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Testing of steel plate shear walls with composite columns and infill plates connected to beams only
Engineering Structures 136 (2017) 165–179
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Testing of steel plate shear walls with composite columns and infill plates connected to beams only Lanhui Guo a, Qin Rong b, Bing Qu c,⇑, Jiepeng Liu d a
School of Civil Engineering, Harbin Institute of Technology, Harbin 150090, China College of Civil Engineering and Architecture, Harbin University of Science and Technology, Harbin 150080, China c Dept. of Civil and Environmental Engineering, California Polytechnic State University, San Luis Obispo, CA 93407, USA d School of Civil Engineering, Chongqing University, Chongqing 00044, China b
a r t i c l e
i n f o
Article history: Received 4 February 2016 Revised 25 August 2016 Accepted 11 January 2017 Available online 20 January 2017 Keywords: Test Steel plate shear walls Buckling restrained Composite columns Seismic performance
a b s t r a c t This paper focuses on a new type of Steel Plate Shear Walls (SPSWs) consisting of Concrete Filled Steel Tube (CFST) columns and infill plates connected to beams only. To evaluate seismic behavior of such a SPSW system, this research team conducted an experimental investigation on three two-story specimens. The specimens were varied to compare performances of the systems consisting of conventional infill plates permitting buckling and infill plates restrained by reinforced concrete panels for prevention of plate buckling. Test results revealed that a system with buckling restrained infill plates tends to have a higher strength but lower ductility than the one having conventional infill plates with similar geometries and material properties. Moreover, hysteretic energy dissipation feature remains the same in the specimens with buckling restrained infill plates regardless of infill plate dimensions and plate edge boundary conditions. Overall, the new SPSWs with conventional infill plates and buckling restrained infill plates both exhibited stable hysteretic responses and acceptable ductility under cyclic loading; hence, they are recommended for future practice. Ó 2017 Elsevier Ltd. All rights reserved.
1. Introduction Steel plate shear walls (SPSWs) are lateral force resisting systems that can be used in regions with moderate and high seismicity [1]. Conventional SPSWs consist of steel infill plates with all edges connected to boundary frame members made of wideflange shapes [1,2]. The infill plates are allowed to buckle in shear and subsequently develop diagonal tension field actions during an earthquake event. Although an SPSW can benefit from the boundary frame moment resisting action if any, its lateral force resistance and hysteretic energy dissipation capacity are primarily achieved through the infill plate tension field actions [3,4]. Over the past decades, investigators in the field have developed improved knowledge in understanding seismic behavior of SPSWs [1,3–23]. In addition, innovative strategies have been proposed and experimentally validated for enhancing seismic performance of SPSWs [24–38]. Currently, seismic design guidelines for conventional SPSWs are available in many countries such as the United States, Canada and China among many others [2,39,40]. Nevertheless, some design challenges still exist, impeding the widespread ⇑ Corresponding author. E-mail address: [email protected] (B. Qu). http://dx.doi.org/10.1016/j.engstruct.2017.01.027 0141-0296/Ó 2017 Elsevier Ltd. All rights reserved.
adoptions of SPSWs in practice. One of the most pressing concerns remaining in the field is design of SPSW columns, particularly the columns in mid-rise and high-rise SPSWs. To achieve the favorable sway mechanism in an SPSW subjected to a severe earthquake, the boundary frame members should be designed using the capacity design approach, i.e., sizing the boundary frame members based upon the tension field actions due to infill plate yielding. Consequently, the columns, particularly those from the lower stories of a multi-story SPSW, cumulate the infill plate yielding forces from upper stories. When the number of stories increases, demands on SPSW columns can become impractically large, resulting in significant challenges in column design. For a possible solution to the above issue, this research team focused on a new type of SPSW system. The new SPSW system consists of a composite boundary frame in which the Concrete Filled Steel Tubes (CFSTs) are used as columns. CFSTs, which have been long recognized as better options for resisting combined compression force and bending moment than conventional wide flange members [41–44], can be uniquely helpful to meet the stiffness and strength requirements typically seen in SPSW column designs. It should be recognized that the concrete infill in a CFST column does not contribute significantly to its tensile capacity and CFST
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columns should be used with caution in the SPSWs subjected to large overturning moments. Additionally, infill plates in the new SPSW system are only connected to beams of the boundary frame. This type of infill plate boundary condition can effectively reduce amplitude of infill plate tension field action [24,26,27,32] and more importantly avoid transferring the infill plate forces directly to columns. Furthermore, conventional infill plates typically exhibit large out-of-plane deformations when buckled. Buckling of the infill plates is aesthetically unfavorable and more importantly reduces compressive strength of infill plates and amplifies plate local strain which in turn triggers tear fractures of infill plates. Recent research has demonstrated that installation of reinforced concrete panels on each side of an infill plate can effectively prevent infill plate buckling. However, a knowledge gap exists for the influences of reinforced concrete panels for restraining infill plate buckling on other aspects of the new SPSW system. As such, the objective of this research was to evaluate seismic behavior of the new SPSW system consisting of CFST columns and infill plates connected to beams only. Through testing a group of specimens, this research team investigated damage progression, failure mechanism, stiffness and strength degradations, and energy absorption features of the new SPSW systems with conventional and buckling restrained infill plates, respectively. This research contributes to the field in that it provides necessary testing data (which is missing in the field) for the new SPSW systems. The test results can form a basis for better understanding fundamental behavior of the new SPSW and help promote its applications in future building constructions. The following sections describe in detail specimen design and construction, material properties, test setup, loading program, and test results.
2. Specimen design and construction The research team tested a group of three specimens designated as Specimens A to C, respectively. Specimen A represented the sys-
tem with conventional steel plate infills while Specimens B and C represented the systems with buckling restrained infill plates. Strengths, thicknesses, widths and lengths of the infill plates of these specimens were selected based upon the available capacities of the actuators and space of the laboratory. The boundary frames members of these specimens were conservatively sized based upon the strength and stiffness requirements for the boundary frame members in conventional SPSWs [2]. Also for better comparison purpose, it was determined to choose the same design for boundary frames of all the specimens. As shown in Fig. 1, each frame consisted of two stories with the same story height. Story height and bay width, H and L, were 1500 mm and 2000 mm, respectively. The beams at the top and the bottom of the frame adopted H 300 150 6 9 sections while the intermediate beam used an H 194 150 6 9 section. Here, the Chinese designation of H Shapes corresponds to the United States designation of W shapes and it reflects member depth, width, web thickness and flange thickness (unit in mm), respectively. The columns were CFSTs. Outer diameter and thickness of the steel tubes were 219 mm and 4 mm, respectively. The beam-to-column connections were designed to have moment resistance based upon the Technical Specification for Design and Construction of Concrete-Filled Steel Tubular Structures [45]. Fig. 2 illustrates details of the beam-tocolumn connections. As shown, at each end of the beam, each beam flange was connected to the adjacent column via a 10 mmthick steel enforced loop. Moreover, beam web was connected to the column through complete joint penetration welds. Table 1 lists dimensions of the infill plates in Specimens A to C. As shown, both stories of Specimen A adopted the same infill plates with the thickness and width, t and L1, of 3 mm and 1100 mm, respectively. Each steel plate was attached to beams of the boundary frame through 6 mm thick fishplates. Length of each fishplate was the same as that of the infill plate. Fillet welds were used to connect the infill plate to the fishplates. Net height of each infill plate in Specimen A measured from the welds along the fish plates, H1, was 1053 mm. As summarized in Table 1, height-to-thickness ratio and width-to-height ratio of the infill plates in Specimen A are 350 and 1.05, respectively.
Second story
Second story
First story
First story
6mm thickness steel plate
6mm thickness steel plate
unit: mm
(a) Specimens A and B Fig. 1. Geometries of tested specimens.
(b) Specimen C
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Enforced loop
134
141
CFST column
Enforced loop CFST 80
275
(a) Plane view (unit:mm)
(b) 3-D view
Fig. 2. Details of steel-beam-to-CFST-column connection.
Table 1 Characteristics of specimens.
a
Specimen
H (mm)
L (mm)
L1 (mm)
t (mm)
L1/H1
H1/t
t1 (mm)
A B C
1500 1500 1500
2000 2000 2000
1100 1100 1560
3 3 2
1.05 1.05 1.50
350 350 525
–a 50 50
Not applicable.
As compared in Table 1, infill plates used in Specimen B were identical to those used in Specimen A except that each infill plate of Specimen B was restrained by a pair of 50 mm-thick reinforced concrete panels to mitigate its out-of-plane buckling deformation. Note in Table 1 that t1 represents thickness of the reinforced concrete panels. The infill-plate-to-beam connection is shown in Fig. 3. The infill plates of Specimen C were the same as those used in Specimen B except that each infill plate of Specimen C had width and thickness of 1560 mm and 2 mm, respectively. As a result, Specimen C had the infill plate height-to-thickness ratio and width-to-height ratio of 525 and 1.50, respectively. Similar to Specimen B, each infill plate in Specimen C was also restrained
by a pair of concrete panels. Each infill plate and its restraining concrete panels in Specimens B and C were fastened via throughthickness bolts with the diameter of 10 mm. For the ease of fabrication, diameter of the bolt holes was determined to be 16 mm. Fig. 4 compares distribution of bolt holes in steel plate infills in Specimens B and C, respectively. Fig. 5 shows rebar distributions in the concrete panels of Specimens B and C, respectively. As shown, rebars with the diameter of 10 mm were used to reinforce the concrete panels. The reinforcement ratios along the horizontal and vertical directions were both kept at 0.79%. During the construction phase, the column tubes were first erected and then plain concrete was filled in the tubes. Vibrators were used to consolidate the concrete. Meanwhile, the concrete panels for mitigation of infill plate buckling were cast and cured. Next, steel beams were welded to the CFST columns through the steel enforced loops followed by installation of steel plate infills. Last, the concrete panels for restraining infill plate buckling were installed in Specimens B and C.
