EXPERIMENTAL INVESTIGATIONS ON STEEL PLATE SHEAR WALLS USING BOX COLUMNS WITH OR WITHOUT INFILL CONCRETE

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1 10NCEE Tenth U.S. National Conference on Earthquake Engineering Frontiers of Earthquake Engineering July 21-25, 2014 Anchorage, Alaska EXPERIMENTAL INVESTIGATIONS ON STEEL PLATE SHEAR WALLS USING BOX COLUMNS WITH OR WITHOUT INFILL CONCRETE C.-H. Li 1, K.-C. Tsai 2, H.-Y. Huang 3, C.-Y. Tsai 4 and C.-H. Lin 1 ABSTRACT Steel plate shear wall (SPSW) is a kind of lateral force resisting system. Unstiffened infill plates are connected to the boundary beams and columns at the all sides of the plates. By allowing the infill panels buckle in shear and develop the tension field action, an SPSW can effectively resist the lateral forces by the tension field action. If the boundary columns in an SPSW are bare or concrete filled steel box columns, the inner column flanges connected to the panels would suffer notable pull-out panel forces. In addition, these boundary columns must resist the axial forces and the moments induced from the sway actions of the SPSW. This study investigates the capacity design methods for the inner flanges of the box columns with or without the infill concrete to prevent the major yielding of the inner flange. Cyclic tests of three full-scale twostory SPSWs using box columns with or without the infill concrete were conducted. Test results show that minor yielding of the inner column flange would not adversely affect the seismic performance of the SPSW while the major yielding in the inner column flange would cause a significant permanent deformation in the flange. 1 Assistant Researcher, National Center for Research on Earthquake Engineering, Taipei, Taiwan 2 Professor, Dept. of Civil Engineering, National Taiwan University, Taipei, Taiwan 3 Former Graduate Student, Dept. of Civil Engineering, National Taiwan University, Taipei, Taiwan 4 Ph. D. Candidate, Dept. of Civil Engineering, National Taiwan University, Taipei, Taiwan Li C-H, Tsai K-C, Huang H-Y, Tsai C-Y and Lin C-H. Experimental Investigations on Steel Plate Shear Walls Using Box Columns with or without Infill Concrete. Proceedings of the 10 th National Conference in Earthquake Engineering, Earthquake Engineering Research Institute, Anchorage, AK, 2014.

2 10NCEE Tenth U.S. National Conference on Earthquake Engineering Frontiers of Earthquake Engineering July 21-25, 2014 Anchorage, Alaska Experimental Investigations on Steel Plate Shear Walls Using Box Columns with or without Infill Concrete C.-H. Li 1, K.-C. Tsai 2, H.-Y. Huang 3, C.-Y. Tsai 4 and C.-H. Lin 1 ABSTRACT Steel plate shear wall (SPSW) is a kind of lateral force resisting system. Unstiffened infill plates are connected to the boundary beams and columns at the all sides of the plates. By allowing the infill panels buckle in shear and develop the tension field action, an SPSW can effectively resist the lateral forces by the tension field action. If the boundary columns in an SPSW are bare or concrete filled steel box columns, the inner column flanges connected to the panels would suffer notable pull-out panel forces. In addition, these boundary columns must resist the axial forces and the moments induced from the sway actions of the SPSW. This study investigates the capacity design methods for the inner flanges of the box columns with or without the infill concrete to prevent the major yielding of the inner flange. Cyclic tests of three full-scale two-story SPSWs using box columns with or without the infill concrete were conducted. Test results show that minor yielding of the inner column flange would not adversely affect the seismic performance of the SPSW while the major yielding in the inner column flange would cause a significant permanent deformation in the flange. Introduction Steel plate shear wall (SPSW) is an effective lateral force resisting structural system. The steel plates are connecting to the boundary beams and boundary columns though the fishplates using welded or bolted connection details. SPSWs can effectively resist horizontal earthquake forces by allowing the development of diagonal tension field action after the infill plates buckle in shear, and then dissipate energy through the cyclic yielding of the panel in tension [1]. For the past two decades, most researches focused on the SPSW using the wide flange boundary beams and columns. However, the studies regarding to SPSWs with box boundary columns are limited so far. On the other hand, the use of box columns is popular in Taiwan practice. Therefore, the present study conducted an experimental program to explore the seismic behaviors of the SPSW using box boundary columns. The flange of the box boundary column connected to the infill panel would be significantly pulled out by the tension field force from infill panels. Thus, this study aims to develop design check methods to prevent the yielding of the column flange. 1 Assistant Researcher, National Center for Research on Earthquake Engineering, Taipei, Taiwan 2 Professor, Dept. of Civil Engineering, National Taiwan University, Taipei, Taiwan 3 Former Graduate Student, Dept. of Civil Engineering, National Taiwan University, Taipei, Taiwan 4 Ph. D. Candidate, Dept. of Civil Engineering, National Taiwan University, Taipei, Taiwan Li C-H, Tsai K-C, Huang H-Y, Tsai C-Y and Lin C-H. Experimental Investigations on Steel Plate Shear Walls Using Box Columns with or without Infill Concrete. Proceedings of the 10 th National Conference in Earthquake Engineering, Earthquake Engineering Research Institute, Anchorage, AK, 2014.

