Arias Intensity Assessment of Liquefaction Test Sites on the East Side of San Francisco Bay Affected by the Loma Prieta, California, Earthquake of 17 October 1989
ROBERT E. KAYEN1 and JAMES K. MITCHELL 2
1Research Civil Engineer, US. Geological Survey, 345 Middlefield Road, Menlo Park, CA 94025 U.S.A. firstname.lastname@example.org
Abstract. Uncompacted artificial-fill deposits on the east side of San Francisco Bay suffered severe levels of soil liquefaction during the Loma Prieta earthquake of 17 October 1989. Damaged areas included maritime-port facilities, office buildings, and shoreline transportation arteries, ranging from 65 to 85 krn from the north end of the Loma Prieta rupture zone. Typical of all these sites, which represent occurrences of liquefaction-induced damage farthest from the rupture zone, are low cone penetration test and Standard Penetration Test resistances in zones of cohesionless silty and sandy hydraulic fill, and underlying soft cohesive Holocene and Pleistocene sediment that strongly amplified ground motions. Postearthquake investigations at five study sites using standard penetration tests and cone penetration tests provide a basis for evaluation of the Arias intensity-based methodology for assessment of liquefaction susceptibility.
Key words: acceleration, arias intensity, earthquake, geotechnical, liquefaction, field test, penetration resistance, seismometer.
Peak horizontal accelerations on sites underlain by rock and stiff alluvium in the east bay generally ranged from 0.08 to 0.12 g, but amplification due to the presence of soft and deep cohesive soil deposits underlying artificial fills produced peak accelerations from 0.11 to 0.29 g at strong-motion recording estimated from strong motion records obtained near the east bay shoreline. On the basis of available strong motion data (Shakal et al., 1989; Kayen et al., 1992), the peak horizontal accelerations on bayshore fills in the vicinity of Oakland International Airport, and Bay Farm Island were apparently about 0.27 g, and at the San Franc isco-Oakland Bay Bridge toll plaza and Port of Oakland's container terminal at Seventh Street were probably about 0.28 to 0.29 g (Figure 1). Arias intensities, integrated from the strong motion records ranged from 0.8 m/s at Oakland International Airport, Bay Farm Island, and Alameda Naval Air Station to 1.71 m/s at the Port of Oakland's container terminal at Seventh Street and the Bay Bridge toll plaza (Kayen and Michell, 1997).
At the Port of Richmond, the most northerly study site, no nearby strong-motion recordings were obtained for sites with similar soil conditions (Figure 1). Approximately 2 km to the north, however, a peak horizontal acceleration of 0. 13 g on the south component was recorded at Richmond's City Hall. We used this record, the CSMIP strong motion recording Station No. 58505, to model the site response at the Port of Richmond study site where soft, deep cohesive-soil deposits underly artificial fill, by means of one-dimensional equivalent-linear seismic site-response analyses, using the computer program SHAKE90 (a modified version of the computer program SHAKE; Schnabel et al., 1972). These results indicate that peak horizontal accelerations at the Port of Richmond were probably about 0.18 g. Synthetic seismograms from the SHAKE90 analysis were integrated and we estimate that the Arias intensity at the Port of Richmond site to be approximately 0.6-0.7 m/s.
This paper describes soil liquefaction, ground deformations and associated liquefaction-damages in artificial fill deposits on the east-bay shoreline from Oakland International Airport to the Port of Richmond, and presents the results of our post earthquake studies at five study sites near the east bay shoreline: (1) Oakland International Airport (Figures 1 and 2), (2) Bay Farm Island and Harbor Bay Plaza (Figures 1 and 2), (3) the Port of Oakland's Marine Container terminal at Seventh Street (Figures 1 and 3), (4) the San Francisco-Oakland Bay Bridge Toll Plaza (Figures 1 and 3), and (5) the Port of Richmond (Figures 1 and 4). The collapse of the Interstate Highway 880 Cypress Street Viaduct occurred near sites 3 and 4, along a stretch of viaduct that is founded on Holocene Bay Mud. Liquefaction was not observed at this site and so we do not discuss damage to the Cypress Freeway here. Detailed discussion of the geotechnical aspects of the Cypress Freeway collapse can be found in Seed et al. (1990).
