Sequential Phased Displacement Tests of Wood-framed Shear Walls with Corners

Transcription

1 VPI&SU Report TE Sequential Phased Displacement Tests of Wood-framed Shear Walls with Corners Virginia Polytechnic Institute and State University Department of Wood Science and Forests Products Brooks Forest Products Research Center Timber Engineering Center 1650 Ramble Road Blacksburg, Virginia Report No. TE By: J.D. Dolan Associate Professor of Wood Engineering C.P. Heine Research Assistant Prepared for: The NAHB Research Center, Inc. Upper Marlboro, MD With Sponsorship from: The National Association of Home Builders Washington, DC and The U.S. Department of Housing and Urban Development Office of Policy Development and Research Washington, DC. September 14, 1997

2 2 VPI&SU Report #TE-1997/-003 INTRODUCTION A conventionally constructed wood frame structure resists lateral wind or earthquake forces by transferring the load to parallel walls which are then primarily loaded in shear. Such walls are commonly referred to as shear walls. Today, the shear resistance in engineered timber framed walls is typically provided by oriented strandboard (OSB) or plywood sheathing. Shear capacity is mainly influenced by size and location of openings, amount of overturning restraint, and nailing patterns. Shear walls in conventionally constructed houses are generally not restrained against uplift by the use of hold-down connectors. Yet, an unquantified amount of restraint is provided by conventional corner framing and anchorage. A remarkable body of literature exists concerning the lateral resistance of sheathed light frame walls. However, most of those walls were tested using a monotonic testing procedure and overturning restraints at the ends that essentially eliminate the overturning failure mode. To the authors knowledge, no investigations have been completed to quantify the capacities of reversed cyclically loaded full-scale shear walls without overturning restraint. This study expands and augments a research project conducted to quantify the effect of overturning restraint on full-scale shear walls. Results of monotonically and reversed cyclically loaded, straight walls with and without tie-down anchors are presented in Dolan and Heine (1997 a and b). To improve the accuracy of designs when no overturning restraints (i.e. hold-down brackets) are provided, the interaction of mutually perpendicular walls is investigated in this study. A detailed description of the complete investigation of the effect of overturning restraint on the performance of timber framed shear walls is presented in Heine (1997).

3 VPI&SU Report TE OBJECTIVES Results of an experimental study of the performance of shear walls with 2 feet and 4 feet corners on each end without overturning restraint are reported. The objectives of this study were: to quantify the effects of corners on uplift restraint of wood frame shear walls tested reversed cyclically. to initiate experiments to quantify an engineering analysis approach for conventional wood-framed shear walls (i.e. without overturning restraint). TEST PROGRAM The test procedure used in this study is a reversed cyclic testing procedure proposed by the Structural Engineers Association of Southern California (SEOSC, 1997). The quasi-static test method incorporates procedures included in the Sequential Phased Displacement Procedure (SPD) proposed by M.L. Porter (1987). This loading procedure is described in detail in a companion report by Dolan and Heine (1997 b). Four walls, 12 feet in length and 8 feet in height were tested using a SPD pattern. Attached to the ends of the walls were 4 feet by 8 feet and 2 feet by 8 feet corner segments, respectively. Each wall configuration was tested twice. The walls were oriented horizontally as displayed in Figure 1.

4 4 VPI&SU Report #TE-1997/-003 Figure 1: Test set up

5 VPI&SU Report TE A steel frame made out of 3 inch by 5 inch steel tubes served as rigid foundation for the corner segments. The top plates of the corner framing were braced using 7/16 inch thick OSB sheets, fastened with 8d common nails (6 in. o.c.) to simulate the assumed lower bound stiffening effect of a floor or ceiling diaphragm (Figure 1). It is important to point out that the walls were tested without gravity load (dead or live load) applied. It was intended to test a conservative condition (i.e. non-load bearing partition walls, or the potential of wind uplift to offset gravity loads). Materials and Fabrication Details Table 1 summarizes materials and construction details used for the wall specimens. Wall framing consisted of double top plates, single bottom plates, and double end studs (Figure 2). Studs were spaced 16 in. on center. All framing consisted of spruce-pine-fir, stud or better grade lumber, purchased from a local lumberyard. Members were arbitrarily chosen when placed in the wall specimens. Exterior sheathing was 7/16 in., OSB sheathing. Panels for the shear wall and the 4 feet corners were 4 feet by 8 feet and oriented vertically. Interior sheathing was 4 feet by 8 feet sheets of 1/2 inch gypsum wallboard, oriented vertically. For the 2 feet corners, the OSB and gypsum wallboard panels were cut in half lengthwise.

