Toward Damage Free Residential Houses Through UniBody Light ...

SEAOC 2014 83rd Annual Convention Proceedings

Toward Damage Free Residential Houses Through UniBody Light-Frame

Construction with Seismic Isolation

Scott Swensen, Cristian Acevedo, Ezra Jampole Eduardo Miranda and Gregory Deierlein

Stanford University Stanford, CA

Amy Hopkins and Benjamin Fell California State University Sacramento, CA

Abstract While light-frame residential house construction generally performs well with respect to life safety, experience from past earthquakes indicates that widespread damage to houses can lead to large financial losses. The damage and losses are not entirely unexpected, owing to prevailing design philosophies that advocate reduced earthquake design forces through the use of ductile seismic systems. In contrast to current approaches, a new seismic design methodology is proposed to reduce deformations and damage through an enhanced strength/stiffness limited ductility light-frame shear wall system. The increased strength and stiffness is achieved by fully engaging all available interior and exterior walls to create a unibody system. Full-scale tests and analyses of component and system assemblies are presented to demonstrate ways to economically achieve enhanced wall strength and stiffness through the use of construction adhesive to attach gypsum wallboard sheathing and improved screw connectors to attach exterior stucco cladding. Component tests utilizing the limited ductility approach have demonstrated ultimate strengths of 1.3 to 1.5 kips/ft for interior walls with two-sided gypsum boards, and as large as 2.4 kips/ft for a

wall with gypsum board (on the interior) and fiberglass mat wallboard with a stucco finish (on the exterior). For improved performance in high seismic regions, the enhanced strength superstructure can be combined with a low-cost seismic isolation system of high-density polyethylene sliding isolators on flat or dish-shaped galvanized steel plates. By taking advantage of inherent strength in the light-frame superstructure, isolator friction and dish resistance on the order of 0.2g to 0.3g limits the isolator travel under maximum considered earthquake motions to about +/-12 inches. The isolators and required displacements are economically accommodated in a wood-framed first-floor isolation platform. Supported by validated analysis models, design methods for the proposed enhanced-unibody light-frame system and seismic isolators will be demonstrated in a shake-table test of a two-story wood frame house at the NEES facility at the University of California at San Diego in the 3rd-quarter of 2014. Introduction Light-frame residential construction constitutes a major portion of the building inventory in the United States and has a large risk exposure to earthquake damage. In


reviewing residential housing damage from several recent disasters, Comerio (1997) describes that the largest contributors to losses are widespread minor to moderate damage. This is counter to common perception that tends to focus on more widely publicized instances of major damage and collapse. In citing statistics from the 1994 Northridge earthquake, Comerio reports that losses due to repair and replacement of residential buildings accounted for more than half of the $25.6 billion in total losses1 from the earthquake. She further notes that of the 500,000 damaged housing units that contribute to the losses, only about 12% were destroyed or severely damaged. The housing damage displaced over 130,000 families, primarily but not exclusively in the heavily damaged units. While somewhat reassuring that the casualties were few in residential houses damaged by the Northridge earthquake (excluding those due to collapse of soft-story light-frame apartment buildings), the financial losses and disruption due to displaced residences were staggering. Since the Northridge earthquake, considerable research and building code changes have been made to improve the seismic performance of light-frame construction. However, most of these changes have been focused on life-safety and collapse prevention under extreme (e.g., Maximum Considered Earthquake, MCE) ground motions, with comparatively less emphasis on reducing damage under more frequent moderate-to-strong ground motions. This is evident, for example, in emphasis on development of ductile light-frame shear panels, which are designed with high seismic response factors (e.g., R equal to 6.5 or 7 for light-frame walls sheathed with wood). While efficient in terms of concentrating seismic design and detailing in specific shear walls, reliance on ductile shear panels with reduced design force requirements can lead to large inelastic deformations and damage under low ground motion intensities with frequent occurrence rates. In contrast to multi-story steel and concrete framed buildings where the structural systems are constructed independently from the architectural cladding and partitions, in low-rise light-frame buildings the cladding