3 Seismic Design for Box Boundary Columns in SPSWs Desired Pushover Responses of SPSW and Capacity Design for Boundary Column Past research [2] proposed the pushover response (Fig. 1) for a properly designed SPSW. There are three important limit states during the anticipated pushover responses of an SPSW. The corresponding performance requirements for each limit state are introduced as follows: (1) Initial Yielding State (State-IY): State-IY is defined as the first member yields during the pushover response. The infill panel is designed to be the first part yielding in an SPSW. As shown in Fig. 1, the yielding is initiated along the tensile diagonal of the panel. State-IY usually occurs at a story drift of 0.3%~0.5% rad. For an economical design, State-IY should occur slightly beyond the state corresponding to the seismic design load. V 2/50 V 10/50 V D V 50/50 elastic range State-IY (2) (1) initial yeilding a (3) A initial yeilding 0.3~0.5% rad. State-UY (4) progressive yielding range full yeilding b B full yeilding 1.0~1.5% rad. mechanism range State-HD (5) SPSW c infill panel C boundary frame 2.5% rad. Disp Uniform Yielding Mechanism strain-hardening In-span Plastic Zone Figure 1. V D (1) Elastic Range (2) State-IY (3) (4) State-UY Plastic hinge Yielding of panel V 2/50 (5) State-HD Strain-hardened Anticipated pushover responses of a properly designed SPSW (2) Uniform Yielding State (State-UY): State-UY is at the time when the SPSW just reaches the uniform yielding mechanism [3]. In this state, all the infill panels fully yield through the entire height of the SPSW, all boundary beams form plastic hinges at the beam ends, and the bottom boundary column form plastic hinges at the column bases. State-UY usually takes place at a story drift of about 1.0% to 1.5% rad. The SPSW is expected to reach State-UY during the design basis earthquake (DBE). It should be noted that the schematic uniform yielding mechanism shown in Fig. 1 has plastic hinges concentrating at the bottom ends of the bottom columns. In fact, for the bottom column which is compressed under the overturning action, the plastic zone could possibly widely spread over the bottom half of the column. This kind of inspan plastic zone is allowed to form at State-UY. (3) Target Hardening State (State-HD): State-HD is defined as the state when the SPSW

