EFFECTS OF GROUND MOTION DURATION ON THE SEISMIC PERFORMANCE OF A TWO-STOREY BALLOON-TYPE CLT BUILDING by Maral Jafari Bachelor’s Degree in Civil Engineering, University of Bojnord, 2011 Master’s Degree in Geotechnical Engineering, Kharazmi University, 2015 THESIS SUBMITTED IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF MASTER OF APPLIED SCIENCE IN ENGINEERING UNIVERSITY OF NORTHERN BRITISH COLUMBIA June 2022 © Maral Jafari, 2022 Abstract The widespread availability of Cross-laminated timber (CLT) provides opportunities to extend the use of wood beyond traditional low-rise residential construction. Although previous studies have shown that ground motion duration impacts the collapse risk of structural systems, duration effects are not explicitly accounted for in current building codes, and information on the impact of ground motion duration on the seismic performance of CLT buildings is not available. This study aims to quantify the effects of long duration ground motions on a newly constructed two-storey balloon-type CLT building located in Vancouver, Canada. A three-dimensional numerical model of the building was developed in OpenSees. The shear wall and connections models were validated with test data. 24 pairs of long and short duration records with approximately the same amplitude, frequency content and the rate of energy build up were used for nonlinear dynamic analyses. The building was subjected to the earthquakes in its long (weak) direction. To assess the building’s collapse capacity, fragility curves were developed based on incremental dynamic analysis. At design intensity level, ground motion duration has been shown to not be a critical factor as the difference between inter-storey drift ratio under the two sets of records was negligible. However, compared with short duration motions, long duration motions increased the probability of collapse by 9% due to the larger number of inelastic cycles. Further research should evaluate the effect of ground motion duration on the seismic performance of taller CLT buildings. i Table of contents Abstract .............................................................................................................................. i Table of contents .............................................................................................................. ii List of tables .................................................................................................................... iv List of figures ................................................................................................................... v Acknowledgements ........................................................................................................ vii 1 2 Introduction .............................................................................................................. 1 1.1 Cross-laminated timber lateral load resisting systems ....................................... 1 1.2 Seismicity........................................................................................................... 1 1.3 Research need .................................................................................................... 2 1.4 Objectives .......................................................................................................... 2 1.5 Thesis organisation and scope ........................................................................... 3 Literature review ....................................................................................................... 4 2.1 Background information .................................................................................... 4 2.2 CLT construction ............................................................................................... 5 2.2.1 CLT as a structural material ....................................................................... 5 2.2.2 CLT lateral load resisting systems ............................................................. 5 2.2.3 Connections for CLT shear walls ............................................................... 7 2.2.4 Seismic performance of CLT structures ................................................... 10 2.3 2.3.1 Seismicity in British Columbia ................................................................ 12 2.3.2 Characteristics of ground motions ............................................................ 13 2.3.3 Effect of ground motion duration on structural performance ................... 15 2.3.4 Seismic analysis and damage measures ................................................... 18 2.3.5 Seismic design of CLT LLRS in Canada ................................................. 19 2.4 3 Seismic performance of structures ................................................................... 12 Summary of literature review .......................................................................... 20 Case study building ................................................................................................ 22 3.1 Building description ......................................................................................... 22 ii 4 5 3.1.1 General...................................................................................................... 22 3.1.2 Gravity load resisting system ................................................................... 24 3.1.3 Lateral load resisting system .................................................................... 26 3.2 Model development ......................................................................................... 30 3.3 Model validation .............................................................................................. 34 3.4 Ground motion selection and scaling ............................................................... 39 3.5 Analyses ........................................................................................................... 43 Results .................................................................................................................... 45 4.1 Nonlinear time history analysis at design level ............................................... 45 4.2 Incremental dynamic analysis .......................................................................... 46 4.3 Fragility assessment ......................................................................................... 48 4.4 Summary .......................................................................................................... 50 Conclusions ............................................................................................................ 52 5.1 Summary of results .......................................................................................... 52 5.2 Outlook ............................................................................................................ 53 References ...................................................................................................................... 54 Appendix A .................................................................................................................... 62 iii List of tables Table 3-1: Structural elements........................................................................................ 25 Table 3-2: HD, SP, LD and WB schedule ...................................................................... 28 Table 3-3: Properties derived from the ETABS model .................................................. 30 Table 3-4: Selected ground motions ............................................................................... 42 iv List of figures Figure 2-1: CLT panel configuration [23] ........................................................................ 5 Figure 2-2: CLT platform-type (left) and balloon-type construction (right) [23] ............ 6 Figure 2-3: Rocking and sliding behavior based on CLT shear walls aspect ratios ......... 7 Figure 2-4: Multi-panel CLT shearwalls [32] .................................................................. 8 Figure 2-5: Nonlinear material model for Pinching4 [37] ................................................ 9 Figure 2-6: Cascadia subduction zone [5] ...................................................................... 13 Figure 3-1: Architectural rendering of Begbie elementary school [81] ......................... 22 Figure 3-2: Plan view of the northern building: a) first floor, and b) second floor ........ 23 Figure 3-3: Elevation views for A-A) Southern, and B-B) Eastern side of the building 26 Figure 3-4: CLT shear wall connections: a) HD1-3, b) HD4, HD5, c) CLT panel to panel joint- Plywood spline, d) CLT panel to panel joint detail, and e) CLT wall base to footing connection WB1-4, f) CLT floor supported by steel ledger at CLT wall. ............................................................................................................................ 29 Figure 3-5: Numerical model of the case study building in OpenSees .......................... 31 Figure 3-6: Calibrated Pinching4 model for HD1 (a), HD2 and HD3 (b) ..................... 32 Figure 3-7: Calibrated Pinching4 model for a) fastener#1 in SP1/2/3, b) fastener #2 in SP1/2/3, and c) SP7/8/9 ...................................................................................... 34 Figure 3-8: (a) Shear wall test specimens , (b) section A-A, and (c) section B-B [47] .. 36 Figure 3-9: The balloon shearwall model ....................................................................... 37 Figure 3-10: Calibrated Pinching4 model for (a)HD (b)SP (c)angle bracket, and (d)Shearwall model validation against the UNBC shearwall test. (Horizontal displacement) ...................................................................................................... 39 v Figure 3-11: Comparison of the (a) response spectra and (b) Husid plots of the SD and LD record pair .................................................................................................... 41 Figure 3-12: Distribution of ground motion duration ..................................................... 43 Figure 3-13: Comparison of response spectra of selected motions with the UHS of Vancouver: (a) SD and (b) LD. .......................................................................... 43 Figure 4-1: Maximum IDR for a) SD and b) LD motions ............................................. 45 Figure 4-2: Mean maximum base shear of the building for SD and LD motions .......... 46 Figure 4-3: IDA curves of a) SD and b) LD motions ..................................................... 47 Figure 4-4: Nonlinear behavior of a critical (a,b) HD and WB connection under the SD record; (c,d) HD and WB connection under the LD record before collapse. ..... 48 Figure 4-5: Fragility curves for collapse. (a) SD and (b) LD motions. .......................... 49 Figure 4-6: CDF results for SD and LD motions ........................................................... 50 vi Acknowledgements This research was partially supported by Fast + Epp structural engineering firm. I feel deeply honored to express my sincere gratitude to my supervisor Prof. Tannert for his great guidance, kind support, and encouragement during my graduate education, from his teaching courses to the completion of my thesis. He is a great person, and it has been an absolute pleasure to work with him. I am deeply grateful to him for giving me freedom to work my way while guiding me and supplying the computational resources required to conduct this work. I owe particular thanks to my supervisory committee member Dr. Yuxin Pan for sharing his knowledge with me and giving me valuable advice and answering to my enquiries during thesis. He was very instrumental in developing OpenSees model and selecting ground motion data. I would also like to express appreciation to my other supervisory committee member Dr. Md Shahnewaz for his kind inspirations, insightful comments and encouragement. He was involved in the design of the case study building and provided me with very useful information. In addition, I would like to thank my dear colleagues and friends, Farideh Rezagholizadeh, Selamawit Dires and Alison Conroy for their novel ideas, love and warm companionship. I would also like to express my gratitude to Dr. Roger Wheate and all UNBC faculty and staff for helping me throughout my graduate studies. Lastly, I would like to express my deepest thanks to my parents and my brothers for their kind support, love and continuous encouragement. Without their love and support, I would not have been able to complete this research. vii I dedicate this work to my parents, who have always loved me unconditionally and whose great examples have taught me to work hard for the goals that I aspire to achieve. viii 1 Introduction 1.1 Cross-laminated timber lateral load resisting systems Cross-laminated timber (CLT) is gaining popularity as a building material due to its biaxial strength and light weight. CLT panels can be used for floor and wall applications as well as shear walls in lateral load resisting systems (LLRS), for both platform-type or balloon-type constructions [1]. The 2020 National Building Code of Canada (NBCC) [2] provides force modification factors for force-based design of CLT shear walls and refers to CSA O86 for design provisions. However, the design provisions for CLT LLRS in the 2020 NBCC [2] and the 2019 Canadian Standard for Engineering Design in Wood (CSAO86) [3] apply only to platform-type construction. Much research on the seismic behavior of CLT LLRS was carried out, with focus on CLT walls, their panel-to-panel shear connections, angle brackets, hold-downs (HDs), and fullscale buildings. These studies contributed toward standard design provisions [4]; one aspect yet to receive attention is the impact of long-duration ground motions. 