1 Note – Others such as Eguchi et al. 1996 have reported total losses from Northridge of $40 billion, half of which would equate to over $20 billion in losses to residential buildings (Kircher et al. 1997)

and partitions are integral with the structure. Thus, there are opportunities to engage these components as part of a ‘unibody’ system, rather than treating the structural and architectural components as independent systems. Rather than ignoring or discounting the structural participation of gypsum board partitions and stucco (or other cladding), an alternative approach is to fully integrate these into the seismic force resisting system. A related important characteristic of light-frame low-rise residential buildings is their relatively low ratios of seismic mass to structural strength and stiffness. Whereas, it is uneconomical to design large multistory buildings with lateral shear strength ratios (V/W) greater than 0.1 to 0.2, low-rise light-frame residential buildings with low seismic mass and configurations with plentiful wall areas make it feasible to design for larger lateral shear ratios. Systems designed with larger strength and stiffness will generally have smaller ductility demands, and hence, require less ductility even under extreme MCE level ground motions. Research Objectives and Scope Motivated by the overall goal to dramatically improve the seismic performance of light-frame residential construction, the specific objectives of this research are to develop and validate:

Practical and cost-effective details to increase the strength and stiffness of light-frame shear walls through improved fasteners for gypsum board, stucco and other sheathing materials to wood and cold-formed steel framing.

Analysis methods to simulate the seismic response of enhanced light-frame unibody construction for routine design and detailed nonlinear assessment.

New seismic design approaches for strength and stiffness enhanced limited-ductility systems to reduce damage under moderate-to-strong ground motions and ensure collapse safety under extreme (rare) motions.

Low-cost seismic isolators that take advantage of the unique attributes of light-frame systems, specifically their relatively low seismic mass (weight) and high inherent strength and stiffness.

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As illustrated in Figure 1, this research entails a series of inter-related studies, which will culminate in the shake table testing and analysis of a full-scale two-story house to validate the proposed design and analysis methods. Referring to the upper tier of activities in Figure 1, development of the enhanced light-frame unibody system began with testing of screw and adhesive fasteners to improve stiffness and strength of the sheathing-to-framing connection. Enhanced screw and adhesive fasteners were then employed in large-scale wall assembly tests. Full-scale room tests were then conducted to further validate the enhanced wall response, including the effect of connections and force transfers through the floor and ceiling diaphragms. Referring to the lower tier of activities in Figure 1, development of the sliding isolator began with characterization of various sliding materials to evaluate friction and wear along with pressure and velocity dependence. Testing and isolator design was guided by parametric analyses sliding isolators for residential house applications. Prototype sliding isolators were then tested under cyclic and dynamic loading on a sliding

isolator platform. Each phase of testing for the fasteners, wall and room assembly tests, and isolator tests is accompanied by companion research to develop, calibrate, and validate computational models. Detailed planning is underway for construction and shake table testing of the two-story house in 3rd quarter 2014. High Strength Limited-Ductility Design of Short Period Structural Systems As illustrated in Figure 2, the key concept of the enhanced strength and stiffness light-frame system lies in providing a lateral system that is significantly stiffer and stronger than a corresponding conventional system. Compared to conventional systems, the enhanced system will remain essentially elastic and experience small drifts under moderate to large ground motions. This stiffening and strengthening is particularly effective at limiting seismic drifts for light-frame low-rise systems with short periods. As shown in Figure 2b, for a given seismic demand, the combination of reducing the period from 0.2s to 0.15s and increasing the yield strength from

Figure 1 –Overview of Research Project Components

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0.4W to 1.0W (where W is the weight associated with the seismic mass) reduces the spectral drift (spectral displacement normalized by an assumed story height of 100 inches), from 1.5% to about 0.3%. This reduction is significant since 0.3% drift is about the threshold of damage to architectural partitions and finishes. While the enhanced unibody system strength drops off quickly from its peak, its residual strength is on the order of the strength provided by conventional ductile systems. This residual strength is important to ensure collapse safety under extreme earthquake ground motions.

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(b) Figure 2 Comparison conventional and enhanced strength and stiffness systems (a) idealized shear versus drift response, and (b) spectral displacement demands.

