PERFORMANCE-BASED SEISMIC DESIGN OF MID-RISE LIGHT-FRAME WOOD BUILDINGS

Transcription

1 PERFORMANCE-BASED SEISMIC DESIGN OF MID-RISE LIGHT-FRAME WOOD BUILDINGS John W. van de Lindt 1, David V. Rosowsky 2, Weichiang Pang 3, and Shiling Pei 4 ABSTRACT: Light-frame wood (woodframe) buildings are typically constructed four stories and below and thus their behaviour in high seismic regions was not studied extensively until the NEESWood Project from During that project, a design approach for direct displacement design (DDD) of mid-rise woodframe buildings was developed and validated experimentally at full-scale on a six-story shake table test in Miki, Japan. In addition to inter-story drift, rod elongation and shear transfer were also key in ensuring the design performance expectations were achieved. This paper systematically presents the design approach ranging from shear wall selection through DDD to free-body diagram calculations for hold down rod diameter selection for shear wall stacks within the building. The six-story Capstone building is used as the example from start to finish including quantitative experimental validation. KEYWORDS: Light-frame wood; Mid-rise building; performance-based seismic design 1 INTRODUCTION 123 Progress in the PBSD of woodframe buildings went little beyond general discussion amongst seismic wood researchers and product suppliers in the 1990 s. The CUREE-Caltech Woodframe Project focused on the improvement of woodframe codes and standards through a detailed and robust experimental, numerical, and documentation effort to examine current practices in the design and construction of woodframe buildings. The project resulted in numerous recommendations based on the results of a large team of researchers and practitioners in the late 1990 s and early 2000 s. Several key papers laying the foundation for the development of PBSD of woodframe buildings were e.g. Rosowsky [1] which developed allowable mass charts based on the desired peak inter-story drift using a numerical model developed within the CUREE-Caltech Woodframe Project (CCWP) entitled SAWS (Seismic Analysis of Woodframe Structures) as well as Filiatrault and Folz [2]who proposed the use of the Direct Displacement Design (DDD) originally developed by Priestley [3] for design of concrete structures. A series of fragility-based analysis and design methods followed from 2003 to John W. van de Lindt, Professor and Drummond Chair, Civil, Construction, and Environmental Engineering, University of Alabama, Tuscaloosa, Alabama, USA. jwvandelindt@eng.ua.edu 2 David V. Rosowsky, Professor and Dean of Engineering, Rensselaer Polytechnic Institute, Troy, New York, USA. rosowd@rpi.edu 3 Weichiang Pang, Assistant Professor, Civil Engineering Department, Clemson University, Clemson, South Carolina, USA. wpang@clemson.edu 4 Shiling Pei, Assistant Professor, Civil and Environmental Engineering, South Dakota State University, Brookings, South Dakota, USA. shiling.pei@sdstate.edu that focused on drift as the performance indicator for woodframe buildings (e.g. Ellingwood et al, [4]; Rosowsky and Ellingwood, [5]; van de Lindt and Walz, [6]). In 2005 van de Lindt introduced a damage-based concept for wood shear walls [6] which was later applied at the wall level by van de Lindt and Gupta [7] and system level by Liang et al. ([8] The NEESWood project [9] sought to develop a performance-based seismic design philosophy for mid-rise woodframe buildings through an extensive series of numerical simulations and full-scale system-level experiments (Filiatrault et al, [10]; van de Lindt et al, [11]). As part of the NEESWood project, Pang and Rosowsky [12] extended the DDD concept introduced by Filiatrault to become a viable performance-based design procedure and applied it to the design (Pang et al, [13] of a fullscale 14,000 square foot six-story woodframe building tested in Miki city, Japan (van de Lindt et al. [16]. To date, no technical documentation on the procedure to design a mid-rise woodframe building for a high seismic zone has been published, thus this paper documents and explains the technical procedure for the performancebased seismic design of mid-rise woodframe buildings from start to finish using the six-story Capstone test conducted in Miki city, Japan as the illustrative example.

