LIQUEFACTION POTENTIAL OF INORGANIC AND ORGANIC SILTS

Christopher D. P. Baxter, Ph.D., P.E. Aaron S. Bradshaw, P.E. George E. Veyera, Ph.D.

University of Rhode Island December 2005

URITC PROJECT NO. 00060

PREPARED FOR UNIVERSITY OF RHODE ISLAND

TRANSPORTATION CENTER

DISCLAIMER

This report, prepared in cooperation with the University of Rhode Island Transportation Center, does not constitute a standard, specification, or regulation. The contents of this report reflect the views of the author(s) who is (are) responsible for the facts and the accuracy of the data presented herein. This document is disseminated under the sponsorship of the Department of Transportation, University Transportation Centers Program, in the interest of information exchange. The U.S. Government assumes no liability for the contents or use thereof.

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Table of Contents 1.0 Introduction…………………………………………………………………................. 1 2.0 Properties of Rhode Island Silts……………………………………………………….. 2 3.0 Evaluation of Sample Quality………………………………………………................. 3

3.1 Field Program………………………………………………………………….. 3 3.1.1 Study Sites…………………………………………………………… 3 3.1.2 Sampling Methods………………………………………................... 4 3.1.3 Results and Conclusions……………………………………………... 8

3.2 Laboratory Program……………………………………………………………. 9 3.2.1 Methods Used to Evaluate Sample Quality………………………….. 9 3.2.2 Laboratory Testing Methods…………………………………………. 11 3.2.3 Results and Conclusions……………………………………………... 11

3.3 Conclusions…………………………………………………………................. 13 4.0 Background on Liquefaction Potential of Silts………………………………………… 15

4.1 Liquefaction Potential of Silts ………………………………………………… 15 4.1.1 Field Observation……………………………………………………. 15 4.1.2 Laboratory Investigations……………………………………………. 15

4.2 Standard of Practice for the Evaluation of Liquefaction Potential……………. 16 4.2.1 Cyclic Resistance of Soil…………………………………................. 16

4.2.1.1 SPT-Based Approach……………………………………… 16 4.2.1.2 CPT-Based Approach…………………………................... 18 4.2.1.3 Magnitude Scaling Factors………………………………… 20

4.2.2 Seismic Demand……………………………………………………... 20 5.0 Laboratory Study of Liquefaction Resistance…………………………………………. 22

5.1 Soils Tested……………………………………………………………………. 22 5.2 Test Method……………………………………………………………………. 23 5.3 Sample Preparation Methods………………………………………………….. 23

5.3.1 Slurry Consolidation…………………………………………………. 24 5.3.2 Moist Tamping……………………………………………................. 24

5.4 Liquefaction Resistance………………………………………………………... 26 5.4.1 Organic Silt…………………………………………………………... 26 5.4.2 Inorganic Silt…………………………………………………………. 28

5.5 Preliminary Evaluation of Earthquake Demand……………………................. 31 5.6 Summary and Conclusions…………………………………………................. 33

6.0 Conclusions……………………………………………………………………………. 34 7.0 Acknowledgements……………………………………………………………………. 34 8.0 References……………………………………………………………………………… 35

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List of Tables Table 2.1 Summary of index properties of inorganic silts compiled from various sites

in the Providence area……………………………………………….............. 2 Table 2.2 Summary of index properties of organic silts compiled from various sites in

the Providence area…………………………………………………….......... 2 Table 3.1 Summary of sampling results from the Fox Point site (organic silts) (data

from Page 2004)……………………………………………………………... 8 Table 3.2 Summary of sampling results from the Old Farmer’s Market site (organic

silts) (data from Page 2004)…………………………………………………. 8 Table 3.3 Criteria for evaluating sample disturbance by Silva (1974)…………………. 10 Table 3.4 Criteria for evaluating sample disturbance by Andresen and Kolstad (1979)

and Terzaghi et al. (1996)…………………………………………………… 10 Table 3.5 Criteria for evaluating sample disturbance by Lunne et al. (1997) for

normally consolidated soils………………………………………………….. 10 Table 3.6 Evaluation of sample disturbance from consolidation test results. Samples

were recovered from Boring URI-1 (data from Page 2004)………………… 12 Table 3.7 Summary of relevant triaxial test parameters……………………………….. 13 Table 5.1 Summary of seismic parameters obtained from the site response analysis

performed for the Washington Bridge by Weston Geophysical (Weston 2000)………………………………………………………………………… 31

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List of Figures Figure 3.2 Site plan showing locations of the two study sites………………………… 4 Figure 3.3 Schematics of the (a) Shelby tube sampler and (b) GUS sampler (Acker