Taking into account room configurations, wall layouts, and seismic masses for typical residential house designs, Swensen (2014) determined that the improvements suggested in Figure 1 are feasible by engaging most of the partition walls in light-frame houses. Through surveys of design layouts for typical 2000 sq.ft. two-story houses, the ratio of floor area of the second-story to the total length of the first-floor walls was found to average about 10 ft2/ft and 20 ft2/ft in the weak and strong framing directions (note - due to garage door and other openings, combined with floor framing layouts, most house plans had distinct weak and strong directions). Assuming an SDS of 1g (high seismic region with R=1) and unit average dead weights of 50 psf for the second floor and 25 psf for the roof (including floor, wall and roof finishes, etc.), the required first-story wall strengths to resist 1g of inertial force range from 0.8 kip/ft to 1.5 kip/ft. The required wall stiffness to limit drifts to about 0.2% (threshold for wall damage) range from 3.5 kip/in/ft to 6.9 kip/in/ft. As described later, these values can be achieved through the proposed unibody design approach that engages the participation of structurally enhanced gypsum board and other partitions and wall finishes (e.g., exterior stucco) to provide lateral strength and stiffness. Enhanced Fastener and Wall Tests The response of light-frame shear walls is generally controlled by deformations of mechanical fasteners between the structural sheathing and the wood or cold-formed steel framing. Previous tests of gypsum wallboard panels have shown that the strengths are controlled by bearing of the nail or screw fasteners against the gypsum board material. Therefore, our approach to improve the strength and stiffness of gypsum wallboard sheathing focuses on improved fasteners. Our initial efforts to improve the strength and stiffness of gypsum wallboard connections involved testing of screw connectors with increased shank areas with larger bearing areas. While these resulted in some improvement in strength and stiffness, the improvements were insufficient to achieve the required performance. We then investigated the use of construction adhesives2

2 After exploring alternative products, we used an interior construction adhesive, LNP-701 by Liquid Nails.

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that resulted in remarkable improvements (Swensen 2014). Based on favorable results from fastener tests, we then tested reduced-scale (4’x4’) and full-scale (8’x8’ and 8’x16’) wall panels constructed with gypsum wallboards attached with adhesive and conventional drywall screws. Shown in Figure 3 is an example of the full-scale wall tests conducted at California State University Sacramento (Hopkins 2013).

Figure 3 – Example of light-frame wall test specimen (a) framing layout, (b) observed damage evolution

Table 1 – Summary of wall panel strength and stiffness (Hopkins 2013) Description Fu Ko

k/ft k/in/ft GB, screws [conventional, baseline] 0.6 2.7 GB, screws/adhesive 1.2 7.2 GB, screws/adhesive, T-return 1.4 11.1 GB + stucco, screws/adhesive, return 2.0 9.8 Plywood, screws/adhesive 1.9 6.9 CFS-GB, screws 0.5 2.5 CFS-GB, screws/adhesive 1.0 4.6 GB – 5/8” Type X Gypsum Board CFS – Cold Formed Steel Framing A total of about 62 fastener tests and 27 wall panel tests were conducted on specimens with either wood or cold-formed steel framing with (1) gypsum board sheathing, (2) gypsum board and stucco, and (3) plywood. Most tests were with two-sided 5/8” thick Type X gypsum wallboard, which was chosen over thinner regular types of gypsum board for its higher stiffness, strength and toughness. Ultimate strength and stiffness (secant stiffness to 0.1% drift) of several representative wall panel tests are reported in Table 1. Compared in Figure 4 are applied load versus drift data for four full-scale wall tests with wood framing and either two-sided gypsum wallboard or gypsum board with stucco (also reported in Table 1). Results in Figure 4a are for a wall constructed with conventional coarse threaded screws, spaced at 7 inches around the perimeter and 12 inches in the field. Results in Figure 4b are for a similar wall, except that construction adhesive is used along with the screws to fasten the gypsum board sheathing. Figure 4c is for a wall with the same adhesive and screw connections as Figure 4b, but with the addition of T-shaped returns at each end of the wall (representing orthogonal walls). Figure 4d shows data of an exterior wall configuration, with gypsum board on the interior and 7/8 inch thick stucco over DensGlass3 sheathing on the exterior. DensGlass was attached to the wood framing with drywall screws and adhesive, and the stucco finish was engaged using ¼ inch diameter lag bolts and lath reinforcement, installed at 4 inches on center around the wall perimeter.