3. Material properties Material properties of the concrete and steel elements used in the specimens were evaluated through tests of material samples. For concrete, tests were conducted on standard cubes (100 mm 100 mm 100 mm). Based upon the test procedure, data recording and result processing requirements in the Standard for Test Method of Mechanical Properties on Ordinary Concrete [46], the compressive strength of concrete, fc,cube, was determined to be 35.8 MPa. For steel, coupons were prepared and tested according to the Chinese National Standard GB/T 228.1-2010 [47]. Table 2 presents yield strength, fy, ultimate strength, fu, modulus of elasticity, Es and the Poisson’s ratio, m, of each type of steel.
4. Test setup loading scheme and instrumentation
Fig. 3. Infill-plate-to-beam connections in Specimens B and C.
Fig. 6 shows the test setup used in this investigation. As shown, each specimen was fixed to a base beam attached to the strong floor. In-plane lateral loads were applied to the specimen through actuators at the top and intermediate floor levels. Moreover, the
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Infill plate Intermediate Beam
Infill plate Bolt hole Fish plate
Column
Bottom Beam
(a) Specimen B Infil plate Intermediate Beam
Infill plate Column
Bolt hole
Fish plate Bottom Beam
(b) Specimen C Fig. 4. Distribution of bolt holes in infill plates and reinforced concrete panels (unit: mm).
top beam of each specimen was restrained along the out-of-plane direction to avoid frame out-of-plane deflection. The in-plane lateral loads were applied according to the Specification of Testing Methods for Earthquake Resistant Building (JGJ 10196) [48]. Specifically, at the initial elastic stage, a force-controlled loading scheme was adopted and the lateral loads from the two actuators were kept identical. With this loading scheme, the base shear was twice the shear at the second story. Once specimen yielding was identified from the base shear versus bottom-story drift curve, the force-controlled loading scheme was replaced with a displacement-controlled loading scheme (in which the bottom story was loaded to a target drift at each step while the actuators at both floor levels were controlled to have the same force output). Each load step of the displacement-controlled loading scheme consisted of two load cycles that had the same peak deformation amplitude. Increment of the peak deformation amplitude from one load step to the next load step was kept as the bottom story yield displacement, Dy (see Table 3). Each specimen was loaded until a base shear degradation of 30% was observed or unstable failures of specimen components appeared likely. The response quantities of interest included story shears, specimen deformations and member strains. Story shears were recorded based upon the measured output from each actuator. Linear Variable Displacement Transducers (LVDTs) were used to capture the deformation quantities. Taking Specimen A for example, Fig. 7 illustrates distribution of the LVDTs. As shown, five LVDTs were arranged along centerlines of all the beams to record horizon-
tal deformations of the specimen and sliding deformation of the base beam relative to the strong floor. In addition, four LVDTs were arranged along the column centerlines to monitor the vertical deformations of the specimen. Furthermore, six LVDTs were used to record the out-of-plane buckling deformation of the infill plates at each story. Fig. 8 depicts distribution of the uniaxial strain gauges and strain rosettes in Specimen A. As shown, both boundary frame and infill plates were instrumented for the strains at selected locations. Boundary frames of Specimens B and C retained the same instrumentations as these of Specimen A; however the infill plates of these two specimens were not instrumented. 5. Test results Tests of all the specimens were conducted as planned. This section reports observed specimen behavior and failure modes followed by interpretation of the recorded test data reflecting global behavior of the specimens. Note that the specimens with infill plates connected to beams only conceptually have reduced demands on their columns. However, quantifying the demand reduction is beyond the scope of this research. It is important to recognize that quantitative information regarding the reduction of column demands obtained from the tested specimens (with identical infill plates from both stories and tested with the assumed seismic force distributions) may not remain the same in other SPSWs with varying infill plate thickness up to the height and subjected to real seismic excitations.
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2
13 5
Hook
13 5 Hook
5
5
5
5
5 5
(a) Specimen B 13 5
Hook
13 5
5
5
5
5
5
5
(b) Specimen C Fig. 5. Reinforcement arrangement in reinforced concrete panels (unit: mm).
Table 2 Steel material properties. Category
Thickness (mm)
fy (N/mm2)
fu (N/mm2)
Es (105 N/mm2)
m
Infill plate
2 3 6 4 9 6
230 290 296 300 275 278
330 421 400 381 426 420
2.13 2.03 1.98 2.08 2.12 2.12
0.29 0.31 0.29 0.31 0.28 0.28
Fishplate Column tube Beam flange Beam web
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Load actuator Reaction wall
Load actuator
Specimen
Base beam
Fig. 6. Test setup (specimen out-of-plane restraints are not shown).
Table 3 Summary of peak responses. Specimen
A B C CSPSW
Maximum base shear (kN)
Drift corresponding to first yielding in system (%)
Bottom-story drift corresponding to maximum base shear (%)
Ductility
Force corresponding to maximum bottomstory drift (kN)
Maximum recorded bottom-story drift (%)
Pull
Push
Pull
Push
Pull
Push
Pull
Push
Pull
Push
Pull
Push
0.64 0.70 0.68 0.70
0.49 0.69 0.61 0.66
573.7 716.2 765.7 1126
581.9 706.9 735.9 1166
2.23 1.45 1.18 1.72
2.18 1.45 1.37 1.67
3.92 2.22 1.87 2.04
3.57 2.08 1.82 1.96
403. 565.3 724.8 1079.8
420.7 564.1 656.0 1029.7
6.12 3.17 2.68 2.91
7.28 3.01 2.98 2.97
LVDT-1 LVDT-8
LVDT-9 LVDT-10 LVDT-10~14
LVDT-15
LVDT-2
B3
B1 B4
LVDT-16
B5 C1 C2 C3
LVDT-17~20
W1 W2
LVDT-21
W3
LVDT-3
LVDT-7
LVDT-4
LVDT-5
LVDT-6 Fig. 7. Distribution of LVDTs.
uniaxial strain gauges strain rosette
Fig. 8. Distribution of strain gauges and strain rosettes.
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(d)
(f)
(b)
(a)
(e) (c)
(a)
(b)
(e)
(f)
(e) damage in infill-plate-tofishplate connection
(f) deformation at beam end
(d)
(a) to (d) damages at column ends.
(c)
Fig. 9. Damages in Specimen A under the bottom-story drift of 3%.
5.1. General behavior and failure mode The loading scheme described in Section 4 together with the fact that identical infill plates were used in both stories of each
specimen lead to a higher degree of damage and more severe nonlinear behavior at the bottom story of each specimen. Therefore, focusing on the bottom stories, the following presents the noteworthy observations during the tests.
Rupture
Rupture
Fig. 10. Rupture in the infill-plate-to-fishplate connections in Specimen A under the bottom-story drift of 3.92%.
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For the system with conventional infill plates (i.e., Specimen A), specimen deflection increased linearly with the base shear at the early stage of the test. When the bottom story of the specimen reached the drift of 0.64% in the pull loading cycle, the specimen begun to exhibit nonlinear response. The corresponding base shear was 429.1 kN. Beyond the bottom-story drift of 0.64%, infill plate out-of-plane deformation quickly developed particularly at each corner of the infill plate. At the bottom-story drift of 1%, infill plate buckling became more pronounced. The maximum recorded infill plate out-of-plane deformation reached 24 mm. Meanwhile, yielding was observed at the ends of the intermediate beam. At the bottom-story drift of 1.3%, the maximum infill plate out-of-plane deformations reached 33.8 mm and 25.8 mm at the infill plate corners and center, respectively. Moreover, the infill plate formed a diagonal tension field and noises were heard when orientation of the diagonal tension field changed due to reversal of the lateral load. At the drift of 2.18% of the push loading cycle, the specimen reached its maximum base shear of 581.9 kN. Further increasing the bottom-story drift to 2.7%, a 15 mm-long rupture developed along the weld connecting the bottom-story infill plate to fishplate. Meanwhile, base shear of the specimen begun to degrade and the maximum infill plate out-of-plane deformation reached 47.7 mm. At the bottom-story drift of 3%, the rupture in the welded infillplate-to-fishplate connection became longer and local buckling was observed at the ends of each CFST column (see Fig. 9). Local buckling at the ends of the steel tubes also triggered formation of ruptures at the same location. At the bottom-story drift of 3.92% of the pull loading cycle, rupture in the welded infill-plate-tofishplate connection grew to 200 mm long (see Fig. 10). Moreover, steel tubes at the ends of the columns ruptured severely as shown in Fig. 11. Consequently, the base shear associated with this drift was degraded to 403 kN (approximately 70% of the maximum base shear of 581.9 kN). It was deemed to be the ultimate state of the specimen and the test was therefore concluded. In comparison with Specimen A, Specimens B and C had reinforced concrete panels installed on each side of the infill plates to mitigate the infill plate out-of-plane buckling deformations. Although infill plates of Specimens B and C had different thicknesses and widths, the two specimens exhibited similar behaviors during the tests. Taking Specimen B for example, the following presents some notable observations. Specimen B exhibited linear elastic behavior up to the bottom-story drift of 0.7% under the pull force. Base shear of the specimen at that drift was 591.4 kN. At the bottom-story drift of 1%, strain data recorded at beam ends suggested occurrence of yielding in web of the intermediate beam.
