Comparison of Solid and Perforated Hybrid Precast Concrete Shear Walls for Seismic Regions

This paper compares the measured lateral load behaviors of two 0.4-scale “hybrid” precast concrete wall test specimens – one wall with solid panels and the other with perforated panels. The test specimens have the same overall geometry and utilize a combination of mild [i.e., Grade 400 (U.S. Grade 60)] steel and high-strength unbonded post-tensioning (PT) steel for lateral resistance across horizontal joints. The mild steel reinforcement is designed to yield in tension and compression, providing energy dissipation. The unbonded PT steel provides self-centering capability, reducing the residual lateral displacements of the wall after a large earthquake. The comparisons between the two walls focus on the applied lateral load versus displacement behavior, energy dissipation, shear deformations, and behavior along the critical horizontal base-panel-to-foundation joint. Both specimens demonstrated excellent re-centering and energy dissipation capabilities as well as ductile behavior over lateral displacements at or greater than the ACI ITG-5.1 (2007) requirement. The perforated specimen was able to achieve greater lateral drift due to improvements made to the details of the panel reinforcement at the wall toes. Ultimately, these results are expected to support the successful code approval of hybrid precast shear walls for moderate and high seismic regions of the United States. INTRODUCTION AND BACKGROUND As described in Smith et al. (2011a), the hybrid precast wall system investigated in this research utilizes a combination of mild [i.e., Grade 400 (U.S. Grade 60)] steel bars and high-strength unbonded post-tensioning (PT) strands for lateral resistance across horizontal joints. Under the application of lateral loads into the nonlinear range, the primary mode of displacement in these walls occurs through gap opening at the horizontal joint between the base panel and the foundation, allowing the wall to undergo large lateral displacements with little damage. Upon unloading, the PT steel provides a restoring force to close this gap, thus reducing the residual lateral displacements of the wall after a large earthquake. The use of unbonded tendons delays the yielding of the PT strands and reduces the tensile stresses transferred to the concrete (thus reducing cracking) as the tendons elongate under lateral loading. The mild steel bars crossing the base joint are designed to yield in tension and compression, providing energy dissipation through the gap opening/closing behavior. A pre-determined length of these energy dissipating bars (or E.D. bars) is unbonded at the bottom of the base panel (by wrapping the bars with plastic sleeves) to reduce the steel strains and prevent low-cycle fatigue fracture. The hybrid precast wall system offers high quality production, relatively simple construction, and excellent seismic characteristics by providing self-centering to the building as well as energy dissipation to control the lateral displacements. Despite these desirable characteristics, hybrid precast walls are classified as nonemulative structures since they do not emulate the behavior of conventional cast-inplace reinforced concrete walls. Thus, experimental validation is required by ACI 318 (2008) and ACI ITG-5.1 (2007) prior to their use in seismic regions of the United States. To date, limited tests and analytical studies are available [(Smith et al. (2011a), Rahman and Restrepo (2000), Holden et al. (2001), Kurama (2002), Kurama (2005)]; however, these studies have not generated the required experimental data to satisfy the ACI ITG-5.1 validation criteria. Furthermore, concrete shear walls often feature perforations to allow for windows and doors to be incorporated into the building system. Previous research on perforated precast concrete walls is extremely limited [Allen and Kurama (2002), Mackertich and Aswad (1997)], and there are currently no results published on hybrid walls with perforations. This paper focuses on these important knowledge gaps. TEST SET-UP AND SPECIMEN PROPERTIES The test set-up and procedure conformed to the requirements of ACI ITG-5.1. Each wall specimen, referred to as Specimens HW3 (solid wall) and HW4 (perforated wall), was designed based on a 4-story prototype wall within a prototype parking garage structure located in Los Angeles, CA. The specimen design satisfied the guidelines and requirements of ACI ITG-5.2 (2009), ACI-318, and ASCE 7 (2005). Schematic drawings of the test set-up and specimen cross-section (depicting the confinement steel geometry) are shown in Figure 1. More information on the overall design of the walls can be found in Smith and Kurama (2009) and Smith et al. (2011a). A detailed discussion of the design of the panel perforations can be found in Smith et al. (2011b). Each test was conducted at 0.40-scale, which satisfies the minimum scaling limit of ACI ITG-5.1. The specimens featured two wall panels: the base panel representing the 1 story of the structure and the upper panel representing the 2 through 4 stories, thereby satisfying the ACI ITG-5.1 requirement for testing multipanel walls (such that an upper panel-to-panel joint as well as the base-panel-tofoundation joint are evaluated). The 0.40-scale wall length, lw, was 243 cm (96 in.), base panel height, hpb, was 145 cm (57.5 in.), and wall thickness, tw, was 15.9 cm (6.25 in.). It was possible to test the upper story panels of the 4-story prototype wall as a single panel since the joints between the upper panels were designed to have no nonlinear behavior and no gap opening. The lateral load was applied 3.66 m (12 ft) from the wall base (near the resultant location of the 1 mode inertial forces), resulting in a wall base moment to shear ratio of Mb/Vb=1.5lw. An external downward axial load of about 325 kN (73 kips) was applied at the center of the top of each specimen to simulate the service-level tributary gravity loads acting on the prototype structure during an earthquake (assumed to be 1.0 times the service gravity load plus 0.25 times the unreduced service live load). For both specimens, the location and material properties of the PT and E.D. steel were the same. The PT steel consisted of two tendons located 28 cm (11 in.) north and south from the wall centerline. The tendons were placed