3 DensGlass is the product name for a moisture resistant fiberglass mat gypsum board for external building walls.


Comparing Figures 4a and 4b (rows 1-2 of Table 1), the adhesive increases the initial stiffness by about 2.6 times, from 2.7 kip/in/ft to 7.2 kip/in/ft, and strength by about 2.0 times, from 0.6 kip/ft to 1.2 kip/ft. Note that after the adhesive failed at about 0.5% drift, the wall maintains higher strength than the wall with screws alone. The T-shaped wall end returns (Figure 4c, row 3 of Table 1) increase the initial strength by about 15% and stiffness by about 50%. In addition, the returns dramatically improve the residual strength to more than 70% of the peak strength. Finally, as shown in Figure 4d, the combination of gypsum board (interior) with DensGlass and stucco (exterior) increases the strength and stiffness to about 3.5 times the conventional wall

(Figure 4a). As reported in Table 1, the gypsum board plus stucco combination yielded about the same strength and stiffness as a wall with 15/32” Structural I plywood sheathing that is attached with 10d nails and adhesive. The adhesive attachment of gypsum wallboard to cold-formed steel framing (rows 6-7 of Table 1) showed somewhat less improvement than wood framing (rows 1-2 of Table 1), primarily because the adhesive did not bond as completely to cold-formed steel as to the wood. Whereas adhesive bond to wood framing was strong enough to force failure into the gypsum wallboard, incomplete adhesion to cold-formed steel framing resulted in failure at the adhesive to steel interface.

Figure 4 – Hysteretic response of wood-frame enhanced wall with (a) 2-sided GB with conventional screw fasteners, (b) 2-sided GB with screw fasteners plus construction adhesive, (c) 2-sided GB with screw fasteners, construction adhesive, and T-shaped end returns, (d) GB interior, Densglass and Stucco exterior with screw fasteners, construction adhesive, and L-shaped end returns

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The initiation and progression of damage in the walls was similar regardless of the type of fastener. As shown in Figure 3b, generally, damage initiated at about 0.2% to 0.3% drift with diagonal cracks around openings, cracks along joint lines, and screw pops. Under increasing drifts, cracking and screw pops would tend to progress until the screw damage was severe enough for the wallboard to become detached at around 0.8 to 1.0% drift. While the damage initiation and pattern was relatively similar for all walls, the enhanced walls with adhesive should perform better as they will sustain much smaller drifts for the same ground motion intensity. Referring back to the previous discussion of system design requirements to minimize damage under design level ground motions, all of the enhanced wood-framed

walls have strengths and stiffnesses that are within the range of the aforementioned targets of 1.5 k/ft (strength) and 6.9 k/in/ft (stiffness). The high stiffness is particularly important to limit damaging drifts under design-level ground motions. Light-Frame Room Assembly Tests To further examine and validate the enhanced strength and stiffness of the enhanced wall systems in a three- dimensional unibody assembly, four large-scale room assemblies where tested. The rooms measured about 16 ft by 7 ft in plan and 9 ft tall, and were tested in the NEES lab at the University of California at Berkeley. In addition to validating the fastener and other enhancements evident from the planar wall tests, the