2003)……………………………………………………………………….. 6 Figure 3.4 Schematics of the (a) fixed piston sampler and (b) SGI sampler (Acker

2003; SGI 2003)……..…………………………………………………….. 7 Figure 3.5 Method for determining degree of disturbance based upon Schmertmann’s

graphical procedure (Silva 1974)………………………………………….. 9 Figure 3.6 Consolidation test results from undisturbed and reconstituted samples

(Page 2004)………………………………………………………………... 12 Figure 3.7 Typical stress versus strain curves obtained from the triaxial tests (Page

2004)……………………………………………………………………….. 13 Figure 4.1 SPT-based liquefaction curves (after Youd and Idriss 2001)……………... 17 Figure 4.2 CPT-based liquefaction curves (after Robertson and Wride 1998)……….. 18 Figure 4.3 Magnitude scaling factors recommended by NCEER for use in

engineering practice (after Youd and Idriss 2001)………………………… 20 Figure 5.1 Grain size distribution of organic and inorganic silts used in this study….. 22 Figure 5.2 Photograph of the cyclic triaxial apparatus used in this study……………. 23 Figure 5.3 Density profile of samples prepared using slurry method (a) and moist

tamping method with 0, 1, and 2 percent undercompaction (b-d)………... 24 Figure 5.4 Dry density verses compaction energy for inorganic silts at two different

molding water contents……………………………………………………. 25 Figure 5.5 Typical cyclic triaxial test results…………………………………………. 27 Figure 5.6 Liquefaction resistance of organic silts prepared by slurry consolidation… 28 Figure 5.7 Liquefaction resistance of inorganic silts prepared by slurry consolidation. 29 Figure 5.8 Liquefaction resistance of inorganic silts prepared by moist tamping at

approximately constant void ratio at different molding water contents and tamping efforts…………………………………………………………….. 29

Figure 5.9 Comparison of liquefaction resistance of samples prepared by slurry consolidation and moist tamping methods at approximately constant void ratio………………………………………………………………………… 30

Figure 5.10 Liquefaction resistance of inorganic silts at two different densities prepared by moist tamping at the optimum water content………………… 30

Figure 5.11 Profiles of average CSR with depth estimated from 1-D site response analysis using four different soil profiles for a (a) 500-year seismic event and a (b) 2,500-year event (after Weston 2000)…………………………… 32

Figure 5.12 Range of average CSRs estimated for the Washington Bridge (Weston 2000) compared to the CRRs of loose organic and inorganic silt determined in the laboratory. Two design levels are shown including a 500-year event (M=5) and a 2,500-year event (M=6)…………………….. 33

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1.0 Introduction

Although Rhode Island is not considered to be a region of high seismicity, by code the liquefaction potential of soil during an earthquake must still be considered in the design of certain highway structures. There is some debate in local engineering practice whether the silty soils that are commonly encountered in Rhode Island are susceptible to liquefaction. These soils are typically characterized as either inorganic silts or organic silts and are sometimes found in a loose condition in the field. The current standard of practice for evaluating liquefaction potential uses insitu tests such as the standard penetration test and engineering correlations, which were developed primarily from tests in sands. Because of this, there is still considerable uncertainty about the dynamic behavior of the Rhode Island silts.

This report presents a laboratory study on the liquefaction potential of silts from Rhode Island. The second section of the report presents a combined field and laboratory study to assess the quality of “undisturbed” samples obtained using standard drilling and sampling techniques. It is extremely important to know whether samples recovered from the field are suitable for liquefaction testing in the laboratory. The third section discusses the properties of silts that are commonly encountered in Rhode Island. The fourth section presents a literature review on liquefaction potential of silts. The purpose of the review is to present field and laboratory evidence of liquefaction in silty soils and to document the standard of practice in evaluating liquefaction potential of silts. The fifth section presents a laboratory study to quantify the liquefaction resistance of organic and inorganic silts. These results are compared to a range of anticipated earthquake demand, thus providing some preliminary assessment of the likelihood of liquefaction of these soils in a design level earthquake in Rhode Island.

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2.0 Properties of the Rhode Island Silts

The soils that were used in this study are commonly referred to as the “Rhode Island Silts” or the “Providence Silts”. As these names suggest, the soils are commonly encountered in Rhode Island and consist primarily of silt-sized particles. These soils, which are considered to be “transitional” soils, have engineering properties that may be similar to sands in some cases yet similar to clay in others. The Providence silts are commonly classified into two categories; inorganic silts and organic silts. The index properties of these soils compiled from geotechnical data at various sites is summarized in Tables 2.1 and 2.2. These properties include the percentage of clay-sized particles, liquid limit (LL), plasticity index (PI), and percent organic content by weight.