1.2 Seismicity LLRS must resist both wind and seismic loads. Seismic activity is particularly high in the southwestern part of British Columbia (BC), Canada. There are three types of earthquakes in this area: i) crustal earthquakes, ii) subduction in-slab earthquakes, and iii) subduction interface earthquakes. Paleoseismic evidence indicates that Cascadia subduction interface earthquakes (the Cascadia subduction zone is a plate boundary that extends from northern Vancouver Island to northern California [5]) occur about every 450 years [6, 7]. Usually, subduction interface records show an exceptionally large magnitude (Mw>9) and long duration (up to several minutes). Amongst the metrics developed for measuring ground motion duration, the 5%-95% significant duration has been found to be the best indicator 1 for evaluating the inelastic performance of structures [8]. Since the last major Cascadia earthquake has been dated to more than 300 years ago (January 1700) [5], the subduction process in this zone may cause an earthquake during the lifetime of current structures. 1.3 Research need The duration of ground motion, along with its intensity and frequency content, plays a significant role on the collapse risk of structural systems [9]. Current building codes across the world account for the intensity and frequency of the ground motion based on the response spectrum of a linear single-degree-of-freedom system (SDOF) [10-12]. However, the duration of ground motion has not been given direct consideration, even though recent studies have shown duration effect on various structural types [13-15], particularly for structures with cyclic degradations [16, 17]. Under long ground motions, components are subjected to more cycles, and after reaching peak, during cyclic loading, they lose more stiffness and strength. There is an increasing worldwide interest in mid- to high-rise wood –in particular CLT– buildings [18]. Also in North America, CLT construction is becoming common, and encapsulated mass timber structures have been incorporated into the 2020 NBCC [2] for buildings up to 12 stories, and the 2021 version of the International Building Code (IBC) for buildings up to 18 stories [19]. However, there has been no research on the evaluation of the influence of ground motion duration on the seismic performance of CLT buildings. 1.4 Objectives The objective of this thesis is to assess the effects of ground motion duration on the seismic performance of a two-storey balloon-framed CLT building. The specific goals are to evaluate drift ratio at various intensity levels to assess the collapse capacity of the building under long and short duration motions. 2 To achieve these objectives, a newly constructed two-storey CLT building, the Begbie Elementary school, in Vancouver BC, was used as a case study. The seismic response of the buildings under two sets of ground motions (long and equivalent short duration records) was evaluated through nonlinear dynamic analysis. The collapse capacities were then identified by developing fragility curves from incremental dynamic analysis. 1.5 Thesis organisation and scope In chapter 2 of this thesis, a literature review of CLT construction and its seismic performance, the seismicity in BC, previous research on the effects of ground motion duration on structural performance, and seismic analyses methods are provided. In chapter 3, the case study building and the development of a 3D model of the building are described. In chapter 4, the effects of ground motion duration on numerical models of the case study buildings are investigated using comprehensive ground motion databases including long and short duration records. Incremental dynamic analyses are performed on the case study buildings and the influences on collapse capacity at various intensity levels are presented. In chapter 5, the findings and contributions of the research work are summarized, and recommendations are offered for future research. The research focused entirely on the seismic performance of the case study building under subduction and crustal ground motions. The pulse effect from near-fault records was not considered. The building was subjected to the motions only in one direction. To evaluate the effects of duration, inter-storey drift ratio was used as a damage measure. Due to the complexities of the case study building, it was not feasible to consider other damage measures. The performance of the building under wind was not studied. 3 2 Literature review 2.1 Background information The world population is increasing by approx. 1% per year, resulting in an additional 83 million people annually. It is estimated that by 2030, the population will reach about 8.6 billion [20]. This, in turn, will increase the demand for living and working spaces and put pressure on natural resources. According to the 2017 global status report, by 2060, buildings floor area will double to 460 billion m2 [10]. From a global perspective, the building industry is responsible for 36% of energy consumption and 39% of energyrelated carbon dioxide emissions. To meet the Paris Agreement targets, the global buildings sector should improve its energy intensity by 30% by 2030 compared to 2015 highlighting the urgent need for sustainable design and construction practices [10]. Among the main building materials, wood is the only renewable one, suited to meet sustainable building objectives. When compared to steel and concrete, wood products have lower environmental impacts over their life cycle as they use less embodied energy in production, release fewer pollutants to the air and water, as well as reduce greenhouse gas emissions [21]. In Europe, wood construction is regarded as a way to meet the climate targets of reducing CO2 emissions by 88-91% by 2050 [10]. Rapid erection –timber structural elements can be prefabricated offsite and transported to the site– and aesthetics of wood structures, as well as the high strength-to-weight ratio are further advantages of using wood as structural material, also in seismic active regions [22]. All these environmental, economic and structural advantages have made wood very popular material in the last two decades. Moreover, the increased use of high efficiency engineered mass timber products such as CLT, glue-laminated timber (Glulam) or Laminated Veneer Lumber (LVL) has made a great contribution to this development [1]. 4 2.2 CLT construction 2.2.1 CLT as a structural material CLT consists of several layers of lumber boards stacked in alternating directions, glued together, and pressed to form a solid panel (Figure 2-1). It was originally introduced in the early 1990s in Austria and Germany [23]. While well established in Europe, the use of CLT is increasing in North America [23]. The large number of buildings built around the world using CLT shows the many advantages it can bring to the construction industry such as faster installation [23]. Moreover, CLT provides dimensional stability, and good thermal and sound insulation. The lower weight, compared to concrete or steel, can result in lower seismic loads and reduced foundation costs. Figure 2-1: CLT panel configuration [23] 2.2.2 CLT lateral load resisting systems CLT panels can be used in a wide range of applications from floors and walls in gravity load resisting systems to shear walls in LLRS. Because of its high in-plane strength and stiffness [23, 24], it can be used as a diaphragm or shear wall to resist lateral wind and seismic loads, in either platform- or balloon type applications (Figure 2-2), even in earthquake-prone regions [25]. 5 Figure 2-2: CLT platform-type (left) and balloon-type construction (right) [23] In platform-type construction, each floor acts as a platform for the floor above. Each wall independently dissipates energy with a rocking motion using both the connections to the floor below and vertical joint connections between individual panels. These walls are connected to the floor below with brackets and HDs and for wall panel connections, plywood splines or half-lap joints wall panel connections are typically used. In this type of building, each wall transfers the gravity load to the floor below. Thus, the maximum height of the building is usually limited by the compression perpendicular to the grain resistance of the CLT floor panels on the lowest floor [23, 26]. In balloon-type construction, walls continue over several floors, and the floor panels are attached to the walls at each story. Thus, compared to platform-type buildings, the number of shear walls is usually lower. There are several advantages to balloon-type CLT shear walls: no compression is applied perpendicularly to grain on the floors because of continuous CLT walls; the elimination of cumulative perpendicular to the grain shrinkage at the building's height. Moreover, this system requires fewer HD and shear bracket connections and offers less flexural deformation [23, 27]. 6 2.2.3 Connections for CLT shear walls CLT panels are almost rigid under in-plane loading; therefore, ductility and energy dissipation in CLT shear walls must be achieved by the connections [28]. This can occur through rocking or sliding behaviour, see Figure 2-3. HDs are designed to resist rocking and the angle brackets are designed to resist sliding. Together with the connection configuration, the panel’s aspect ratio (height-to-length) determines the governing kinematic. Wall panels with aspect ratios between 2:1 and 4:1 are more likely to act in rocking, whereas the panels with aspect ratios less than 2:1 tend to slide. Panels with large aspect ratios facilitate rocking, which is more beneficial as sliding failure is very dangerous in terms of structural stability. In multi-panel shear walls, as shown in Figure 2-4, individual panels are coupled with vertical joints such as spline joints, half-lap joints, and butt joints [28]. Figure 2-3: Rocking and sliding behavior based on CLT shear walls aspect ratios Extensive experimental research has been conducted on CLT walls with different types of connections. Gavric et al. [29] studied nailed HD and bracket connections and proposed overstrength factors of 1.3. Subsequently, Gavric et al. [30] evaluated energy dissipation and ductility ratio in screwed wall-to-wall, floor-to-floor and wall-to-floor connections and reported ductile behaviour as long as requirements for end and edge distances were 7 satisfied. Hossain et al. [31] performed monotonic and cyclic tests on double-angled butt joints with STS and reported that such joints were moderate to highly ductile. Figure 2-4: Multi-panel CLT shearwalls [32] For numerical modeling of LLRS, an accurate constitutive model of the connections is necessary. Some hysteretic models that can simulate timber joints are available in the OpenSees platform [33]. OpenSees, the Open System for Earthquake Engineering Simulation, is an open-source software framework that allows users to create finite element computer applications for simulating the response of structural and geotechnical systems subjected to earthquakes and other hazards [34]. OpenSees has advanced capabilities of modeling and analyzing the nonlinear response of systems by using a wide range of elements, material models, and solution algorithms. Moreover, the collapse analysis methods it provides allow to obtain the seismic and progressive collapse analyses of structures. One commonly used hysteretic model in OpenSees is named Pinching4, which accounts for strength and stiffness degradation under cyclic loading. Originally developed for analysis of beam-column joints in reinforced concrete frame structures [35], it is now widely used for analyzing timber and steel joints. Pinching4 model incorporates 22 parameters to describe the load-displacement response of a material [36], 8 and 17 damage parameters that accounts for stiffness and strength degradation. The definition and material properties are shown in Figure 2-5. In this model, ePf and ePd are force and deformation points on the positive response envelope, while eNf and eNd refers to force and deformation points on the negative response envelope. rDispP and rDispN correspond to the ratio of the deformation at which reloading occurs to the maximum and minimum historic deformation demand, respectively. fForceP and fForceN refers to the ratio of the force at which reloading begins to the force corresponding to the maximum and minimum historic deformation demand, respectively. uForceP and uForceN represent the ratio between the strength developed upon unloading from a negative load and the maximum strength developed under monotonic loading, respectively. Moreover, there are three cyclic degradation models controlling unloading stiffness degradation (gK), reloading stiffness degradation (gD) and strength degradation (gF). Type of damage can be defined as cyclic or energy. Figure 2-5: Nonlinear material model for Pinching4 [37] 9 2.2.4 Seismic performance of CLT structures The seismic performance of multi-story CLT structures has been the subject of previous experimental and numerical research with focus on developing design provisions, collapse capacity, drift demand, and damage measures in CLT structures. These studies can be classified into platform and balloon-type CLT structures. To date, much research has been carried out on platform-type constructions. van de Lindt et al. [38] conducted a series of shake-table tests on a full-scale two-story mass-timber building with platform-type CLT shear walls. They reported that adding transverse walls on the building did not affect the ability of the structural panels to rock, but it improved the building structural system performance. Amini et al. [39] tested different wall configurations and concluded that wall panel aspect ratio, boundary constraints and gravity loading are very influential on the wall performance. Deng et al. [40] reported that the panel aspect ratio is the most influential parameter in the sliding and rocking behavior of the CLT shear wall. Sustersic et al. [41] test results on a 4-storey CLT building showed that smaller wall segments connected by vertical joints dissipate more energy and are more seismically efficient. Latour and Rizzano [42] used an innovative steel bracket “XL-Stub” instead of hold-downs and reported that the XL-Stub connectors had higher energy dissipation and displacement capacity compared to traditional hold-downs. The most comprehensive experimental study was the SOFIE project [43-45] where different CLT buildings were tested in the lab. The results showed that these structures were strong enough to withstand 15 consecutive destructive earthquakes without severe damage. Full-scale shake table tests on a 5-story building subjected to 100% of Kobe ground motion and a 3-story building subjected to 140% [46] showed that while the 3story building was severely damaged, the 5-story building suffered some damage such as yielding anchor bolts and compression rupture at the corner of CLT wall panels. 