Moreover, noises were heard due to friction and slip between the infill plates and the reinforced concrete panels. When the bottom-story drift increased to 1.45% of the pull loading cycle, base shear of the specimen reached its maximum of 716.2 kN. Up to the bottom-story drift of 1.9%, cracks initiated from the corners of the reinforced concrete panels and progressively extended to the areas around bolt holes. However, no visible out-of-plane deformations were observed in the infill plates. In addition, ruptures were found at the ends of CFST columns (in the proximity of the welds between steel tubes and the enforced loops) and along the welds connecting the infill plates to fishplates. These ruptures reduced base shear of the specimen to 680 kN. Fig. 12 shows the damages in the specimen at that bottom-story drift. When the bottomstory drift reached 2.22% of the pull loading cycle, cracks in the reinforced concreted panels extended diagonally and maximum width of the cracks reached about 0.5 mm. Fall-off failures of the reinforced concrete panels were judged to be likely and then the test was concluded. Base shear associated with the ultimate drift was degraded to 565.3 kN (approximately 79% of the maximum base shear of 716.2 kN). Fig. 13 further shows the infill plates of both stories after removal of the reinforced concrete panels. As shown, no significant buckling deformations or ruptures occurred in the upper-story infill plate. For the bottom-story infill plate, local buckling was found at the corners where concrete panels were severely damaged. Moreover, ruptures were found along the welded infill-plate-to-fishplate connections. 5.2. Representative strain data and yielding progression The strain histories were recorded at some selected locations for each specimen (see Fig. 8). Focusing on Specimen A, this section discusses strain data recorded at the locations where significant inelastic responses were observed. Specimen A was selected since it was instrumented for the strain data of both boundary frame and infill plates. Note that strain data of Specimens B and C are discussed in detail elsewhere [49]. Based upon the data gleaned from Strain Rosettes B1 and B3 (see Fig. 8), Fig. 14a compares shear strains in the web of the intermediate beam. Note that the data recorded at the peak bottomstory drift of each loading cycle were extracted for comparison. Moreover, the bottom-story drifts due to the pull and push loads exerted by the actuators are assigned positive and negative signs in the figure, respectively. As shown, when the bottom story is pulled up to 2.0%, shear strain at the right end of the intermediate beam is higher. This is because over such a drift range the infill
Fig. 11. Ruptures at column ends in Specimen A under the bottom-story drift of 3.92%.
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173
(b)
(d)
(f)
(c)
(a) (e)
(a) facture at bottom right corner
(b) facture at top right corner
(c) fracture at the base of the left column
fracture fracture
(d) fracture at the top of the left column (e) fracture at the base of the right column
(f) diagonal cracks in the concrete panel
Fig. 12. Damage in Specimen B under the bottom-story drift of 1.9%.
plate tension field at the bottom story was more complete compared with that of the upper story due to the loading scheme described in Section 4, which introduced extra downward loads along the intermediate beam. Moreover, the shear force components respectively due to the boundary frame moment resisting action and the extra downward loads are additive at the right end of the intermediate beam while they are subtractive at the left end of the intermediate beam (see Fig. 15), resulting in a higher shear strain at the right end of the beam. When the bottom story is pulled beyond the drift of 2.0%, shear strain at the right end of the intermediate beam degrades to be lower than that of the left end. This is due to the progressive development of the rupture along the infill-plate-to-fishplate connection around the right end of the intermediate beam (see the top right corner of the bottom-story infill plate in Fig. 10). Similarly, when the specimen
is pushed to relatively small drift levels, shear strain at the left end of the intermediate beam is higher than that at the right end due to the stronger infill plate tension field action at the bottom story. When the specimen is pushed to larger drift levels, shear strain at the right end of the intermediate beam becomes higher and even exceeds the shear yielding. Fig. 14b compares the flange axial strains obtained from Strain Gauges B4 and B5 (see Fig. 8). As shown, when the bottom-story drift increases beyond the yielding limit of the entire system, flange yielding occurs at both instrumented locations due to the combined boundary frame moment resisting action and infill plate forces. Based upon the measurements from Strain Rosettes C1, C2 and C3, Fig. 14c compares the axial strains at the top end of the right column of the bottom story. As shown, axial strain at the outer
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A
B
C
D
(a) infill plate of the bottom story
(c) rupture at corner A
(d) rupture at corner B
(b) infill plate of the upper story
(e) rupture at corner C
(f) rupture at corner D
Fig. 13. Deformation of steel panels after removal of reinforced concrete panels in Specimen B.
fibers tend to be higher than that of the inner fiber. Moreover, the presented strain data reveal that yielding occurred in all the three instrumented locations particularly when the bottom-story drift exceeds 1.5%. When the bottom story is loaded beyond the drift of 2.5% in the pull loading cycle, strains at the outer fibers (i.e., Strain Rosettes C1 and C3) reduce significantly due to the rupture of steel tube (see Fig. 11) which redistributed the concentrated local strains and stresses. Fig. 14d compares shear strains of the bottom-story infill plate. As shown, shear yielding occurred at all the three selected locations (i.e., strain Rosettes W1, W2 and W3, see Fig. 8 for their instrumented locations). It is also found that amplitude and sign of the infill plate shear strain changes at higher drift levels. This is primarily due to the fact that the infill plate buckling mode changes as the drift increases which in turn changes the local shear strains in the infill plate. Based upon the entire strain database, Fig. 16 illustrates yielding progression in the specimen under the yielding base shear and certain bottom-story drifts. Note that the yielded locations are highlighted. As shown, yielding first occurred in the infill plate at the bottom story and then progressed into the intermediate beam and the infill plate at the upper story. When the bottom story reached the drift of 2.0%, yielding was also observed at the ends of the top beam and the ends of the columns at the bottom story.
ever infill plates in Specimen CSPSW were connected to both beams and columns of the boundary frame. Considering that Specimen CSPSW provides reference results for comparing the influence of infill plate boundary conditions on specimen hysteretic behavior and other aspects of system performance, test data of Specimen CSPSW are included in the following result discussions. More detailed information about Specimen CSPSW is available elsewhere [25]. Hysteretic curves of all the specimens are presented in Fig. 17. As depicted, compared with the top stories, the bottom stories of all the specimens consistently sustained larger inelastic loading due to the loading scheme described in Section 4. As such, the following discussions focus on the bottom-story hysteretic responses. As shown, the specimens tested in this investigation all exhibit strength and stiffness degradations from cycle to cycle, particularly in the cycles with larger peak inter-story drifts. However, in-cycle strength degradation which may cause collapse of a system is not observed in these specimens, suggesting that all systems can exhibit stable hysteretic behavior under earthquake loading. Compared with the specimens with buckling restrained infill plates (i.e., Specimens B, C and CSPSW), hysteretic curves of the one with conventional infill plates (i.e., Specimen A) exhibit remarkably pinching due to buckling of the infill plates. Overall, hysteretic curves of Specimens B and C are similar to those of Specimen CSPSW except that more severe strength degradations can be observed in Specimens B and C.
5.3. Discussion of hysteretic curves Prior to this investigation, the research team also tested another specimen with a composite boundary frame (referred to herein as Specimen CSPSW) [25]. Dimensions, materials and loading scheme of Specimen CSPSW were the same as those of Specimen C; how-
5.4. Comparison of maximum responses Envelops of the hysteretic curves, known as backbone curves, are compared for all the specimens in Fig. 18. The maximum base
175
4
4
2
2
με (x103)
με (x103)
L. Guo et al. / Engineering Structures 136 (2017) 165–179
0
-2
0
-2 B1 B3
-4 -4
-2
2
0
B4 B5
-4
4
-4
-2
Inter-story drift (%) shear yielidng strain yielding inter-story drift
4
yielidng strain yielding inter-story drift
(a) Beam web shear strain
(b) Beam flange axial strain
7.5
7.5
5.0
5.0
2.5
2.5
με (x103)
με (x103)
2
0
Inter-story drift (%)
0.0
0.0 -2.5
-2.5 C1 C2 C3
-5.0
W1 W2 W3
-5.0 -7.5
-7.5 -4
-2
0
2
4
-4
-2
2
4
shear yielidng strain yielding inter-story drift
yielidng strain yielding inter-story drift
(c) Column axial strain
0
Inter-story drift (%)
Inter-story drift (%)
(d) Inifll plate shear strain
Fig. 14. Strain data from Specimen A.
shear and the corresponding bottom-story drift, and the maximum bottom-story drift and the corresponding base shear identified from the backbone curves are summarized in Table 3 together with the yielding drift and ductility of each system. Comparing the maximum base shear resistances of Specimen A (573.7 kN and 581.9 kN under the pull and push cycles, respectively) with those of Specimen B (716.2 kN and 706.9 kN under the pull and push cycles, respectively), it is found that the reinforced concrete panels can increase strength of a system up to 24.8%. Note that the observed strength increase in Specimen B was due to two factors: prevention of infill plate buckling and friction between the infill plates and concrete panels. However, the respective contributions of these two factors were not explicitly measured. Additionally, it is found that ductility capacity of the SPSW system may reduce up to 59% when adopting the buckling restrained infill plates (7.28 of Specimen A versus. 3.01 of Specimen B, see Table 3). Note that test of Specimen B was concluded essentially due to the potential fall-off failures of the concrete panels. Therefore, enhancing flexural strength of the concrete panels may help improve ductility of the system with bucking restrained infill plates.
Comparison of the results from Specimens C and CSPSW reveal that connecting the infill plates to beams only, although help reduce demands on columns, may cause a lower system strength and more severe strength degradations. The lower system strength observed in Specimen C is primarily due to the fact that the infill plates in Specimen C failed to form the tension field actions as complete and uniform as those of Specimen CSPSW. The more severe strength degradations in Specimen C were primarily due to ruptures at the ends of the infill-plate-to-fishplate connections. This observation indicates that free edges of an infill plate can cause stress concentration and amplified local strain at the ends of the infill-plate-to-fishplate connection which can trigger formation of ruptures in the connections. As listed in Table 1, infill plates in Specimens B and C had different thicknesses and widths; however, cross-section areas of their infill plates (i.e., product of thickness and width) were similar (3300 mm2 from Specimen B vs. 3120 mm2 from Specimen C). Although differences in infill plate width and thickness can cause different infill plate buckling modes and post-buckling strengths, the reinforced concrete panels enabled the two specimens to exhibit comparable ultimate strength and ductility as shown in Table 3.