We found that the overburden-stress-correction factor, Cn , can be resonably expressed as
where N1, is the overburden- stress-corrected standard penetration resistance, vo is the effective overburden-stress, and ref is a reference stress of 1 atm. Likewise, to evaluate qc values measured during CPT tests we adjust the measured qc, values as follows:
We found that Cq can be reasonably expressed as
Liquefiable layers at the Toll Plaza, Port of Oakland, Bay Farm Island, and Oakland International Airport study sites can be characterized as clean sand with a mean grain diameter (D50) of at least 0.25 mm and a fine-particle content of less than 5%. The liquefiable layer at the Port of Richmond study site has a D50 value of approximately 0.06 to 0.07 mm and a fine-particle content of >35%. Silty sands, like those at the Port of Richmond site, with the same liquefaction resistance to that of a corresponding reference 'clean' sand are typically observed to have distinctly lower measured penetration resistance values. Seed et al. (1984) presented boundary curves that effectively convert penetration resistance values of silty sands to that for clean sands with no fine particles, and defined a SPT-fines content correction factor (Seed and de Alba, 1986):
The fines-correction factor N1 is a function of the fines content, FC, in percent. Following the guidelines presented in the consensus statement from the NCEER Workshop on Evaluation of Liquefaction Resistance, 4-5 January 1996, Salt Lake City, Utah, we used the following equations for correcting SPT-values of low-plasticity silty sands to equivalent values for clean sand
We applied these correction factors to the Port of Richmond site SPT data. Detailed description of the borehole logs from all of the liquefaction test sites, including those discussed herein, can be found in Kayen et al. (1992, 1997), Kayen (1993) and Mitchell et al. (1994).
Extensive soil liquefaction occurred in the western section of the airport fill, damaging the northwestern 900 m of the 3000 m-long main runway. The adjacent taxiway pavement was also heavily damaged. Most of the runway damage was repaired within 4 weeks, and as a result, the airport was able to resume essentially full operations with a shortened operational runway 2700 m long on 20 November 20 1989.
In addition to liquefaction, settlement, and lateral spreading of the main runway fill, settlement and lateral spreading also occurred in several places in the surrounding perimeter dikes at the west end of the runway fill. The maximum observed levee settlement of the perimeter dikes was approximately 0.5 to 0.7 m, and lateral deformations were similar in magnitude. Liquefaction was also observed near the main terminal buildings.
The logs of three boreholes drilled at the Airport show little or no surficial crust between 0 and 2 m depth, where the water table was penetrated (log ACPT7 is shown in Figure 5). From 2 to 4.5-5.8 m depth, is an extremely loose deposit of fine-sand hydraulic fill that almost certainly was responsible for the observed liquefaction and lateral spreading. qc, values within this deposit range from 2 to 14 MPa, and N values from 1 and 4 blows/ft. The material in this layer correlates with sand boil material observed at the surface.
Bay Farm Island and Harbor Bay Plaza: At Bay Farm Island, immediately north of Oakland International Airport, considerable liquefaction occurred at the northwest corner and at points along the west edge of the fill (Figures 1 and 2). Most of the western part of Bay Farm Island consists of sandy hydraulic fill, underlain by bay mud and deeper, stiffer alluvium and estuarine sediment. Soil liquefaction on the island was largely confined to undeveloped, artificially filled land. Numerous sand boils, many relatively large (with diameters more than 3 m) occurred in this area. Fill in a perimeter dike that bounds the western side of Bay Farm Island was densified by dynamic compaction, which appears to have successfully prevented soil liquefaction of the dike during the earthquake.
The logs of three boreholes drilled at the Bay Farm Island study sites suggest a complex stratigraphy of relatively thin (less than 1 m thick) interbeds of sand and finer material. A total of 20 CPT logs were also taken at one of the liquefaction test sites to observe any changes in strength over time; a representative log (HBI-CPT 1) is presented in Figure 6. The logs in the unimproved fill indicate a noncohesive layer of sand and silty sand at 2 to 3 m, with qc values of typically 2 to 16 MPa and N values of 8 to 18 blows/ft. The water table is within this layer at between 1.5 and 2 m depth. Below this layer, to approximately 6.5 m depth, is a set of finer interlayered deposits exhibiting decreasing qc values, and elevated Bq and FR values. Below 6.5 m depth, silty sand qc defines the rest of the logged soil column. Liquefaction probably occurred in the fine sand and silty sand at 2 to 3 m depth. on the basis of correlation of sand boil materials with recovered samples.
