experiments were then used in the evaluation of the analytical model.

    The test parameter was the vertical load applied at the top boundary columns. Test specimens are one-storey walls. The height of the specimen was 0.75 m, and the width was 1.1 m. The columns were 1 m apart from center to center. Figure 1 shows the size and configuration of the specimens. The plate thickness was 2.1 mm Q235 steel with the yield strength of 255 MPa. The size of the infill plate was 600 mm × 900 mm. The frame members are built-up sections made of Q345 steel with the yield strength of 460 MPa. The boundary columns, i.e., H-overall depth (d) × flange width (bf) × web thickness (tw) × flange thickness (tf), have, respectively, the dimensions of 100 mm × 100 mm × 6 mm × 8 mm. The top beam, connected to the actuator, has the corresponding dimensions of 150 mm × 100 mm × 6 mm × 9 mm. This beam was stiff to ensure a smooth transfer of the load to the tension field occurred below the beam. Moment connections were used at all beam-to-column joints. Connection of the beam flanges to the columns was constructed using complete penetration groove welds. The beam webs were welded to the column flange by two-sided fillet. The infill high strength plate was connected to the boundary beams and columns using the fishplates. Figure 2 shows the fishplates of 50 mm width and 3 mm thickness. Continuous fillet welds on both sides of the fishplates were used. The infill steel plate is fitted to the fishplates with a lap of approximately 20 mm all around.

    Figure 1 shows a sketch of the test specimen with the vertical load and the lateral cyclic load. The constant vertical load of specimens A, B, C, and D was, respectively, 300 kN, 600 kN, 900 kN, and 1200 kN. It was kept steady during the whole test. The beam that distributed the vertical load was hinged to the top of the two boundary columns. A hydraulic jack generated the vertical load at the top of the load distribution beam. It was supported by a stiff steel frame. To avoid any shear force, a roller was placed between the steel frame and the hydraulic jack.

    Horizontal cyclic load acted at the center line of the top beam. A hydraulic jack, supported by a laboratory reaction wall, generated the horizontal load. Up to the first yield in the steel plate, a force-controlled load was applied. Depending on the load combination, the first yielding stage was achieved by increasing the horizontal load from ±50 kN to ±200 kN with an increment of 50 kN. In the subsequent load cycles, a displacement-controlled loading was performed until the specimen failure by increasing the displacement after each three cycles from ±4 mm.

    Eighteen strain gauges were attached along the boundary columns to measure the axial strain, and 8 strain rosettes were attached at the surface of the infill spring steel (see Figure 1). Two linear variable differential transformers (LVDTs) were installed at the base of the boundary column and at the center line of the top beam.

    Table 1 lists the results of the coupon tests of the two materials Q345 and Q235. Three coupons were tested for each material, and the average value is used for the subsequent analytical analyses.

    Under pure gravity loads, no buckling occurred in specimens A, B, and C. However, in the case of specimen D, horizontal buckling of the infill steel plate took place. The following described the behaviour of the four specimens.

    In the case of specimen A, until the top displacement reached 4 mm, there was no buckling in the infill steel plate. The tension strips, formed from the lower left corner to the upper right corner, have an inclination angle near to 45°. During the first displacement cycle of 8 mm, the first loud bangs occurred. In the subsequent cycles, these noises continued to occur. With an increase of the top displacement, various parts of the infill steel plate progressed to yield. A large residual deformation formed at the end of each pull or push loading with a further increase of residual deformation in the subsequent cycles. The first tear was detected in the upper left corner between the fishplate and the infill steel plate during the first 12 mm displacement load. The tear gradually increased to nearly 30 mm at the end of the first 12 mm cycles, as shown in Figure 3. At the end of the second 12 mm displacement cycle, new tears were detected at the two lower corners. All the tears extended with an increase of the top displacement, but no new tear was observed. During the first 20 mm displacement cycle, the shear resistance of the specimen did not decrease, but the tears grew faster and the specimen was pushed over. The ultimate deformation at the top of the specimen reached more than 50 mm. The force resistance only dropped by about 15%. At the end of the test, all the tears extended to be more than 60 mm. The tension boundary column failed due to the rupture of the weld at the bottom of the column, as shown in Figure 4. Meanwhile, the compression column experienced local buckling in the flange, and only a slight out-of-plane deformation was observed.

    In the case of specimen B, i.e., under 600 kN vertical load, no buckling or yielding was detected. Prior to reaching the yield displacement, four load cycles with increasing magnitude from ±50 kN to ±200 kN were necessary to cause the first yielding. At the cyclic load of 150 kN, the first loud bang occurred. These noises also occurred during the unloading phase in the following cycles. However, tension strips first only occurred during the second cycle of the 2 mm level, as indicated by the diagonal lines in Figure 5. The first tear was detected at the upper left corner at the end of the third cycle of 4 mm. The length of the tear was about 10 mm, and this tear gradually grew in the following cycles. At the first 6 mm displacement cycle, new tears of about 15 mm length occurred at the upper right and lower left corners. The length of the tear at the upper left corner extended to about 20 mm. At the end of the second 10 mm displacement cycle, a slight buckling at the support of the boundary column under compression was observed. The length of tear at the upper right corner extended to about 50 mm. At the second 12 mm displacement cycle, as shown in Figure 6, the upper left tear extended to nearly 60 mm, and the shear force resistance did not decrease. During the first 16 mm displacement pull load, the specimen reached a maximum resistance at the top displacement of 13 mm. The shear resistance then began to decrease. At the second 16 mm displacement cycle, the shear resistance decreased rapidly. The failure started at the support of the column under compression where a significant buckling and yield occurred. An out-of-plane deformation increased very fast, resulting in a loss of the in-plane shear resistance.

    In the case of specimen C, soft noises occurred during the application of the vertical load of 900 kN, even prior to the horizontal load. However, no buckling and yielding were detected. Prior to reaching the yield displacement, three load cycles with increasing magnitude from ±50 kN to ±150 kN were necessary to cause the first yielding. The first loud bang occurred at the cycle of 150 kN. The tension strips occurred while pushing. After unloading, a large residual deformation was observed. At the first 6 mm displacement cycle, a first vertical tear occurred at the upper left corner. In the following load cycle, a new tear occurred at the upper right corner at the compression direction. At the first 8 mm displacement cycle, local buckling of the compression column appeared. The shear resistance did not decrease. At the second 8 mm displacement cycle, the tear at the upper left corner extended to about 40 mm. The global out-of-plane buckling occurred, and the shear resistance of the specimen decreased rapidly to less than 100 kN. Specimen C failed because of the global out-of-plane buckling, as shown in Figure 7.

    In the case of specimen D, a soft noise occurred when the vertical load reached 900 kN, and no buckling and yielding were observed in the specimen. The vertical load gradually increased to 1200 kN. The noises increased, but there was no loud bang. Slight horizontal waves due to buckling of the infill tower plate occurred. The first bang occurred at the 50 kN horizontal load cycle. These noises occurred several times during each cycle in the following load cycles. After the 150 kN load cycle, a displacement loading was applied. The shear resistance capacity was stable at 2 mm and 4 mm displacement cycles, and there were no tear. At the 6 mm displacement cycles, the shear resistance began to decrease during the pulling. The reason was the anchorage of the column under compression failed (Figure 8(a)). The failure mode is shown in Figure 8(b).