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B1
ki = equivalent secant stiffness of the system under the ith loading step;
B3
and V push = base shear of the system under the pull and V pull i i push loads of the ith loading step, respectively; and Unbalanced vertical component of infill plate forces
Reactions at left end
Dpull and Dpush = inter-story displacement of the bottom story of i i the system under the pull and push loads of the ith loading step, respectively. Fig. 19 compares the equivalent secant stiffness of each considered system. As shown, at the small drift with elastic response, Specimens A, B and C exhibited the same equivalent secant stiffness, indicating that the concrete panels do not seem to help improve initial stiffness of the system. However, as the drift progressively increases up to 2%, Specimen A appears to have a lower equivalent secant stiffness value due to buckling of the infill plates. Beyond the drift of 2%, Specimens B and C end up with the same secant stiffness as that of Specimen A due to the cracks in the reinforced concrete panels. Overall, Specimens B and C exhibited very similar equivalent secant stiffness, suggesting that the infill plates with similar cross-section areas tend to exhibit the same lateral stiffness when the concrete panels are used to prevent infill plate buckling. Moreover, it is found that the secant stiffness of Specimen CSPSW is consistently higher than the other specimens. Specifically, at the bottomstory drift of 1.75%, secant stiffness of Specimen C is about 61% of that of Specimen CSPSW.
Reactions at right end
(a) shear strains at B1 and B3 due to infill plate forces B1
B3
Rotation at left end
Rotation at right end
(b) shear strains at B1 and B3 due to frame sway-pull loading cycle B1
B3
Rotation at left end
Rotation at right end
(c) shear strains at B1 and B3 due to frame sway-push loading cycle Fig. 15. Illustration of shear strains at the web of intermediate beam.
Another important observation from Table 3 is that ductility of the systems with buckling restrained infill plates overall have similar ductility factors on the order of 3.0 regardless of infill plate geometries and boundary conditions along plate edge. Such ductility factors are lower than that of the system with conventional infill plates permitting buckling.
5.5. Lateral stiffness degradation As shown in Fig. 17, stiffness of each system (i.e., slope of the hysteretic curve) varies at different drift levels. For a more convenient comparison of stiffness degradation as drift in a system increases, the following equivalent secant stiffness is defined.
pull push V i þ V i ki ¼ push pull Di þ Di
ð1Þ
where
4 29 kN
1% drift
5.6. Energy absorption features The infill plates installed in the composite frames were intended to dissipate the earthquake energy released into the systems. Specimens A and B provided a unique opportunity to compare the energy absorption features of the infill plates without and with plate buckling restraining actions. Fig. 20 compares the hysteretic loops of the two specimens at a selected load step. Note that the area enclosed by a hysteretic loop corresponds to the energy absorbed in the system during the load cycles. As shown, when the peak drifts remain about the same, Specimen B absorbs more energy than Specimen A due to the beneficial contribution of reinforced concrete panels to prevention of infill plate buckling. Note that the area enclosed by a hysteretic loop is affected by strength of the system and the peak deformation it achieves. To eliminate the influences of these two parameters and achieve direct comparisons among all the considered systems, the energy absorption coefficient, E, defined in the Code for Seismic Design of
1.33% drift
1.5% drift
Fig. 16. Yielding progression in Specimen A.
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800
Specimen B 2nd story
600
Story shear (kN)
Story shear (kN)
800
Specimen A 2nd story
600 400 200 0 -200 -400
400 200 0 -200 -400 -600
-600
-800
-800 -4
-2
4
2
0
-3
-2
800
400 200 0 -200 -400
2
3
400 200 0 -200 -400
-800
-800 -4
-2
4
2
0
-3
-2
800
1200
Story shear (kN)
Specimen C 2nd story
600
-1
0
1
2
3
Inter-story drift (%)
Inter-story drift (%)
Story shear (kN)
1
-600
-600
400 200 0 -200 -400
800
Specimen CSPSW 2nd story
400 0 -400 -800
-600 -800
-1200 -3
-2
-1
0
1
2
-3
3
-2
Inter-story drift (%) 800
-1
0
1
2
3
2
3
Inter-story drift (%) 1200
Story shear (kN)
Specimen C 1st story
600
Story shear (kN)
0
Specimen B 1st story
600
Story shear (kN)
Story shear (kN)
800
Specimen A 1st story
600
-1
Inter-story drift (%)
Inter-story drift (%)
400 200 0 -200 -400
800
Specimen CSPSW 1st story
400 0 -400 -800
-600
-1200
-800 -3
-2
-1
0
1
2
3
-3
-2
-1
0
1
Inter-story drift (%)
Inter-story drift (%) Fig. 17. Comparison of hysteretic curves.
Buildings (GB50011-2010) [40] was calculated for each specimen. Fig. 21 schematically shows the definition of E. As illustrated, E represents ratio of the energy absorbed by the system in a loading cycle (i.e., the area enclosed by a hysteretic loop, which is calculated as the sum of SABC and SCDA, to the sum of the areas enclosed by the two triangles, SOBE and SODF). Based upon this definition, a system with a larger value of E tends to have a better energy absorption capacity. Fig. 22 compares the energy absorption coefficient of each considered system. As shown, the energy absorption coefficient of
Specimen A is consistently lower than the other specimens over the drift range of interest, indicating that the energy absorption capacity of conventional infill plates is inferior to that of the buckling restrained infill plates. Another important observation from Fig. 22 is that all the specimens with the buckling restrained infill plates exhibit almost identical trend in energy absorption coefficient as the drift increases. This indicates that geometries and boundary conditions of a buckling restrained infill plate, although influence system strength and stiffness, do not affect its favorable feature in absorbing hysteretic energy.
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V B
F
D
A
O
C
E
E=
S ABC +S CDA SOBE +S ODF
Fig. 21. Definition of energy absorption coefficient. Fig. 18. Comparison of backbone curves.
Fig. 22. Comparison of energy absorption coefficient. Fig. 19. Comparison of equivalent secant stiffness values.
800 600
Story shear (kN)
400 200 0 -200 -400 Specimen A -600 -800
Specimen B -2
-1
0
1
2
Bottom-story drift (%) Fig. 20. Comparison of hysteretic energy absorption: Specimen A versus Specimen B.
6. Conclusions Based upon the test results obtained from this investigation, the following significant conclusions were drawn:
Overall, the SPSWs consisting of CFST columns and infill plates connected to beams only exhibited stable responses and acceptable ductility under cyclic loading. Therefore, such systems can be used in seismic design of multi-story buildings. When connected to beams only, a buckling restrained infill plate tends to offer a higher strength but lower ductility than a conventional infill plate with similar geometry and material property. Nevertheless, the systems with buckling restrained infill plates were able to achieve the ultimate drift on the order of 2% and the ductility of 3.0. Moreover, stiffness degradations in the systems with buckling restrained infill plates are not as severe as the one with conventional infill plates. Therefore, the buckling restrained infill plates are recommended for future practice. Failures of the systems consisting of buckling restrained infill plates are due to severe cracks and possible fall-off failures of the concrete panels. Therefore, increasing flexural strength of the concrete panels may help enhance ductility of the system and expand its range of applicability. Hysteretic energy dissipation feature remains the same in all the specimens with buckling restrained infill plates regardless of infill plate dimensions and plate edge boundary conditions. Inelastic behavior may occur in the CFST columns when the new SPSW system is loaded to large inter-story drifts. Specifically, the ends of each CFST column (in the proximity of beam-toCFST-column connections) may exhibit steel tube local buckling
L. Guo et al. / Engineering Structures 136 (2017) 165–179
followed by ruptures. However, this may be due to the loading scheme adopted in this experimental investigation. Future research opportunities exist to develop improved design models for the CFST columns and better detailing requirement for the beam-to-CFST-column connections.
