SEISMIC DESIGN OF A MULTI-STORY CROSS LAMINATED TIMBER BUILDING BASED ON COMPONENT LEVEL TESTING

Transcription

1 244 SEISMIC DESIGN OF A MULTI-STORY CROSS LAMINATED TIMBER BUILDING BASED ON COMPONENT LEVEL TESTING Shiling Pei 1, Marjan Popovski 2, John W. van de Lindt 3 ABSTRACT: Cross laminated timber (CLT) is a new type of timber structural system that has just been introduced in North America, but has been used successfully in Europe for over 20 years. There have not been any notable data sets developed on the performance of tall CLT buildings during major earthquakes and there are no seismic design criteria currently in place for CLT structures. Shake table testing of walls, assemblies and multi-storey CLT structures have been performed, and development of dedicated nonlinear numerical models has been initiated.. In order to provide the necessary information to develop practical seismic design procedure for mid-rise CLT buildings, a number of CLT walls with different geometry and connector configurations were tested at FPInnovations, Canada to obtain the subassembly level hysteresis. Then, a simplified numerical model to predict the reverse cyclic behaviour of CLT walls was developed and calibrated with the test results. Using the wall component capacity predicted by the model, a 10-story CLT building was designed using a performance-based seismic design procedure known as direct displacement design (DDD) to enable the structure to remain under the predefined drift limits at various seismic hazard levels. This design procedure was adopted in a recent research effort to successfully design a 6-story light frame wood building tested at Japan s E-Defense shake table in In the present study described in this paper, the 10-story CLT building was designed with 80% non-exceedance probability of remaining below 4% inter-story drift when subjected to a maximum credible earthquake intensity level (2500 year return period) for the City of Los Angeles, California. The DDD of the building was refined and verified with nonlinear time history simulation using a suite of bi-axial ground motions scaled to the predefined hazard levels. Based on the performance-based design results and laboratory testing of individual CLT shear walls, a response modification factor (R-factor) is proposed for structures with CLT wall components according to current force-based design approach (i.e. ASCE 7), thus providing quantitative insight into CLT design using traditional design procedures in North America. KEYWORDS: Cross laminated timber, performance based seismic design, multi-story, numerical model 1 INTRODUCTION 123 In Europe, a structural system known as Cross Laminated Timber (CLT) was introduced approximately 20 years ago for panelized construction of mixed use and light-commercial buildings. The CLT building components can be fully prefabricated in a factory environment. CLT can be made with lumber as small as 1x4 inch nominal. When small lumber is layered together to form the CLT panels, the final thickness of the panel is such that the building has better fire 1 Shiling Pei, Assistant professor, South Dakota State University, Brookings SD, U.S.A. shiling.pei@sdstate.edu 2 Marjan Popovski, Principal Scientist, FP Innovations, Vancouver, Canada. marjan.popovski@fpinnovations.ca 3 John W. van de Lindt, Professor and Garry Neil Drummond Endowed Chair in Civil Engineering, University of Alabama, Tuscaloosa AL, U.S.A. jwvandelindt@eng.ua.edu resistance than light-frame wood, essentially similar to heavy timber construction. The CLT system could be classified as Heavy Timber (HT) in the International Building Code [1] and can be a viable option for midrise buildings particularly.. However, most of the existing tall CLT buildings do not consider seismic load effects as they are not located in active seismic regions. In fact the seismic resistance of CLT components and systems has only been studied by a handful of researchers around the world. The current approach used to construct multi-story CLT buildings relies on mechanical connectors to rigidly connect the wall and floor panels together. Results from quasi-static tests on CLT wall panels showed that the connection layout and design has a strong influence on the overall behavior of the wall [2] and the resulting system can be very stiff [3]. At the component level, quasi-static monotonic and cyclic tests have been carried out on CLT walls to study the influence of boundary