10 Balloon-type constructions have also received some attention in recent years. Shahnewaz et al. [47] tested a two-story balloon-type CLT shear wall with different ledgers under monotonic and cyclic loading and reported that the Canadian standard specifications for platform type construction can be used to design balloon-type CLT shear walls. Zhang et al. [48] tested full-scale 3-story CLT structures under quasi-static cyclic loading. These structures included platform-type constructions with either narrow or wide wall panels, a balloon-type structure, and a balloon-type structure with glulam beams. Their results showed that the latter provided the highest lateral load resistance and its performance depended on moment resisting performance of glulam beam and CLT wall joint. Numerical studies have verified that component-based modelling of CLT walls using connection test data is a reasonable approach in which the panel is modelled using elastic shell elements. Most numerical studies on CLT buildings have been focused on platformtype buildings. Using incremental dynamic analysis, Shahnewaz et al. [49] measured the damage states of a six-story CLT platform-type building at various intensity levels and concluded that the building can safely be built in a high seismic zone if appropriately designed. Sun et al. [50] determined the drift limitations for frequent, medium, and rare seismic hazard levels of multi-story platform-type CLT buildings; for mid-rise or highrise buildings, these values were estimated as 0.25%, 0.70%, and 1.30%, respectively, whereas for low-rise buildings, the values were 0.30%, 0.75%, and 1.40%. Sun et al. [51] reported that while the maximum inter-storey drift of the post-tensioned platform-type CLT structures is approximately twice that of the conventional structures, the maximum residual inter-storey drift is 50%. Using Pushover and time-history dynamic analysis, Sun et al. [52] evaluated the seismic performance of 8-storey energy-dissipating posttensioned platform-type CLT shear wall structures with UFP dissipaters or friction dissipaters. Their results showed that the maximum inter-storey drift of the UFP 11 dissipaters is similar to that of the friction dissipaters, but friction dissipaters typically perform better in mitigating absolute response accelerations than UFP dissipaters. In addition to platform-type buildings, a number of numerical studies have been conducted on balloon-type CLT constructions in recent years. Zhang et al. [32] investigated the effect of hold-downs, vertical and horizontal shear connections between the CLT panels on the period and stiffness of tall balloon-type CLT buildings. Their results showed that the horizontal shear connections have the greatest impact on the overall stiffness of the building and this influence decreased as the building height increased. Hashemi et al. [53] conducted nonlinear static pushover and nonlinear dynamic time–history analyses on a 5-story balloon-type CLT structure with resilient slip friction joints. They reported that the system can be considered as a resilient seismic solution for timber structures and presented a preliminary design procedure of this system. 2.3 Seismic performance of structures 2.3.1 Seismicity in British Columbia The southwestern part of BC is in a seismically active area subjected to three types of earthquakes: i) crustal earthquakes within the North America plate (<20 km deep), ii) subduction in-slab earthquakes (45-65 km deep) in the Juan de Fuca plate, and iii) subduction interface earthquakes between the Juan de Fuca Plate subducting and the North American continent [5]. As shown in Figure 2-6, the intersection of these plates occurs in the Cascadia Subduction Zone off the Pacific coast of Northern California, Oregon, Washington, and BC [54]. 12 Figure 2-6: Cascadia subduction zone [5] Paleoseismic evidence indicates that Cascadia subduction interface earthquakes occur about every 450 years [6, 7]. The most recent event is dated to January 26, 1700, when an earthquake that occurred along the Cascadia subduction zone caused a tsunami in Japan [5]. The time passed since then has raised some concerns regarding the subduction process in this zone; it is estimated that the probability of the next great Cascadia earthquake in the next 50 years is 23% [55]. While there is no record for the Subduction interface earthquakes in Cascadia subduction zone (written records in the region are too recent [54]), it has been reported that the Cascadia subduction earthquakes are similar to subduction earthquakes around the Pacific Ocean, such as the 2011 Japan, the 1960 and 2010 Chile, the 2004 Indonesia, the 1964 Alaska, and the 1985 Mexico earthquakes [17, 54]. 2.3.2 Characteristics of ground motions Two important characteristics of subduction interface earthquakes are their large magnitude and long duration. Considering a total rupture area of 60,000 km2 in the 13 boundary between the North American and Juan de Fuca plates [56], 500 years of plate motion can generate enough seismic moment for a magnitude 9 earthquake [57]. The impact of ground motion duration on structural damage and economic losses became evident in recent earthquakes, such as those at Hokkaido, Japan (Mw 8.3, 2003), Sumatra, Indonesia (Mw 9.1, 2004), Wenchuan, China (Mw 7.9, 2008), Maule, Chile (Mw 8.8, 2010), and Tohoku, Japan (Mw 9.0, 2011) [9]. Because of the large rupture areas in these subduction earthquakes, they have significantly longer durations than earthquakes involving smaller rupture areas such as those in California [54]. The duration of the 1964 Alaska earthquake was estimated to be 4 minutes and the strong shaking 2 minutes [58], a 3.5 minutes duration was estimated for the earthquake that occurred in Southern Chile in 1960 [59], and the shaking that was felt in the 1985 Michoacan earthquake in Mexico City lasted 5 minutes [60]. Earthquake intensity and frequency are accounted for in the current design codes across the world, based on the response spectrum of a linear single-degree-of-freedom system (SDOF) and the corresponding ductility demand [10-12]. However, the duration of ground motion has not been given direct consideration, even though it has proven to have a substantial impact on the seismic performance of various types of structures [13-15]. One of the major reasons for this is inconsistency in the findings of the previous research on this topic [61]. The main challenges to estimating the duration effects on structural response include conducting bidirectional nonlinear dynamic analysis as it is impossible to have a pair of long and short duration motions whose spectra are equivalent for both horizontal components, the isolation of the duration effect from other characteristics of ground motion such as the rate of energy build-up by using the latest Zengin et al. [13] method, and the possibility of overlooking some other unknown ground motion characteristics. 14 There are more than 30 metrics for measuring ground motion duration in the literature [62]. The significant duration is defined as the time in which 90% of the Arias intensity (IA) is accumulated. Arias Intensity is a measure of the strength of ground motion that determines the intensity of shaking by measuring the acceleration of seismic waves [63]: = ∫ ( ) 2-1 Where IA is the Arias Intensity, a(t) is the recorded ground acceleration, tmax the length of record, and g the acceleration due to gravity. The accumulation times are determined for 5% to 95% of the total IA. Having examined the 5%-95% significant duration, Foschaar et al. [8] concluded that this is the best indicator of the inelastic performance of structures. 2.3.3 Effect of ground motion duration on structural performance Research since early 1960’s investigated the effect of ground motion duration on structural performance on a wide range of structures, including steel [8, 13, 64, 65], concrete [15, 17, 54, 66], and light-frame wood buildings [14, 67, 68]. The impacts of ground motion duration were shown to depend on the material models (with or without degradation), intensity level (design or collapse) and damage measures (e.g., drift, energy demand or damage index) used for assessment. Zengin et al. [13] showed that the displacement and the inter-story-drift ratios (IDR) of steel frame buildings, were not strongly affected by ground motion duration. A similar result was shown by Raghunandan and Liel [17] for reinforced concrete structures. However, Han et al. [66] reported for a 4-story reinforced concrete structure that longer duration may result in a larger IDR and a larger residual displacement but only for intensity level high enough to produce nonlinear deformation. They asserted that fully capturing the structural components’ strength and stiffness degradation is of paramount importance. Barbosa et al. [64], used the modified Ibarra–Medina–Krawinkler 15 deterioration model on steel buildings in the OpenSees platform and reported that the duration effect on peak IDR is higher for larger spectral acceleration. Pan et al. [68] evaluated light-frame wood structures with pinching and degradation and obtained a 17% higher probability of exceeding a 3% design drift limit for long-durations. Energy-based damage indices, such as hysteretic energy dissipation, have a relatively good correlation with ground motion duration [9, 69-71]. E.g., Zengin et al. [13] showed that the dissipated hysteretic energy is sensitive to duration. Fairhurst et al. [15] studied 6- to 30-storey MDOF reinforced concrete models with pinching and degradation, and reported that energy demand is greatly increased by earthquake duration. Collapse capacity has been reported to be affected by duration [15] [66]. The collapse capacity in a mid-rise wood building was shown to decrease by 18% under longer duration [14]. Similar findings were reported for steel moment frames [13, 16] and reinforced concrete structures [17]. While no correlation between duration and maximum displacement-based damage measures was found for non-degraded structures [61, 72], good correlation between duration and maximum response measures was found for degraded systems [16, 17]. When the ground motion duration is longer, components are subjected to more cycles and after reaching peak, they lose more stiffness and strength during cyclic loading. According to this, the lateral strength of structures should be large enough to accommodate an increased number of cycles from long duration motions. Pan et al. [14] studied mid- and low-rise wood-frame buildings with the Residual Strength Hysteresis material model, and reported that at the maximum intensity level, longer duration increased the median Park–Ang damage index by 36%. Barbosa et al. [64] reported that the duration effect on Park–Ang and Reinhorn–Valles damage indices is higher for larger 16 values of spectral acceleration. Zengin et al. [13] showed the Modified Park–Ang damage index is affected by duration. Effectively isolating duration effect from other ground-motion characteristics is a challenge. Some researchers decoupled the duration from the amplitude and frequency content of the ground motions by using spectrally matched accelerograms [73] and spectrally equivalent record pairs [16]. Spectral matching eliminates spectral-amplitude differences in the ground motion records, whereas the spectrally equivalent method decouples the differences in amplitude and frequency between the record sets. It is commonly assumed that after decoupling the duration from the amplitude and frequency content, the remaining effects can be attributed to the duration [13]. However, the above methods can only partially decouple the duration influence from other ground motions characteristics as some studies have shown that the rate of energy build-up in the accelerogram can also affect structural response. Based on the studies conducted by Kennedy et al. [74] and Trifunac [75], accelerograms that reach their final energy quickly are more likely to damage and collapse structures than those that reach their final energy slowly. The rate of energy build-up of ground motions can be roughly determined by the slope of Husid plot [76], based on the rate of Arias Intensity (IA) build-up over the significant duration [13]. Trifunac [75] reported that accelerogram damage capability is affected by not only duration but also by the slope of the Husid plot. Similarly, it has been reported that records with steeper slopes of Husid plot have higher energy dissipation demands [77]. To date, Zengin et al. [13] have been the only researchers that considered the rate of energy build-up of earthquake as a control parameter in addition to amplitude and frequency content to avoid its possible influence on the evaluation of duration effects on structural response. 