4 deforms to the target story drift limit for the maximum considered earthquake (MCE). The SPSW shall have adequate ductility to be pushed to the target story drift. At State-HD, the plastic zones which had formed at State-UY have experienced a certain amount of strain-hardening, referred to as target strain-hardening. Except for those plastic zones, the remaining parts in the SPSW shall be designed to remain elastic at State-HD in order to avoid the occurrence of any undesirable plastic mechanism, such as soft-story mechanism. Past pseudo-dynamic test results [4] show that, under the MCE ground motions, the peak inter-story drift of a properly-designed SPSW reached about 2.5% rad. Therefore, the target story drift for MCE is set as 2.5% rad. Tsai et al. [2] proposed a capacity design method for the bottom boundary column in SPSWs. The design objective is to ensure the top ends of the bottom columns remaining elastic even when a SPSW is laterally deformed to State-HD. The prevention of the yielding at the top end of the bottom boundary column in an SPSW is experimentally verified to be necessary for avoiding the soft-story mechanism at the bottom of an SPSW or instability problems on the bottom columns. Except for the plastic zones developed in the lower halves of the bottom boundary columns, the remaining boundary columns should be designed to remain elastic throughout the pushover of an SPSW [5]. Design Checks for Local Pull-out Deformation of the Flange of the Box Column Figure 2 shows the representative free-body diagram of a segment of a box boundary column in an SPSW under lateral forces. The flange connected with the infill panel is designated as interior flange. The horizontal component of tension field forces of the infill panel would pull the interior flange to bend outward, which is hereinafter referred as to pull-out deformation of the flange. It can be expected that, if the flange is too slander, the pull-out action possibly cause the flange to yield. This could result in permanent pull-out deformation on the flange. It is considered that significant pull-out inelastic deformation of the interior flange should be prevented. Thus, this study develops several design check methods for the pull-out action. In this study, two simplified analytical models for estimating the pull-out responses of the interior flange are proposed. As shown in Fig. 3(a), the fixed-end beam model is proposed for estimating the pull-out action of the interior flange in the box column with infill concrete considering that the presence of the infill concrete would considerably restraint the rotational deformation of the corners between the interior flange and the webs. Consider a segment of the column having a unit height, the interior flange of the segment can be represented by a fixed-end beam having a rectangular cross section and subjected to a concentrative transverse load, which is from the tension field force of infill panel, at the mid-span of the beam. The transverse load, T p, can be determined by ω ch (1 unit height) and Ω p ω ch (1 unit height) for State-UY and State-HD. The ω ch = t p F yp sin 2 α [1] is the horizontal component of the tension field forces of a fully yielding panel acting on the boundary columns, where t p, F yp and α are the thickness, yield stress and tension field angle [5] of the infill panel, respectively. The Ω p is the factor accounting the strain hardening effect of the infill panel at State-HD. According to the fixed-end beam model, the moment demands on the interior flange per unit column height, M d-uy and M d-hd, which are respectively corresponding to the State-UY and State-HD, can be estimated by:

5 M = ω B/8; M =Ω ω B/8 (1) d HD ch d HD p ch In addition, the pull-out deflections of the flange, Δ UY and Δ HD, which are respectively corresponding to State-UY and State-HD can be estimated by: ω Ω ω ch Δ UY = 4x + 3 Bx ; Δ = 4x + 3Bx 48EI 48EI 3 2 p ch 3 2 HD (2) where beam span B is taking as the center lines of the two column webs and x is the distance from the web center to the mid-span of the interior flange. Figure 2. Free body diagram of a box boundary column in SPSWs Figure 3. Analytical models for estimating the pull-out action: (a) portal frame model and (b) fixed end beam model For the pure steel box column, the corners between the interior flange and the webs are considered to have somewhat rotational flexibility under the pull-out action. As shown in Fig. 3(a), the portal frame model is proposed for estimating the pull-out action of the interior flange. The interior flange and the column webs are respectively considered as the beam and the two columns of a portal frame. The tension field force is represented by concentrative transverse load acting at the mid-span of the beam of the portal frame. According to the portal frame model, the moment demands on the interior flange per unit column height, M d-uy and M d-hd, which are respectively corresponding to the State-UY and State-HD, can be estimated by: M = ω B/6; M =Ω ω B/6 (3) d HD ch d HD p ch Moreover, the pull-out deflections of the flange, Δ UY and Δ HD, which are respectively corresponding to State-UY and State-HD can be estimated by: ω Ω ω ch Δ UY = 4x + 2 Bx + B x ; Δ = 4x + 2Bx + B x 48EI 48EI p ch HD (4) Considering the interior flange per unit column height as a beam having a rectangular cross-section with a unit width and a length equal to the column flange thickness, the moment