CPT and SPT logs were taken along the improved perimeter dike, which did not liquefy during the earthquake. At the elevated dike structure, the water table is somewhat deeper in the soil profile at 2.5 m depth. However, within the sandy deposits at 2.5 to 7 m depth, qc values are noticeably higher, ranging from 10 to 3 5 MPa, and N values of 29 to 58 blows/ft.
The Port of Oakland: Immediately south of the San Francisco-Oakland Bay Bridge, soil liquefaction caused considerable damage to marine-container facilities at several places in the Port of Oakland, adjacent to the Oakland Outer Harbor, as shown in Figures 1 and 3. Much of the extreme western part of the area near the Oakland shoreline, south of the mole, is filled land underlain by a relatively thin layer of soft, normally consolidated Holocene marine clay (bay mud). These shallow surficial units are, in turn, underlain by upwards of 180 m of stiffer Pleistocene glacial and interglacial deposits.
Most of the surface fill at the Port of Oakland was hydraulically placed to sea level, above which a combination of hydraulic and dumped fill was placed. After placement of the unimproved hydraulic fill, a relatively thick asphalt-cement pavement was laid in many of the terminal areas to support heavy vehicles and shipping containers. All of the Port terminals have pile-supported concrete wharves at the loading-dock edge of the terminal fill. The most severe liquefaction-induced damage to Port facilities occurred at the Marine Container Terminal at Seventh Street. Liquefaction of the hydraulic fill resulted in settlement, lateral spreading, and cracking of the pavement over large areas of the terminal. Maximum settlements of the paved container yards inboard of the wharves were about 0.3 m. Several cranes which operate along the edges of the fill, traverse laterally along the wharves on heavy rails. The outboard rail, which was pile supported on the concrete wharf, did not settle appreciably. The inboard crane rail, however, which was supported on fill throughout the entire terminal, underwent differential settlement, damaging several of the loading cranes and the dock facility was inoperable after the earthquake.
The logs of six boreholes drilled at the Port of Oakland study site (Figure 2) near the Marine Container Terminal at Seventh Street were used to assess the liquefaction susceptibility of the Port. We note that evidence of liquefaction of the subroad base material appears to have extended across essentially the entire western part of the Marine Container Terminal, as evidenced by settlement, sand boils, and tension cracking of the asphalt. The logs of boreholes P007-2 and P007-3 are shown in Figures 7 and 8. These sites typically have a resistant surface layer in the upper 3 to 4 m with qc values typically 25 to 35 MPa and N values 25 to 36 blows/ft. The water table within this layer averages approximately 2 to 2.3 m depth and fluctuates with tidal action. Below this surface layer, the fill consists of looser deposits of fine marine sand with qc values of 8 to 15 MPa, and N values of 10 and 25 blows/ft. Liquefaction appears to have occurred in the materials with in the depth range 4-8 m, on the basis of correlations of SPT samples with surface sand-boil material recovered.
San Francisco-Oakland Bay Bridge Toll Plaza: The San Francisco-Oakland Bay Bridge mole (peninsula-approach fill), immediately south of Emeryville, was extensively damaged by soil liquefaction (Figures 1 and 3). Appreciable settlement (max. 40 cm) occurred over most of the mole. Lateral spreading also was significant along most of the fill, causing numerous longitudinal fissures in the pavement parallel to the fill edges. Settlements of the mole and the elevated structures of Interstate Highway 580 and West Grand Avenue were severe, resulting, at each site, in pavement collapse and open fissures. as much as 0.3 m wide, at the soil-structure interface. The approach mole also settled below the bridge- and ramp-structure road level by as much a 0.5 m. The fissures and uneven pavement surface interrupted ground transportation in the interchange areas after the earthquake.