Figure 5 – Results of Large-Scale Room Assembly Tests


three-dimensional room assemblies included more realistic boundary conditions and force transfer through the floor diaphragms. Shown in Figure 5 are the overall load versus drift responses for two of the tests, where response curves in Figure 5a are for a room with two-sided 5/8 inch Type X gypsum board sheathing and curves in Figure 5b are for a combination of gypsum board on the interior with 5/8 inch DensGlass and 7/8 inch stucco on the exterior. With normalized strength and stiffness response similar to the wall tests, the room assemblies further validated the high stiffness and strength achieved by the design enhancements. Most remarkably, the tests demonstrated how the walls can maintain about 1/3 to 1/2 of their peak strengths out to drift ratios in excess of 10%. The tests further confirmed that the diaphragm force transfer details worked as expected with minimal deformation or other influence on the global response. Computational Modeling of Unibody System In conjunction with the wall testing, three types of computational models were developed and calibrated to simulate the response of wall panels and complete building systems. These included two types of models to simulate the nonlinear response and an equivalent elastic model for use in design. A detailed continuum finite element model was developed to simulate the nonlinear response based on fundamental properties of the wall sheathing, framing, connectors, and holddowns. As illustrated in Figure 6, the model combines flexural beam-column elements to simulate wood or steel framing members, plane stress elements to simulate the sheathing, nonlinear elastic-gap springs for the hold-downs, and nonlinear spring elements to simulate the sheathing-to-framing connection. The framing, sheathing, and hold-down members are typically modeled as elastic, since most of the nonlinearity occurs in the sheathing-to-framing connectors. We calibrated the nonlinear connector models from screw and adhesive fastener tests (Figure 6b), where the connector elements simulate a tributary length of adhesive connection, typically taken as equal to the screw spacing. When assembled into a complete model of the wall panels using ABAQUS (Figure 6c), the combined model compared well with the wall test data (Figure 6d). Since it is based on fundamental wall

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Figure 6 – Detailed finite element model (a) idealization, (b) calibration of screw and adhesive connector springs, (c) validation against planar wall panel test.

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Figure 7 – Calibration of phenomenological diagonal strut model for gypsum board wall with screw and adhesive fastener components, the detailed model has the resolution to examine wall geometries and boundary conditions beyond those investigated in the wall testing program. The second type of model is an equivalent strut, shown in Figure 7, which is calibrated to wall tests and detailed models to simulate the overall wall response. While its reliability depends on the data against which it is calibrated, the model enables simulation of the nonlinear dynamic response of entire buildings under multiple ground motions. While the strut model is calibrated to simulate the nonlinear wall panel shear response, destabilizing P- effects and flexibility in the vertical hold-down and foundation can be readily included in the model. The wall panel strut has been implemented and calibrated to parameters of the nonlinear IMK (Ibarra et al., 2005) and CUREE (Folz and Filiatrault, 2001) models in OpenSees. Shown in Figure 7 is an example of the model calibration out to large drift levels.

Finally, the third type of model is one comprised of elastic shell finite elements, which is readily implemented in commercial structural analysis software such as ETABS. Referring ahead to Figure 14c, this model is envisioned as providing a practical way to analyze internal forces and deformations in house configurations with complicated wall configurations, considering wall openings and in-plane and out-of-plane diaphragm flexibility. Guidelines for defining equivalent elastic properties, based on deformations and secant stiffnesses at design-level ground motions have been developed (see Swensen, 2014). High-Friction Sliding Seismic Isolation Beyond the improvements in performance provided by enhancing the strength and stiffness of conventional house construction, further performance improvements can be realized by combining the enhanced unibody system with a low-cost seismic isolation system. Shown in Figure 8 is a schematic layout of an isolated house, where the isolators are located on foundations beneath the first-floor framing. This system will require design enhancements to the first-floor isolation diaphragm, but otherwise, the design and construction is similar to conventional light-framed houses. In spite of the attractiveness of seismic isolation, conventional approaches to isolation have not been widely adopted for residential house construction due to a number of reasons, primarily high cost and practical difficulties of implementation. To help overcome these obstacles, the proposed isolation design is devised to take advantage of the unique aspects of light-frame residential construction. The relatively high inherent strength to weight ratio of the enhanced light-frame construction suggests that the isolation system can develop relatively large base shears (e.g, V/W ratios of 0.2 to 0.3g) while still limiting damage to the house; and the relatively light weight of houses opens up opportunities to use isolator materials that would not be candidates for existing isolator designs for heavy concrete, masonry and steel structures. After considering alternative options, the proposed isolation system employs sliding isolators that utilize readily available materials with moderately high friction coefficients, on the order of 0.15 to 0.25. High- friction