2 Figure 1: DDD Flowchart for shear wall selection 2 PERFORMANCE EXPECTATIONS Initially, the performance expectations for mid-rise woodframe buildings was felt to be able to be approximated using the results of low-rise woodframe shake table tests. Although this was expected to be an approximation, the absence of damage at lower drift levels was notable during mid-rise woodframe tests. As a result of the NEESWood Capstone tests the performance expectation in Table 1 were proposed. The Performance-Based Seismic Design (PBSD) procedure explained herein and outlined in the NEESWood project is a multi-objective design, in which each performance objective is defined by a probability of not exceeding (NE t ) a target inter-story drift limit (θ lim ), conditioned on the occurrence of a specified level of seismic hazard (H), expressed as PNE ( θ < θlim H ) NEt (1) where θ is the inter-story drift and the term P NE (.) denotes the non-exceedance probability of the inter-story drift at a prescribed hazard level (H). In the PBSD framework, multiple design objectives can be specified. For example, a 1% drift limit was specified in the NEESWood project design Level 1 (Table 2) to minimize the damage to the designed building in a low intensity earthquake event. Since this is not a life threatening limit state the non-exceedance probability was set to 50%, essentially attempting to make this the performance at this hazard level, on average. In the NEESWood Capstone building the performance requirements were coupled with seismic hazard and Table 2 presents the hazard levels considered. It was assumed that the site was not adjacent to a fault and only hazard levels 1, 2, and 3 were considered in the design. Note that the non-exceedance probability for the level 3 design was 80%, whereas the median was targeted for levels 1 and 2. This 80% NE was selected to reflect the consequences of exceeding a drift at MCE intensity. Level 3 eventually governed the design for this building which is typical for relatively short period buildings. 3 SHEAR WALL SELECTION The determination of shear wall nailing schedules is performed using the DDD procedure [13] to meet four performance expectations (damage limitation, life-safety, far-field collapse prevention, and near-fault collapse prevention). Unlike the force-based design procedure where the vertical distribution factor is proportional to the story heights, vertical distribution of base shear in DDD is proportional to the design story displacements

3 relative to the ground. An overview of the DDD design process is shown in Figure 1. Table 1: Proposed Performance Expectations for Mid- Rise Woodframe Buildings Table 2: Design Levels for Mid-Rise Woodframe Buildings The DDD procedure begins with the specification of the target performance expectations. For the NEESWood Capstone building, the performance expectations shown in Table 2 were used. As previously stated, the nonexceedance (NE) probability for performance Level 3 (far field collapse prevention) was 80%. In order to design for a target NE probability of inter-story drift greater than the median, the design spectral value was adjusted upward to reflect the increase in the design NE probability using the non-exceedance probability adjustment factor, C NE. The C NE factor is assumed to be log-normally distributed with a median value of 1.0 and a logarithmic standard deviation, β R, which accounts for the uncertainty in the ground motions as well as the uncertainty associated with the design procedure. The β R for the Capstone Building was estimated to be 0.75 and the corresponding C NE factor, computed using the following equation, for Level 3 was 1.88[13]. 1 CNE = exp Φ ( NEt ) β R (2) 1 where Φ (.) is the inverse CDF of the standard normal distribution. Next, the design inter-story drift limit was also adjusted for the NE probability using the C NE factor. The target drift limit for the Capstone building at Level 3 was 4%. The design inter-story drift, adjusted to 50% NE probability, was 2.13% (4%/CNE). Using the equivalent 50% NE inter-story drift, the vertical distribution factors for base shear were determined as a function of the story seismic weight and floor displacement relative to the ground. Note that in force-based design, the vertical distribution factors for base shear are a function of the story weight and story height relative to the ground. Next, an equivalent single degree of freedom (SDOF) system was determined by computing the effective seismic weight and height. Using the effective height, a target displacement at the effective height, eff, was computed assuming a first-mode deformation. The DDD procedure is based on an