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).