2 245 conditions, magnitudes of vertical load, and the type of anchoring systems (e.g. [4] [5] [6]) on performance and capacity. It was shown that CLT panel walls can exhibit significant levels of ductility if the boundary condition is set up to allow rocking of the wall panels. At the system level, a shake table test of a seven-story CLT building was conducted by Ceccotti and colleagues at Japan s E- Defense facility in Miki, Japan, showing that the structural panel members remained essentially elastic and the accelerations in the higher levels were on the order of 4g, compared to the earthquake input acceleration of less than 1g [6]. Finally, a recent handbook published by FPInnovations in Canada summarizes recent development and practice in CLT design and construction [7]. Figure 1 shows an example of the CLT wall test conducted. Detailed description of the FPI test program and specimen can be found in [6]. The impact of Table 1: CLT wall tests conducted at FPInnovations The objective of the study summarized in this paper was to (1) perform a performance-based seismic design for a 10-story CLT apartment building which utilizes the ductility of CLT panel walls observed in wall level component tests, and (2) derive an appropriate strength reduction factor (R-factor) for force-based design procedures to achieve damage free performance during a MCE (Maximum Credible Earthquake) level event. In this study, a PBSD procedure termed the simplified Direct Displacement Design (DDD) developed in the NSF funded NEESWood project was used [8]. The load resisting characteristics of the CLT walls during cyclic loading were evaluated based on component tests performed at FPInnovations. Based on the PBSD result, a response modification factor (R factor) for force-based design procedure is calibrated approximately to obtain a design equivalent to the PBSD. 2 WALL LEVEL CALIBRATION 2.1 CLT WALL TESTS In order to perform PBSD, the nonlinear deformationresistance characteristics of CLT walls must be obtained. In this study, a nonlinear CLT wall model was developed and calibrated based on wall component data from FPInnovations. Figure 1: Behaviour of Wall 11 during the testing

3 246 multiple factors including panel geometry, bracket type, fastener type, and gravity loading on the lateral response of CLT walls was evaluated using those tests. A brief summary of the tested CLT walls and which test data was used in this study is shown in in Table KINEMATIC MODEL Because the shear deformation of the CLT wall panel itself is not significant compared to the deformation at the panel-to-floor or inter-panel connections, a simplified model was developed that assumes all CLT wall lateral deformation is a result of the CLT panels rotating as a rigid body about the corners, as shown in Figure 2. The lateral resistance of a CLT wall is essentially a scaled summation of the load-slip resistance of all the connectors engaged in the rocking movement of the wall. The scale factor for each connector is a function of their location and the geometry of the panel. The resistance F at lateral displacement D may be calculated as: (1) Figure 2: CLT wall kinematics model Figure 3: Hysteretic model for CLT connections 2.3 PARAMETER CALIBRATION The back-calibration procedure used in this study can be described as follows. The tested CLT walls were modelled numerically using the simplified model and were subjected to the same displacement protocols used in the experimental tests. Then the model hysteresis was compared with the experimental measurements. The parameters for nails, screws, and hold-down connections were adjusted systematically until the model closely matched the observed experimental response. As a result from this calibration, a group of connector parameters were obtained. These connector parameters were then used to develop the hysteretic responses for any given CLT wall configurations (including configurations different than those tested) based on the kinematic assumption illustrated in Figure 2. These models were also used to develop the design resistance values for the CLT design tables. Figures 4 and 5 illustrated the comparison between calibrated model prediction and test results for two different wall configurations. Note the effect of gravity is considered in the calibrated model as well (Figure 5). All parameters in equation (1) can be obtained from the wall configuration (as shown in Figure 2) except for the connector resistance. In this study, the connector hysteresis was assumed to follow the CUREE 10- parameter model, which has been widely adopted for wood frame shear wall and connection modelling [9]. The behavior of the model and each control parameter is shown in Figure 4. The parameters for each type of connector were computed in this study by backcalibration using the wall test results. Figure 4: Calibrated model compared with test (no gravity)