17 2.3.4 Seismic analysis and damage measures LLRS can be analyzed using different methods: i) linear static analysis, ii) non-linear static analysis, iii) linear time history analysis, and iv) non-linear time history analysis (NLTHA). In linear static analysis, the dynamic loading of an earthquake is substituted by a static force applied laterally to a structure to achieve an approximate similar effect. In the non-linear static analysis, the seismic structural deformations can be estimated. Static procedures are appropriate when the effects of higher modes are not significant. In the linear time history analysis, the effects of higher modes are considered, but since this analysis is based on linear elastic response, their applicability reduces with increasing nonlinear behavior. NLTHA while being more complex and time-consuming, is the most accurate and realistic method. Complex high-rises, irregular buildings, and high-profile structures need to be analyzed using NLTHA. A series of earthquake records is applied to the structure, generating response histories for the quantity of interest, e.g., displacements or forces. It is important to note that NLTHA is highly sensitive to both the characteristics of the selected records and their scaling. NBCC 2020 [2] specifies that at least eleven ground motion time histories are required for dynamic analysis. If two or more seismic scenarios are considered, at least five ground motions must be selected for each scenario. Spectrum scaling is the commonly used ground motion scaling approach for the dynamic analysis of earthquake acceleration records. In this method, the spectrum derived from the earthquake records is not less than the equivalent design spectrum in a given period. As specified by the NBCC 2020 [2], this period range should have an upper bound [2.0T1, 1.5s] and a lower bound [0.15T1, T90%] where T1 is the first mode period of the structure and T90% is the periods of the modes necessary to achieve 90% mass participation. 18 NLTHA can be applied to incremental dynamic analysis (IDA) , a parametric method that estimates the structural collapse capacity under seismic loads [78]. A structural model is subjected to earthquake records, and each record is linearly scaled to multiple levels of intensity until the structure collapses. The purpose is to determine the damage measures (DMs) of the structural model at each intensity level of the ground motion, and the resultant values are often plotted versus the intensity as continuous curves. In most cases, the spectral acceleration at the fundamental period of the structure, Sa(T1), is considered as intensity measure and the maximum IDR is used to monitor structural response. IDA has been widely adopted for studying the effects of ground motion duration as its linear scaling process does not affect the motion’s spectral shape and significant duration. In addition to IDR, which is a good metric for assessing the impact of ground motion duration when approaching collapse, other damage measures have been proposed, e.g. maximum response measures, energy measures, cyclic fatigue measures, and combined measures [61]. 2.3.5 Seismic design of CLT LLRS in Canada NBCC 2020 [2] included platform-type CLT shear walls into its acceptable seismic forceresisting systems, and provided force modification factors R0=1.5, and Rd=2.0. The overstrength-related force modification factor, R0, represents the dependable portion of reserve strength in CLT shear walls while the ductility-related force modification factor, Rd, reflects the ability of CLT shear walls to dissipate energy via reversed cyclic inelastic behavior. NBCC specifies that timber structures should be designed according to CSA O86 [79, 80]. The corresponding design provisions for CLT shear walls in CSA O86 apply only to platform-type construction and the commentary emphasizes that balloontype applications are beyond the scope of the standard. 19 There are some differences between the 2016 supplement to CSA O86-2014 [80] and the 2019 [3] standard: the standard no longer accepts the combination of rocking and sliding. To meet this requirement, the allowable aspect ratio limit was increased from 1:1 to 2:1. The most significant standard change consisted in defining which connection can account for the energy dissipation. While the 2016 supplement allowed discrete HDs to dissipate energy, CSA O86-2019 specifies that all non-linear deformations and energy dissipation should occur in: i) wall-to-foundation or wall-to-floor panels below, ii) vertical joints between wall panels. Both versions of CSA O86 specify that energy dissipative connections must have sufficient ductility and deformability. To this end, the connections must be designed such that a yielding mode governs the resistance, be at least moderately ductile in all nonrestricted directions of the CLT panels’ kinematic modes and be able to accommodate sufficient deformation capacity to allow the CLT panels to develop their deformation behavior. Non-dissipative connections, e.g., those not expected to undergo plastic deformations such as the connections between perpendicular walls or between floor panels or between the roof and the walls below, must be capacity-protected with sufficient overstrength to remain linear elastic under the force and displacement demands induced on them when the energy-dissipative connections reach their 95th percentile of ultimate resistance or target displacement under reversible cyclic loading, but the seismic design force need not exceed the force determined using Rd R0 = 1.3. CLT panels that are part of a LLRS must be capacity-protected in the same way. 2.4 Summary of literature review Timber in general and CLT in particular is gaining popularity in taller residential and nonresidential construction. Easier on-site delivery, faster installation, and high dimensional stability are some of the advantages of these products. The light weight and the dissipative 20 response shear walls make it viable to build multi-story CLT structures, in either platform- or balloon-type applications, also in seismic-prone regions. The southwestern part of BC is a seismically active region. It is reported that Cascadia subduction interface earthquakes occur about every 450 years and the last major event has been dated to almost 300 years ago, raising some concerns about an occurrence in the near future. Subduction earthquakes have a long duration in addition to their large magnitude. Previous studies investigated the effects of ground motion duration on structural performance of mid- to high-rise buildings, including steel, concrete, and lightframe wood buildings. According to these studies, long-duration earthquakes could potentially cause more damage to the structure, increase energy demand, and reduce collapse capacity. Nevertheless, seismic codes and guidelines do not consider the effect of duration on the seismic response of structures. CLT buildings have been the center of many studies for the last 20 years and the works on the seismic performance of these structures are still ongoing. Recent effort has focused on developing standardized design provisions for CLT buildings. Many studies have been dedicated to the ductile behavior of connections. Several studies performed an IDA to evaluate the damage states at various intensity levels. However, no research has been done on the effects of ground motion duration on the performance of CLT buildings. Hence, this research is intended to investigate the seismic performance of a two-storey balloon-type CLT building under long duration earthquakes in southwestern BC. 21 3 Case study building 3.1 Building description 3.1.1 General The Begbie Elementary School, in Vancouver BC, was used as a case study. This twostorey, 3,400 m2, elementary school building was designed by Fast + Epp and is the first in the Vancouver School Board's district to use CLT as the primary construction material. Throughout the building, large areas are exposed to timber on both walls and ceilings. Figure 3-1 shows an architectural rendering of the building. Figure 3-1: Architectural rendering of Begbie elementary school [81] This building consists of two parts, north and south. Since the northern and southern part of the building are not connected, each building can be studied separately. This research is focused on the northern building. The plan dimensions for the northern building are 25.5m by 35.1m. Figure 3-2 shows the first and second floor plans of the building. 22 a) b) Figure 3-2: Plan view of the northern building: a) first floor, and b) second floor 23 The gravity loads considered in the design of the building are 2.4 kPa live load and a superimposed dead load of 2.5 kPa which includes non-structural architectural topping, partitions, pavers and finishes, and the self weight of structural members was computed by the software ETABS. Wind load is 0.45 kPa based on q (1/50) considering and importance factor of 1.15. The design procedures followed part 4 of the British Columbia Building Code 2018 (BCBC) [82] and the Vancouver Building By-Law. For LLRS of the building, CLT shear walls, in both platform and balloon type construction, are used. Seismic design of the building followed the 2015 NBCC. Force modification factors Rd and R0 applied in the building are 2.0 and 1.5, respectively. The soil type of the site is class C. For elementary school building, in a high importance category, the importance factor (IE) is considered as 1.3. 3.1.2 Gravity load resisting system The building’s gravity load resisting system (GLRS) are floors, columns and walls. Floors are responsible for resisting and transmitting gravity loads to vertical framing systems, walls and columns. Floor systems used in the building are CLT floor panels or a combination of CLT panels with glulam beams. In the southeastern part of the floor plan, due to the 17 m long span, CLT panels are supported on glulam beams that were spaced at 800 mm and connected to the CLT panels by screws. Steel HSS (Hollow Structural Sections) columns are used in the building. For walls, CLT panels with three different thicknesses 139mm, 191mm and 105 mm and stress grades V2.1, V2.1 and V2M1.1, respectively, were used. Table 3-1 provides an overview of the main GLRS components. 24 Table 3-1: Structural elements Mark Type Remarks B1 SPF 20f-E beam 130X684 Glulam B6 SPF 20f-EX beam screwed to CLT panel 215x380 Glulam B11 350W steel beam W410X67 B12 350W steel beam W610X82 B13 350W steel beam 750 deep built-up w-section B15 350W steel beam W200x42 Beam Floor panel FP2 245 deep. 7-ply SPF V2-XL or E1 7-ply CLT FP3 105 deep. 3-ply SPF V2 3-ply CLT FP4 139 deep. 5-ply SPF V2 5-ply CLT C1 HSS152X152X8 steel col. C2 HSS114Øx9.5 steel col. C3 W150x37 steel col. Column As can be seen in Figure 3-2, northern and western side of the building have out of plane offsets and external walls discontinues from first level to second level; two typical elevation views of the southern and eastern part of the building are presented in Figure 3-3. The storey heights for the first and second floor are 4.35 m and 4.00 m, respectively. 25 Figure 3-3: Elevation views for A-A) Southern, and B-B) Eastern side of the building 3.1.3 Lateral load resisting system The building’s LLRS are mainly provided by two-story continuous coupled CLT panels forming shear walls. The red highlighted square in Figure 3-3 presents a typical view of these shear walls. These shear walls dissipate energy through HDs and vertical panel-topanel joints while wall base connections (WB) were capacity protected. As shown in Figure 3-4(a), for HDs 1-3, a typical Rothoblasas WHT angle plates with different numbers of LBA threaded anker nails were used. Another type of HD used in the building (HD4, HD5) is shown in Figure 3-4(b), where an internal steel plate is tight fitted with 12Ø pins to ensure high tension resistance. Table 3-2 provides detailed information on these HDs fasteners. 26 Plywood spline and lap spline were used for panel-to-panel joints. For the panel-to-panel vertical spline connections, surface mounted 25 mm by140 mm D. Fir plywood pieces were used to attach the two panels using screws and smooth shank nails at different spacings, see Figure 3-4(c). Details on the fastener’s schedule can be found in Table 3-2. For lap spline connections, half-lap joint connections with 80 mm lap length and maximum 2 mm gap were connected using partially threaded screws, shown in Figure 3-4(d). More information about these screws can be found in Table 3-2. For WB connections, a notched glulam sill plate was anchored to the footing to support the CLT wall and 10Ø 140 partially threaded screws were used to connect the wall to the sill plate, see Figure 3-4(e). Table 3-2 presents details about this connector. The shear walls are 8.35m high supporting intermediate floor at mid-height using steel ledgers. As shown in Figure 3-4(f), the steel ledger was attached to the CLT wall and to the floor below using 10Ø×120 and 8Ø×80 PSC partially threaded screws, respectively. Details for fastener schedule are presented in Table 3-2. 