6 6 VPI&SU Report #TE-1997/-003 Table 1: Wall materials and construction data Component Framing Members Sheathing: Exterior Interior Anchor Bolts Fabrication and Materials Stud, Spruce-Pine-Fir, 2 x 4 inch (nom.) Structural Oriented Strand Board, 7/16 in., 4 ft. x 8 ft. sheets installed vertically. Gypsum wallboard, 1/2 in., installed vertically, joints taped 5/8 in. diameter A307 bolt with 3 in. square x 1/4 in. steel plate washers. Figure 2: Construction detail corner All joints in the interior sheathing were taped and covered with drywall compound. Compound drying times complied with manufacturer s recommendations, and were adjusted to ambient temperature and humidity. Both exterior and interior sheathing were able to rotate past the test fixture at the top and bottom (i.e. the steel test fixture was narrower than the wood framing used for top and bottom plates).

7 VPI&SU Report TE Three different types of nails were used in constructing the wall specimens (Table 2). All framing connections used 16d (0.162 in. diameter and 3.5 in. length) brite common nails. Brite common 8d (0.131 in. diameter and 2.5 in. length) nails attached the OSB sheathing to the framing, and 0.1 inch in diameter, 1-1/2 inch long drywall nails attached gypsum wallboard to the framing. A nail spacing of 6 inch around the perimeter and 12 inch for intermediate framing was used for the OSB sheathing, and 7 inch perimeter and 10 inch field for the gypsum wallboard. Table 2: Fastener schedule Connection Description Framing No. and Type of Connector Connector Spacing Top Plate to Top Plate (Face-nailed) 16d common per foot Top / Bottom Plate to Stud (End-nailed) 2-16d common per stud Stud to Stud (Face-nailed) 2-16d common 24 in. o.c. Stud to Sill (End-nailed) 2-16d common per stud Anchor bolts 1-A307 5/8 in. dia. bolt 24 in. o.c. and within 1 ft. of wall ends or outside corners Sheathing: OSB 8d 6 in. edge / 12 in. field (2 rows for end stud) Gypsum wall board 0.1 in. x 1½ in. (3/8 in. head) 7 in. edge / 10 in. field Wall Attachment to Test Frame and Instrumentation The walls were raised approximately 16 inches above the ground to allow instruments and load cell sufficient clearance to be attached to the wall. A307 or SEA grade 2 bolts were used to make all attachments to the steel structural tube test fixture. All bolts were 5/8 inch diameter National Coarse thread. The bottom plate was bolted to

8 8 VPI&SU Report #TE-1997/-003 a fixed steel structural tube 24 inches on center, beginning one foot from the end or outside corners of the wall, with 3 inch x 3 inch square, 1/4 inch thick steel plate washers. Consequently, the 2 feet corner segments were anchored only with a single 5/8 inch diameter bolt to the bottom plate. A hydraulic actuator, with a range of ±6 inches and capacity of 55,000 lbs., displaced the top corner of each shear wall. A steel tube was used to distribute the loading to the wall s double top plate. The steel tube and the double top plate were attached in the same manner as the bottom plate. Four casters allowed horizontal motion, as shown in Figure 1. The casters were fixed parallel to loading and did not rotate during testing. The data were corrected for friction generated by the casters. Figure 3 depicts the location of four linear variable differential transducers (LVDT) that were attached to the frame of each wall to measure wall displacements. LVDT #0 was fixed to the steel tube that was attached to the hydraulic actuator and measured the slip between tube and wall. LVDT #1 and LVDT #2 measured the compression and uplift displacement of the end studs relative to the foundation. These sensors determined the amount of crushing in the sill plate, or uplift of the end stud, depending on which corner of the wall was in compression or tension, respectively. The data were corrected to compensate for amplifications caused by the geometry of the LVDT fixtures.