equivalent linear elastic SODF system. To account for the nonlinear effect, a damping reduction factor, B ξ, was computed as a function of the ratio of secant stiffness at target displacement to initial stiffness: 4 B ζ = (3) 5.6 ln(100( ζ int + ζhyst )) where ζ int is the intrinsic damping (assumed to be 5%) andζ is the hysteretic damping is given by ; hyst ζ hyst 1.38K s Ko = 0.32e (4) The design base shear coefficient was then computed using the capacity spectrum approach: CNESXS B ξ Cc = min 2 (5) g CNES X 1 2 4π eff B ξ where S XS and S X1 are the design spectral acceleration values at short-period and 1-second period, respectively. For the hazard Level 3, the S XS and S X1 were taken as 1.5g and 0.9g, accordingly. Using Eqn. (3), the base shear coefficient for Level 3 was determined to be Once the base shear coefficient was obtained, the base shear, the equivalent static lateral forces, the required story shears, the overturning moments and the required story secant stiffness at each story were computed based on basic engineering principles [13]. The equivalent 50% NE inter-story drifts and the associated required story shears were used to define the demand points (or design points). A shear wall database which contained the shear wall backbone curve (shear strength versus drift) of shear walls with different panel perimeter nail spacing was used to determine the nail schedules for the wall lines. A portion of the shear wall database is shown in Figure 2. The nailing patterns for the shear walls on each floor were determined such that the story backbone curve was above the demand points. It should be noted that the effect of torsion was assumed to be minimal for the Capstone building of rectangular floor plan; hence, direct summation of the equivalent stiffness of full-height shear wall segments was used to generate the backbone curves.

4 Wall Height (m) 2.74 Backbone Force at Different Drift Levels (kn per m) Figure 2: A selection of displacement-based shear wall design table for unit wall width (per m). Secant Stiffness, K s (kn/mm per m) Wall Edge Nail K Type/ o F Spacing u (kn/mm Sheathing (kn per m) (mm) per m) Wall Drift Wall Drift Layer 0.5% 1.0% 2.0% 3.0% 4.0% 0.5% 1.0% 2.0% 3.0% 4.0% Standard (a) Midply (b) GWB (c) HOLD DOWN AND SHEAR TRANSFER With the shear wall configuration selected with DDD, the numerical model for the entire building system was constructed in SAPWood (a software program developed during the NEESWood project for nonlinear time history analysis of wood buildings) and subjected to a suite of biaxial earthquake ground motions scaled to the desired hazard levels. The numerical simulation yielded the maximum shear demand for each wall. The 80 th percentile values of the maximum demand from the all the simulations were then used to design the shear transfer for each wall component. The continuous anchor tie-down system was designed based on the 80 th percentile tie-down forces needed in the time history simulation for shear wall stacks to achieve simplified free-body-diagram equilibrium. The tie-down rods were sized so that the elongation of the rods for each story will not exceed 6 mm under the 80 th percentile demand. Note that a deformation requirement rather than strength requirement was imposed here because of the purpose of these tie-down rods, which is to provide adequate boundary restraint for shear walls to engage in shear deformation instead of rocking. A flow chart of this process is illustrated in Figure 3. Figure 3 presents a detailed example of seismic force demand calculation for a single ATS rod at the end of a shear wall. The conceptual free-body diagram (FBD) for an N-story shear wall stack is shown below in Figure 4 which was developed to generate tensile force demand for the ATS (rod) based on shear demands on the wall stack, which was obtained through NLTHA. Certainly this is a simplified approach to obtain the ATS forces, but it has been shown to be effective based on the test results. Other more advanced techniques (e.g. FEM) can be used if desired. Figure 4: Simple FBD to obtain ATS tension from shear forces Figure 3: Flowchart for sizing the ATS components Shear transfer requirements between the shear walls and floor diaphragms were calculated using a similar procedure to that of the continuous rod hold down system, but based directly on shear wall forces from the NLTHA without a FBD. This could also be accomplished using basic structural analysis techniques or a structural analysis software package. Strapping throughout the floor diaphragms was based on the same 80% non-exceedance shear values and then NDS nailing requirements were computed based on those demands.