27 Table 3-2: HD, SP, LD and WB schedule Mark Type Fasteners HD1 Rothoblaas WHT440 30 - nails LBA 4Ø 60 mm long. HD2 Rothoblaas WHT620 55 - nails LBA 4Ø 60 mm long. HD3 Rothoblaas WHT740 75 - nails LBA 4Ø 60 mm long. HD4 Custom 6-12Ø stainless steel tight fit pins HD5 Custom 10-12Ø stainless steel tight fit pins Mark Type Fasteners#1 Fasteners#2 SP1 Plywood Spline PSW 8Ø 120 @ 600 mm Shank Nails 4Ø 60 @ 250 mm SP2 Plywood Spline PSW 8Ø 120 @ 600 mm Shank Nails 4Ø 60 @ 200 mm SP3 Plywood Spline PSW 8Ø 120 @ 600 mm Shank Nails 4Ø 60 @ 150 mm SP7 Half-lap Joint PSH screws 8Ø 140 @ 250 mm - SP8 Half-lap Joint PSH screws 8Ø 120 @ 200 mm - SP11 Half-lap Joint w/ Steel Plate 2 rows PSL screws 8Ø 120 @ 200 mm - Type Fasteners#1 Fasteners#2 LD1 2 rows PSC 10Ø 120 @450 mm PSC 8Ø 80 @200 mm LD2 2 rows PSC 10Ø 120 @250 mm PSC 8Ø 80 @200 mm LD3 2 rows PSC 10Ø 120 @175 mm PSC 8Ø 80 @200 mm Type Fasteners WB1 PSW screws 8Ø 140 @300 mm WB2 PSW screws 10Ø 140 @150 mm WB3 PSW screws 10Ø 140 @100 mm WB4 PSW screws 10Ø 140 @65 mm 28 Figure 3-4: CLT shear wall connections: a) HD1-3, b) HD4, HD5, c) CLT panel to panel joint- Plywood spline, d) CLT panel to panel joint detail, and e) CLT wall base to footing connection WB1-4, f) CLT floor supported by steel ledger at CLT wall. 29 3.2 Model development A linear elastic ETABS model of the building, designed by Fast + Epp, was used as benchmark for developing the nonlinear model. The 3D model with six degrees of freedom of the building, shown in Figure 3-5, was developed in OpenSees [33]. Table 3-3 presents the properties derived from the ETABS model to develop the OpenSees model. The first two fundamental periods of the OpenSees model in E-W and N-S directions are 0.46 s and 0.43 s, respectively, which are very close to the periods from the ETABS model (0.49 s and 0.43 s) with less than a 6% difference. In accordance with the ETABS model, the first mode of the model is in E-W direction and the second mode is in N-S direction, showing that the building is weaker in the E-W direction. Thus, the building was subjected to earthquakes in E-W direction. Floors were modeled as rigid diaphragm with lumped mass at each story. The seismic masses for first and second floor were defined the same as the ETABS model as 3,006 kN and 1,890 kN, respectively. Floors were constrained in all degrees of freedom except translation along X, Y and rotation about Z axis. Supports of the building were modeled as fully fixed. CLT walls with two different thicknesses of 191 mm and 139 mm were modeled and defined as elastic isotropic material with elastic modulus of 9,500 MPa and Poisson's ratio of 0.01, the same as the ETABS model. Table 3-3: Properties derived from the ETABS model CLT panel properties Wall corner Seismic mass Modulus of elasticity Poisson's ratio Stiffness Story 1 Story 2 30 9,500 0.01 60,000 3,006 1,890 MPa kN/m kN kN Figure 3-5: Numerical model of the case study building in OpenSees Energy dissipative HDs (HD1, HD2, and HD3) were modeled such that they resist tension and compression. The HD was not designed for shear resistance. For the tension behavior, Pinching4 material was used whereas for the compression, an elastic-no-tension (ENT) material model with an elastic modulus of 1,500 MPa was used and then a parallel material was applied to combine them. Test data on balloon CLT shear walls conducted at UNBC [47] was used to calibrate Pinching4 material for HD1, HD2 and HD3. Figure 3-6 presents the UNBC test data compared to Pinching4 model for these HDs. To compare the energy dissipation between the model and the test, the area under their backbone curves was calculated using the AREA command in the software AutoCAD 2021 and then they were scaled to obtain the actual area value. The difference in the energy dissipation of the positive response envelope between the model and the test for the HDs was only 2% with 978 kN-mm and 1,000 kN-mm for HD1, and 2,255 kN-mm and 2,302 kN-mm for HD2-3, respectively. Thus, the calibrated Pinching4 model simulated 31 nonlinear behavior of the HD connections well. Since the numbers of nails used in the building for the HDs are greater than that of the test, the calibration was first carried out for one single screw and then its corresponding strength capacity was scaled up according to the number of the screws with the same deformation capacity. The strength of steel connections in timber construction is almost linearly proportional to the number of screws. This simplification has been widely used and accepted in connection modeling. [26],[83]. HD4 and HD5 were modeled elastically with a stiffness in accordance with the design data of 44,270 kN/m and 73,780 kN/m, respectively. Due to the main focus on the balloon frame, HD8, HD10 and HD11, located on the second floor, were not modeled. The ETABS model also did not consider these HDs. In out of plane direction, all HDs were defined elastically, based on the study by [84]. 160 80 Model Test 60 Force (kN) Force (kN) 40 20 0 -20 -40 80 40 0 -40 -10 -5 a) Model Test 120 0 5 10 15 20 25 30 -10 -5 Displacement (mm) b) 0 5 10 15 20 25 30 Displacement (mm) Figure 3-6: Calibrated Pinching4 model for HD1 (a), HD2 and HD3 (b) For panel-to-panel joints, Pinching4 material was used in shear direction (Z axes). For other directions (X and Y axes), in plane and out of plane, the joints were modeled elastically. For SP1, SP2 and SP3, which includes both nails and screws, each fastener was separately modeled with a Pinching4 material and then were combined into one material using a parallel material. Pinching4 materials for these joints were calibrated 32 using test data from Sullivan et al. [85] and UNBC. Figure 3-7(a,b) shows the test data compared to Pinching4 model for these connections. For SP7 and SP8 which have only screw fasteners, Pinching4 materials were used to model the screws and were calibrated using test data from Gavric et al. [30]. Figure 3-7(c) presents the test data compared to Pinching4 model for these connections. The difference in the energy dissipation of the positive response envelope between the model and the test for SP1, SP2/3 and SP7/8/9 are 5%, 5% and 6%, respectively, with 4,161 kN-mm and 4,391 kN-mm for SP1, and 990 kN-mm and 947 kN-mm for SP2/3, and 732 kN-mm and 692 kN-mm for SP7/8/9, indicating a good simulation of the nonlinear behavior. Unlike these connections, SP11 was modeled elastically due to its very rigid design. The stiffness for this connection was considered as 60,000 kN/m based on the design data. SP9, SP10 and SP12 were not modeled as they were used for walls which were not considered as shear walls. For modeling the orthogonal wall-to-wall connections at a corner, a rigid connection with vey large stiffness was assumed for in plane and out of plane directions. However, in Z direction, a stiffness of 60,000 kN/m was used, based on the design data. For modeling WB connections, the number of screws along the shear walls was counted and then based on these numbers, some zero length elements were defined in which a shear bracket was modeled to represent the WB connection. These shear brackets were then modeled such that they resist tension and shear. It has been reported that conventional wall base angle bracket connections have similar strength in tension and shear [49, 86], so the same Pinching4 parameters were used for the two directions. Since there is no test data available for the designed WB connections, it was simplified that the elastic region of the Pinching4 material was determined using the design strength of the WB connections (3.8 kN per screw) and their post-yielding and hysteresis parameters were calibrated using 33 TCN240 angle bracket cyclic test data. For out of plane directions, a rigid connection with vey large stiffness was assumed. 300 50 Model 100 0 10 -10 -100 -30 -200 a) Model Test 30 Test Force (kN) Force (kN) 200 -50 -300 -60 -40 -20 0 20 Displacement (mm) 40 60 b) -40 -30 -20 -10 0 10 20 30 40 Displacement (mm) 30 Force (kN) 20 Model Test 10 0 -10 -20 -30 c) -50 -40 -30 -20 -10 0 10 20 30 40 50 Displacement (mm) Figure 3-7: Calibrated Pinching4 model for a) fastener#1 in SP1/2/3, b) fastener #2 in SP1/2/3, and c) SP7/8/9 3.3 Model validation Since the design provisions in CSA O86 only apply to platform-type constructions, lab test results [47] were used to investigate the seismic behavior of the balloon CLT shear wall system designed to provide rocking behavior. The lateral performance of the CLT shear wall system was tested under quasi-static monotonic and reversed cyclic loading. The tested system consisted of two CLT panels with hold-downs, steel angle brackets, plywood surface splines and nails as fasteners. The CLT panels were strength grade 191V2: 7-ply, 191 mm thick. The two coupled CLT panels are 1,219 mm wide and 3,658 34 mm tall with an aspect ratio of 3:1, representing a half-scale two-story shear wall, see Figure 3-8. All the other parameters of the shear wall were full-size as designed for the actual building. The shear walls were anchored to a steel beam on the outer edges with two WHT740 HDs that were attached to the panel with 75 4Ø×60 mm anker nails; two TCN240 angle brackets were attached to each panel with 36 4Ø×60 mm anker nails. For the panel-to-panel vertical spline connections, surface mounted 25×140 mm D. Fir plywood pieces were used, spliced at one-third height of the wall, and attached to the panel with 4Ø×60 mm anker nails at 200 mm on center with one additional 10Ø×120 mm screw at the top and the bottom of each piece of the spline. An approximation of equal forces was applied to the roof and to the second floor based on the equivalent static analysis for story shears on the two-story building. Four different types of ledgers were tested where they were attached at mid-height of CLT shear walls to represent the floor at the first level. The test setup did not allow to apply gravity loading to the ledgers. Thus, the test did not entirely represent the actual loading situation for a balloon-type structure. However, it was believed that the difference has no significant effects on the wall behavior. Tests were stopped when the load carrying capacity dropped to 80% of the maximum load, defined as failure. Strength, stiffness, deformation, ductility and energy dissipation were evaluated in the tests. For the cyclic tests, both maximum forces for positive (Fmax+) and negative (Fmax-) cycles were recorded as well as their corresponding deformations (dFmax+, dFmax-) at the top right corner of the shear walls. Energy dissipation, E, was derived from the area under the loading and unloading cycles of the load-deflection hysteresis loops using the trapezoidal rule. 35 Figure 3-8: (a) Shear wall test specimens , (b) section A-A, and (c) section B-B [47] Test results showed that panel rocking dominated the behavior of shear wall system and shear brackets remained elastic. All energy dissipation occurred in HDs and panel-topanel joints. The shear wall displacement was due to the rocking of CLT panels which themselves act as rigid bodies with minor in-plane deformation. Additionally, the ledgers did not prevent the walls from rocking, and the rocking of the walls had very little effect on the ledgers' gravity load bearing capacity. The obtained results indicated that the twostory balloon-type CLT shear wall system meets the design demands from CSA-O86. The results of the tests were consequently used in the design of the balloon frame school building. 36 In this study, before assembling for the entire building, the modeling approach was validated with the nonlinear response of a half-scale balloon-type CLT shear wall tested at UNBC. Figure 3-9 shows the modeled balloon shear wall. Figure 3-9: The balloon shearwall model In the shear wall model, two HDs were modeled on the outer ages of the shear walls calibrated with UNBC WHT740, 38-4Ø×60 test data. Four angle brackets, two on each panel, were modeled on the base wall in the same location as the angle brackets used in the test and were calibrated with angle bracket 200, 30-4Ø×60, using test data by Tomasi et al. [87]. Moreover, a vertical connection was modeled between panels representing panel-to-panel spline connections. This connection was calibrated with panels relative vertical displacement test data from UNBC. The calibrated connections are presented in Figure 3-10(a-c). HD and angle bracket connection models were then scaled up to reach the stiffness of connections in the tested shear wall. CLT wall panels were modeled with a thickness of 191 mm and defined as elastic isotropic material with modulus of elasticity of 9,500 MPa, 37 and Poisson's ratio of 0.01. A 20-kN/m vertical gravity load was applied at the top of the wall, the same as the test setup. A Lateral load was applied at the top left corner of the wall panel. Considering this, the shear wall model was developed. The horizontal deformations were then determined at the top-right corner of the shear wall, the same as the test. Figure 3-10d shows the shear wall model validation against the UNBC shear wall test. The difference in the energy dissipation of the positive response envelope between the model and the test for HD, SP, angle bracket and the shear wall are 2%, 2%, 3% and 2%, respectively, with 2,255 kN-mm and 2,302 kN-mm for HD, 4,097 kN-mm and 4,191 kN-mm for SP, and 1,159 kN-mm and 1,127 kN-mm for angle bracket, and 14,130 kNmm and 14,348 kN-mm for the shearwall. These results show that shearwall hysteresis behavior was well predicted by the numerical model. 38 150 80 Model Test 40 20 0 0 -50 -150 -40 -10 a) 0 10 20 Displacement (mm) 30 -60 -40 b) 60 -20 0 20 40 Displacement (mm) 60 150 Model Test Model Test 100 Force (kN) 40 20 Force (kN) 50 -100 -20 0 -20 50 0 -50 -40 -100 -150 -60 c) Model Test 100 Force (kN) Force (kN) 60 -40 -20 0 20 Displacement (mm) 40 d) -200 -100 0 100 Displacement (mm) 200 Figure 3-10: Calibrated Pinching4 model for (a)HD (b)SP (c)angle bracket, and (d)Shearwall model validation against the UNBC shearwall test. (Horizontal displacement) 3.4 Ground motion selection and scaling In seismic design of structures that use a target spectrum, the first step is to select and scale ground motions that represent the seismic hazard at the structure site. Many countries have commonly used the uniform hazard spectrum (UHS) as their target spectrum. In Canada, new buildings are designed for 2% in 50-year probability of exceedance shaking levels (UHS). This spectrum is derived from the probabilistic seismic hazard analysis (PSHA), where a combination of probabilities of earthquake scenarios with different magnitudes and distances, as well as predictions of ground motion intensity, is used to compute the seismic hazard at a site [88]. 