The logs of five boreholes drilled at the San Francisco-Oakland Bay Bridge Toll Plaza study site indicate a complex stratigraphy. Five sets of logs were taken at locations along the median strip and open space between the east bound and west bound lanes of Interstate Highway 80. The log of borehole SFOBB-1 was taken at the easternmost end of the Toll Plaza parking lot. The log is marked by a moderately dense surface crust between 0 and 2.8 m depth, followed by an apparently interbedded sequence of cohesionless silty sand and cohesive clay to 4.5 m depth, as noted by the local variations in FR and qc values. The water table during sampling was in the upper part of the section at approximately 2 m depth. Between 4.5 and 7.5 m lies cohesionless material with low qc values typically 5 MPa and N values of typically 5 to 17 blows/ft with a maximum N value of 28 blows/ft at 5.2 m, and a minimum N value of 0 blows/ft (drill rod and hammer sank under their own weight) at 6.6 m depth. This layer appears to have been responsible for the observed liquefaction, on the basis of both inplace testing and correlation with surface sand-boil material.
The logs taken of two boreholes SFOBB-3 and SFOBB-4 (not shown), drilled in the central part of the median strip directly south of the Toll Plaza and maintenance buildings, show similar soil profiles. The upper 4 m sections of these two logs suggest moderately loose sand (qc = 5-15 MPa) interbedded with finer materials. Below, there is a somewhat uniform, more sand rich deposit to 8-m, with qc, values of typically 4 to 15 MPa. This layer appears to be the principal liquefied layer. No SPT measurements were made at boreholes SFOBB-3 or SFOBB-4. The logs of two boreholes, SFOBB-2 and SFOBB-5 (Figure 10), in a wooded area at the west end of the median strip show gravelly sand to 6 m depth beneath a dense crust. From 6 to 8.5 m depth is a layer of silty and sandy soil, qc values of 5 to 10 MPa and N values of 7 to 19 blows/ft, within which liquefaction apparently occurred.
Port of Richmond: Soil liquefaction occurred at a study site in the western part of Richmond Inner Harbor (Figures 1 and 4). The area that liquefied is a zone approximately 75 m wide and 300 m long at the foot of Harbor Way Road (10th Street). Approximately 85 km north of the Loma Prieta rupture zone, this site represents the most distant point from the zone of energy release to undergo soil liquefaction sufficient to damage structures. Much of the land at the edge of the Richmond Inner Harbor had been created by placement of uncompacted sandy hydraulic fill. The liquefaction zone overlies deposits of soft bay mud, which are underlain, in turn, by deeper deposits of stiffer, overconsolidated noncohesive and cohesive soils that naturally fill a deep fluvial channel at the west end of the harbor. Four large and a dozen smaller sand boils vented fine sand and silty sand from the underlying fill. In addition, minor settlements of approximately 2 to 8 cm and fissures of similar magnitude occurred at the edge of the harbor.
Three boreholes (POR-2 through POR-4) drilled in the open space north of the Harbor seawall (Figures 1 and 4). SPT's were performed in shallow boreholes adjacent to each CPT. The logs of the boreholes drilled at the Port of Richmond study site (Figure 2) are all similar and are characterized in the upper part of the section by an oxidized, tan-brown crusted silty sand to 0.8 m depth; oxidized tan-brown silty sand to 1.8 m depth; and oxidized, tan-brown sandy clay to reduced sandy clay to 4.1 m depth. The water table was approximately at a depth of 2.5 m during sampling. This section has low qc values of approximately 4 to 5 MPa, FR values greater than 3% and pore pressure generation, and N values ranging from 1 to 6 blows/ft. Below this section, at 4.1 to 7.8 m depth, is a hydraulic-fill layer of olive gray fine-silty sand containing shell fragments. This layer almost certainly was responsible for the observed liquefaction at the surface, and surface sand-boil material is identical to SPT samples of this layer. This layer has extremely low qc values 1 to 3 MPa, average FR values of 0.3 to 1.0%, and minor pore pressure generation during cone penetration, and N values of 2 to 11 blows/ft, but typically, 2 to 5 blows/ft. Below this layer is a thin deposit of soft bay mud to 9.5 m depth, underlain by dense sand deposits.
where amax is the peak ground acceleration, g is the gravitational acceleration (9.81 m/s2), vo is the total overburden stress, vo is the effective overburden stress, and rd is a depth-reduction factor, which can be estimated in the upper 10 m of the soil column by the following relation:
where z is the depth in meters (Kayen et al., 1992).