6 capacities M y and M p corresponds to the limit states associated with the yielding only at the surfaces of the interior flange and the fully yielding across the whole flange thickness, respectively. In this research, the limit states corresponding to the abovementioned initial yielding and fully yielding of the interior flange are respectively designated as minor yielding and major yielding. For the design check purpose, four demand-to-capacity ratios (DCRs), DCR y-uy, DCR p-uy, DCR y-hd and DCR p-hd are defined as the follow: Md UY Md UY Md HD Md HD DCRy UY = ; DCRp UY = ; DCRy HD = ; DCRp HD = (5) M M M M y p y p The subscripts of UY and HD for the DCRs represent the State-UY and State-HD, respectively. The subscripts of y and p for the DCRs respectively represent the DCRs associated with the minor yielding and major yielding. Once the value of a DCR exceeds 1, it represents that the related yielding state (minor or major yielding) takes place at the associated drift level (State-UY or State-HD). On the other hand, when a DCR value is smaller than 1, it means that the related yielding is not reached at the associated drift level. Specimen Design Test Program As shown in Fig. 4, three two-story SPSW specimens were 3420 mm wide by 7640 mm high (3820 mm at each story). The thickness of infill panels and the sizes of the boundary beams were all the same for the three specimens. The panels were 2.6 mm thick low yield strength (LYS) steel with a measured yield stress, F yp = 210 MPa. The panels were welded at all edges to the boundary elements using 6-mm thick and 50-mm wide fishplate connections. All the boundary elements and stiffener plates used in the three specimens were A572Gr50 steel. H-shape H , H and H beams were employed as the top beam (TB), the middle beam (MB) and the bottom beam (BB), respectively. The four numbers identified above for Taiwan H-shape section respectively represent the depth, flange width, web and flange thicknesses (in mm). As shown in Fig. 4, reduced beam section (RBS) details were adopted for the beam-to-column connections of TBs and MBs. The selection of the boundary beams follows the recommendation by Vian et al. [6], which intends to avoid the plastic hinge forming within the beam span. For the boundary beams, the all-welded beam-to-column connection was adopted. The shear tab was fillet welded at three edges to the beam web and a groove weld was added between the beam web and the column flange. The key differences among the specimens are the size of the column and the use of infill concrete. The three specimens are respectively named NSB (normal steel box column), NCB (normal concrete-infill box column) and WCB (weak concrete-infill box column). For each specimen, the same columns were used for the 1 st and 2 nd stories. BOX , BOX and BOX columns were used for Specimens NSB, NCB and WCB, respectively. Moreover, the columns in Specimens NCB and WCB were infilled by self-consolidating concrete having an expected minimum compressive strength of 280 kgf/cm 2.

7 Figure 4. Test specimens Figure 5. Test setup Table 1. Design checks of the pull-out actions of the interior flanges for the three specimens Specimen NSB NCB WCB Simplified model portal frame model fixed-end beam model fixed-end beam model DCR y-uy DCR p-uy DCR y-hd DCR p-hd Table 1 lists the design checks for the pull-out actions of the interior column flange for the three specimens. For Specimen NSB and NCB, both of the DCR y-uy and DCR y-hd exceed 1 while both of the DCR p-uy and DCR p-hd are smaller than 1. These mean that the interior flange would only suffer minor yielding at State-UY but would not reach major yielding up to State-HD. For Specimen WCB, all DCRs are larger than 1, these represent that the major yielding of the interior flange would take place at State-UY. Test Setup, Key Instrumentations and Loading Procedures Figure 5 shows the test setup. Each SPSW specimen was mounted on the strong floor in National Center for Research on Earthquake Engineering (NCREE). A lateral support frame was erected to restrain the out-of-plane displacement of the specimen. A stiff lateral support element (16 total) was connected to each continuity plate in the beam-to-column joints of TB and MB though bolts. The support elements were also designed to contact with the side edges of the TB and MB flanges, providing a sliding surface between the specimen and lateral support frame to transmit the out-of-plan restraining forces from the lateral support frame to the specimen. The instrumentation focused on measuring the deformation responses of the boundary elements. As shown in Fig. 6a, two reference columns fixed on the strong floor at the north and south sides of the specimen. A total of 17 string-pot displacement transducers, DN1 to DN9 and DS1 to DS8 connected the reference columns to the boundary columns, respectively. As shown in Fig. 6b, a total of 26 tilt-meters were installed on the webs of boundary beams and columns in