39 Although UHS-based ground motion selection and scaling directly consider the effects of spectral shape and amplitude of ground motion, it rarely suggests including the duration as one selection criteria. It has been reported that using only short duration ground motions from the PEER NGA-West2 database or the FEMA P695 far-field sets may lead to an underestimation of collapse risk for structures [89, 90]. Although the latest NBCC [2] has incorporated the Cascadia subduction earthquake into a probabilistic manner, the duration effects are not considered in the selection process. In order to accurately quantify the effects of ground motion duration on structural responses, the duration must be decoupled from other characteristics of the ground motion. In this study, a new methodology by Zengin et al. [13] was applied to isolates the duration from the amplitude, frequency content, and rate of energy build-up of the ground motion. This can be achieved by selecting long- and short-duration record pairs that have similar spectral shape and slope of the Husid plot. In this way, the sole effect of the duration on the structural response of the building can be examined. For this purpose, first, long duration records, from S2GM ground motion data base, were scaled and spectrally matched to Vancouver UHS (with 2% probability of exceedance in 50 years) with Site Class C soil condition over a period range of 0.15T to 1.5s, where T is the fundamental period of the building [10]. Next, each long duration motion was a target spectrum for PEER NGA-West2 database meaning that for each selected long duration record, the corresponding short duration record with inherently similar response spectrum was chosen from the database. After constructing a subset of spectrally equivalent ground-motion pairs, a new methodology by Zengin et al. [13] were applied to identify the record pairs having a similar rate of energy build-up. To the best of my knowledge, this is the second study that uses the rate of energy build up as a control parameter. Figure 3-11 compares the response spectra and Husid plots of the ground 40 motions pair #3, respectively. As can be seen in the figure, both records display a very similar spectral shape and amplitude over a wide period range (Figure 3-11a), and the rate of energy build-up between the two records is similar over the significant duration (Figure 3-11b). The ground motion at Ciencias Argronomicas station during Maule earthquake is longer, with 38.6 seconds significant duration, than that from Chi-Chi (TCU053) earthquake, with 22.3 seconds significant duration. 3 Long: Maule (Chile) 1.2 Short: Chi-Chi (TCU053) 1 0.8 0.6 0.4 Short: Chi-Chi (TCU053) 2 1.5 1 0.5 0.2 0 0 0 a) Long: Maule (Chile) 2.5 5-95% IA [m/s] Spectral Acceleration [g] 1.4 1 2 Period [s] 3 0 4 b) 20 40 60 Time [s] Figure 3-11: Comparison of the (a) response spectra and (b) Husid plots of the SD and LD record pair Finally, a database that consists of 24 pairs of long duration (LD) and short duration (SD) motions was generated. Long and short duration motions were classified based on their significant duration. For comparison purpose with previous studies, the 5-95% significant duration D5-95 is used here. The records with a D5-95 smaller than 30 s are considered SD records, whereas records with D5-95 equal or greater than 30 s are considered LD records, based on recommendations by Hou and Qu [9]. Table 3-4 presents the complete information of all 24 pairs of selected ground motions. Most of the LD records were selected from subduction events, such as the 2003 and 2011 Japan earthquakes, the 1985 Mexico earthquake, and the 2010 Chile earthquake. 41 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 Event Hokkaido (Japan) Tohoku (Japan) Maule (Chile) Hokkaido (Japan) Tohoku (Japan) Tohoku (Japan) Hokkaido (Japan) Tohoku (Japan) Tohoku (Japan) Tohoku (Japan) Tohoku (Japan) Maule (Chile) Maule (Chile) Maule (Chile) Tohoku (Japan) Tohoku (Japan) Tohoku (Japan) Michoacan (Mexico) Hokkaido (Japan) Hokkaido (Japan) Tohoku (Japan) Tohoku (Japan) Tohoku (Japan) Tohoku (Japan) Year 2003 2011 2010 2003 2011 2011 2003 2011 2011 2011 2011 2010 2010 2010 2011 2011 2011 1985 2003 2003 2011 2011 2011 2011 Station Horokeshi Hiratsuka-st5 Ciencias Agronomicas Urakawa Hachiohji Takasaki Akan Okudo Koganei Shinozaki Nakoso Santiago Center Colegio Las Americas La Florida Gyoutoku Nishiaidu Chiba Villita Hobetsu Oiwake Koga Hiratsuka-st1 Kawagoe Tatsumi Pair Long duration subduction motions Mw 8.30 9.00 8.80 8.30 9.00 9.00 8.30 9.00 9.00 9.00 9.00 8.80 8.80 8.80 9.00 9.00 9.00 8.10 8.30 8.30 9.00 9.00 9.00 9.00 R [km] 121 270 167 65 216 197 92 177 219 177 60 176 183 174 178 176 174 16 134 161 155 273 190 188 D5-95 [s] 32.90 118.10 38.60 39.20 75.00 71.60 35.70 108.90 67.80 107.70 88.10 35.20 37.10 39.90 102.00 90.90 93.50 44.10 40.90 44.60 93.90 127.60 73.40 120.10 42 Event Chi-Chi Chi-Chi Chi-Chi Loma Pierta Chi-Chi Manjil Kalamata, Greece-01 Taiwan Smart-01 Taiwan Smart-01 Taiwan Smart-01 Imperial Valley-06 Imperial Valley-06 Taiwan Smart-01 Chi-Chi Loma Prieta Imperial Valley-06 Superstition Hills-02 Loma Prieta Superstition Hills-02 Northridge-01 Loma Prieta Chi-Chi Imperial Valley-06 Imperial Valley-06 Year 1999 1999 1999 1989 1999 1990 1986 1986 1986 1986 1979 1979 1986 1999 1989 1979 1987 1989 1987 1994 1989 1999 1979 1979 Station TCU054 TCU029 TCU053 Hollister Differential Array CHY035 Abhar Kalamata SMART1-I01 SMART1-O02 SMART1-I02 El Centro Array #3 Brawley Airport SMART1-E01 TCU051 Oakland-Outer Harbor Holtville Post Office Kornbloom Road Palo Alto - SLAC Lab Poe Road Santa Monica City Hall Anderson Dam (L Abut) CHY028 El Centro Array #1 EC County Center FF Short duration crustal motions Table 3-4: Selected ground motions Mw 7.60 7.60 7.60 6.90 6.30 7.40 6.20 7.30 7.30 7.30 6.50 6.50 6.30 7.60 6.90 6.50 6.50 6.90 6.50 6.70 6.90 6.30 6.50 6.50 R [km] 5 28 6 25 42 76 6 56 57 56 13 10 56 8 74 8 18 31 11 26 20 34 22 7 D5-95 [s] 24.50 23.60 22.30 12.40 13.20 21.10 5.00 20.40 16.30 21.60 11.90 14.10 8.70 24.20 8.70 11.80 14.00 11.60 13.60 10.70 12.70 12.10 15.00 10.40 Figure 3-12 presents the distribution in duration among the selected ground motions. The median significant duration D5-95 for the LD records and the spectrally equivalent SD Number of Records records are 71.95 s and 15.00 s, respectively. 16 14 12 10 8 6 4 2 0 14 7 7 3 3 3 3 1 3 1 1 2 D5-95 [s] Figure 3-12: Distribution of ground motion duration Figure 3-13 shows the response spectrum of each individual long and short duration motions matched to the UHS of Vancouver. 2.5 ith SD motion 2 Spectral Acceleration [g] Spectral Acceleration [g] 2.5 2015 Vancouver UHS 1.5 ith LD motion 2 2015 Vancouver UHS 1.5 Mean Spectrum 1 Mean Spectrum 1 0.5 0.5 0 0 a) 1 2 Period [s] 3 0 0 4 b) 1 2 Period [s} 3 4 Figure 3-13: Comparison of response spectra of selected motions with the UHS of Vancouver: (a) SD and (b) LD. 3.5 Analyses Since it is impossible to have a pair of long and short duration motions that have equivalent spectra for both horizontal components, in this study, unidirectional NLTHA was used to evaluate the response of the building under the two sets of motions at the 43 design intensity level. Since the long direction of the building was identified as the weak direction (E-W), the building was subjected to the earthquakes in this direction. Next, the effects of ground motion duration on the collapse capacity of the building were evaluated by incremental dynamic analysis (IDA). To perform IDA, each ground motion scaled to multiple intensity levels to approach collapse. At the intensity level 480% of UHS, the building collapsed under all the scaled motions. Thus, in this study, the scaling range 100% to 480% of UHS design intensity was selected with an increment of 20% to cover a wide range of intensity levels and the maximum IDR was used to monitor structural response. In total, 960 NLTHA (2 sets of 24 motions at 20 intensity levels) were carried out. Afterwards, based on the IDA results, the fragility curves for the building under the two sets of motions were generated. The fragility curves were constructed by assuming a lognormal distribution. The cumulative probability of occurrence of damage equal to or higher than the specified drift limit can be calculated as Equation 3-1 [91]: ( | = )=Ф Where ( | 3-1 ( ⁄ ) = ) is the probability that a ground motion record with IM equals to x will cause the structure to collapse. Φ () is the standard normal cumulative distribution function (CDF). θ is the median of the fragility function and β is the standard deviation of ln IM [91]. 44 4 Results 4.1 Nonlinear time history analysis at design level The response of the building subjected to the two sets of records were calculated using unidirectional NLTHA at design intensity level. Figure 4-1shows the IDR of first (1F) and second floor (2F) of the building under SD and LD motions in E-W direction. The mean maximum IDR for the two record sets is close to 0.7%, which is well below the 2% drift limit specified in the 2015 NBCC for high importance buildings. The difference in IDR under the two sets of motions at the design intensity level is only 0.02%. This is consistent with results from previous studies conducted on concrete and steel frame structures [17, 64, 66]. 2 0.52% 1 ith LD motion Mean Mean Drift limit Drift limit 0.72% 0 1 0.70% 0 0% a) 0.51% ith SD motion Floor Floor 2 1% Maximum IDR 2% 0% b) 1% 2% Maximum IDR Figure 4-1: Maximum IDR for a) SD and b) LD motions Figure 4-2 shows the mean maximum base shear (BS) forces of the building in E-W direction under SD and LD motions. There is a very small difference, only 21kN, between BS values of SD and LD motions. 45 Mean base shear (kN) 2500 2052 2031 SD LD 2000 1500 1000 500 0 Figure 4-2: Mean maximum base shear of the building for SD and LD motions 4.2 Incremental dynamic analysis Figure 4-3 shows IDA curves for the building model subjected to long and short duration motions, 48 records, at different intensity levels in E-W direction. In this figure, the 100% scaling level refers to the 2% in 50 year shaking level (the code design level according to the 2015 NBCC). Each black dot on the curve in Figure 4-3 represents the onset of collapse, which means the building will collapse at the next intensity level. It should be noted that there is no defined collapse drift limit for balloon-type CLT buildings (FEMA P695 [92] defines 7% for light-frame wood buildings), so the collapse under each record is judged to occur directly from lateral dynamic instability (excessive lateral displacement). At the design intensity level (100% of UHS), the building performs quite well and maximum IDRs were less than 1% under both sets of ground motions. However, with the increase of intensity, the influence of duration can be observed. The IDA curves indicated that the SD motions on average reach a maximum IDR of 4.8% before causing collapse. In contrast, the LD motions reach a maximum IDR of 3.6% before collapse. This shows that when large inelastic deformations begin to occur in the structure, caused by the resulting deterioration in strength and stiffness, a LD motion is more likely to lead to structural collapse. 46 650% ith SD motion Shaking level of UHS 600% 550% 0nset of collapse 500% Mean peak IDR before collapse 450% 400% 350% 300% 250% 200% 150% 100% 50% 0% 0% 2% 4% a) 6% Maximum IDR 8% 10% 12% 8% 10% 12% 650% ith LD motion 600% 550% 0nset of collapse 500% Mean peak IDR before collapse Shaking level of UHS 450% 400% 350% 300% 250% 200% 150% 100% 50% 0% 0% b) 2% 4% 6% Maximum IDR Figure 4-3: IDA curves of a) SD and b) LD motions Figure 4-4 shows the nonlinear behavior of a critical HD and WB under a SD (Imperial Valley-06) and a LD (Tohoku-Japan at Koga station) record (red highlighted circles in 47 Figure 4-3) before collapse. The Nonlinear behavior of these connections under all the records at collapse and before collapse are presented in Appendix A. These results show that the critical connections could perform at onset of collapse (black dots shown in Figure 4-3). However, at the next intensity level, they showed failure which resulted in lateral dynamic instability of the building. Critical HD-Tension 250 150 150 -150 -150 Force (kN) -50 Force (kN) 250 50 Critical WB-Shear 50 -50 -250 -250 -10 0 a) 10 20 Displacement (mm) 30 -400 b) Critical HD-Tension Critical WB-Shear 250 150 150 -150 -150 Force (kN) -50 Force (kN) 250 50 50 -50 -250 -250 -10 c) -200 0 200 Displacement (mm) 0 10 20 Displacement (mm) 30 -200 d) -100 0 Displacement (mm) 100 Figure 4-4: Nonlinear behavior of a critical (a,b) HD and WB connection under the SD record; (c,d) HD and WB connection under the LD record before collapse. 4.3 Fragility assessment To calculate the collapse rate of the building under both long and short motions, collapse fragility curves were derived from the IDA results. Figure 4-5 shows the fragility curves for the two sets of motions. In this figure, the black dots refer to the empirical collapse 48 data points and the red line is the CDF by fitting a lognormal distribution through these empirical points. 100% Probability of collapse Empirical CDF 80% Lognormal fit CDF 60% 40% 20% 0% 0% 100% 200% 300% 400% 500% 400% 500% Shaking level of UHS a) 100% Empirical CDF Probability of collapse 80% Lognormal fit CDF 60% 40% 20% 0% 0% b) 100% 200% 300% Shaking level of UHS Figure 4-5: Fragility curves for collapse. (a) SD and (b) LD motions. Figure 4-6 presents the comparison between the CDF results of the two ground motion suites. As shown in the figure, the building had 0% probability of collapse at the design intensity level under the two sets of motions. However, with increasing intensity, the building starts to exhibit a higher probability of collapse under LD motions. For the purpose of ensuring a sufficient margin of safety against collapse, FEMA P695 uses the collapse margin ratio (CMR), which is the ratio between the median collapse shaking level and the 2% in 50 year shaking level. The median collapse capacity (50% probability 49 of collapse) for LD motions were determined at 292% of UHS intensity level. In contrast, the median collapse capacity for SD motions were determined at 318% of UHS. This implies that the SD suite, on average, required scaling factors 9% (318% to 292%) greater than the LD suite to induce structural collapse. It should be noted that uncertainty was not considered in CMR evaluation. These results show that ground motion duration can affect the collapse risk and reduce the margin of safety for CLT buildings. Probability of collapse 100% LD SD 75% CMR-Short CMR-Long 50% 25% 0% 0% 100% 200% 300% 400% 500% Shaking level of UHS Figure 4-6: CDF results for SD and LD motions 4.4 Summary This chapter studied the effects of ground motion duration on the seismic response of the case study building. Based on the results, at design intensity level, the 2% in 50 year shaking level, there was not much difference in the building’s response under the long and short duration motions, the maximum difference in mean IDR was only 0.02%. The mean maximum IDRs for the two record sets were almost 0.7%, which is well below the 2% drift limit specified in the 2015 NBCC and the building had 0% probability of collapse under the two records at the design level. However, with increasing intensity, the influence of duration became more significant. The mean maximum IDRs before collapse under LD and SD motions were 3.6% and 4.8%, respectively, indicating that a LD ground 50 motion is more likely to induce structural collapse. Similarly, it was shown that LD motions reduced the median collapse capacity of the building by 9% (318% to 292%) compared to SD motions. These results indicated that ground motion duration may not be a critical factor at design levels of shaking. Nevertheless, at higher levels of shaking, where larger levels of damage are expected, duration becomes a relevant factor. The conclusion that ground motion duration can reduce the median collapse scaling level of structures is consistent with other studies carried out on wood frame buildings (e.g., Pan et al. [14, 67, 68]). 