The second approach utilizes a quantitative measure of earthquake shaking intensity, termed Arias intensity (Arias, 1970). It is the sum of the energy absorbed by an evenly-spaced population of idealized undamped simple oscillators in response to the earthquake motion. For the two horizontal components of motion, the Arias intensity Ih is calculated as follows:
where Ixx is the Arias intensity measured in the x direction; Iyy is the Arias intensity in the y direction; g is the acceleration due to gravity; t0 is the duration of earthquake shaking; and ax and ay are the transient accelerations in the x and y directions. The Arias intensity measure has the dimensional units of velocity.
As in the case of amax, we find that Ih typically diminishes with depth in the soil column. For liquefaction assessment we need to know the Ih profile within the soil column. We estimated the attenuation of Ih with depth through a parametric study of synthetic accelerograms propagated through shear modulus representations of soil columns at our east bay study sites. Synthetic seismograms were generated for the surface and depth nodes of soil column input files with the ground-response computer-program SHAKE90, a modified version of the program presented by Schnabel et al. (1972). Strong motion records from eight earthquakes were propagated through a soil column representing loose sandy fill (Vs = 150 m/s.) overlying a cohesive soil with shear wave velocity profiles for San Francisco bayshore sites (Sun et al., 1989; Golesorkhi, 1989; Kayen, 1993). Each profile was underlain by an elastic half-space with a shear wave velocity of 2500 m/s, representative of the Mesozoic Franciscan Formation underlying the east bay sites. Output synthetic acceleration-time histories for layer nodes at depth in the model soil profiles were integrated to calculate Ih and then normalized by the value of Ih at the surface. This normalization process allows us to evaluate the depth-dependency of Ih over a broad range of earthquake and site conditions by collapsing the profiles to a common reference value (unity) at the ground surface. We define the Arias intensity depth-of-burial reduction parameter, rb, as the ratio of the buried-to- surface Arias intensity.
The rb parameter is analagous to rd and can be calculated from either one- or two-component horizontal Arias Intensity. The mean and ±1 responses calculated from the suite of SHAKE90 runs are plotted in Figure 12. The mean attenuation of rb falls somewhat linearly from 1.0 to a value of 0.58 at 6 m depth, then decreases to 0.46 at 10 m. The value of rb remains essentially constant at depths below 10 meters (30.5 ft). We calculated the Arias intensity in the soil column, Ihb, as the product of Ih measured from recorded seismograms in the east bay, and the reduction parameter rb estimated from the SHAKE90 study.
In our study, we associate CSR and Ihb profiles with the liquefaction field performance data (N1)60, and qc1 from our study sites (Figure 13). Detailed discussion of the Arias intensity method for liquefaction potential assessment is presented in Kayen (1993), Kayen and Mitchell (1997), and Kayen et al. (1997). CSR and Ihb values used in the analyses of critical soil layers at the east bay study sites (Figure 1) are summarized in Table 1.
The SPT boundary curve for the assessment of soil liquefaction potential based on Arias intensity is presented in detail in Kayen and Mitchell (1997) and is presented here for soils with low-plasticity fines content of less than 5 % (Figure 13a). An equivalent cyclic stress ratio based boundary for assessment of liquefaction potential was presented by Seed and de Alba (1986). The insensitivity of CSR to earthquake magnitude and duration requires that magnitude correction factors be applied to CSR-values to scale for the severity of the earthquake shaking. This results in the suite of magnitude dependent boundary curves. Recent work questions the CSR boundary curves presented in Seed and de Alba (1986) as derived from the scaling factors of Seed and Idriss (1982) (Ambraseys, 1988; Arango, 1996). Arias intensity incorporates magnitude, frequency and earthquake shaking duration elements of strong motion and therefore needs no magnitude correction factors.
