## About

The Civil and Environmental Engineering Department at UCLA (CEE-UCLA) is in the Henry Samueli School of Engineering and Applied Science and was formed in 1982. Within CEE-UCLA, teaching and research activities occur within a number of inter-disciplinary research units and centers involving world-renowned faculty, undergraduate and graduate students, research staff, and post-doctoral and visiting scholars. You are invited to peruse this site to learn more about these research activities.

Jonathan P. Stewart, Professor and Vice Chair

University of California, Los
Angeles

Civil and Environmental Engineering Department

5731 Boelter Hall

Los
Angeles, CA 90095-1593

## UCLA Civil and Environmental Engineering

## Earthquake Engineering (26)

### Assessment of soil-structure interaction modeling strategies for response history analysis of buildings

A complete model of a soil-foundation-structure system for use in response history analysis requires modification of input motions relative to those in the free-field to account for kinematic interaction effects, foundation springs and dashpots to represent foundation-soil impedance, and a structural model. The recently completed ATC-83 project developed consistent guidelines for evaluation of kinematic interaction effects and foundation impedance for realistic conditions. We implement those procedures in seismic response history analyses for two instrumented buildings in California, one a 13-story concrete-moment frame building with two levels of basement and the other a 10-story concrete shear wall core building without embedment. We develop three-dimensional baseline models (MB) of the building and foundation systems (including SSI components) that are calibrated to reproduce observed responses from recorded earthquakes. SSI components considered in the MB model include horizontal and vertical springs and dashpots that represent the horizontal translation and rotational impedance, kinematic ground motion variations from embedment and base slab averaging, and ground motion variations over the embedment depth of basements. We then remove selected components of the MB models one at a time to evaluate their impact on engineering demand parameters (EDPs) such as inter-story drifts, story shear distributions, and floor accelerations. We find that a “bathtub” model that retains all features of the MB approach except for depth-variable motions provides for generally good above-ground superstructure responses, but biased demand assessments in subterranean levels. Other common approaches using a fixed-based representation can produce poor results.

### Site response in NEHRP Provisions and NGA models

Site factors are used to modify ground motions from a reference rock site condition to reflect the influence of geologic conditions at the site of interest. Site factors typically have a small-strain (linear) site amplification that captures impedance and resonance effects coupled with nonlinear components. Site factors in current NEHRP Provisions are empirically-derived at relatively small ground motion levels and feature simulation-based nonlinearity. We show that NEHRP factors have discrepancies with respect to the site terms in the Next Generation Attenuation (NGA) ground motion prediction equations, both in the linear site amplification (especially for Classes B, C, D, and E) and the degree of nonlinearity (Classes C and D). The misfits are towards larger linear site factors and stronger nonlinearity in theNEHRP factors. The differences in linear site factors result largely from theirnormalization to a reference average shear wave velocity in the upper 30 m of about 1050 m/s, whereas the reference velocity for current application is 760 m/s. We show that the levels of nonlinearity in the NEHRP factors are generally stronger than recent simulation-based models as well as empirically-based models.

### Site effects in parametric ground motion models for the GEM-PEER Global GMPEs Project

We review site parameters used in ground motion prediction equations (GMPEs) for various tectonic regimes and describe procedures for estimation of site parameters in the absence of site-specific data. Most modern GMPEs take as the principal site parameter the average shear wave velocity in the upper 30 m of the site (V_{s30}) either directly or as the basis for site classification into categories. Three GMPEs developed for active regions also use basin depth parameters. We review estimation procedures for V_{s30} that utilize surface geology, terrain-based site categories, ground slope, or combinations of these. We analyze the relative efficacy of those procedures using a profile data set from California assembled in a recent NGA project. The results indicate that no single procedure is most effective and that prediction dispersion is lower for young sediments than for stiff soils or rock.