9 VPI&SU Report TE Figure 3: Wall instrumentation LVDT #3 measured horizontal displacement of the bottom plate relative to a fixed point. This measurement allows rigid body translation of the wall to be subtracted from the global displacement to obtain inter-story drift. LVDT #4 was built into the hydraulic actuator, measuring the relative displacement of the actuator. In order to obtain the actual

10 10 VPI&SU Report #TE-1997/-003 displacement of the top of the wall the data of LVDT #4 were corrected using similar triangles to account for the depth of the steel load distribution beam. This provides the correct displacement at the top of the wall. The actuator also contained a sensor recording the load resisted by the wall. Inter-story drift was calculated as the displacement readings of LVDT #4 - ( LVDT #0 + LVDT #3). A tension bolt was inserted close to the end stud were the load was applied (Figure 3). The data acquisition system recorded the data 35 times per second. PROPERTY DEFINITIONS Initial and stabilized load envelope (or backbone) curves, similar to the ones shown in Figure 4, were determined for each wall. A typical initial load envelope curve consists of positive and negative peak loads and corresponding displacements of the first cycle for each phase of SPD loading. Positive and negative peak loads and corresponding displacements of the last cycle for each phase form a stabilized envelope curve. The highest average of the absolute values of peak positive and negative load occurring in the first cycle of each phase determined the initial capacity, F max,init. Likewise, the average value of the corresponding interstory drifts determined the displacement at initial capacity, denoted as max,init. Stabilized capacity, F max,stab, is the highest average load resistance occurring in the last cycle of each phase of SPD loading. The corresponding average interstory drift was denoted as max,stab.

11 VPI&SU Report TE Fmax, init. Load (lbs) max, init. Initial envelope curve Stabilized envelope curve Interstory drift (in) Figure 4: Typical initial and stabilized load envelope curves Failure of the walls was defined as a significant drop in load resistance between two successive phases or 0.8 F max, depending on what occurred first. Elastic stiffness, k e, was defined as the slope of the line passing through the origin and the point on each load envelope curve where the load equals 40 percent of F max. This stiffness represents a good estimate of the stiffness that shear walls will exhibit after being loaded a number of times at low to moderate magnitudes. An equivalent energy elastic-plastic curve, used for comparison purposes, was determined for the initial cycle and stabilized cycle envelope curves for each wall. This artificial curve, shown in Figure 5, depicts how an ideal perfectly elastic-plastic wall would perform and dissipate an equivalent amount of energy until failure. Therefore, the area under the equivalent elastic-plastic curve (EEPC) equals the area under the loaddisplacement envelope curve from zero drift to drift at failure, failure. The elastic portion of the EEPC contains the origin and has a slope equal to the elastic stiffness, k e. The

12 12 VPI&SU Report #TE-1997/-003 plastic portion of the EEPC is a horizontal line equal to F yield. Integrating the obtained load displacement curve for each wall, F yield was determined according to: F yield = ± failure 1 ke 2 2 failure A ke (1) where A is the area under the respective true load-displacement curve. Displacement at yield, yield, was defined as the displacement at the intersection of the elastic and plastic lines of the EEPC. This is the same definition used by Dolan and Johnson (1996a and b). The ratio of failure and yield provides the ductility D for each wall: failure D = yield (2) This definition of EEPC was also used in the monotonic tests, and is similar to that proposed in the sequential phased displacement test developed for ASTM (Dolan, 1994). The Structural Engineers Association of Southern California (SEAOSC, 1996) uses an alternate variable to quantify this behavior. Unit shear was determined using the equation: g F L = max (3)