4 247 and DA stands for Double All, meaning all brackets in the panel are double sided. Based on Figure 6, the designer can specify, for example, the 3DE configuration for the entire story, etc. It should be noted that for the case of the wall panel with only 2 brackets, configurations DE and DA are identical. The maximum number of 16d nails that can be put in a single bracket is 6, due to the limitation of the bracket to floor connection strength. These typical walls will be used later in the design of the 10-story CLT building. Figure 5: Calibrated model compared with test (with gravity) The calibrated hysteretic parameters for three types of connectors are listed in Table 2, including the Simpson Strong-tie HTT-16 hold-downs installed at corners of CLT walls and 16d spiral nails with D=3.9mm and L=89mm with Simpson Strong Tie 90mm x 105mm x 105mm Bracket commonly used in CLT construction in Europe. The parameters listed for the 16d nails are for a single nail connection. The 16d with step joint is the equivalent nail parameter to be used for multiple panelled walls with step joints. With the numerical model and connector parameters calibrated, the loaddeformation curve or the hysteresis curve for any CLT wall configuration with specific connectors can be developed. This provides a useful tool to generate CLT wall backbone curves that will be used in PBSD and tables with lateral load design values that are needed for force-based design of CLT structures. Table 2: Calibrated CLT connector parameters Connector Type Hysteretic Parameters (N, mm) K0 r1 r2 r3 r4 HTT D-SN D-SN+step joint F0 F1 X a b HTT D-SN D-SN + step joint TYPICAL WALL CONFIGURATIONS Several typical CLT wall configurations were considered in this study as shown in Figure 6. It is assumed that a structural CLT wall in the multi-story building will have 2, 3, or 4 brackets attached at the bottom of the wall, providing connection between the wall and the floor panel. The size of a single panel can vary from 0.96 m (4 ft) to 1.83 m (6 ft). It is assumed that for walls longer than 1.83m, multiple1.22 m (4 ft) panels are to be combined together and each panel will have bracket configuration as shown in Figure 6. The notation S stands for Single sided brackets for each location, DE stands for Double sided brackets at the End of the panel, Figure 6: Typical wall configurations 3 DIRECT DISPLACEMENT DESIGN (DDD) OF CLT BUILDING 3.1 CAPSTONE BUILDING FLOOR PLAN The floor plan of the NEESWood Capstone structure, a six-story light frame wood apartment building tested at Japan s E-Defense shake table in 2009 [10], served as the model to develop the 10-story CLT structure, i.e. similar floor plan. The elevation and floor plan are presented in Figure 7. The building foot print is about 12 x 18 meters, with total height of 27.4 meter. All floor plans are identical except for the top story, where a penthouse unit may be integrated. The seismic weight of the building was assumed to be 2.2kN/m 2 for the first story, 1.4kN/m 2 for the roof, and 2.1kN/m 2 for all other stories. The total building weight is 4,537kN (462 metric tons). Given the floor plan, the wall selection is constrained in that only a limited amount of wall segments can be placed in each story. The design began by identifying the total usable wall segment length in each direction from the architectural floor plan. The numbers in Figure 7 show the maximum amount of CLT wall panel segments one can put in any particular line of the floor plan, which may or may not be fully utilized in the design process as the designer may choose to select some walls as non-structural partition and only apply minimal connections. The exterior walls are shown with window openings removed from the wall line. However, for CLT panels the windows are typically pre-cut into

5 248 the wall panel, so the window opening will not significantly affect the strength of the outside wall segments. For the interior, the door openings do interrupt the CLT walls (due to the height of the door openings) so these walls are broken into smaller segments. Note that the X direction is the longitudinal direction in the floor plan while the Y direction is the shorter direction. curve. The backbone curve for CLT walls used in the design can be readily obtained using the simplified model and calibrated connector parameters obtained earlier. Thus DDD can be utilized to design the CLT building presented in this paper. Detailed description of the DDD procedure can be found in [12]. The basic philosophy of DDD is to identify the required story lateral resistance at prescribed drift levels enabling the designer to select shear walls that satisfy this requirement. Three performance objectives were adopted in the design of the CLT building, limiting the maximum inter-story drift under different hazard levels. Because the inter-story drift level correlates well with seismically induced damage to a building, the performance objectives outlined in Table 3 will also ensure