Acknowledgements This research was financially supported by the National Natural Science Foundation of China (Awards Nos. 51578187 and 50808053) and the National key research and development program (No. 2016YFC0701201). The third author was also supported by the Tom and Lucia Chou Fund. The authors wish to acknowledge the sponsors. However, any opinions, findings, conclusions and recommendations presented in this paper are those of the authors and do not necessarily reflect the views of the sponsors. References [1] Bruneau M, Uang CM, Sabelli R. Ductile design of steel structures. McGraw Hill; 2011. [2] AISC. Seismic provisions for structural steel buildings. ANSI/AISC 341–10. American Institute of Steel Construction; 2010. [3] Qu B, Bruneau M. Design of steel plate shear walls considering boundary frame moment resisting action. ASCE J Struct Eng 2009;135:1511–21. [4] Qu B, Bruneau M. Behavior of vertical boundary elements in steel plate shear walls. AISC Eng J 2010;47:109–22. [5] Berman JW, Bruneau M. Plastic analysis and design of steel plate shear walls. ASCE J Struct Eng 2003;129:1448–56. [6] Berman JW, Bruneau M. Experimental investigation of light-gauge steel plate shear walls. ASCE J Struct Eng 2005;131:259–67. [7] Caccese V, Elgaaly M, Chen R. Experimental study of thin steel-plate shear walls under cyclic load. ASCE J Struct Eng 1993;119:573–87. [8] Elgaaly M. Thin steel plate shear walls behavior and analysis. Thin Wall Struct 1998;32:151–80. [9] Elgaaly M, Caccese V, Du C. Postbuckling behavior of steel-plate shear walls under cyclic loads. ASCE J Struct Eng 1993;119:588–605. [10] Guo L, Li R, Zhang S. Hysteretic analysis and modified strip model of steel plate shear walls. Adv Struct Eng 2012;15:1751–64. [11] Lubell AS, Prion HGL, Ventura CE, Rezai M. Unstiffened steel plate shear wall performance under cyclic loading. ASCE J Struct Eng 2000;126:453–60. [12] Montgomery CJ, Medhekar M. Discussion on unstiffened steel plate shear wall performance under cyclic loading. J Struct Eng 2001;127. [13] Qu B, Bruneau M. Seismic behavior and design of boundary frame members in steel plate shear walls. Technical report MCEER-08-0012: Multidisciplinary Center for Earthquake Engineering Research, State University of New York at Buffalo, Buffalo, NY.; 2008. [14] Qu B, Bruneau M. Capacity design of intermediate horizontal boundary elements of steel plate shear walls. ASCE J Struct Eng 2010;136:665–75. [15] Qu B, Bruneau M. Plastic moment of intermediate horizontal boundary elements of steel plate shear walls. AISC Eng J 2011;48:49–64. [16] Qu B, Bruneau M, Lin CH, Tsai KC. Testing of full-scale two-story steel plate shear wall with reduced beam section connections and composite floors. ASCE J Struct Eng 2008;134:364–73. [17] Qu B, Guo X, Chi H, Pollino M. Probabilistic evaluation of effect of column stiffness on seismic performance of steel plate shear walls. Eng Struct 2012;43:169–79. [18] Qu B, Guo X, Pollino M, Chi H. Effect of column stiffness on drift concentration in steel plate shear walls. J Constr Steel Res 2013;83:105–16. [19] Zhao Q, Astaneh-Asl A. Cyclic behavior of traditional and innovative composite shear walls. ASCE J Struct Eng 2004;130:271–84. [20] Purba R, Bruneau M. Experimental investigation of steel plate shear walls with in-span plastification along horizontal boundary elements. Eng Struct 2015;95.
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Contents lists available at ScienceDirect
Engineering Structures journal homepage: www.elsevier.com/locate/engstruct
Testing of steel plate shear walls with composite columns and infill plates connected to beams only Lanhui Guo a, Qin Rong b, Bing Qu c,⇑, Jiepeng Liu d a
School of Civil Engineering, Harbin Institute of Technology, Harbin 150090, China College of Civil Engineering and Architecture, Harbin University of Science and Technology, Harbin 150080, China c Dept. of Civil and Environmental Engineering, California Polytechnic State University, San Luis Obispo, CA 93407, USA d School of Civil Engineering, Chongqing University, Chongqing 00044, China b
a r t i c l e
i n f o
Article history: Received 4 February 2016 Revised 25 August 2016 Accepted 11 January 2017 Available online 20 January 2017 Keywords: Test Steel plate shear walls Buckling restrained Composite columns Seismic performance
a b s t r a c t This paper focuses on a new type of Steel Plate Shear Walls (SPSWs) consisting of Concrete Filled Steel Tube (CFST) columns and infill plates connected to beams only. To evaluate seismic behavior of such a SPSW system, this research team conducted an experimental investigation on three two-story specimens. The specimens were varied to compare performances of the systems consisting of conventional infill plates permitting buckling and infill plates restrained by reinforced concrete panels for prevention of plate buckling. Test results revealed that a system with buckling restrained infill plates tends to have a higher strength but lower ductility than the one having conventional infill plates with similar geometries and material properties. Moreover, hysteretic energy dissipation feature remains the same in the specimens with buckling restrained infill plates regardless of infill plate dimensions and plate edge boundary conditions. Overall, the new SPSWs with conventional infill plates and buckling restrained infill plates both exhibited stable hysteretic responses and acceptable ductility under cyclic loading; hence, they are recommended for future practice. Ó 2017 Elsevier Ltd. All rights reserved.
1. Introduction Steel plate shear walls (SPSWs) are lateral force resisting systems that can be used in regions with moderate and high seismicity [1]. Conventional SPSWs consist of steel infill plates with all edges connected to boundary frame members made of wideflange shapes [1,2]. The infill plates are allowed to buckle in shear and subsequently develop diagonal tension field actions during an earthquake event. Although an SPSW can benefit from the boundary frame moment resisting action if any, its lateral force resistance and hysteretic energy dissipation capacity are primarily achieved through the infill plate tension field actions [3,4]. Over the past decades, investigators in the field have developed improved knowledge in understanding seismic behavior of SPSWs [1,3–23]. In addition, innovative strategies have been proposed and experimentally validated for enhancing seismic performance of SPSWs [24–38]. Currently, seismic design guidelines for conventional SPSWs are available in many countries such as the United States, Canada and China among many others [2,39,40]. Nevertheless, some design challenges still exist, impeding the widespread ⇑ Corresponding author. E-mail address: [email protected] (B. Qu). http://dx.doi.org/10.1016/j.engstruct.2017.01.027 0141-0296/Ó 2017 Elsevier Ltd. All rights reserved.
adoptions of SPSWs in practice. One of the most pressing concerns remaining in the field is design of SPSW columns, particularly the columns in mid-rise and high-rise SPSWs. To achieve the favorable sway mechanism in an SPSW subjected to a severe earthquake, the boundary frame members should be designed using the capacity design approach, i.e., sizing the boundary frame members based upon the tension field actions due to infill plate yielding. Consequently, the columns, particularly those from the lower stories of a multi-story SPSW, cumulate the infill plate yielding forces from upper stories. When the number of stories increases, demands on SPSW columns can become impractically large, resulting in significant challenges in column design. For a possible solution to the above issue, this research team focused on a new type of SPSW system. The new SPSW system consists of a composite boundary frame in which the Concrete Filled Steel Tubes (CFSTs) are used as columns. CFSTs, which have been long recognized as better options for resisting combined compression force and bending moment than conventional wide flange members [41–44], can be uniquely helpful to meet the stiffness and strength requirements typically seen in SPSW column designs. It should be recognized that the concrete infill in a CFST column does not contribute significantly to its tensile capacity and CFST
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columns should be used with caution in the SPSWs subjected to large overturning moments. Additionally, infill plates in the new SPSW system are only connected to beams of the boundary frame. This type of infill plate boundary condition can effectively reduce amplitude of infill plate tension field action [24,26,27,32] and more importantly avoid transferring the infill plate forces directly to columns. Furthermore, conventional infill plates typically exhibit large out-of-plane deformations when buckled. Buckling of the infill plates is aesthetically unfavorable and more importantly reduces compressive strength of infill plates and amplifies plate local strain which in turn triggers tear fractures of infill plates. Recent research has demonstrated that installation of reinforced concrete panels on each side of an infill plate can effectively prevent infill plate buckling. However, a knowledge gap exists for the influences of reinforced concrete panels for restraining infill plate buckling on other aspects of the new SPSW system. As such, the objective of this research was to evaluate seismic behavior of the new SPSW system consisting of CFST columns and infill plates connected to beams only. Through testing a group of specimens, this research team investigated damage progression, failure mechanism, stiffness and strength degradations, and energy absorption features of the new SPSW systems with conventional and buckling restrained infill plates, respectively. This research contributes to the field in that it provides necessary testing data (which is missing in the field) for the new SPSW systems. The test results can form a basis for better understanding fundamental behavior of the new SPSW and help promote its applications in future building constructions. The following sections describe in detail specimen design and construction, material properties, test setup, loading program, and test results.
2. Specimen design and construction The research team tested a group of three specimens designated as Specimens A to C, respectively. Specimen A represented the sys-
tem with conventional steel plate infills while Specimens B and C represented the systems with buckling restrained infill plates. Strengths, thicknesses, widths and lengths of the infill plates of these specimens were selected based upon the available capacities of the actuators and space of the laboratory. The boundary frames members of these specimens were conservatively sized based upon the strength and stiffness requirements for the boundary frame members in conventional SPSWs [2]. Also for better comparison purpose, it was determined to choose the same design for boundary frames of all the specimens. As shown in Fig. 1, each frame consisted of two stories with the same story height. Story height and bay width, H and L, were 1500 mm and 2000 mm, respectively. The beams at the top and the bottom of the frame adopted H 300 150 6 9 sections while the intermediate beam used an H 194 150 6 9 section. Here, the Chinese designation of H Shapes corresponds to the United States designation of W shapes and it reflects member depth, width, web thickness and flange thickness (unit in mm), respectively. The columns were CFSTs. Outer diameter and thickness of the steel tubes were 219 mm and 4 mm, respectively. The beam-to-column connections were designed to have moment resistance based upon the Technical Specification for Design and Construction of Concrete-Filled Steel Tubular Structures [45]. Fig. 2 illustrates details of the beam-tocolumn connections. As shown, at each end of the beam, each beam flange was connected to the adjacent column via a 10 mmthick steel enforced loop. Moreover, beam web was connected to the column through complete joint penetration welds. Table 1 lists dimensions of the infill plates in Specimens A to C. As shown, both stories of Specimen A adopted the same infill plates with the thickness and width, t and L1, of 3 mm and 1100 mm, respectively. Each steel plate was attached to beams of the boundary frame through 6 mm thick fishplates. Length of each fishplate was the same as that of the infill plate. Fillet welds were used to connect the infill plate to the fishplates. Net height of each infill plate in Specimen A measured from the welds along the fish plates, H1, was 1053 mm. As summarized in Table 1, height-to-thickness ratio and width-to-height ratio of the infill plates in Specimen A are 350 and 1.05, respectively.
Second story
Second story
First story
First story
6mm thickness steel plate
6mm thickness steel plate
unit: mm
(a) Specimens A and B Fig. 1. Geometries of tested specimens.
(b) Specimen C
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Enforced loop
134
141
CFST column
Enforced loop CFST 80
275
(a) Plane view (unit:mm)
(b) 3-D view
Fig. 2. Details of steel-beam-to-CFST-column connection.