13 VPI&SU Report TE where g is the unit shear, F max stands for wall capacity, and L is the sum of the widths of full-height sheathing panels. Consequently, the theoretical uplift is obtained by multiplying the unit shear g times the normal distance between applied load and foundation, which equals the wall height in this investigation. Fmax Load-displacement envelope curve Equivalent elastic-plastic curve Fyield Load (lb) F 0.8 F yield m a x 0.4 Fmax yield max Interstory drift (in) failure Figure 5: Typical equivalent energy, elastic-plastic curve RESULTS AND DISCUSSION Strength Table 3 lists load resistance values obtained from initial and stabilized envelope curves at capacity and at interstory drifts of 1.6 inch, 0.96 inch, and 0.32 inch. The walls with 2 feet corners attached show a fairly high variation. The sample size of two walls is not sufficient to make inferences whether the variation is due to statistical error. More

14 14 VPI&SU Report #TE-1997/-003 tests will be needed to obtain more reliable results. However, the values obtained from walls with 4 feet corners are almost equivalent. On average, walls with 4 feet corners reached higher ultimate capacities and higher loads at the inter-story drifts listed in Table 3 than walls with shorter, 2 feet long, perpendicular segments. The average reduction between initial and stabilized capacity was 17% (Std. Dev. =1.1%) and practically constant. Dolan and Heine (1997 b) found the same trend when investigating different anchorage conditions on full-scale timber framed shear walls loaded equivalently. The damage experienced by the walls is fairly uniform regardless of the length of perpendicular wall segments attached, implying that the performance of timber framed shear walls is primarily governed by the sheathing nails. Unit shear at ultimate capacity along with average unit shear values at ultimate capacity of the specimens with corner framing are presented in Table 4. Dolan and Heine (1997 b) tested 40 feet long and 8 feet tall timber framed walls constructed similarly to the walls presented in this report. However, no perpendicular wall segments were attached to these walls. Instead, the walls were tested with and without overturning restraints of type Simpson HTT22 tie-down anchors attached at each wall end. As with the tests presented here, no gravity load was applied.

15 VPI&SU Report TE Table 3: Initial cyclic and stabilized cyclic load resistance data Capacity 2ft corners Wall Specimen 4ft corners Wall 1 Wall 2 average Wall 1 Wall 2 average Initial SPD (kips) Stabilized SPD (kips) Stabilized/Initial Unit shear Load 1.60in. interstory drift Initial SPD (kips) - (1) - (1) - (1) (1) - (1) Stabilized SPD (kips) - (1) - (1) - (1) - (1) - (1) - (1) Load 0.96in. interstory drift Initial SPD (kips) Stabilized SPD (kips) Load 0.32in. interstory drift Initial SPD (kips) Stabilized SPD (kips) (1) Failure occurred before interstory drift was reached It should be noted that unit shear values were determined by equation (3) using actual test results. This equation does not account for the effects of wall length on unit shear. In other words, the assumption was made that unit shear remains constant and uniform with changing wall length, and the data reflects the observed unit shears. This is in fact not true since the shear is distributed to the wall in proportion to the wall segment stiffness, with the end segment, next to the corner, carrying a higher portion of the lateral load. The values from Table 4 are depicted in Figure 6 which illustrates the effect of corner framing on unit shear. Walls with 4 feet corners show somewhat higher ultimate unit shear values than walls without corners and overturning restraint. However, this

16 16 VPI&SU Report #TE-1997/-003 assumes that the shear load is distributed uniformly to the top of the wall, when in fact the shear load is distributed according to stiffness. Therefore, for this study to be complete, the results should be compared to straight shear walls that are 12 ft long. This indicates that the interaction of two mutually perpendicular walls provides restraint against overturning forces that in turn increases shear capacity. However, the average ultimate unit shear obtained from walls with 2 feet corners is somewhat lower than the unit shear for walls without overturning restraint. This is partly attributed to the difference in length between the specimens. In addition, the high variation of the values obtained from the walls with 2 feet corner framing and the small sample size may also be contributing for the observed difference. It is interesting to note the tested unit shear varied by only 115 plf or about 17%. Table 4: Cyclic capacities and unit shear values of walls with corner framing and long walls with and without tie-down anchors Capacity 12 ft Walls with corner framing Wall Specimen 2 ft corner (2) 4 ft corner (2) overturning no restraint (3) 40 ft walls no corner framing (1) overturning restraint (2) Initial SPD (kips) Unit shear (1) From Dolan and Heine (1997 b) (2) Average values out of two specimens 3) Values obtained from one specimen only