minimal damage during these earthquake events. Table 3: Performance Objectives Seismic Hazard Performance Expectations Inter-story Drift Non-exceedance 50%/50yr 1% 50% 10%/50yr 2% 50% 2%/50yr (MCE) 4% 80% Based on the performance objectives, the target point for the backbone curve for each story was identified using DDD. Design of the CLT building was conducted by choosing the CLT wall configuration for each story to produce a backbone curve that will satisfy (be larger than) the corresponding target point. The CLT walls selected for each story in both directions are listed in Table 4. The resulted backbone curves for all stories were plotted in Figure 8, together with the target points for these stories resulting from DDD. Table 4: Wall selection based on DDD Figure 7: CLT building architectural plan and wall segments location 3.2 PERFORMANCE OBJECTIVE AND DIRECT DISPLACEMENT DESIGN Direct Displacement Design (DDD [11]) was the design approach employed in the NEESWood research project to design the NEESWood Capstone building for prescribed drift limits under different levels of seismic intensity. It was validated through full scale shake table test of a six-story light frame wood building. In fact, the DDD method can also be applied to any lateral resistance system which has a clearly defined backbone CLT walls in X-direction 3.66 m 4.88 m 6.1 m ST # Con. # Con. # Con DA 2 4DA 8 4S 2 2 4DA 2 4DA 8 4S 3 2 4DA 2 4DA 7 4S 4 2 3DA 2 4DA 7 4S 5 2 3DA 2 3DA 7 3S 6 2 3DA 2 3DA 7 3S 7 2 4S 2 3DA 7 3S 8 2 2S 2 2S 7 3S 9 2 2S 2 2S 7 2S S 5 2S CLT walls in Y-direction 1.53 m 1.83 m 2.44 m 6.1 m ST # Con. # Con. # Con. # Con DA 2 4DA 14 4S 1 4DA 2 4 4DA 2 4DA 14 4S 1 4DA 3 4 4DA 2 4DA 14 4S 1 4DA 4 4 3DA 2 3DA 14 3S 1 4DA 5 4 3DA 2 3DA 16 3S 1 3DA 6 4 3DA 2 3DA 14 3S 1 3S 7 4 3DA 2 3S 14 3S 1 3S 8 4 2S 2 2S 14 2S 1 3S 9 2 2S 2 2S 10 2S 1 2S S 8 2S 0 -- In Figure 8, note that only targets for the level 3 design are shown in the Figures since this case controls the

6 249 design. From the plots, one can see that all of the backbone curves exceeded (are higher) the DDD target points associated with them, both in the X and Y directions. The design of the CLT building following DDD is completed once all CLT walls are selected. Practically, the other details of the building, such as hold-down details, still need to be designed. Additional PBSD procedures will be needed to design these details [13]. In this study, it is assumed that there will be adequate tie-down and over-turning restraints for each story thus the design will only focus on the lateral force resisting component. Generally, it can be used to predict non-linear building system seismic responses given the hysteretic response of the components, which makes it suitable for the purposes of this study. The model for as-designed CLT buildings was constructed in SAPWood and subjected to a suite of earthquake ground motion records scaled to the predefined hazard levels. Maximum inter-story drift within the building from each simulation was extracted to develop the distribution of the building maximum drift for that corresponding hazard level. The probability of exceedance associated with the design target can be evaluated by plotting the empirical cumulative distribution function curve based on all drift samples within that hazard level. The ground motion suite recommended for use in the U.S. Federal Emergency Management Agency Document P-695 [15] was adopted in this study, which includes 22 bi-axial far-field ground motions scaled to three target hazard levels. The response spectra of all these ground motions scaled to the Maximum Credible Earthquake (MCE, level 3) hazard level is shown in Figure 9. These bi-axial ground motions were also rotated by 90-degrees and applied to the model building. The building natural period estimated through the numerical model is 1.12 sec. At each performance level, the building was subjected to a total of 44 ground motions. The maximum inter-story drifts of the structure at any story and in either direction were recorded and rank-ordered as empirical cumulative distribution function curves shown in Figure 10. The PBSD objectives are also shown in Figure 10 as square shaped points for all three hazard levels. It can be seen that the building satisfied all performance requirements, i.e. the PNE (probability of non-exceedance) value for the designated drift level is equal to or higher than the performance requirements. Figure 8: DDD target points and design backbone curves 4 PERFORMANCE EVALUATION OF THE BUILDING WITH A NUMERICAL MODEL 4.1 SAPWood PROGRAM The CLT building designed using simplified DDD was subjected to different levels of seismic hazard through numerical simulation to verify that the design objectives have been achieved. The analysis was conducted using the software program Seismic Analysis Package for Woodframe Structures (SAPWood) [14]. SAPWood is a numerical tool specially developed to conduct nonlinear time history analysis for wood frame buildings but uses general enough hysteresis models that it has been used for modelling other types of system as is being done in the present case. It was part of the deliverables from the NEESWood project and has been validated by numerous component and system level shaking table tests, including the 6-story NEESWood Capstone building. Figure 9: Response spectra of ATC63 ground motions scaled to MCE hazard level