## Geotechnical Engineering (4)

### Estimating Undrained Strength of Clays from Direct Shear Testing at Fast Displacement Rates

When the direct shear test is performed in accordance with ASTM guidelines, the measured shear stresses at failure estimate drained strength parameters. We investigate the possibility of estimating undrained strength using direct shear testing at variable shear displacement rates on specimens composed of various combinations of kaolinite and bentonite. Even at fast displacement rates, constant volume conditions are not achieved in the direct shear device because of changes in specimen height that are large relative to allowable ASTM thresholds for constant volume simple shear testing. However, undrained strengths established by constant volume simple shear testing at slow strain rates are well approximated by direct shear tests conducted at fast shear displacement rates (time to failure < t_{50}/8, where t_{50}=time to 50% consolidation in a conventional oedometer test). Because of the simplicity of direct shear testing, such estimates of undrained strength may be useful in engineering practice when access to a simple shear device is limited. Nevertheless, fast direct shear tests have shortcomings, including lack of control of rate effects, and constant volume testing is recommended for critical projects.

### Laboratory investigation of the pre- and post-cyclic volume change properties of Sherman Island peat

We investigate through laboratory testing the volume change characteristics of peaty organic soil from Sherman Island, California under static conditions (consolidation, secondary compression) and post-cyclic conditions. Incremental consolidation tests indicate the material to be highly compressible (C_{c} = 3.9, C_{r} = 0.4) and prone to substantial ageing from secondary compression (C_{a}/C_{c} = 0.05 following virgin compression). Strain-controlled cyclic triaxial testing of the peat finds the generation of cyclic pore pressures for cyclic shear strain levels beyond approximately 0.5-1.0%, with the largest residual pore pressure ratios r_{ur} (cyclic residual pore pressure normalized by pre-cyclic consolidation stress) being approximately 0.2-0.4. Post cyclic volume change occurs from pore pressure dissipation and secondary compression. The level of post-cyclic secondary compression increases with r_{ur}. Many of these phenomena have not been documented previously and suggest the potential for seismic freeboard loss in levees due to mechanisms other than shear failure.

### Full Scale Cyclic Testing of Foundation Support Systems for Highway Bridges. Part II: Abutment Backwalls

This research involved analysis and field testing of numerous foundation support components for highway bridges. Two classes of components were tested - cast-in-drilled-hole (CIDH) reinforced concrete piles (drilled shafts) and an abutment backwall. The emphasis of this document (Part II of the full report) is abutment backwall elements.

The backwall test specimen was backfilled to a height of 5.5 up from the base of the wall with well-compacted silty sand backfill material (SE 30). The wall is displaced perpendicular to its longitudinal axis. Wing walls are constructed with low-friction interfaces to simulate 2D conditions. The backfill extends below the base of the wall to ensure that the failure surface occurs entirely within the sand backfill soil, which was confirmed following testing. The specimen was constructed and tested under boundary conditions in which the wall was displaced laterally into the backfill and not allowed to displace vertically.

A maximum passive capacity of 497 kips was attained at a wall displacement of about 2.0 in, which corresponds to a passive earth pressure coefficient Kp of 16.3. Strain softening occurs following the peak resistance, and a residual resistance of approximately 460 kips is achieved for displacements > 3.0 inch. The equivalent residual passive earth pressure coefficient is Kp = 15.1 and the equivalent uniform passive pressure at residual is approximately 5.1 ksf, which nearly matches the value in the 2004 Seismic Design Criteria of 5.0 ksf. The average abutment stiffness K50 was defined as a secant stiffness through the origin and the point of 50% of the ultimate passive force. For an abutment wall with a backfill height H of 5.5 ft, this stiffness was found to be K50 = 50 kip/in per foot of wall. The load-deflection behavior of the wall-backfill system is reasonably well described by a hyperbolic curve.

The passive pressure resultant is under predicted using classical Rankine or Coulomb earth pressure theories. Good estimates of capacity are obtained using the log-spiral formulation and the method-of-slices. The method-of-slices approach is implemented with a log-spiral hyperbolic method of evaluating backbone curves that provides a good match to the data.