8 which plastic rotations were likely significant. In addition, as shown in Fig. 6c, three dial gauges were utilized to measure the pull-out deformation of the interior flange at the mid-height of each 2F column. Each specimen was painted overall using whitewashes. The locations where the plastic deformations took place can be recognized by observing the flaking of the whitewashes. Figure 6. Key instrumentations: (a) string pot; (b) tilt-meter and (c) dial gauge Prescribed cyclically increasing displacements, consisting of two cycles of roof drifts of 0.1%, 0.25%, 0.5%, 0.75%, 1%, 1.5%, 2%, 2.5%, 3% and 4% rad., were applied onto the south column at the top beam elevation using two 980-kN actuators. Both actuators were driven using the same displacement commands. Before applying the cyclic lateral displacements, a vertical load of about 30% of the column nominal axial yield capacity was applied at each column top using the post-tension rods through a 3440 mm long loading beam. These vertical loads were maintained and recorded throughout the test by using a hydraulic pressure stabilizing system. Test Results and Numerical Simulations Numerical Models A finite element shell model was developed for each specimen by using ABAQUS program [7]. The steel part of the specimen was represented using the reduced integration and a large-strain formulation for the 4-node, quadrilateral, stress/displacement shell elements (ABAQUS S4R Element). All the fishplates were not included in the shell models. The panels were directly tied to the boundary beams and columns. Both columns were fixed at their bases. The material properties adopted for shell elements were bilinear incorporating the coupon yield stresses and a post-yield stiffness of 0.01E (E is Young's modulus). The infill concrete was modeled by the 3D solid element (ABAQUS 3D8I Element), the details, including the stress-strain relationship and yield criterion for the infill concrete and the interface between the infill concrete and steel box column, followed the suggestions made by Hu et al. [8] The out-of-plane displacement was restrained at the nodes for each beam-to-column joint at the 2nd floor and roof level. In order to properly simulate the buckling of the infill panels, initial imperfection was specified on the panels. The imperfection distribution was determined from the superposition of the first two buckle mode shapes for each panel obtained from an ABAQUS buckle analysis. Two steps of nonlinear static analyses were performed for each model: (1) applying vertical loads on the top of each 2F column and (2) conducting the

9 displacement control pushover analysis. Figure 7. Experimental cyclic force versus displacement relationships and analytical pushover responses for the three specimens (a) (b) Figure 7 (a) Photo of the specimens after tests and (b) analytical plastic zone distributions at 2.5% rad. roof drift Cyclic Responses of the Overall Specimens Figure 7 show the cyclic force versus displacement relationships for the three specimens and the analytical pushover results from the corresponding ABAQUS models. It can be found that the analytical pushover curves have a good agreement with the envelopes of experimental hysteresis loops. All the specimens exhibited ductile behaviors. All the specimens had a deformation capacity exceeding a roof drift of 4% rad. Figure 8a shows the conditions of the thee specimens after tests. Based on the flaking of whitewash on the specimens, the plastic zones developed on the specimens can be recognized. It can be found that the buckling waves remained on the infill panel, implying the tension field action had taken place. Plastic hinges formed at all the ends of the top and middle boundary beams. In addition, a widespread plastic zone formed over the lower half of each 1F column. Nevertheless, the remaining parts of the beams and columns remained elastic after the tests. These confirm that the three specimen developed desired plastic mechanism during the tests. Figure 8b shows the distributions of the plastic zones at a roof drift of 2.5% rad. obtained by the ABAQUS models. The dark regions indicated the plastic zones. It can be found that the analytical results quite match the test results.