7 250 design, i.e. the higher factor one uses, the higher the computed R factor. Since at this point the design values for CLT walls are not defined neither in US nor in Canada, it was decided to do the analyses with the design level equal to the ultimate divided by a factor of 2.5; that is utilizing only 40% of the wall ultimate strength in the design. The resulting design resistance values for standard CLT wall configurations using the 16d spiral nails is shown in Table 5. Table 5: CLT wall design resistance values (kn) Figure 10: Performance objective target points and simulated structural performance 5 APPROXIMATE RESPONSE MODIFICATION FACTOR R FOR THE CLT BUILDING The equivalent lateral force procedure (ELFP) is one of the most commonly used design procedures for seismic design of multi-story buildings in many force based design codes (e.g. ASCE7-10 [16], NBCC [17]). The lateral force demand for each story is obtained by calculating the total base shear, and re-distributing it to each story. This study utilized the recommended ELFP procedure in the 2010 ASCE7 to calculate the level of story shear forces required in traditional force-based design for the CLT Capstone building. The objective of this analysis is to identify a suitable response modification factor (R-factor in ASCE7) to be used in the ELFP design calculation, so that the force-based design will result in a final design will be similar to what has been obtained from PBSD procedure earlier. As the design using PBSD has been verified by the nonlinear time history simulation using the state-of-the-art computer model, it is assumed that the codified design using the calibrated R-factor will lead to satisfactory building performance. In other words, when subjected to an MCE event in Los Angeles, California, the CLT building will have an 80% chance of not exceeding 4% inter-story drift (experience only minor damage) as was also observed in the NEESWood Capstone Building shaking table tests. For light frame wood buildings, the selection of wood shear walls can be conducted based on the codified shear wall design capacity tables (NDS Wind and Seismic Supplement) once the story shear demand is obtained using ELFP. The design resistance values for CLT walls in this study were developed in a similar way to those for wood-frame shear walls, which is by taking the ultimate strength of the CLT wall backbone curve under monotonic pushover test (which in this case was generated using the simplified model and the parameters obtained earlier), and dividing it by a factor. Note that this factor will eventually affect the R factor used in Bracket# Config S DE DA S DE DA S DE DA 0.92 m 1.22 m 1.53 m 1.83 m 2x1.22 m 3x1.22 m 4x1.22 m 5x1.22 m With the CLT wall selection for the building already determined from PBSD, the equivalent force-based design storey shear resistance can be obtained by adding the resistance of all the walls at that storey based on individual wall resistance listed in Table 5. This step can be easily performed for each story in both the X and Y directions. The resulting minimum story shear resistance (minimum between the two directions) is the demand that should be produced by force-based design procedure, i.e. ELFP with the appropriate R factor. The required story resistance is listed in Table 6 in the Target column. Following the ELFP in section 12.8 of ASCE 7-10, the Importance Factor (I) was taken as 1.0; the building period was calculated based on empirical formula (ASCE7) to be 0.58 sec. The seismic hazard map design values for the city of Los Angeles were obtained from the USGS Design maps application (conforms to ASCE 7-10) as S DL =0.57g and S DS =1.62g. The R value was changed manually until the final resistance distribution matched the target resistance. The calibrated R factor and the detailed calculation results are listed in Table 6. Table 6: ELFP design of the 10-story CLT building Cs I Tn SDL SDS