The Ihb at the liquefaction boundary in Figure 13A represents the minimum threshold seismic intensity (Ihb,1) required to induce liquefaction in soil of a given (N1)60 or qc1 value. A similar threshold in terms of CSR can be determined based on the boundaries presented by Seed and de Alba (1986). CPT-based boundary curves for CSR have been proposed by Robertson and Campanella, 1985; Seed and de Alba, 1986; Shibata and Teparaksa (1988) and Mitchell and Tseng (1990). The proposed boundaries of Robertson and Campanella (1985) and Seed and De Alba (1986) are based on qc/N ratios rather than direct field investigations of liquefaction test sites. The method of Shiibata and Teparaksa (1988) is based on a limited data set of direct measurements at liquefaction sites for earthquakes in Japan, China, and the United States. Mitchell and Tseng (1990) presented boundary curves developed from qc - Dr, (relative density)-CSR relationship at a given effective confining stress which are based on cavity expansion theory and cone chamber test results.
We compared CSR1 and Ihb ,1 with the corresponding seismic-shaking intensity function induced by the earthquake (CSReq and Ihb , eq) using Equations (7) through (10) to determine profiles of factor of safety against liquefaction. The Arias intensity based factor of safety FIhb against liquefaction occurrence is defined as the ratio of the Arias intensity required to cause liquefaction to the Arias intensity imparted by the earthquake.
This Arias intensity definition of factor of safety is analogous to that defined by Seed and his colleagues based on cyclic stress (FCSR).
We converted the N profiles from Figures 5-11 into factor of safety profiles for both FIhb and FCSR in Figure 14, based in the calculated CSR and Ihb for the study sites. At Bay Farm Island both methods accurately predict no-liquefaction at the improved-dike and liquefaction at between 3.5 and 4.5 m at HBI-CPT1, and at Oakland International Airport both methods predict liquefaction in a zone between approximately 2 and 5 m. The profiles at the Port of Oakland for POO7-2 are in agreement and identify the zone of liquefaction between 5 and 8 m, and POO7-3 indicates thins seams of loose potentially liquefiable material at 6.5 and 7.5 m, but otherwise a profile of non-liquefaction susceptibility. Profiles of FCSR and FIhb for the San Francisco-Oakland Bay Bridge site correctly indicate that liquefaction occurred principally at a depth of between approximately 5.5 and 7.5 m at SFOBB-1 and between 6.0 and 9.0 m at SFOBB-5, based on field observations. At the Port of Richmond, the profiles accurately identify the zone of liquefaction between 4 and 5 m in the profile POR-2. At POR-Hall St., both profiles accurately indicate no liquefaction.
The CPT data used to construct the Ihb - qc1 association in Figure 13b are derived almost exclusively from the Loma Prieta dataset. Thus, FIhb profiles when presented in a manner similar to Figure 14, are entirely consistent and in agreement with the results presented in Figure 13b where soil layers are summarized as individual Ihb / qc1 points. The new Arias intensity boundary for liquefaction assessment, presented here, is therefore based on a limited data set from the Loma Prieta earthquake. This boundary will enhanced by new liquefaction field data that become available.
A direct comparison was made of the cyclic stress and Arias intensity-based methods for assessing the potential of soil liquefaction using the SPT field test. We found that when using SPT data, the two methodologies are largely in agreement in predicting the occurrence or nonoccurrence of liquefaction in soils layers at our study sites during the earthquake. Field data from the Loma Prieta study sites allow us to construct a new Arias intensity-cone penetration resistance boundary for liquefaction assessment.
For liquefaction assessment, the advantages in using Arias intensity over peak ground acceleration (amax), used in the cyclic-stress approach, are that (1) Arias intensity is derived from the acceleration records of both horizontal components of motion over the entire duration of motion, whereas amax utilizes a single arbitrarily selected value; (2) Arias intensity incorporates the intensity of motions over the full range of recorded frequency, whereas, amax is often associated with high-frequency motion; and (3) the breakdown of soil structure that results in liquefaction is fundamentally more dependent upon input-energy than on a single level of acceleration (Liang, and others., 1996).
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For further information PLEASE CONTACT: Robert Kayen
last modified 6 January 2005
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Department of the Interior U.S. Geological Survey Geologic Division Coastal & Marine Geology Program Western Region Coastal & Marine Geology