51 5 Conclusions 5.1 Summary of results While previous research indicated that LD motions may be more damaging than SD motions, current design codes do not explicitly take duration effects into account. Further, the effect of ground motion duration on CLT buildings has not been studied. The findings of this study could help to better understanding of the effects of ground motion duration on structural performance of CLT buildings. In this thesis, the effects of ground motion duration on a newly constructed two-storey balloon-type CLT building, designed for Vancouver, BC, were evaluated. The balloon shear walls were tested at UNBC, and the test results were used to validate an OpenSees balloon shear wall model with calibrated connector properties. A database including 24 pairs of long and short duration records (48 records, in total) were established to isolate the effect of duration from other ground motion’s characteristics. These record pairs were equated with respect to amplitude, frequency content and the rate of energy build up so that the only difference between them is duration. Then the seismic response of the building was investigated using unidirectional incremental dynamic analysis. The methodology presented by Zengin et al. [13], in which the rate of energy build up of the ground motions is considered in addition to amplitude and frequency, was used to isolate duration from other characteristics of ground motions. Based on the IDA curves, the average maximum IDR before collapse for LD and SD motions is 3.6% and 4.8%, respectively, meaning that a LD motion is more likely to cause structural collapse when large inelastic deformations take place. However, at design intensity level, ground motion duration has been shown to not be a critical factor as the difference between inter-storey drift ratio under the two sets of records was negligible. 52 It should be noted that the results presented in this study were for an elementary school, which was designed for high importance category. Moreover, the building design is very conservative as the mean maximum IDRs for the two record sets was only 0.72%, which is well below the 2% drift limit specified in the 2015 NBCC. Thus, for a taller and normal importance category CLT building, the effects of duration at design intensity level might be different. 5.2 Outlook This study focused on a high importance category two-storey balloon-framed CLT building. The influence of ground motion duration on normal importance category CLT buildings, high-rise CLT buildings and hybrid timber buildings, where wood is used in combination with concrete or steel, should also be investigated. In this study, due to the complexities of the case study building, inter-storey drift ratio was used as a damage measure to evaluate the effects of duration. Further research is needed to consider other damage measures (e.g., energy demand or damage index) to assess the duration effects. 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Council, A.T., Quantification of building seismic performance factors. 2009: US Department of Homeland Security, FEMA. 61 Appendix A Kalamata, Greece-01 (300% UHS) Critical HD-tension 300 150 200 100 Force (kN) Force (kN) Critical WB-shear 200 100 0 -100 50 0 -50 -100 -200 -150 -300 -200 -10 0 10 20 -40 30 -20 0 20 40 Displacement (mm) Displacement (mm) Taiwan Smart-01 -SMART1-E01 (260% UHS) Critical WB-shear Critical HD-tension 200 300 150 100 Force (kN) Force (kN) 200 100 0 -100 50 0 -50 -100 -200 -150 -200 -300 -10 0 10 20 -100 30 -50 0 50 100 Displacement (mm) Displacement (mm) Loma Prieta-Oakland-Outer Harbor Wharf (260% UHS) Critical WB-shear Critical HD-tension 200 300 150 200 Force (kN) Force (kN) 100 100 0 -100 50 0 -50 -100 -200 -150 -300 -200 -10 0 10 20 30 -40 Displacement (mm) -20 0 20 Displacement (mm) 62 40 Imperial Valley-06-EC County Center FF (320%UHS) Critical HD-tension Critical WB-shear 200 300 150 200 Force (kN) Force (kN) 100 100 0 -100 50 0 -50 -100 -200 -150 -200 -300 -10 0 10 20 -150 30 Displacement (mm) -100 -50 0 50 Displacement (mm) Northridge-01-Santa Monica City Hall (280% UHS) Critical WB-shear 300 150 200 100 50 100 Force (kN) Force (kN) Critical HD-tension 0 -100 0 -50 -100 -200 -150 -300 -200 -10 0 10 -40 20 -20 0 20 40 Displacement (mm) Displacement (mm) Loma Prieta-Palo Alto - SLAC Lab (280%UHS) Critical HD-tension 100 Force (kN) 200 Force (kN) Critical WB-shear 150 300 100 0 -100 -200 50 0 -50 -100 -150 -300 -200 -10 0 10 20 -40 Displacement (mm) -20 0 Displacement (mm) 63 20 Imperial Valley-06 (340%UHS) Critical HD-tension 300 150 200 100 100 Force (kN) Force (kN) Critical WB-shear 200 0 -100 50 0 -50 -100 -200 -150 -300 -200 -10 0 10 20 30 -200 Displacement (mm) 0 200 400 Displacement (mm) Imperial Valley-06-El Centro Array #3 (360%UHS) Critical HD-tension Critical WB-shear 300 100 Force (kN) Force (kN) 200 0 -100 -200 -300 -10 0 10 20 30 200 150 100 50 0 -50 -100 -150 -200 -100 Displacement (mm) 0 100 200 Displacement (mm) CHICHI_CHY028-N (280% UHS) Critical WB-shear Critical HD-tension 300 200 200 150 100 Force (kN) Force (kN) 100 0 -100 50 0 -50 -200 -100 -300 -150 -5 0 5 10 15 -50 20 0 50 100 Displacement (mm) Displacement (mm) 64 150 Loma Pierta-Hollister Differential Array (280% UHS) Critical HD-tension Critical WB-shear 200 150 100 50 0 -50 -100 -150 -200 300 Force (kN) Force (kN) 200 100 0 -100 -200 -300 -10 0 10 20 30 -150 -100 Displacement (mm) -50 0 50 100 Displacement (mm) Loma Prieta-Anderson Dam (L Abut) (300% UHS) Critical WB-shear 500 0 -500 -1,000 -1,500 -2,000 -2,500 -3,000 -3,500 -4,000 200 150 100 50 0 -50 -100 -150 -200 Force (kN) Force (kN) Critical HD-tension -40 -20 0 20 40 -150 -100 -50 0 50 100 150 Displacement (mm) Displacement (mm) CHICHI_CHY035-N (180% UHS) Critical WB-shear 200 200 150 100 Force (kN) Force (kN) Critical HD-tension 300 0 -100 -200 100 50 0 -50 -100 -150 -300 -5 0 5 10 -20 15 0 20 40 Displacement (mm) Displacement (mm) 65 60 Superstition Hills-02-Poe Road (440% UHS) Critical HD-tension Critical WB-shear 200 300 150 100 100 Force (kN) Force (kN) 200 0 -100 50 0 -50 -100 -200 -150 -200 -300 -10 0 10 20 -200 -100 Displacement (mm) 0 100 200 Displacement (mm) Superstition Hills-02 (340% UHS) Critical WB-shear 300 200 200 150 100 100 Force (kN) Force (kN) Critical HD-tension 0 -100 50 0 -50 -100 -200 -150 -300 -200 -10 0 10 20 30 -200 Displacement (mm) -100 0 100 200 300 Displacement (mm) Imperial Valley-06-Brawley Airport (440% UHS) Critical WB-shear Critical HD-tension 200 300 150 Force (kN) Force (kN) 200 100 0 -100 100 50 0 -50 -100 -200 -150 -300 -10 0 10 20 -200 30 -400 Displacement (mm) 66 -200 0 Displacement (mm) 200 Imperial Valley-06-El Centro Array #1 (320% UHS) Critical WB-shear Critical HD-tension 300 200 150 200 Force (kN) Force (kN) 100 100 0 -100 50 0 -50 -100 -200 -150 -300 -200 -10 0 10 20 30 -100 -50 0 50 Displacement (mm) Displacement (mm) Taiwan Smart-01-SMART1-O02 (340% UHS) Critical HD-tension 300 150 200 100 100 Force (kN) Force (kN) Critical WB-shear 200 0 -100 -200 50 0 -50 -100 -300 -150 -10 0 10 20 30 -40 Displacement (mm) -20 0 20 40 Displacement (mm) Taiwan Smart-01-SMART1-I01 (380% UHS) Critical WB-shear Critical HD-tension 300 200 200 150 100 Force (kN) Force (kN) 100 0 -100 50 0 -50 -100 -200 -150 -200 -300 -20 0 20 40 -100 60 0 100 200 Displacement (mm) Displacement (mm) 67 300 MANJIL_184327 (260% UHS) Critical HD-tension Critical WB-shear 300 200 200 150 100 Force (kN) Force (kN) 100 0 50 0 -50 -100 -100 -200 -150 -300 -200 -10 0 10 20 Displacement (mm) 30 -200 -100 0 100 Displacement (mm) 200 Taiwan Smart-01-SMART1-I02 (300% UHS) Critical HD-tension Critical WB-shear 200 300 150 Force (kN) Force (kN) 200 100 0 -100 100 50 0 -50 -100 -200 -150 -300 -200 -10 0 10 20 30 40 -150 -100 Displacement (mm) -50 0 50 100 100 150 Displacement (mm) CHICHI_TCU053-E (340% UHS) Critical WB-shear 200 200 150 100 Force (kN) Force (kN) Critical HD-tension 300 0 -100 100 50 0 -50 -200 -100 -300 -150 -10 0 10 20 30 -50 0 50 Displacement (mm) Displacement (mm) 68 CHICHI_TCU029-E (340% UHS) Critical HD-tension 300 150 200 100 100 Force (kN) Force (kN) Critical WB-shear 200 0 -100 50 0 -50 -100 -200 -150 -200 -300 -10 0 10 20 -150 30 Displacement (mm) -100 -50 0 50 Displacement (mm) CHICHI_TCU051-E (200% UHS) Critical WB-shear 150 200 100 50 100 Force (kN) Force (kN) Critical HD-tension 300 0 -100 -200 0 -50 -100 -150 -200 -300 -5 0 5 10 15 -40 20 Displacement (mm) -20 0 20 40 Displacement (mm) CHICHI_TCU054-E (180% UHS) Critical WB-shear 200 200 150 Force (kN) Force (kN) Critical HD-tension 300 100 0 100 50 0 -100 -50 -200 -100 -150 -300 -10 0 10 20 -40 30 -20 0 20 Displacement (mm) Displacement (mm) 69 40 Hokkaido (Japan)-Horokeshi (280% UHS) Critical HD-tension Critical WB-shear 200 300 150 Force (kN) Force (kN) 200 100 0 -100 100 50 0 -50 -100 -200 -150 -300 -200 -10 0 10 20 Displacement (mm) 30 -50 0 50 100 Displacement (mm) Maule (Chile)-Santiago Center (420% UHS) Critical WB-shear 300 200 200 150 Force (kN) Force (kN) Critical HD-tension 100 0 -100 100 50 0 -50 -100 -200 -150 -300 -200 -10 0 10 20 30 -200 -150 -100 -50 0 50 Displacement (mm) Displacement (mm) Hokkaido (Japan)-Akan (420% UHS) Critical HD-tension Critical WB-shear 300 200 150 Force (kN) 200 Force (kN) 100 0 -100 100 50 0 -50 -100 -200 -150 -200 -300 -10 0 10 20 -150 30 Displacement (mm) -100 -50 0 Displacement (mm) 70 50 Maule (Chile)-Colegio Las Americas (460% UHS) Critical HD-tension Critical WB-shear 200 300 150 200 Force (kN) 100 Force (kN) 100 0 -100 50 0 -50 -100 -200 -150 -200 -300 -10 0 10 20 30 -150 40 -100 Displacement (mm) -50 0 50 100 Displacement (mm) Maule (Chile)-Ciencias Agronomicas (240% UHS) Critical HD-tension Critical WB-shear 100 50 Force (kN) 100 Force (kN) 200 0 -100 -200 0 -50 -100 -150 -300 -200 -5 0 5 10 15 -40 Displacement (mm) -30 -20 -10 Displacement (mm) 0 10 Hokkaido (Japan)-Urakawa (220% UHS) Critical WB-shear 150 200 100 100 Force (kN) Force (kN) Critical HD-tension 300 0 -100 50 0 -50 -100 -200 -150 -200 -300 -5 0 5 10 -40 15 Displacement (mm) -20 0 Displacement (mm) 71 20 Maule (Chile)-La Florida (360% UHS) Critical HD-tension Critical WB-shear 300 200 150 200 100 Force (kN) Force (kN) 100 0 -100 50 0 -50 -100 -200 -150 -300 -200 -10 0 10 20 -200 30 -150 -100 -50 0 50 Displacement (mm) Displacement (mm) Hokkaido (Japan)-Hobetsu (380% UHS) Critical HD-tension Critical WB-shear 300 200 200 150 100 Force (kN) Force (kN) 100 0 -100 50 0 -50 -100 -200 -150 -300 -200 -5 0 5 10 15 20 25 -100 -50 0 50 Displacement (mm) Displacement (mm) Michoacan (Mexico) (320% UHS) Critical HD-tension Critical WB-shear 200 300 150 200 Force (kN) 100 Force (kN) 100 0 -100 50 0 -50 -100 -200 -150 -300 -200 -10 0 10 20 -250 30 Displacement (mm) -200 -150 -100 -50 Displacement (mm) 72 0 50 Hokkaido (Japan)-Oiwake (300% UHS) WB No. 107 Shear 200 200 150 100 Force (kN) Force (kN) Critical HD-tension 300 0 -100 100 50 0 -50 -200 -100 -150 -300 -10 0 10 20 -40 30 -20 Displacement (mm) 0 20 40 Displacement (mm) Tohoku (Japan)-Koganei (140% UHS) Critical HD-tension Critical WB-shear 100 200 150 100 50 0 -50 -100 -150 -200 -250 -300 Force (kN) Force (kN) 50 0 -50 -100 -150 -5 0 5 10 15 -15 Displacement (mm) -10 -5 0 5 10 Displacement (mm) Tohoku (Japan)-Takasaki (260% UHS) Critical WB-shear Critical HD-tension 300 150 200 100 Force (kN) Force (kN) 100 0 -100 50 0 -50 -100 -200 -150 -300 -200 -5 0 5 10 15 -100 Displacement (mm) -50 0 Displacement (mm) 73 50 Tohoku (Japan)-Kawagoe (300% UHS) Critical HD-tension Critical WB-shear 300 200 200 150 100 Force (kN) Force (kN) 100 0 -100 50 0 -50 -100 -200 -150 -300 -200 -10 0 10 20 -40 Displacement (mm) -20 0 20 40 Displacement (mm) Hokkaido (Japan)-Hachiohji (300% UHS) Critical HD-tension Critical WB-shear 300 200 200 150 100 Force (kN) Force (kN) 100 0 -100 50 0 -50 -200 -100 -300 -150 -200 -400 -10 0 10 20 -50 30 Displacement (mm) 0 50 100 Displacement (mm) Tohoku (Japan)-Nakoso (260% UHS) Critical HD-tension 300 Critical WB-shear 200 150 200 Force (kN) Force (kN) 100 100 0 -100 50 0 -50 -100 -200 -150 -300 -200 -10 0 10 20 30 -50 Displacement (mm) 0 50 Displacement (mm) 74 100 Tohoku (Japan)-Nishiaidu (260% UHS) Critical HD-tension 300 100 200 Force (kN) 100 Force (kN) Critical WB-shear 150 0 -100 50 0 -50 -100 -200 -150 -300 -200 -5 0 5 10 15 -40 Displacement (mm) -20 0 20 Displacement (mm) Tohoku (Japan)-Chiba (160% UHS) Critical HD-tension Critical WB-shear 200 100 100 50 Force (kN) 150 Force (kN) 300 0 -100 0 -50 -200 -100 -300 -150 -5 0 5 10 -10 15 Displacement (mm) 0 10 20 Displacement (mm) Tohoku (Japan)-Koga (300% UHS) Critical HD-tension Critical WB-shear 200 300 150 200 100 Force (kN) Force (kN) 100 0 -100 50 0 -50 -100 -200 -150 -200 -300 -10 0 10 20 -200 30 Displacement (mm) -100 0 Displacement (mm) 75 100 Tohoku (Japan)-Gyoutoku (220% UHS) Critical HD-tension Critical WB-shear 300 150 200 100 50 Force (kN) Force (kN) 100 0 -100 0 -50 -100 -200 -150 -300 -200 -5 0 5 10 -100 15 Displacement (mm) -50 0 50 Displacement (mm) Tohoku (Japan)-Shinozaki (280% UHS) Critical HD-tension Critical WB-shear 200 300 150 