Table 1 Characteristics of specimens.
a
Specimen
H (mm)
L (mm)
L1 (mm)
t (mm)
L1/H1
H1/t
t1 (mm)
A B C
1500 1500 1500
2000 2000 2000
1100 1100 1560
3 3 2
1.05 1.05 1.50
350 350 525
–a 50 50
Not applicable.
As compared in Table 1, infill plates used in Specimen B were identical to those used in Specimen A except that each infill plate of Specimen B was restrained by a pair of 50 mm-thick reinforced concrete panels to mitigate its out-of-plane buckling deformation. Note in Table 1 that t1 represents thickness of the reinforced concrete panels. The infill-plate-to-beam connection is shown in Fig. 3. The infill plates of Specimen C were the same as those used in Specimen B except that each infill plate of Specimen C had width and thickness of 1560 mm and 2 mm, respectively. As a result, Specimen C had the infill plate height-to-thickness ratio and width-to-height ratio of 525 and 1.50, respectively. Similar to Specimen B, each infill plate in Specimen C was also restrained
by a pair of concrete panels. Each infill plate and its restraining concrete panels in Specimens B and C were fastened via throughthickness bolts with the diameter of 10 mm. For the ease of fabrication, diameter of the bolt holes was determined to be 16 mm. Fig. 4 compares distribution of bolt holes in steel plate infills in Specimens B and C, respectively. Fig. 5 shows rebar distributions in the concrete panels of Specimens B and C, respectively. As shown, rebars with the diameter of 10 mm were used to reinforce the concrete panels. The reinforcement ratios along the horizontal and vertical directions were both kept at 0.79%. During the construction phase, the column tubes were first erected and then plain concrete was filled in the tubes. Vibrators were used to consolidate the concrete. Meanwhile, the concrete panels for mitigation of infill plate buckling were cast and cured. Next, steel beams were welded to the CFST columns through the steel enforced loops followed by installation of steel plate infills. Last, the concrete panels for restraining infill plate buckling were installed in Specimens B and C.
3. Material properties Material properties of the concrete and steel elements used in the specimens were evaluated through tests of material samples. For concrete, tests were conducted on standard cubes (100 mm 100 mm 100 mm). Based upon the test procedure, data recording and result processing requirements in the Standard for Test Method of Mechanical Properties on Ordinary Concrete [46], the compressive strength of concrete, fc,cube, was determined to be 35.8 MPa. For steel, coupons were prepared and tested according to the Chinese National Standard GB/T 228.1-2010 [47]. Table 2 presents yield strength, fy, ultimate strength, fu, modulus of elasticity, Es and the Poisson’s ratio, m, of each type of steel.
4. Test setup loading scheme and instrumentation
Fig. 3. Infill-plate-to-beam connections in Specimens B and C.
Fig. 6 shows the test setup used in this investigation. As shown, each specimen was fixed to a base beam attached to the strong floor. In-plane lateral loads were applied to the specimen through actuators at the top and intermediate floor levels. Moreover, the
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Infill plate Intermediate Beam
Infill plate Bolt hole Fish plate
Column
Bottom Beam
(a) Specimen B Infil plate Intermediate Beam
Infill plate Column
Bolt hole
Fish plate Bottom Beam
(b) Specimen C Fig. 4. Distribution of bolt holes in infill plates and reinforced concrete panels (unit: mm).
top beam of each specimen was restrained along the out-of-plane direction to avoid frame out-of-plane deflection. The in-plane lateral loads were applied according to the Specification of Testing Methods for Earthquake Resistant Building (JGJ 10196) [48]. Specifically, at the initial elastic stage, a force-controlled loading scheme was adopted and the lateral loads from the two actuators were kept identical. With this loading scheme, the base shear was twice the shear at the second story. Once specimen yielding was identified from the base shear versus bottom-story drift curve, the force-controlled loading scheme was replaced with a displacement-controlled loading scheme (in which the bottom story was loaded to a target drift at each step while the actuators at both floor levels were controlled to have the same force output). Each load step of the displacement-controlled loading scheme consisted of two load cycles that had the same peak deformation amplitude. Increment of the peak deformation amplitude from one load step to the next load step was kept as the bottom story yield displacement, Dy (see Table 3). Each specimen was loaded until a base shear degradation of 30% was observed or unstable failures of specimen components appeared likely. The response quantities of interest included story shears, specimen deformations and member strains. Story shears were recorded based upon the measured output from each actuator. Linear Variable Displacement Transducers (LVDTs) were used to capture the deformation quantities. Taking Specimen A for example, Fig. 7 illustrates distribution of the LVDTs. As shown, five LVDTs were arranged along centerlines of all the beams to record horizon-
tal deformations of the specimen and sliding deformation of the base beam relative to the strong floor. In addition, four LVDTs were arranged along the column centerlines to monitor the vertical deformations of the specimen. Furthermore, six LVDTs were used to record the out-of-plane buckling deformation of the infill plates at each story. Fig. 8 depicts distribution of the uniaxial strain gauges and strain rosettes in Specimen A. As shown, both boundary frame and infill plates were instrumented for the strains at selected locations. Boundary frames of Specimens B and C retained the same instrumentations as these of Specimen A; however the infill plates of these two specimens were not instrumented. 5. Test results Tests of all the specimens were conducted as planned. This section reports observed specimen behavior and failure modes followed by interpretation of the recorded test data reflecting global behavior of the specimens. Note that the specimens with infill plates connected to beams only conceptually have reduced demands on their columns. However, quantifying the demand reduction is beyond the scope of this research. It is important to recognize that quantitative information regarding the reduction of column demands obtained from the tested specimens (with identical infill plates from both stories and tested with the assumed seismic force distributions) may not remain the same in other SPSWs with varying infill plate thickness up to the height and subjected to real seismic excitations.
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2
13 5
Hook
13 5 Hook
5
5
5
5
5 5
(a) Specimen B 13 5
Hook
13 5
5
5
5
5
5
5
(b) Specimen C Fig. 5. Reinforcement arrangement in reinforced concrete panels (unit: mm).
Table 2 Steel material properties. Category
Thickness (mm)
fy (N/mm2)
fu (N/mm2)
Es (105 N/mm2)
m
Infill plate
2 3 6 4 9 6
230 290 296 300 275 278
330 421 400 381 426 420
2.13 2.03 1.98 2.08 2.12 2.12
0.29 0.31 0.29 0.31 0.28 0.28
Fishplate Column tube Beam flange Beam web
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Load actuator Reaction wall
Load actuator
Specimen
Base beam
Fig. 6. Test setup (specimen out-of-plane restraints are not shown).
Table 3 Summary of peak responses. Specimen
A B C CSPSW
Maximum base shear (kN)
Drift corresponding to first yielding in system (%)
Bottom-story drift corresponding to maximum base shear (%)
Ductility
Force corresponding to maximum bottomstory drift (kN)
Maximum recorded bottom-story drift (%)
Pull
Push
Pull
Push
Pull
Push
Pull
Push
Pull
Push
Pull
Push
0.64 0.70 0.68 0.70
0.49 0.69 0.61 0.66
573.7 716.2 765.7 1126
581.9 706.9 735.9 1166
2.23 1.45 1.18 1.72
2.18 1.45 1.37 1.67
3.92 2.22 1.87 2.04
3.57 2.08 1.82 1.96
403. 565.3 724.8 1079.8
420.7 564.1 656.0 1029.7
6.12 3.17 2.68 2.91
7.28 3.01 2.98 2.97
LVDT-1 LVDT-8
LVDT-9 LVDT-10 LVDT-10~14
LVDT-15
LVDT-2
B3
B1 B4
LVDT-16
B5 C1 C2 C3
LVDT-17~20
W1 W2
LVDT-21
W3
LVDT-3
LVDT-7
LVDT-4
LVDT-5
LVDT-6 Fig. 7. Distribution of LVDTs.
uniaxial strain gauges strain rosette
Fig. 8. Distribution of strain gauges and strain rosettes.
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(d)
(f)
(b)
(a)
(e) (c)
(a)
(b)
(e)
(f)
(e) damage in infill-plate-tofishplate connection
(f) deformation at beam end
(d)
(a) to (d) damages at column ends.
(c)
Fig. 9. Damages in Specimen A under the bottom-story drift of 3%.
5.1. General behavior and failure mode The loading scheme described in Section 4 together with the fact that identical infill plates were used in both stories of each
specimen lead to a higher degree of damage and more severe nonlinear behavior at the bottom story of each specimen. Therefore, focusing on the bottom stories, the following presents the noteworthy observations during the tests.
Rupture
Rupture
Fig. 10. Rupture in the infill-plate-to-fishplate connections in Specimen A under the bottom-story drift of 3.92%.