Wall Shear Strut

Floor Diaphragm

Boundary Column Strut


Figure 8 – Schematic configuration of seismically isolated house isolators can be effectively paired with the inherent strength of the enhanced unibody superstructure to reduce isolator displacement demands. In addition, the light weight of the house results in relatively low bearing stresses on the sliding isolators, on the order of 500 psi. Thus, many of the challenges associated with designing isolators with high bearing stresses are avoided. Eight combinations of sliding isolator and plate materials were investigated to evaluate their dynamic friction resistance for the sliding isolator (Jampole et al, 2014). This included tests of four sliding materials: unfilled (UF) and glass-filled (GF) Polytetrafluoroethylene4 (PTFE), high-density polyethylene (HDPE), and Kevlar-filled Nylon (KFN). Two plate materials were tested: #8 mirror finish stainless steel, which is often used in low-friction bearings due to its smooth surface and corrosion resistance, and a less expensive galvanized steel. Tests were conducted at bearing pressures of 570 psi, which is representative of conditions expected for the house

4 Teflon by DuPont Company is a widely known brand name of PTFE.

isolators but much lower that some typical applications such as for expansion slider bearings used in bridges. Results of dynamic sliding tests are shown in Figure 9. In general, all of the materials exhibited a friction coefficient that varied with velocity, which tended to saturate at velocities of about 6 in/s. The friction coefficients tend to be higher than reported previously in the literature for some of these materials, which is likely due to the difference in bearing stresses (where lower bearing stresses are known to give higher friction coefficients). Interestingly, the friction coefficients were similar for the mirror finish stainless steel and the galvanized plate, and the coefficient was sometimes even lower for the galvanized plate. Shown in Figure 10 are results of a parametric study of peak and residual sliding isolator displacements based on a prototype house subjected to ground motions scaled to MCE levels for a high seismic site in Los Angeles (SMS equal to 2.1g). These analyses were conducted using a sliding friction isolator model that incorporates the post-slip stiffness provided by a dish-shaped sliding surface. The curved dish provides a restoring (post-slip) stiffness of W/r, where W is the weight of the supported structure and r is the radius. Thus, the dish adds to the post-slip resistance of the isolator and provides a restoring effect. The values in the figure are median peak (Figure 10a) and residual (Figure 10b) isolator displacements, reported as a function of sliding friction coefficient and dish radius, where the large dish radius of 500 inches essentially represents a flat isolator.

Figure 9 – Measured friction coefficient response of sliding isolator materials tested through harmonic loading

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(a) (b) Figure 10 – Calculated peak and residual sliding isolator displacement for MCE level ground motions in high seismic region As evident from Figure 10a, the friction coefficient has a significant influence on the peak sliding displacement, ranging from 24 inches for a low friction coefficient of 0.05 down to about 6 inches for a coefficient of 0.25. Of course, the reduced isolator displacements come with increased base shear demands on the superstructure, but the reduction in sliding displacements can dramatically simplify the isolator system design (e.g., by allowing smaller cantilever overhangs in the isolator diaphragm design, Figure 8). The sliding friction of 0.15 (which is in the mid-range of the tested materials) has a median peak displacement of 12 inches for a flat isolator and (by interpolation) about 10 inches for an 80 inch radius isolator. At 10 inches peak displacement, the dish isolator exerts an effective sliding resistance base shear ratio of 0.275 (0.15 + 10/80) versus 0.15 for the flat isolator. Perhaps more significant than the modest (20%) reduction in peak displacement is the reduction in residual displacement. As shown in Figure 10b, the 80 inch dish radius (with friction equal to 0.15) reduces the residual displacement from about 3 inches, for the flat isolator, to less than 1 inch. The re-centering capability of the 80-inch dish isolator is illustrated further in the comparison of response traces under an earthquake ground motion (KJM-000) in Figure 11. Shown in Figures 12 and 13 are the schematic design and a lab test of the proposed design for a sliding dish isolator beneath a wood-framed isolator plane. Similar in concept to friction pendulum systems (Zayas, et al 1989), the galvanized steel-HDPE isolator has a dish radius of 80 inches and range of +/- 15 inches.