8 251 Figure 11 illustrates the impact of the variation of the R- factor to the storey shear distribution and how will the shear resistance profile compare to the target one. The story resistance of the PBSD design configuration for stories 1 through 10 was plotted along with the resistance calculated based on the ELFP, with different R values. The dashed lines approximately represents the appropriate lower and upper bounds for the R values equal to 3 and 5.5, with the line in between representing the calibrated chosen R-factor (4.3) in this study. Note that the bounds presented here are not an indication that the building should be designed using these R factors, but just to illustrate the sensitivity of story demand distribution to the R factor. In fact, the numerical simulation indicated that the controlling drift level occurs at higher stories. It is critical for force based design to capture the PBSD demand at higher stories. If the design only satisfies the lower stories, such as the case when designing with R=5.5, such design will not satisfy the drift requirement based on numerical simulation results at the upper floors. Figure 11: Story shear demand distribution with different R factors 6 CONCLUSIONS AND DISCUSSION Based on the analysis conducted in this study, CLT is a viable option for mid-rise wood buildings up to ten stories in moderate to high seismic regions, given the buildings are correctly designed. It may also be reasonable to expect only limited damage under MCE level earthquake excitation in high seismic regions. By adjusting the response modification factor, equivalent force-based design that will meet the PBSD performance objectives can be developed. Based on the results of this study a value of R=4.3 can be assigned to the analysed building and similar symmetrical buildings in ASCE7 as examples of structures with CLT wall panels that utilize spiral nails in the brackets, provided that the design values for such walls implemented in the material design standard are similar to the values included in Table 5. The holddowns and adequate overturning restraints should always be installed at both ends of a wall in order to ensure the desired performance. From the comparison between the story shear demand from ELFP and PBSD, it appears that over-design for lower stories in the force-based design procedure will result if the PBSD performance objectives need to be satisfied using current ELFP method. It should be kept in mind that both the design resistance value and the actual as-built resistance value were simulated indirectly in this study through a numerical model calibrated with limited number of wall level tests. However, the conclusions on the R factors drawn from this study are believed to be representative of a typical CLT system. Also, it should be noted that the results in this study are based on the assumption that the walls in the system will be installed in a way that enables them to rotate as in the tests conducted and generate similar backbone curves. If the boundary conditions of the walls differ significantly to the ones used in the analyses (the walls are not allowed to rock), the backbone curves of such wall configurations may be different and the results of this study might not fully apply. However, the kinematic assumption used in this study should be valid for the inter-story range (0~3%) of interest in this study. Thus the results are reliable unless one needs to extrapolate the conclusions to very large inter-story drift levels. In the future, if component testing is conducted with different boundary conditions and backbone curves are obtained for CLT walls at large deformation range, the same procedure utilized in this study can be used to derive the R-factor for that situation. Finally, the response modification factor computed in this study was based on one ten-story building with a given floor plan and one type of fasteners used in the brackets and the hold-downs. A variation of the R value may exist if different fasteners are used in the brackets and the hold-downs. Such variation may also exist as a function of the floor plan and the building height. If buildings with different heights are analysed, it is unlikely that the R-factor will change significantly, provided that the building floor plans remain symmetrical in both directions. However, it is recommended that further studies with a wider scope look into issues related to R-factors for structures with different archetypes and non-symmetrical floor plans according to FEMA P-695 guidelines. REFERENCES [1] International Code Council, (2009). International Building Code International Code Council. [2] Ceccotti, A., Follesa, M., Lauriola, M.P., Sandhaas, C. (2006) Sofie Project Test Conference on Earthquake Engineering and Seismicity, Geneva, Switzerland. [3] Lauriola, M.P., Sandhaas, C. (2006) Quasi-Static and Pseudo-Dynamic Tests on XLAM Walls and Buildings COST E29 International Workshop on Earthquake Engineering on Timber Structures, Coimbra, Portugal.