5 5 VALIDATION BY EXPERIMENT The approach described briefly above was applied to the design of a six-story full-scale apartment building and the building (Figure 5) was constructed in Miki, Japan, home to the world s largest shake table facility known as E-Defense. The building was subjected to a series of shakes, which are described in detail in [14]. Table 3 shows the seismic intensity levels for the three tri-axial shakes. The test level 1 and level 2 were done on the same day and test level 3 was done separately. Recall from the earlier discussion that level 3 controlled the DDD. This largest shake represents a 2,500 year scaling of the Canoga Park, 1994 Northridge California earthquake, which resulted in the responses shown in Figure 3. The peak deformation for level 3 is shown in the bottom plots of Figure 3, with a peak of 211 mm. This is just over 1% drift for the building, which had an approximate 20 m height. The response to the level 2 and level 1 seismic intensity are shown in the middle and upper windows of Figure 6, respectively. Damage was limited to only the gypsum wall board (GWB) near the building openings as can be seen in Figure 7. No structural damage occurred and only damage to the GWB was found during post-shake inspection, meaning the building design resulted in performance that satisfied the PBSD procedure expectations. It should be noted that the damage was felt to be in-line with the new Table 1. Table 3: Seismic Intensity Levels usedduring the ExperimentalValidation. Figure 5: Six-story, 1350 m 2 Capstone building on the world s largest shake table (E-Defense) in Miki, Japan. Figure 6: Central roof displacement as a result of seismic intensity level 3. Figure 7: Typical damage observed near openings after the 2500 year shake 6 SUMMARY AND CONCLUSIONS A performance-based seismic design (PBSD) procedure to select shear walls, size hold down rods, and ensure shear transfer was summarized herein using the NEESWood Capstone building as an illustrative example. Experimental validation demonstrated that a mid-rise woodframe building can be designed to achieve target performance at high seismic intensity levels, thereby making it a realistic option in high seismic regions worldwide. The procedure is available for PBSD of woodframe buildings, but some basic improvements and modifications are still needed to ensure it can be implemented by practicing engineers. While significant progress toward robust and practical PBSD of mid-rise woodframe construction was made over the last five years, several key challenges to widespread implementation remain. In 2008 at a Wood Research Needs Workshop in Vancouver, Canada, van de Lindt outlined some challenges associated with development and particularly implementation of general performance-based design for woodframe buildings. Even in light of progress made in the last few years, two major challenges remain: (1) the need to further improve nonlinear time history models, and (2) the need to package PBSD of mid-rise woodframe buildings in a more designer-friendly format, i.e. with parallels to the National Design Specification (NSD) for Wood. Work is in progress in the United States to implement PBSD of woodframe buildings into earthquake engineering

6 practice as a working group within the Design of Wood Structures subcommittee of the American Society of Civil Engineers. As mentioned, one challenge to the implementation and widespread adoption of PBSD for mid-rise woodframe buildings is to present it in a format easily understood and applied by engineers. For example, non-exceedance probability at the 4% drift limit can be plotted against building height and design charts developed for each value of the non-exceedance probability adjustment factor, C NE, shown earlier in Figure 3. Given the building height and desired non-exceedance probability, engineers/designers could select the appropriate minimum value of C NE using these charts. Using design charts of this type, engineers/designers are able to specify a target drift limit as well as a target nonexceedance probability when using the simplified DDD procedure giving flexibility. The results could also be easily tabulated to be more consistent with the National Design Specification (NDS) for wood. ACKNOWLEDGEMENT The material presented in this presentation is based upon work