200 100 Force (kN) Force (kN) 100 0 -100 50 0 -50 -100 -200 -150 -300 -10 0 10 20 -200 30 -100 Displacement (mm) -50 0 50 100 30 50 Displacement (mm) Tohoku (Japan)-Okudo (240% UHS) Critical WB-shear Critical HD-tension 200 150 100 100 Force (kN) 200 Force (kN) 300 0 -100 50 0 -200 -50 -300 -100 -150 -400 -10 0 10 20 -30 30 Displacement (mm) -10 10 Displacement (mm) 76 Tohoku (Japan)-Hiratsuka-st5 (260% UHS) Critical HD-tension 300 150 200 Force (kN) 100 Force (kN) Critical WB-shear 200 0 -100 100 50 0 -50 -200 -100 -300 -150 -10 0 10 20 Displacement (mm) 30 -40 -20 0 20 40 Displacement (mm) Tohoku (Japan)-Tatsumi (220% UHS) Critical HD-tension Critical WB-shear 200 100 100 50 Force (kN) 150 Force (kN) 300 0 -100 0 -50 -100 -200 -150 -300 -10 0 10 20 -40 30 -20 Displacement (mm) 0 20 40 Displacement (mm) Tohoku (Japan)-Hiratsuka-st1 (180% UHS) Critical HD-tension 200 100 Force (kN) 100 Force (kN) Critical WB-shear 150 0 -100 -200 50 0 -50 -100 -300 -150 -5 0 5 10 15 -15 Displacement (mm) 77 -10 -5 0 5 Displacement (mm) 10 15 Kalamata, Greece-01- (320% UHS-Failure) Critical HD-tension Critical WB-shear 200 1E+10 150 0 Force (kN) 100 Force (kN) -1E+10 -2E+10 -3E+10 50 0 -50 -100 -4E+10 -5E+10 -4E+08 -150 -2E+08 0 200000000 -200 -1,500 -1,000 -500 0 500 Displacement (mm) Displacement (mm) Taiwan Smart-01 -SMART1-E01 (280% UHS- Failure) Critical HD-tension 200 1E+10 Critical WB-shear 150 0 Force (kN) Force (kN) 100 -1E+10 -2E+10 -3E+10 50 0 -50 -100 -4E+10 -5E+10 -4E+08 -150 -2E+08 0 -200 -10,000 200000000 Displacement (mm) -5,000 0 Displacement (mm) 5,000 Loma Prieta-Oakland-Outer Harbor Wharf (280% UHS- Failure) Critical WB-shear Critical HD-tension 1E+10 200 150 Force (kN) Force (kN) 0 -1E+10 -2E+10 -3E+10 100 50 0 -50 -100 -4E+10 -5E+10 -4E+08 -150 -2E+08 0 200000000 Displacement (mm) -200 -1,000 -500 0 Displacement (mm) 78 500 Imperial Valley-06-EC County Center FF (340%UHS- Failure) Critical WB-shear Critical HD-tension 0 100 -1E+10 50 Force (kN) Force (kN) 1E+10 150 -2E+10 -3E+10 0 -50 -4E+10 -100 -5E+10 -150 -6E+10 -4E+08 -200 -2E+08 0 200000000 -500 0 500 Displacement (mm) 1,000 1,500 2,000 Displacement (mm) Northridge-01-Santa Monica City Hall (300% UHS- Failure) Critical HD-tension Critical WB-shear 200 150 100 Force (kN) Force (kN) 100 0 -100 50 0 -50 -100 -200 -150 -200 -3,000 -300 -5 0 5 10 15 Displacement (mm) -2,000 -1,000 0 1,000 Displacement (mm) Loma Prieta-Palo Alto - SLAC Lab (300%UHS- Failure) Critical HD-tension 150 0 100 Force (kN) 1E+10 Force (kN) -1E+10 -2E+10 -3E+10 50 0 -50 -100 -4E+10 -5E+10 -4E+08 Critical WB-shear -150 -2E+08 0 200000000 Displacement (mm) -200 -500 0 Displacement (mm) 79 500 Imperial Valley-06 (360% UHS – Failure) 1E+10 Critical WB-shear Critical HD-tension 200 150 100 -1E+10 Force (kN) Force (kN) 0 -2E+10 -3E+10 50 0 -50 -100 -4E+10 -5E+10 -4E+08 -150 -200 -2E+08 0 -400 -200 0 Displacement (mm) 200 400 600 Displacement (mm) Imperial Valley-06-El Centro Array #3 (380% UHS- Failure) Critical HD-tension Critical WB-shear 1E+10 150 0 100 Force (kN) Force (kN) 50 -1E+10 -2E+10 -50 -100 -3E+10 -150 -4E+10 -5E+10 -4E+08 0 -200 -2E+08 0 200000000 -250 -1E+08 Displacement (mm) 0 100000000 200000000 Displacement (mm) CHICHI_CHY028-N (300% UHS- Failure) Critical WB-shear 200 0 150 -1E+10 Force (kN) Force (kN) Critical HD-tension 1E+10 -2E+10 -3E+10 -4E+10 -5E+10 -4E+08 100 50 0 -50 -100 -150 -2E+08 0 200000000 Displacement (mm) -500 0 Displacement (mm) 80 500 Loma Pierta-Hollister Differential Array (300%UHS- Failure) Critical HD-tension Critical WB-shear 150 100 50 Force (kN) Force (kN) 5E+09 0 -5E+09 -1E+10 -1.5E+10 -2E+10 -2.5E+10 -3E+10 -3.5E+10 -4E+10 -4.5E+10 -5E+10 -4E+08 0 -50 -100 -150 -2E+08 0 200000000 -200 -1500 -1000 Displacement (mm) -500 0 500 Displacement (mm) Loma Prieta-Anderson Dam (L Abut) (320% UHS- Failure) 1E+10 Critical HD-tension 150 -1E+10 100 Force (kN) 0 Force (kN) -2E+10 -3E+10 -4E+10 50 0 -50 -5E+10 -100 -6E+10 -150 -7E+10 -6E+08 Critical WB-shear 200 -200 -2E+08 -400 200000000 -200 Displacement (mm) 0 200 Displacement (mm) CHICHI_CHY035-N (200% UHS- Failure) Critical HD-tension Critical WB-shear 1E+10 150 0 100 Force (kN) Force (kN) -1E+10 -2E+10 50 0 -3E+10 -50 -4E+10 -100 -5E+10 -4E+08 -2E+08 0 200000000 Displacement (mm) -150 -1000 0 1000 2000 Displacement (mm) 81 3000 5E+12 0 -5E+12 -1E+13 -1.5E+13 -2E+13 -2.5E+13 -3E+13 -3.5E+13 -4E+13 -4.5E+13 -5E+13 -4E+08 Critical HD-tension Critical WB-shear 200 150 100 Force (kN) Force (kN) Superstition Hills-02-Poe Road (460% UHS- Failure) 50 0 -50 -100 -150 -2E+08 0 200000000 -200 -1000 0 Displacement (mm) 1000 2000 Displacement (mm) Superstition Hills-02 (360% UHS- Failure) Critical HD-tension 150 1E+10 100 0 -1E+10 Force (kN) Force (kN) Critical WB-shear -2E+10 -3E+10 50 0 -50 -100 -4E+10 -150 -5E+10 -4E+08 -2E+08 0 200000000 Displacement (mm) -200 -1000 0 1000 2000 Displacement (mm) 3000 Imperial Valley-06-Brawley Airport (460% UHS- Failure) Critical HD-tension Critical WB-shear 300 200 150 200 Force (kN) Force (kN) 100 100 0 -100 50 0 -50 -100 -200 -150 -200 -300 -10 0 10 20 -400 30 -200 0 Displacement (mm) Displacement (mm) 82 200 Imperial Valley-06-El Centro Array #1 (340% UHS- Failure) 1E+10 Critical HD-tension Critical WB-shear 200 150 0 Force (kN) Force (kN) 100 -1E+10 -2E+10 50 0 -50 -3E+10 -100 -4E+10 -5E+10 -4E+08 -150 -2E+08 0 200000000 -200 -1500 -1000 -500 0 500 Displacement (mm) Displacement (mm) Taiwan Smart-01-SMART1-O02 (360% UHS- Failure) Critical HD-tension 150 100 -1E+10 Force (kN) Force (kN) 0 -2E+10 -3E+10 50 0 -50 -100 -4E+10 -5E+10 -4E+08 Critical WB-shear 200 1E+10 -2E+08 0 200000000 -150 -4000 -2000 0 2000 Displacement (mm) Displacement (mm) Taiwan Smart-01-SMART1-I01 (400% UHS- Failure) 1E+10 Critical HD-tension 150 100 -1E+10 Force (kN) Force (kN) 0 -2E+10 -3E+10 50 0 -50 -100 -4E+10 -5E+10 -4E+08 Critical WB-shear -150 -2E+08 0 200000000 -200 -2000 -1000 0 Displacement (mm) Displacement (mm) 83 1000 MANJIL_184327 (280% UHS- Failure) Critical HD-tension Critical WB-shear 200 1E+10 150 0 Force (kN) Force (kN) 100 -1E+10 -2E+10 0 -50 -3E+10 -100 -4E+10 -5E+10 -4E+08 50 -150 -2E+08 0 200000000 -200 -2000 -1000 0 1000 2000 Displacement (mm) Displacement (mm) Taiwan Smart-01-SMART1-I02 (320% UHS- Failure) 200 0 150 -5E+09 100 -1E+10 Force (kN) Force (kN) Critical HD-tension 5E+09 -1.5E+10 -2E+10 50 0 -50 -2.5E+10 -100 -3E+10 -150 -3.5E+10 -4E+10 -3E+08 Critical WB-shear -1E+08 100000000 -200 -2000 -1000 0 1000 Displacement (mm) Displacement (mm) CHICHI_TCU053-E (360% HUS- Failure) Critical WB-shear 200 0 150 -1E+10 100 Force (kN) Force (kN) Critical HD-tension 1E+10 -2E+10 -3E+10 50 0 -4E+10 -50 -5E+10 -100 -6E+10 -4E+08 -2E+08 0 200000000 -150 -1000 -500 0 500 Displacement (mm) Displacement (mm) 84 1000 CHICHI_TCU029-E (360% UHS- Failure) Critical HD-tension Critical WB-shear 200 1E+10 150 100 Force (kN) Force (kN) 0 -1E+10 -2E+10 -3E+10 50 0 -50 -100 -4E+10 -150 -5E+10 -4E+08 -200 -2E+08 0 -400 200000000 -200 0 200 400 Displacement (mm) Displacement (mm) CHICHI_TCU051-E (220% UHS- Failure) Critical WB-shear Critical HD-tension 150 100 50 Force (kN) Force (kN) 5E+09 0 -5E+09 -1E+10 -1.5E+10 -2E+10 -2.5E+10 -3E+10 -3.5E+10 -4E+10 -4.5E+10 -5E+10 -4E+08 0 -50 -100 -150 -200 -2E+08 0 -500 200000000 Displacement (mm) 0 500 Displacement (mm) 1000 CHICHI_TCU054-E (200% UHS- Failure) Critical WB-shear 150 0 100 -1E+10 Force (kN) Force (kN) Critical HD-tension 1E+10 -2E+10 -3E+10 0 -50 -100 -4E+10 -5E+10 -4E+08 50 -150 -2E+08 0 200000000 -200 -1000 0 1000 2000 Displacement (mm) Displacement (mm) 85 3000 Hokkaido (Japan)-Horokeshi (300% UHS- Failure) Critical HD-tension Critical WB-shear 200 300 150 200 Force (kN) Force (kN) 100 100 0 -100 50 0 -50 -100 -200 -150 -300 -200 -50 0 50 100 -400 -200 0 200 400 Displacement (mm) Displacement (mm) Maule (Chile)-Santiago Center (440% UHS- Failure) 1E+10 Critical HD-tension Critical WB-shear 200 150 0 Force (kN) Force (kN) 100 -1E+10 -2E+10 -3E+10 50 0 -50 -100 -4E+10 -5E+10 -4E+08 -150 -200 -2E+08 0 -500 200000000 Displacement (mm) 0 500 1000 Displacement (mm) 1500 Hokkaido (Japan)-Akan (440% UHS- Failure) 200 0 150 -5E+09 100 -1E+10 Force (kN) Force (kN) Critical HD-tension 5E+09 -1.5E+10 -2E+10 -2.5E+10 50 0 -50 -100 -3E+10 -150 -3.5E+10 -4E+10 -3E+08 Critical WB-shear -1E+08 100000000 Displacement (mm) 86 -200 -2000 -1000 0 1000 Displacement (mm) 2000 Maule (Chile)-Colegio Las Americas (480% UHS- Failure) Critical HD-tension Critical WB-shear 150 1E+10 100 Force (kN) Force (kN) 0 -1E+10 -2E+10 50 0 -50 -3E+10 -100 -4E+10 -150 -5E+10 -4E+08 -2E+08 0 -200 -1000 200000000 -500 0 500 1000 Displacement (mm) Displacement (mm) Maule (Chile)-Ciencias Agronomicas (260% UHS- Failure) Critical WB-shear Critical HD-tension 150 Force (kN) 100 Force (kN) 5E+09 0 -5E+09 -1E+10 -1.5E+10 -2E+10 -2.5E+10 -3E+10 -3.5E+10 -4E+10 -4.5E+10 -5E+10 -4E+08 50 0 -50 -100 -150 -2E+08 0 200000000 -200 -2000 -1000 0 1000 2000 Displacement (mm) Displacement (mm) Hokkaido (Japan)-Urakawa (240% UHS- Failure) Critical HD-tension Critical WB-shear 150 Force (kN) 100 Force (kN) 5E+09 0 -5E+09 -1E+10 -1.5E+10 -2E+10 -2.5E+10 -3E+10 -3.5E+10 -4E+10 -4.5E+10 -4E+08 50 0 -50 -100 -150 -2E+08 0 200000000 Displacement (mm) 87 -200 -1500 -1000 -500 0 Displacement (mm) 500 Maule (Chile)-La Florida (380% UHS- Failure) Critical HD-tension 0 100 -5E+09 -1E+10 Force (kN) Force (kN) Critical WB-shear 150 5E+09 -1.5E+10 -2E+10 -2.5E+10 -3E+10 50 0 -50 -100 -3.5E+10 -4E+10 -3E+08 -1E+08 -150 -1000 100000000 Displacement (mm) 0 1000 Displacement (mm) 2000 Hokkaido (Japan)-Hobetsu (400% UHS- Failure) Critical HD-tension 1E+10 Critical WB-Shear 200 150 0 100 Force (kN) Force (kN) -1E+10 -2E+10 50 0 -50 -3E+10 -100 -4E+10 -150 -5E+10 -4E+08 -2E+08 0 200000000 -200 -1000 0 1000 2000 Displacement (mm) Displacement (mm) Michoacan (Mexico) (340% UHS- Failure) Critical HD-tension 1000 0 -1000 -2000 -3000 -4000 -5000 -6000 -7000 -8000 -9000 -10000 Critical WB-shear 200 150 Force (kN) Force (kN) 100 50 0 -50 -100 -150 -100 -50 0 50 -200 -1000 100 -500 0 Displacement (mm) Displacement (mm) 88 500 Hokkaido (Japan)-Oiwake (320% UHS- Failure) Critical HD-tension Critical WB-shear 200 0 150 -1E+10 100 Force (kN) Force (kN) 1E+10 -2E+10 -3E+10 50 0 -4E+10 -50 -5E+10 -100 -6E+10 -4E+08 -2E+08 0 -150 -2000 200000000 Displacement (mm) -1000 0 Displacement (mm) 1000 Tohoku (Japan)-Koganei (300% UHS- Failure) Critical WB-shear 150 0 100 -1E+10 50 Force (kN) Force (kN) Critical HD-tension 1E+10 -2E+10 -50 -3E+10 -4E+10 -5E+10 -4E+08 0 -100 -2E+08 0 -150 -2000 200000000 Displacement (mm) -1000 0 Displacement (mm) 1000 Tohoku (Japan)-Takasaki (280% UHS- Failure) Critical WB-shear Critical HD-tension 150 Force (kN) 100 Force (kN) 5E+09 0 -5E+09 -1E+10 -1.5E+10 -2E+10 -2.5E+10 -3E+10 -3.5E+10 -4E+10 -4.5E+10 -5E+10 -4E+08 50 0 -50 -100 -150 -2E+08 0 -200 -1500 -1000 200000000 -500 0 500 Displacement (mm) Displacement (mm) 89 1000 Tohoku (Japan)-Kawagoe (320% UHS- Failure) Critical WB-shear Critical HD-tension 150 Force (kN) 100 Force (kN) 5E+09 0 -5E+09 -1E+10 -1.5E+10 -2E+10 -2.5E+10 -3E+10 -3.5E+10 -4E+10 -4.5E+10 -5E+10 -4E+08 50 0 -50 -100 -150 -2E+08 0 -200 -1000 200000000 0 1000 2000 3000 4000 Displacement (mm) Displacement (mm) Hokkaido (Japan)-Hachiohji (320% UHS- Failure) Critical WB-shear Critical HD-tension 200000 1E+13 150000 0 Force (kN) 100000 -1E+13 Force (kN) 50000 -2E+13 0 -50000 -3E+13 -100000 -4E+13 -150000 -5E+13 -4E+08 -2E+08 0 -200000 -2000 200000000 0 2000 4000 Displacement (mm) Displacement (mm) Tohoku (Japan)-Nakoso (280% UHS- Failure) Critical WB-shear 150 100 Force (kN) Force (kN) Critical HD-tension 250 200 150 100 50 0 -50 -100 -150 -200 -250 -300 50 0 -50 -100 -5 0 5 10 -150 -3000 15 -2000 -1000 0 Displacement (mm) Displacement (mm) 90 1000 Tohoku (Japan)-Nishiaidu (280% UHS- Failure) Critical WB-shear 150 0 100 Force (kN) Force (kN) Critical HD-tension 1E+10 -1E+10 -2E+10 -3E+10 0 -50 -100 -4E+10 -5E+10 -4E+08 50 -150 -2E+08 0 -200 -1000 200000000 0 1000 2000 3000 Displacement (mm) Displacement (mm) Tohoku (Japan)-Chiba (180% UHS- Failure) Critical WB-shear 1E+10 150 0 100 Force (kN) Force (kN) Critical HD-tension -1E+10 -2E+10 50 0 -3E+10 -50 -4E+10 -100 -5E+10 -4E+08 -2E+08 0 -150 -1000 200000000 0 1000 2000 Displacement (mm) Displacement (mm) Tohoku (Japan)-Koga (320% UHS- Failure) 1E+10 Critical HD-tension Critical WB-shear 150 100 -1E+10 50 Force (kN) Force (kN) 0 -2E+10 0 -3E+10 -50 -4E+10 -100 -5E+10 -4E+08 -2E+08 0 200000000 -150 -4000 -3000 -2000 -1000 Displacement (mm) Displacement (mm) 91 0 1000 Tohoku (Japan)-Gyoutoku (240% UHS- Failure) Critical WB-shear 200 0 150 Force (kN) Force (kN) Critical HD-tension 1E+10 -1E+10 -2E+10 -3E+10 50 0 -50 -4E+10 -5E+10 -4E+08 100 -100 -150 -2E+08 0 -600 200000000 Displacement (mm) -400 -200 0 200 Displacement (mm) Tohoku (Japan)-Shinozaki (300% UHS- Failure) Critical WB-shear 200 0 150 100 -1E+10 Force (kN) Force (kN) Critical HD-tension 1E+10 -2E+10 -3E+10 50 0 -50 -4E+10 -100 -5E+10 -150 -6E+10 -4E+08 -2E+08 0 -200 -1000 200000000 -500 0 500 Displacement (mm) Displacement (mm) Tohoku (Japan)-Okudo (260% UHS- Failure) Critical WB-shear 200 0 150 -5E+09 Force (kN) Force (kN) Critical HD-tension 5E+09 -1E+10 -1.5E+10 50 0 -50 -2E+10 -2.5E+10 -2E+08 100 -100 -150 -1E+08 0 100000000 -800 -600 -400 -200 0 Displacement (mm) Displacement (mm) 92 200 400 Tohoku (Japan)-Hiratsuka-st5 (280% UHS- Failure) 150 0 100 -1E+10 Force (kN) Force (kN) Critical HD-tension 1E+10 -2E+10 -3E+10 Critical WB-shear 50 0 -50 -100 -4E+10 -150 -5E+10 -4E+08 -200 -2E+08 0 -200 200000000 0 200 400 600 Displacement (mm) Displacement (mm) Tohoku (Japan)-Tatsumi (240% UHS- Failure) Critical WB-shear 200 0 150 -1E+10 100 Force (kN) Force (kN) Critical HD-tension 1E+10 -2E+10 -3E+10 50 0 -4E+10 -50 -5E+10 -100 -6E+10 -4E+08 -2E+08 0 -150 -1500 -1000 200000000 -500 0 500 1000 Displacement (mm) Displacement (mm) Tohoku (Japan)-Hiratsuka-st1 (200% UHS- Failure) 150 0 100 -1E+10 50 Force (kN) Force (kN) Critical HD-tension 1E+10 -2E+10 0 -3E+10 -50 -4E+10 -100 -5E+10 -4E+08 -2E+08 0 200000000 Displacement (mm) 93 Critical WB-shear -150 -1000 0 1000 Displacement (mm) 2000 94