17 VPI&SU Report TE Unit Shear Unit Shear (lbs/ft) no tie-down anchors (40ft wall) 2ft corner no tie-down 4ft corner no tie-down anchors (12ft wall) anchors (12ft wall) tie-down anchors (40ft wall) Figure 6: Unit shear of walls with no tie-down anchors, tie-down anchors at the end and 2ft and 4ft corner segments, respectively Stiffness and Ductility Table 5 depicts elastic-plastic parameters obtained from the equivalent energy elastic-plastic curve. The same variation that occurred between the ultimate capacity values of the two walls with 2 feet perpendicular segments is reflected in the F yield values. As with the recorded ultimate capacities, F yield increased with increasing corner length. Elastic stiffness, defined as the secant slope passing through the origin and a point on the load-drift envelope curve where the resisted load equals 40 percent of capacity, was not significantly influenced by the length of the corner segments. Figure 7 shows clearly that within the elastic range the four initial envelope curves almost coincide. Stabilized stiffness was consistently higher than initial stiffness (Table 5). The reason is found in the definition of elastic stiffness. In the elastic range stabilized and initial envelope curves coincide. Since stabilized capacity will always be lower than initial capacity the secant of stabilized elastic stiffness will pass through a point on the envelope curve closer to origin where the curve has a steeper slope. Consequently, the slope of the secant is also steeper

18 18 VPI&SU Report #TE-1997/-003 and therefore, the stiffness is higher. The ratio of stabilized to initial elastic stiffness is therefore similarly constant as the ratio of stabilized to initial capacity. Table 5: Equivalent elastic-plastic curve parameters Wall Specimen 2ft corners 4ft corners Wall 1 Wall 2 average Wall 1 Wall 2 average 40ft Wall (1) No anchors F yield Initial SPD (kips) Stabilized SPD (kips) Stabilized/Initial Elastic Stiffness Initial SPD (kips/in) Stabilized SPD (kips/in) Stabilized/Initial Ductility Initial SPD ductility Stabilized SPD ductility) Stabilized/Initial (1) From Dolan and Heine (1997 b) Average initial envelope curves Load (lbs) ft one 4ft two 2ft one 2ft two Interstory drift (in) Figure 7: Average initial envelope curves of all four specimens

19 VPI&SU Report TE Ductility values were determined by dividing the interstory drift at failure by the interstory drift at yield load (Equation 2). Since drift at yield load is decisively dependent upon initial stiffness, stabilized ductility values in Table 5 are consistently higher than pertinent initial values, as expected. The average initial ductility values of both wall configurations are equivalent. However, considering Figure 7, the walls with 4 feet corner framing sustained higher loads beyond failure than the walls with the shorter corner segments, indicating a tougher (or more robust) structural response. The ductility values of both walls are significantly higher than the values of the 40 feet long straight wall without overturning restraint tested by Dolan and Heine (1997 b). This indicates that the interaction between two mutually perpendicular walls improve the overall seismic performance of the structure when loaded cyclically. Vertical End Stud Displacement The recorded compression and uplift displacement of the stud in each corner was corrected to compensate amplifications caused by the geometry of the LVDT fixture. The walls were assumed to rotate about the outer bottom edge of either wing wall depending on the direction of the load exerted on the wall. Table 6 lists the total displacements of both corner studs at ultimate capacity and at failure of the walls with 2 feet and 4 feet corner framing and of the 40 feet long straight walls tested by Dolan and Heine (1997 b).