Figure 11 – Sliding isolator displacement for flat and dish isolators under earthquake ground motion

Figure 12 – Sliding Dish Isolator Detail

Figure 13 – Sliding Dish Isolator Test

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Cyclic harmonic and dynamic testing of the dish isolator and comparative flat isolators were conducted on an ‘isolator sleigh’ setup (Figure 1) at the NEES laboratory at the University of California, Berkeley. As shown in Figure 13 the test isolator assemblies included a load cell to measure vertical and shear loads in the isolator (see Jampole et al., 2014 for further details on these tests). System Response and Validation Work is currently underway to implement and validate the isolated unibody system design approach in a shake

table test of a two-story house at the University of California, San Diego (UCSD) NEES laboratory. The floor plans of the house are shown in Figures 14a-b and measure 36 x 22 ft in plan, comprising about 1580 sq.ft. on two floors. The house specimen is configured to represent a typical four bedroom house with an enclosed garage. The garage and room areas are somewhat reduced from a full size house configuration to fit on the UCSD shake table, but overall, the specimen framing and design details are constructed at full scale. The experimental program includes three phases of

Figure 14 - Full-scale wood frame house test (a) first floor plan, (b) second floor plan, (c) ETABS elastic finite element model, (d) OpenSees nonlinear strut model

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testing, including two phases with sliding isolators (dish and flat isolators) and a final fixed base condition. The isolation diaphragm will be configured similar to the illustration in Figure 8, where twelve isolators will support the house with isolator-to-isolator framing spans up to about 12 ft and cantilever overhangs of about 2 ft. The walls and wall openings in the first story (Figure 14a) provide a design shear strength of V/W equal to about 1g. Wall design strengths were based on design strengths of 1.2 kip/ft for the two-sided gypsum board walls and 1.4 kip/ft for the exterior gypsum-stucco walls. These design strengths were based on median wall resistances measured at 0.2% drift, which corresponds to the initiation of damage. Wall hold downs are designed to have mean ultimate strengths equal to the calculated design strengths of the enhanced wall design strengths (where the ultimate strengths of the hold down devices are assumed to be about three times their specified working design strength values). The unibody house system is designed using the elastic ETABS model, shown in Figure 14c, using equivalent elastic properties for the shell finite elements. The nonlinear response is evaluated using an OpenSees model with equivalent struts (see Figure 14d), as described previously (Figure 7). The house has a calculated fundamental period of vibration of about 0.2s in each direction. To evaluate the nonlinear response, several sets of ground motions (20 record pairs in each set) were developed using the Conditional Spectrum approach to represent ground shaking at increasing earthquake intensities (Baker 2010). The design level ground motions are scaled to a short period Sa equal to 1g, and the MCE level motions are scaled to 1.5g. Shown in Figure 15a is a plot of the probability of the house exceeding the damage initiation drift of 0.2% under increasing ground motions. At the design ground motion there is about a 40% chance of exceeding the damage threshold, whereas at the MCE ground motion (at 1.5 times the intensity of the design level) there is a high probability of exceeding the damage threshold. For comparison, fragility plots of a comparable house designed according to current design standards (e.g., ASCE 7-10 with R equal to 6.5) are shown. The more fragile of the two plots (furthest to the left) corresponds to a conventional house that is analyzed with the minimum structural shear walls. The second conventional plot is based on analyses that consider the

stiffening and strengthening effect of the wall finishes (i.e., gypsum board sheathing and stucco enclosure with standard attachment details). At design level shaking (1g), the conventional system including finishes has about double the risk of exceeding the damage threshold. Figure 15b shows comparable plots for collapse risk. While the fragility curves are all within expectations for modern seismically designed structures (i.e., less than 10% probability of collapse under MCE level ground motions), the enhanced unibody system has about half the risk of collapse as the conventional systems.