near the wall centerline to reduce the strand strains and also keep the PT ducts away from the critical confined toe regions of the wall. Each tendon contained three 1.3-cm diameter (0.5-in.) strands [design ultimate strength, fpu=1862 MPa (270 ksi)] with an unbonded length from the top of the wall to the bottom of the foundation beam of approximately 5.5 m (18 ft). The average initial tendon stress, calculated from the measured individual strand forces prior to the application of the lateral load, was fpi=0.54fpu. The E.D. steel crossing the base joint consisted of four 19-mm diameter (U.S. No. 6) bars [measured yield strength, fsy=448 MPa (65 ksi), and yield strain, εsy=0.0023 cm/cm], with one pair of bars placed 19 cm (7.5 in.) north and south from the wall centerline and the other pair 8.9 cm (3.5 in.) north and south from the centerline. The E.D. bars were unbonded over a length of 38 cm (15 in.) at the bottom of the base panel. Across the upper panel-to-panel joint, only two 22-mm (U.S. No. 7) diameter bars were used, with one bar located 10 cm (4.0 in.) from each end of the wall. This reinforcement was designed not to yield; thereby limiting any gap opening along this joint. To prevent strain concentrations in the panel-to-panel joint reinforcement, a short 7.6 cm length (3.0 in.) of the bars was unbonded at the bottom (A) (B) Figure 1. Test configuration and specimen details: (a) HW3 (solid wall); (b) HW4 (perforated wall). of the upper panel. The design unconfined concrete strength for the walls was 41 MPa (6.0 ksi) and the design confined concrete strength (at the toes of the base panel) was 62 MPa (9.0 ksi). On the day each wall was tested, the measured unconfined concrete strength for the base panel was 55 MPa (8.0 ksi) for Specimen HW3 and 50 MPa (7.3 ksi) for Specimen HW4. At the horizontal joints, fiber-reinforced grout (with polypropylene microfilament fibers at 0.065% by volume) was used. The testday strength of the base joint grout was 58 MPa (8.4 ksi) for both the solid and perforated walls. As shown in Figure 1b, the panels of Specimen HW4 featured two rectangular openings, each with a length, lo, of 36 cm (14 in.) and a height, ho, of 51 cm (20 in.). The perforations were placed in a symmetrical layout with respect to the wall centerline, with the exterior edges of the perforations located 51 cm (20 in.) from the panel edge and 36 cm (14 in.) from the panel base. The perforations in the upper panel represented those in the 2 story. The perforations in the 3 and 4 stories of the prototype wall were not modeled since they would be less critical than the lower story perforations. Based on the observed performance of the Specimen HW3 (discussed in more detail later in the paper), the concrete confinement detailing at the toes of the structure was modified in the Specimen HW4. Figure 1a depicts the confinement reinforcement in Specimen HW3, which consisted of a large hoop [11.4 cm (4.5 in.) by 40.6 cm (16 in.)] with an intermediate crosstie located at the mid-length of the hoop. The confinement hoop and crosstie spacing was 7.6 cm (3.0 in.). The horizontal distributed 10-mm (U.S. No. 3) deformed bars at each face of the base panel were placed outside of the confinement cages for ease of construction. As shown in Figure 1b, the confinement reinforcement in Specimen HW4 was modified to consist of two smaller hoops [each 11.4 cm (4.5 in.) by 25.4 cm (10 in.)] that were overlapped to create a larger overall confinement steel geometry [11.4 cm (4.5 in.) by 45.7 cm (18 in.)]. This modification was made primarily to reduce confined concrete damage, which partially resulted from bowing-out of the longer legs of the bottom hoops observed in the solid specimen (as discussed in more detail later, the cross-ties were not effective in confining the concrete), and secondarily to allow for the entire exterior vertical chord of the perforated base panel to be reinforced. The confinement hoop spacing between the two walls remained the same [7.6 cm (3.0 in.)]. As an additional modification, the horizontal distributed 10-mm (U.S. No. 3) deformed bars were placed inside of the confinement cages in an attempt to limit the concrete cover spalling in the perforated base panel. MEASURED BEHAVIOR OF SPECIMENS Figure 2 shows the reversed-cyclic lateral displacement history used in the testing of the wall specimens, with three repeated cycles at each displacement increment. The wall drift, ∆w (positive with the wall displaced southward), was measured as the relative lateral displacement of the wall between the lateral load location and the foundation divided by the height to the lateral load. For the given wall dimensions, the validation-level drift per ACI ITG-5.1 is ∆w=2.30% for each specimen. Specimen HW3 was able to sustain two cycles at this validation-level drift followed by a greater drift cycle of ∆w=±2.95%. In comparison, Specimen HW4 was able to sustain three cycles at the validation-level drift followed by an additional set of three cycles at a maximum drift of ∆w=±3.05%. Figure 3a shows the solid wall specimen at ∆w=+2.95% (note the gap opening along the base joint at the north end). Minor concrete cracking in the base panel and no cracking in the upper panel was observed (note that the cracks visible in Figure 3 were highlighted with markers during the test for enhanced viewing). Crushing of the confined concrete at the wall toes (see Figure 3b) and slight bowing of the longer legs of the bottom confinement hoops occurred during the large displacement cycles. It was possible for the longer legs of the confinement hoops to bow outwards (which reduced the confinement effectiveness) since the crossties did not directly engage the hoop steel [note that the crossties did engage the longitudinal (vertical) reinforcement within the confinement cages as required by ACI 318]. The total strength loss at the completion of the drift history was less than 20%, thus satisfying the ACI ITG-5.1 validation requirement. Additional loading of the wall beyond the required displacement history resulted in further strength loss and subsequent failure of the specimen due to concrete crushing (defined by ACI ITG-5.1 as the drift at which the total strength loss exceeds 20%). No concrete crushing in the upper panel and no significant gap opening or slip in the upper panel-to-panel joint was observed during