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For the system with conventional infill plates (i.e., Specimen A), specimen deflection increased linearly with the base shear at the early stage of the test. When the bottom story of the specimen reached the drift of 0.64% in the pull loading cycle, the specimen begun to exhibit nonlinear response. The corresponding base shear was 429.1 kN. Beyond the bottom-story drift of 0.64%, infill plate out-of-plane deformation quickly developed particularly at each corner of the infill plate. At the bottom-story drift of 1%, infill plate buckling became more pronounced. The maximum recorded infill plate out-of-plane deformation reached 24 mm. Meanwhile, yielding was observed at the ends of the intermediate beam. At the bottom-story drift of 1.3%, the maximum infill plate out-of-plane deformations reached 33.8 mm and 25.8 mm at the infill plate corners and center, respectively. Moreover, the infill plate formed a diagonal tension field and noises were heard when orientation of the diagonal tension field changed due to reversal of the lateral load. At the drift of 2.18% of the push loading cycle, the specimen reached its maximum base shear of 581.9 kN. Further increasing the bottom-story drift to 2.7%, a 15 mm-long rupture developed along the weld connecting the bottom-story infill plate to fishplate. Meanwhile, base shear of the specimen begun to degrade and the maximum infill plate out-of-plane deformation reached 47.7 mm. At the bottom-story drift of 3%, the rupture in the welded infillplate-to-fishplate connection became longer and local buckling was observed at the ends of each CFST column (see Fig. 9). Local buckling at the ends of the steel tubes also triggered formation of ruptures at the same location. At the bottom-story drift of 3.92% of the pull loading cycle, rupture in the welded infill-plate-tofishplate connection grew to 200 mm long (see Fig. 10). Moreover, steel tubes at the ends of the columns ruptured severely as shown in Fig. 11. Consequently, the base shear associated with this drift was degraded to 403 kN (approximately 70% of the maximum base shear of 581.9 kN). It was deemed to be the ultimate state of the specimen and the test was therefore concluded. In comparison with Specimen A, Specimens B and C had reinforced concrete panels installed on each side of the infill plates to mitigate the infill plate out-of-plane buckling deformations. Although infill plates of Specimens B and C had different thicknesses and widths, the two specimens exhibited similar behaviors during the tests. Taking Specimen B for example, the following presents some notable observations. Specimen B exhibited linear elastic behavior up to the bottom-story drift of 0.7% under the pull force. Base shear of the specimen at that drift was 591.4 kN. At the bottom-story drift of 1%, strain data recorded at beam ends suggested occurrence of yielding in web of the intermediate beam.
Moreover, noises were heard due to friction and slip between the infill plates and the reinforced concrete panels. When the bottom-story drift increased to 1.45% of the pull loading cycle, base shear of the specimen reached its maximum of 716.2 kN. Up to the bottom-story drift of 1.9%, cracks initiated from the corners of the reinforced concrete panels and progressively extended to the areas around bolt holes. However, no visible out-of-plane deformations were observed in the infill plates. In addition, ruptures were found at the ends of CFST columns (in the proximity of the welds between steel tubes and the enforced loops) and along the welds connecting the infill plates to fishplates. These ruptures reduced base shear of the specimen to 680 kN. Fig. 12 shows the damages in the specimen at that bottom-story drift. When the bottomstory drift reached 2.22% of the pull loading cycle, cracks in the reinforced concreted panels extended diagonally and maximum width of the cracks reached about 0.5 mm. Fall-off failures of the reinforced concrete panels were judged to be likely and then the test was concluded. Base shear associated with the ultimate drift was degraded to 565.3 kN (approximately 79% of the maximum base shear of 716.2 kN). Fig. 13 further shows the infill plates of both stories after removal of the reinforced concrete panels. As shown, no significant buckling deformations or ruptures occurred in the upper-story infill plate. For the bottom-story infill plate, local buckling was found at the corners where concrete panels were severely damaged. Moreover, ruptures were found along the welded infill-plate-to-fishplate connections. 5.2. Representative strain data and yielding progression The strain histories were recorded at some selected locations for each specimen (see Fig. 8). Focusing on Specimen A, this section discusses strain data recorded at the locations where significant inelastic responses were observed. Specimen A was selected since it was instrumented for the strain data of both boundary frame and infill plates. Note that strain data of Specimens B and C are discussed in detail elsewhere [49]. Based upon the data gleaned from Strain Rosettes B1 and B3 (see Fig. 8), Fig. 14a compares shear strains in the web of the intermediate beam. Note that the data recorded at the peak bottomstory drift of each loading cycle were extracted for comparison. Moreover, the bottom-story drifts due to the pull and push loads exerted by the actuators are assigned positive and negative signs in the figure, respectively. As shown, when the bottom story is pulled up to 2.0%, shear strain at the right end of the intermediate beam is higher. This is because over such a drift range the infill
Fig. 11. Ruptures at column ends in Specimen A under the bottom-story drift of 3.92%.
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(b)
(d)
(f)
(c)
(a) (e)
(a) facture at bottom right corner
(b) facture at top right corner
(c) fracture at the base of the left column
fracture fracture
(d) fracture at the top of the left column (e) fracture at the base of the right column
(f) diagonal cracks in the concrete panel
Fig. 12. Damage in Specimen B under the bottom-story drift of 1.9%.
plate tension field at the bottom story was more complete compared with that of the upper story due to the loading scheme described in Section 4, which introduced extra downward loads along the intermediate beam. Moreover, the shear force components respectively due to the boundary frame moment resisting action and the extra downward loads are additive at the right end of the intermediate beam while they are subtractive at the left end of the intermediate beam (see Fig. 15), resulting in a higher shear strain at the right end of the beam. When the bottom story is pulled beyond the drift of 2.0%, shear strain at the right end of the intermediate beam degrades to be lower than that of the left end. This is due to the progressive development of the rupture along the infill-plate-to-fishplate connection around the right end of the intermediate beam (see the top right corner of the bottom-story infill plate in Fig. 10). Similarly, when the specimen
is pushed to relatively small drift levels, shear strain at the left end of the intermediate beam is higher than that at the right end due to the stronger infill plate tension field action at the bottom story. When the specimen is pushed to larger drift levels, shear strain at the right end of the intermediate beam becomes higher and even exceeds the shear yielding. Fig. 14b compares the flange axial strains obtained from Strain Gauges B4 and B5 (see Fig. 8). As shown, when the bottom-story drift increases beyond the yielding limit of the entire system, flange yielding occurs at both instrumented locations due to the combined boundary frame moment resisting action and infill plate forces. Based upon the measurements from Strain Rosettes C1, C2 and C3, Fig. 14c compares the axial strains at the top end of the right column of the bottom story. As shown, axial strain at the outer
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A
B
C
D
(a) infill plate of the bottom story
(c) rupture at corner A
(d) rupture at corner B
(b) infill plate of the upper story
(e) rupture at corner C
(f) rupture at corner D
Fig. 13. Deformation of steel panels after removal of reinforced concrete panels in Specimen B.
fibers tend to be higher than that of the inner fiber. Moreover, the presented strain data reveal that yielding occurred in all the three instrumented locations particularly when the bottom-story drift exceeds 1.5%. When the bottom story is loaded beyond the drift of 2.5% in the pull loading cycle, strains at the outer fibers (i.e., Strain Rosettes C1 and C3) reduce significantly due to the rupture of steel tube (see Fig. 11) which redistributed the concentrated local strains and stresses. Fig. 14d compares shear strains of the bottom-story infill plate. As shown, shear yielding occurred at all the three selected locations (i.e., strain Rosettes W1, W2 and W3, see Fig. 8 for their instrumented locations). It is also found that amplitude and sign of the infill plate shear strain changes at higher drift levels. This is primarily due to the fact that the infill plate buckling mode changes as the drift increases which in turn changes the local shear strains in the infill plate. Based upon the entire strain database, Fig. 16 illustrates yielding progression in the specimen under the yielding base shear and certain bottom-story drifts. Note that the yielded locations are highlighted. As shown, yielding first occurred in the infill plate at the bottom story and then progressed into the intermediate beam and the infill plate at the upper story. When the bottom story reached the drift of 2.0%, yielding was also observed at the ends of the top beam and the ends of the columns at the bottom story.
ever infill plates in Specimen CSPSW were connected to both beams and columns of the boundary frame. Considering that Specimen CSPSW provides reference results for comparing the influence of infill plate boundary conditions on specimen hysteretic behavior and other aspects of system performance, test data of Specimen CSPSW are included in the following result discussions. More detailed information about Specimen CSPSW is available elsewhere [25]. Hysteretic curves of all the specimens are presented in Fig. 17. As depicted, compared with the top stories, the bottom stories of all the specimens consistently sustained larger inelastic loading due to the loading scheme described in Section 4. As such, the following discussions focus on the bottom-story hysteretic responses. As shown, the specimens tested in this investigation all exhibit strength and stiffness degradations from cycle to cycle, particularly in the cycles with larger peak inter-story drifts. However, in-cycle strength degradation which may cause collapse of a system is not observed in these specimens, suggesting that all systems can exhibit stable hysteretic behavior under earthquake loading. Compared with the specimens with buckling restrained infill plates (i.e., Specimens B, C and CSPSW), hysteretic curves of the one with conventional infill plates (i.e., Specimen A) exhibit remarkably pinching due to buckling of the infill plates. Overall, hysteretic curves of Specimens B and C are similar to those of Specimen CSPSW except that more severe strength degradations can be observed in Specimens B and C.
5.3. Discussion of hysteretic curves Prior to this investigation, the research team also tested another specimen with a composite boundary frame (referred to herein as Specimen CSPSW) [25]. Dimensions, materials and loading scheme of Specimen CSPSW were the same as those of Specimen C; how-
5.4. Comparison of maximum responses Envelops of the hysteretic curves, known as backbone curves, are compared for all the specimens in Fig. 18. The maximum base
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4
4
2
2
με (x103)
με (x103)
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0
-2
0
-2 B1 B3
-4 -4
-2
2
0
B4 B5
-4
4
-4
-2
Inter-story drift (%) shear yielidng strain yielding inter-story drift
4
yielidng strain yielding inter-story drift
(a) Beam web shear strain
(b) Beam flange axial strain
7.5
7.5
5.0
5.0
2.5
2.5
με (x103)
με (x103)
2
0
Inter-story drift (%)
0.0
0.0 -2.5
-2.5 C1 C2 C3
-5.0
W1 W2 W3
-5.0 -7.5
-7.5 -4
-2
0
2
4
-4
-2
2
4
shear yielidng strain yielding inter-story drift
yielidng strain yielding inter-story drift
(c) Column axial strain
0
Inter-story drift (%)
Inter-story drift (%)
(d) Inifll plate shear strain
Fig. 14. Strain data from Specimen A.
shear and the corresponding bottom-story drift, and the maximum bottom-story drift and the corresponding base shear identified from the backbone curves are summarized in Table 3 together with the yielding drift and ductility of each system. Comparing the maximum base shear resistances of Specimen A (573.7 kN and 581.9 kN under the pull and push cycles, respectively) with those of Specimen B (716.2 kN and 706.9 kN under the pull and push cycles, respectively), it is found that the reinforced concrete panels can increase strength of a system up to 24.8%. Note that the observed strength increase in Specimen B was due to two factors: prevention of infill plate buckling and friction between the infill plates and concrete panels. However, the respective contributions of these two factors were not explicitly measured. Additionally, it is found that ductility capacity of the SPSW system may reduce up to 59% when adopting the buckling restrained infill plates (7.28 of Specimen A versus. 3.01 of Specimen B, see Table 3). Note that test of Specimen B was concluded essentially due to the potential fall-off failures of the concrete panels. Therefore, enhancing flexural strength of the concrete panels may help improve ductility of the system with bucking restrained infill plates.