9 252 [4] Dujic B., Aicher S. Zarnic R. (2006) Testing of Wooden Wall Panels Applying Realistic Boundary Conditions Proceedings of the 9th World Conference on Timber Engineering, Portland, Oregon, USA. [5] Dujic, B., Pucelj, J., Zarnic, R. (2004) Testing of Racking Behavior of Massive Wooden Wall Panels Proceedings of the 37th CIB-W18 Meeting, Edinburgh, Scotland. [6] Popovski, M., Schneider, J., Schweinsteiger, M. (2010) Lateral load resistance of cross-laminated wood panels World Conference on Timber Engineering 2010, Trentino, Italy. [7] Popovski, M., Karacabeyli, E., Ceccotti, A. (2011) Seismic Performance of Cross-Laminated Timber Buildings - Chapter 4 CLT Handbook - Cross- Laminated Timber, FPInnovations Special Publication SP-528E, Canadian Edition. [8] Pei, S., van de Lindt, J.W., Pryor, S.E., Shimizu, H., and Isoda, H. (2010) Seismic testing of a full-scale six-story light-frame wood building: NEESWood Capstone test, NEESWood Report NW-04. [9] Filiatrault, A. and Folz, B. (2002). Performancebased seismic design of wood framed buildings, ASCE Journal of Structural Engineering, 128(1): [10] van de Lindt, J.W., Pei, S., Pryor, S.E., Shimizu, H., and Isoda, H. (2010) Experimental seismic response of a full-scale six-story light-frame wood building ASCE Journal of Structural Engineering, 136(10): [11] Pang, W., Rosowsky, D.V., van de Lindt, J.W., and Pei S.(2009) Simplified Direct Displacement Design of Six-story NEESWood Capstone Building and Pre-Test Seismic Performance Assessment, NEESWood Report NW-05. [12] Pang, W., Rosowsky, D.V., Pei, S., and van de Lindt, J.W. (2010) Simplified Direct Displacement Design of Six-story Woodframe Building and Pretest Seismic Performance Assessment. ASCE Journal of Structural Engineering. 136(7): [13] van de Lindt, J.W., Rosowsky, D.V., Pang, W., and Pei, S. (2012) Performance-Based Seismic Design of Mid-Rise Woodframe Buildings Journal of Structural Engineering, Accepted. [14] Pei, S., and van de Lindt, J.W. (2011) Seismic Numerical modeling of a six-story light-frame wood building: Comparison with experiments Journal of Earthquake Engineering, 15(6): [15] FEMA (2009) Quantification of building seismic performance factors: FEMA P695 Federal Emergency Management Agency. [16] ASCE (2010) Minimum design loads for buildings and other structures American Society of Civil Engineers. [17] NBCC (2010). National Building Code of Canada. National Research Council of Canada.