supported by the National Science Foundation under Grant No. CMS (NEES Research) and CMS (NEES Operations). Any opinions, findings, and conclusions or recommendations expressed in this material are those of the author(s) and do not necessarily reflect the views of the National Science Foundation. The authors are grateful to the rest of the overall NEESWood project team including Andre Filiatrault, Rachel A. Davidson, and Michael D. Symans. Steven Pryor of Simpson Strong Tie collaborated on the test planning and test execution and is gratefully acknowledged. Thank you to NSF REU s Doug Allen and Kathryn Pfrefzschner, researchers Izumi Nakamura, Chikahiro Minowa, and Mikio Koshihara at the Uniersity of Tokyo. Two graduate students, Kazaki Tachibana and Tomoya Okazaki, contributed to the construction and instrumentation of the test specimen. Thank you also to Tim Ellis of Simpson Strong Tie Co. and David Clyne of Maui Homes USA. Technical collaborators beyond the authors affiliation included the U.S. Forest Product Laboratory, FP Innovations-Forintek Division, Maui Homes U.S.A, and Structural Solutions Inc Financial and in-kind product and personal donations were provided by Simpson Strong Tie, Maui Homes, B.C. Ministry of Housing and Social Development, Stanley Bostitch, Strocal Inc., Structural Solutions Inc., Louisiana Pacific Corp., Natural Resources Canada, Forestry Innovation Investment, APA-The Engineered Wood Association, American Forest and Paper Association, Howdy, Ainsworth, and Calvert Glulam. REFERENCES [1] Rosowsky, D.V. (2002). Reliability-based seismic design of wood shear walls, ASCE Journal of Structural Engineering, 128(11): [2] Filiatrault A. and Folz B. (2002). Performancebased seismic design of wood framed buildings, ASCE Journal of Structural Engineering, 128(1): [3] Priestley, M.J.N. (1998). Displacementbasedapproachesto rational limitstates design ofnewstructures. KeynoteAddress, Proceedingsofthe 11th European Conference on Earthquake Engineering, Paris, France. [4] Ellingwood, B. R., Rosowsky, D. V., Yue, Li S., and Kim, J. H. (2004). Fragility Assessment of Light- Frame Wood Construction Subjected to Wind and Earthquake Hazards, ASCE Journal of Structural Engineering, 130(12): [5] Rosowsky, D.V. and Ellingwood, B.R. (2002). Performance-based engineering of wood frame housing: a fragility analysis methodology, ASCE Journal of Structural Engineering, 128(1): [6] van de Lindt, J.W. (2005). Damage-Based Seismic Reliability Concept for WoodframeStructures, ASCE Journal of Structural Engineering, 131(4): [7] van de Lindt, J.W. and R. Gupta. (2006). Damage and Damage Prediction for Wood Shearwalls Subjected to Simulated Earthquake Loads, ASCE Journal of Performance of ConstructedFacilities, 20(2): [8] Liang, H., Y-K Wen, and G.C. Foliente. (2011). Damage Modeling and Damage Limit State Criterion for Wood-frame Buildings Subjected to Seismic Loads. Journal of Structural Engineering, 137(1), [9]van de Lindt, J.W., D.V. Rosowsky, A. Filiatrault, M.D. Symans, R.A. Davidson. (2006). The NEESWood Project: Progress on the Development of a Performance-Based Seismic Design Philosophy for Mid-Rise Woodframe Construction. Proc of the 2006 World Conference on Timber Engineering, Portland, OR. [10]Filiatrault, A, I. Christovasilis, A. Wanitkorkul, and J.W. van de Lindt. (2010). ExperimentalSeismic Response of a Full-Scale Light-Frame Wood Building, ASCE Journal of StructuralEngineering, 136(3): [11]van de Lindt, J.W., S. Pei, S.E. Pryor, H. Shimizu, and H. Isoda. (2010). Experimental SeismicResponse of a Full-Scale Six-Story Lightframe Wood Building, ASCE Journal of StructuralEngineering, 136(10): [12] Pang W. and Rosowsky D. (2008). Performancebased seismic design of six-story woodframe

7 structure, Structural Engineering International: Journal of the International Association for Bridge and Structural Engineering (IABSE), 18(2): [13] Pang, W., Rosowsky D.V., S, Pei, and J.W. van de Lindt. (2010). Simplified Direct Displacement Design of a Six-Story Woodframe Building and Pre- Test Performance Assessment, ASCE Journal of Structural Engineering, 136(7): [14] Pei, S., J.W. van de Lindt, S.E. Pryor, H. Shimizu, H. Isoda, and D. Rammer. (2010). "Seismic Testing of a Full-Scale Mid-rise Building: The NEESWood Capstone Test., NEESWood Project Report NW-04, 532pp.Available on-line at