10 Pull-out Responses of the Interior Flange Figure 9 shows the experimental and analytical pull-out deflections on the interior flanges at the mid-height of the 2F column at a roof drift of 2.5% rad. The black dots represent the test result measured by the dial gauges while the dashed lines and solid lines are the deflections predicted by the proposed fixed-end beam model and the portal frame model, respectively. For the fixedend beam model, the span B can be taken as the distance between the center lines of the column webs or the clear distance between the two column webs. The short dashed lines and long dashed lines respectively denote the pull-out deflections from the models based on clear flange width and web centreline distance. It can be found that predictions from the fixed-end beam models utilizing the two kinds of span are about the same. Furthermore, Fig. 9 shows that the fixed-end beam models can satisfactorily predict the interior flange pull-out deflections for the Specimens NCB and WCB, which used the concrete filled steel box columns. For the Specimen NSB, which used the pure steel box columns, the fixed-end beam model somewhat underestimates the pullout deflections. This implicates that the use of the fixed-end beam model for design check of the yielding on the interior flange in the pure steel box column could be unsafe. On the other hand, Fig. 9 also shows that the portal frame model obviously overestimates the pull-out deflections in the steel box column. This indicates that the use of the portal frame model to design the pure steel box column would lead to a conservative result. State-HD 16%6 16%1 1&%6 1&%1 :&%6 :&%1 Figure 9 Experimental and analytical pull-out deflections at 2.5% rad. roof drift Figure 10 Analytical distributions of the plastic zones on the inner and outer surfaces of the interior flanges of 2.5% rad. roof drift Figure 10 shows the analytical distributions of the plastic zones on the inner surfaces and outer surfaces of the column interior flanges at a roof drift of 2.5 % rad. for the three specimens.

11 The dark regions represent the plastic zones. It can be found that, except for some column ends where the axial forces and moments are significant, the interior flanges of the remaining parts of the columns for Specimens NSB and NCB remain essentially elastic at 2.5 % rad. roof drift. However, for Specimen WCB, the yielding of the interior column flanges take place at the center and the both ends of the flanges over the entire columns. These abovementioned analytical responses of the interior column flanges quite match the predictions made by the proposed design checks for the major yielding as listed in Table 1. Figure 9 also shows that the interior flange pull-out deflections in Specimen WCB were much larger than those in Specimens NSB and NCB. This suggests that the major yielding of the interior flange would result in significant pull-out deflections. In order to avoid severe permanent pull-out deflections on the interior flange, the major yielding of the flange shall be prevented. Conclusions This study conducted the cyclic tests of three full-scale two-story SPSWs having box boundary columns with or without infill concrete. All the three SPSWs exhibited satisfactorily ductile behaviours even though the interior column flanges in Specimen WCB developed significant pull-out deformations by the tension field forces of the panels. Test results confirm that the proposed fixed-end beam model can effectively predict the pull-out actions of the interior flange of the box column with infill concrete. For the pure steel box column, the fixed-end beam model slightly underestimates the pull-out actions while the portal frame model notably overestimates. Moreover, the effectiveness of the proposed design checks in predicting the major yielding of the interior flange has been verified experimentally and numerically. References 1. Thorburn, L.J., Kulak, G.L. and Montgomery C.J. Analysis of Steel Plate Shear Walls. Structural Engineering Report No. 107, Dept. of Civil Engineering, Univ. of Alberta, Edmonton, Alberta, Canada, Tsai, K.C., Li, C.H., and Lee, H.C. Seismic Design and Testing of the Bottom Vertical Boundary Elements in Steel Plate Shear Walls. Part 1: Design Methodology. Journal of Earthquake Engineering and Structural Dynamics (accepted for publication) 3. Berman, J.W. and Bruneau, M. Plastic analysis and design of steel plate shear walls. Journal of Structural Engineering (ASCE) 2003; 129(11): AISC (American Institute of Steel Construction). ANSI/AISC : Seismic Provisions for Structural Steel Buildings. American Institute of Steel Construction, Chicago, Illinois, Lin, C.H., Tsai, K.C., Qu, B., and Bruneau, M. Sub-structural Pseudo-dynamic Performance of Two Full-scale Two-story Steel Plate Shear Walls. Journal of Constructional Steel Research 2010; 66(12): Vian, D., Bruneau, M., Tsai, K.C. and Lin, Y.C. Special Perforated Steel Plate Shear Walls with Reduced Beam Section Anchor Beam I: Experimental Investigation, Journal of Structural Engineering (ASCE) 2009; 135(3): HKS (Hibbit Karlsson and Sorenson, Inc.) ABAQUS/Standard 6.7 User s Manual, Hibbit, Karlsson and Sorenson, Inc., Pawtucket, RI, Hu, H.T., Huang, C.S. and Chen Z.L. Finite element analysis of CFT columns subjected to an axial compressive force and bending moment in combination. Journal of Constructional Steel Research 2005; 61(12):