20 20 VPI&SU Report #TE-1997/-003 Table 6: End stud displacement between positive and negative peak drifts during initial cycle of max load and failure 12 ft Walls with corner framing Wall Specimen 2 ft corner (2) 4 ft corner (2) overturning no restraint (3) 40 ft walls no corner framing (1) overturning restraint (2) Left end stud (LVDT max load Right end stud (LVDT max load Left end stud failure Right end stud (LVDT failure (4) 0.26 Theoretical capacity (1) From Dolan and Heine (1997 a) (2) Average values out of two specimens (3) Values obtained from one specimen only (4) Data not available due to failure mode The values suggest that perpendicular wall segments reduce the uplift by providing some hold-down effect. On average, total stud movement at capacity was reduced by 36 percent and 41 percent for walls with 2 feet and 4 feet corner framing, respectively, compared to a straight wall with no overturning restraint. General Observations It is remarkable that there were no typical damage signs of racking of the sheathing panels with corner framing. The taped joints between the drywall panels experienced no damage. The corner segments hindered the free rotation of the sheathing with respect to the framing. Basically the walls rotated as rigid bodies and eventually separated from the bottom plate. This behavior may be a result of the relatively short wall length of 12 feet. However, due to the hold-down effect provided by the corner framing the separation of the wall from the bottom plate started to occur at relatively high loads.

21 VPI&SU Report TE Only at the bottom plate nails tore and pulled through the sheathing. Nails simply withdrew along the bottom of the wing walls. There was no nail fatigue observed in the corner specimens, which is a significant change from the behavior observed by Dolan and Heine (1997 b) during SPD tests of the long and straight walls. The wing walls rotated about the nail furthest away from the corner of the specimens. CONCLUSIONS Four walls with different corner framing were tested under reversed cyclic loading. The measured wall capacities were compared with long straight walls tested by Dolan and Heine (1997 b). Based on these results the following conclusions can be drawn: 1) Corner framing generally provides a hold-down effect that increases wall capacity and ductility when compared to straight walls with no overturning restraint and no perpendicular walls attached. 2) The hold-down effect provided by corner framing is sufficient to provide for development of unit shear slightly less than, but comparable to, straight walls with hold-down devices. Straight shear walls that are 12 ft in length should be tested to quantify the effect of the corner walls better. 3) Walls with corner framing showed no apparent racking of the sheathing. The walls responded mainly through rigid body rotation until complete separation from the bottom plate occurred. No nail fatigue was observed and there were no signs of damage at joints between drywall panels.

22 22 VPI&SU Report #TE-1997/-003 RECOMMENDATIONS The following recommendations are given: Expand testing and develop an engineered approach for conventionally framed wood walls without overturning restraint devices (i.e. hold-downs). Perform tests combining gravity loads and corner framing effects while confirming a modified perforated shear wall design approach. REFERENCES Dolan, J.D., Proposed Test Method for Dynamic Properties of Connections Assembled with Mechanical Fasteners. ASTM Journal of Testing and Evaluation 22(6): Dolan, J.D. and Heine, C.P., 1997 a. Monotonic Tests of Wood-frame Shear Walls with Various Openings and Base Restraint Configurations. Virginia Polytechnic Institute and State University Timber Engineering Report TE Dolan, J.D. and Heine,C.P., 1997 b. Sequential Phased Displacement Cyclic Tests of Wood-frame Walls with Various Openings and Base Restraint Configurations. Virginia Polytechnic Institute and State University Timber Engineering Report TE Heine, C.P., Effect of Overturning Restraint on the Performance of Fully Sheathed and Perforated Timber Framed Shear Walls, thesis submitted in partial fulfillment of Master's of Science Degree in Timber Engineering. Virginia Polytechnic Institute and State University, Blacksburg, Virginia. Porter, M.L., "Sequential Phased Displacement (SPD) Procedure for TCCMAR Testing." Proceedings of the Third Meeting of the Joint Technical Coordinating Committee on Masonry Research, U.S. - Japan Coordinated Earthquake Research Program, Tomamu, Japan. Structural engineers Association of Southern California (SEOSC), Standard Method of Cyclic (Reversed) Load Test for Shear Resistance of Framed Walls for Buildings. SEAOSC, Whitten, CA.