Comparison of the results from Specimens C and CSPSW reveal that connecting the infill plates to beams only, although help reduce demands on columns, may cause a lower system strength and more severe strength degradations. The lower system strength observed in Specimen C is primarily due to the fact that the infill plates in Specimen C failed to form the tension field actions as complete and uniform as those of Specimen CSPSW. The more severe strength degradations in Specimen C were primarily due to ruptures at the ends of the infill-plate-to-fishplate connections. This observation indicates that free edges of an infill plate can cause stress concentration and amplified local strain at the ends of the infill-plate-to-fishplate connection which can trigger formation of ruptures in the connections. As listed in Table 1, infill plates in Specimens B and C had different thicknesses and widths; however, cross-section areas of their infill plates (i.e., product of thickness and width) were similar (3300 mm2 from Specimen B vs. 3120 mm2 from Specimen C). Although differences in infill plate width and thickness can cause different infill plate buckling modes and post-buckling strengths, the reinforced concrete panels enabled the two specimens to exhibit comparable ultimate strength and ductility as shown in Table 3.
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B1
ki = equivalent secant stiffness of the system under the ith loading step;
B3
and V push = base shear of the system under the pull and V pull i i push loads of the ith loading step, respectively; and Unbalanced vertical component of infill plate forces
Reactions at left end
Dpull and Dpush = inter-story displacement of the bottom story of i i the system under the pull and push loads of the ith loading step, respectively. Fig. 19 compares the equivalent secant stiffness of each considered system. As shown, at the small drift with elastic response, Specimens A, B and C exhibited the same equivalent secant stiffness, indicating that the concrete panels do not seem to help improve initial stiffness of the system. However, as the drift progressively increases up to 2%, Specimen A appears to have a lower equivalent secant stiffness value due to buckling of the infill plates. Beyond the drift of 2%, Specimens B and C end up with the same secant stiffness as that of Specimen A due to the cracks in the reinforced concrete panels. Overall, Specimens B and C exhibited very similar equivalent secant stiffness, suggesting that the infill plates with similar cross-section areas tend to exhibit the same lateral stiffness when the concrete panels are used to prevent infill plate buckling. Moreover, it is found that the secant stiffness of Specimen CSPSW is consistently higher than the other specimens. Specifically, at the bottomstory drift of 1.75%, secant stiffness of Specimen C is about 61% of that of Specimen CSPSW.
Reactions at right end
(a) shear strains at B1 and B3 due to infill plate forces B1
B3
Rotation at left end
Rotation at right end
(b) shear strains at B1 and B3 due to frame sway-pull loading cycle B1
B3
Rotation at left end
Rotation at right end
(c) shear strains at B1 and B3 due to frame sway-push loading cycle Fig. 15. Illustration of shear strains at the web of intermediate beam.
Another important observation from Table 3 is that ductility of the systems with buckling restrained infill plates overall have similar ductility factors on the order of 3.0 regardless of infill plate geometries and boundary conditions along plate edge. Such ductility factors are lower than that of the system with conventional infill plates permitting buckling.
5.5. Lateral stiffness degradation As shown in Fig. 17, stiffness of each system (i.e., slope of the hysteretic curve) varies at different drift levels. For a more convenient comparison of stiffness degradation as drift in a system increases, the following equivalent secant stiffness is defined.
pull push V i þ V i ki ¼ push pull Di þ Di
ð1Þ
where
4 29 kN
1% drift
5.6. Energy absorption features The infill plates installed in the composite frames were intended to dissipate the earthquake energy released into the systems. Specimens A and B provided a unique opportunity to compare the energy absorption features of the infill plates without and with plate buckling restraining actions. Fig. 20 compares the hysteretic loops of the two specimens at a selected load step. Note that the area enclosed by a hysteretic loop corresponds to the energy absorbed in the system during the load cycles. As shown, when the peak drifts remain about the same, Specimen B absorbs more energy than Specimen A due to the beneficial contribution of reinforced concrete panels to prevention of infill plate buckling. Note that the area enclosed by a hysteretic loop is affected by strength of the system and the peak deformation it achieves. To eliminate the influences of these two parameters and achieve direct comparisons among all the considered systems, the energy absorption coefficient, E, defined in the Code for Seismic Design of
1.33% drift
1.5% drift
Fig. 16. Yielding progression in Specimen A.
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800
Specimen B 2nd story
600
Story shear (kN)
Story shear (kN)
800
Specimen A 2nd story
600 400 200 0 -200 -400
400 200 0 -200 -400 -600
-600
-800
-800 -4
-2
4
2
0
-3
-2
800
400 200 0 -200 -400
2
3
400 200 0 -200 -400
-800
-800 -4
-2
4
2
0
-3
-2
800
1200
Story shear (kN)
Specimen C 2nd story
600
-1
0
1
2
3
Inter-story drift (%)
Inter-story drift (%)
Story shear (kN)
1
-600
-600
400 200 0 -200 -400
800
Specimen CSPSW 2nd story
400 0 -400 -800
-600 -800
-1200 -3
-2
-1
0
1
2
-3
3
-2
Inter-story drift (%) 800
-1
0
1
2
3
2
3
Inter-story drift (%) 1200
Story shear (kN)
Specimen C 1st story
600
Story shear (kN)
0
Specimen B 1st story
600
Story shear (kN)
Story shear (kN)
800
Specimen A 1st story
600
-1
Inter-story drift (%)
Inter-story drift (%)
400 200 0 -200 -400
800
Specimen CSPSW 1st story
400 0 -400 -800
-600
-1200
-800 -3
-2
-1
0
1
2
3
-3
-2
-1
0
1
Inter-story drift (%)
Inter-story drift (%) Fig. 17. Comparison of hysteretic curves.
Buildings (GB50011-2010) [40] was calculated for each specimen. Fig. 21 schematically shows the definition of E. As illustrated, E represents ratio of the energy absorbed by the system in a loading cycle (i.e., the area enclosed by a hysteretic loop, which is calculated as the sum of SABC and SCDA, to the sum of the areas enclosed by the two triangles, SOBE and SODF). Based upon this definition, a system with a larger value of E tends to have a better energy absorption capacity. Fig. 22 compares the energy absorption coefficient of each considered system. As shown, the energy absorption coefficient of
Specimen A is consistently lower than the other specimens over the drift range of interest, indicating that the energy absorption capacity of conventional infill plates is inferior to that of the buckling restrained infill plates. Another important observation from Fig. 22 is that all the specimens with the buckling restrained infill plates exhibit almost identical trend in energy absorption coefficient as the drift increases. This indicates that geometries and boundary conditions of a buckling restrained infill plate, although influence system strength and stiffness, do not affect its favorable feature in absorbing hysteretic energy.
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V B
F
D
A
O
C
E
E=
S ABC +S CDA SOBE +S ODF
Fig. 21. Definition of energy absorption coefficient. Fig. 18. Comparison of backbone curves.
Fig. 22. Comparison of energy absorption coefficient. Fig. 19. Comparison of equivalent secant stiffness values.
800 600
Story shear (kN)
400 200 0 -200 -400 Specimen A -600 -800
Specimen B -2
-1
0
1
2
Bottom-story drift (%) Fig. 20. Comparison of hysteretic energy absorption: Specimen A versus Specimen B.
6. Conclusions Based upon the test results obtained from this investigation, the following significant conclusions were drawn:
Overall, the SPSWs consisting of CFST columns and infill plates connected to beams only exhibited stable responses and acceptable ductility under cyclic loading. Therefore, such systems can be used in seismic design of multi-story buildings. When connected to beams only, a buckling restrained infill plate tends to offer a higher strength but lower ductility than a conventional infill plate with similar geometry and material property. Nevertheless, the systems with buckling restrained infill plates were able to achieve the ultimate drift on the order of 2% and the ductility of 3.0. Moreover, stiffness degradations in the systems with buckling restrained infill plates are not as severe as the one with conventional infill plates. Therefore, the buckling restrained infill plates are recommended for future practice. Failures of the systems consisting of buckling restrained infill plates are due to severe cracks and possible fall-off failures of the concrete panels. Therefore, increasing flexural strength of the concrete panels may help enhance ductility of the system and expand its range of applicability. Hysteretic energy dissipation feature remains the same in all the specimens with buckling restrained infill plates regardless of infill plate dimensions and plate edge boundary conditions. Inelastic behavior may occur in the CFST columns when the new SPSW system is loaded to large inter-story drifts. Specifically, the ends of each CFST column (in the proximity of beam-toCFST-column connections) may exhibit steel tube local buckling
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followed by ruptures. However, this may be due to the loading scheme adopted in this experimental investigation. Future research opportunities exist to develop improved design models for the CFST columns and better detailing requirement for the beam-to-CFST-column connections.
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