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Figure 15 – Light-frame house with conventional and enhanced construction (a) probability of exceeding 0.2% story drift, and (b) probability of collapse.


The addition of sliding isolators to the house is expected to result in little or no damage to the drift sensitive structural and architectural components under design and even MCE level ground shaking. The expected MCE level isolator displacements of 12 inches are comfortably accommodated within the 15 inch isolators; and the shear force transferred through the isolator assembly is much less than the shear capacity of the first-story walls. The performance expectations shown in Figure 15 are probabilistic estimates of the fixed-base house under series multiple ground motion pairs. Testing on the uni-axial NEES-UCSD shake table will be conducted for a select number of individual ground motions parallel to the longitudinal axis of the building. Initially, the house will be tested in the isolated condition, first on dish isolators and then on flat isolators, to help validate predictions for the isolator displacements and house response. The isolated house will be subjected to motions that are scaled up progressively to MCE level motions. Following the isolated house tests, the first-floor diaphragm will be locked to the table to evaluate the fixed base condition for ground motions. Drifts of about 0.2% to 0.4% are expected under design and MCE level ground motions, respectively. Motions will then be scale up to twice MCE level intensity, where drifts in excess of 10% (i.e., near collapse) are expected. Summary and Future Work This research aims to develop and validate design approaches that make more effective use of the inherent robustness of residential light-frame construction. Fundamentally, the proposed approach relies on improved fasteners and other details to more completely engage the inherent strength and stiffness of exterior and partition walls and finish materials to create a unibody earthquake resistant system. This is in contrast to prevailing approaches that concentrate seismic resistance in ‘ductile’ seismic systems, which can lead to large deflections and damage. With the ultimate aim towards more damage resistance and resilience, pairing of the new unibody system with low-cost seismic isolation offers tremendous potential for dramatically improved performance. Beyond the work reported in this paper, research is ongoing to address some of the remaining challenges:

Development of more complete and transparent design guidelines and criteria for enhanced unibody systems that will permit implementation through alternative construction technologies, such as SIPs and other systems that can provide high strength with limited ductility.

Comprehensive validation of enhanced strength/stiffness unibody approach, including a FEMA P695 assessment to confirm collapse safety.

Improved models to characterize sliding displacements of high-friction isolators (e.g., friction coefficient ~0.2 to ~0.3), considering their unique stick-slip properties over more conventional low-friction (e.g., friction coefficient ~0.05) isolators.

Continued development of reliable low-cost sliding isolators, including research to characterize influence of aging/deterioration, construction tolerances, foundation settlement, and other factors on isolator performance.

Evaluation of foundation uplift of light-frame construction and its relationship to seismic performance and its influence on design and detailing of seismic holddowns.

Acknowledgments This material is based upon work supported by the National Science Foundation (NSF) under CMMI-NEES Grant No. 1135029. Additional financial support was provided by NSF through student fellowships, an EERI Graduate Student Fellowship, and the Blume Earthquake Engineering Center at Stanford University. Donation of materials and supplies is greatly appreciated from Simpson Strong-Tie, PPG Industries, Alabama Metal Industries Corporation (AMICO), and Albion Engineering. Any opinions, findings, and conclusions or recommendations expressed in this material are those of the authors and do not necessarily reflect the views of the National Science Foundation or other sponsors. The authors would also acknowledge the technical input and advice of a talented group of professional consultants including: K. Cobeen of WJE; J. Bomba, A. Roufegarinejad, R. Vignos and M. Walters of Forell/Elsesser Inc.; J. Osteraas of Exponent; D. Mar of Tipping-Mar; G. Luth of GPL&A; S. Pryor of Simpson

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Strong-Tie; and R. Serrette of Santa Clara University. The authors further acknowledge the assistance of several undergraduate research students, K. deLaveaga, C. Fong, M. Hardy, R. Khan, C. McEniry, R. McNerney, and G. Oren, as well as the support of the staff at NEES@Berkeley and NEES@UCSD. Finally, the authors are indebted to Ben Schmid, who inspired, guided and supported the authors to pursue research to improve the seismic performance of residential house construction. References

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