Table 2.1. Summary of index properties of inorganic silts compiled from

various sites in the Providence area.

Site % Clay LL PI %

Organics Reference

Old Farmers Market, Providence <5 NP NP 0.6-1.3 Page (2004)

Wellington Ave, Cranston 2-33 NP-30 NP-5 0.1-1.3 DMJM Harris (2002)

I-95, Cranston 2-3 23-29 7-11 - D’Andrea (1966)

Table 2.2. Summary of index properties of organic silts compiled from various sites in the Providence area.

Site % Clay LL PI % Organics Reference

Quonsett Point, Davisville 13-15 33-74 25-40 1.2-2.0 Baxter et al. (2004);

Silva et al. (2003) Fox Point, Providence 11-16 55-73 24-35 2.4-4.2 GZA (2002) Civic Center, Providence 7-13 36-47 - 2.5-3.2 GZA (1998)

Wilkes Barre Pier, Providence 13-19 60-75 24-42 6.7-7.1 Nacci (1980)

The inorganic silts are typically characterized as ML in the Unified Soil Classification

System. These soils were deposited as outwash material during the last glacial recession and are found in a loose to dense condition insitu. They tend to be highly stratified or varved, having alternating layers of fine and coarse-grained material. These soils are typically non-plastic or have very low plasticity (PI<5) and low organic contents. The grain size can vary greatly from site to site ranging from sandy silt to clayey silt. In terms of engineering behavior, the inorganic silts could be thought of as very fine sand having a relatively low permeablility.

The organic silts, on the other hand are more analogous to clay soils and are typically found in coastal areas surrounding Narragansett Bay. These soils are characterized by higher clay contents and significantly higher plasticity (PI = 24-42). The organics contained within the soil are what give these soils their plastic nature classifying them as OH.

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3.0 Evaluation of Sample Quality

In laboratory liquefaction studies it is necessary to test soil samples that are representative of their in situ condition. This can be accomplished by obtaining high quality undisturbed samples from the field. Previous drilling experience in Rhode Island suggests that silt samples can be recovered using conventional sampling equipment. However, it is uncertain if these samples would be suitable for liquefaction testing. This section presents an investigation of the sample quality of inorganic and organic silts. 3.1 Field Program

A field program was conducted to recover undisturbed samples of organic and inorganic silts for laboratory testing. The selected study sites, sampling methods, and results and conclusions are presented below. 3.1.1 Study Sites

A total of four geotechnical borings were performed as part of this study at two sites in

Providence, RI to obtain samples of inorganic and organic silts. The location of the study sites are shown on the site plan in Figure 3.2. Organic silts were obtained from a site located near Fox Point, at the corner of South Water St. and India St. It is also the same location as the new approach ramp on the eastern side of the Providence River Bridge that is currently being constructed as part of the I-195 rerouting project. Two borings (URI-1 and URI-2) were performed at Fox Point on December 8 and 9, 2003 to maximum depths of 11.6 and 12.2 m. Existing geotechnical information (GZA 2002) was used to select the boring locations and sampling depths. This information indicated that the organic silts at the sampling location were at a depth of about 5.5 m below the ground surface, had a thickness of about 9.8 m, and were in a soft condition.

Inorganic silts were obtained from the former Farmers Market located to the west of the Providence Place Mall and just north of the Amtrak rail lines. Two borings (URI-3 and URI-4) were performed at the site on January 29 and 30, 2004 to maximum depths of 15.2 and 15.9 m. Existing boring logs (GZA 1998) indicated that stratified layers of inorganic silt and clayey silt were encountered at a depth of about 5.8 m at the sampling location. The silt stratum had a thickness of about 40 m and was primarily in a loose condition.

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Figure 3.2. Site plan showing locations of the two study sites.

3.1.2 Sampling Methods

Geotechnical sampling was performed at various depths during drilling operations to obtain specimens for laboratory testing. Drilling was performed using the rotary wash boring method. In this method a hydraulic motor mounted on the drill rig powers a drill bit attached to the end of a drill pipe. Fluid is pumped down the center of the drill pipe which washes the cuttings to the ground surface. Samples were recovered by drilling to a specified depth, attaching a tube sampler to the end of the drill pipe, lowering the sampler to the bottom of the hole, and pushing the sampler into the soil. To improve recovery of the sample, some time is given to allow the sample to “set-up” or expand within the tube. Either water or a weighted drilling mud was used in the borehole to maintain the vertical overburden stress and limit soil disturbance at the bottom of the hole. Four different types of sampling equipment were used including a thin-walled Shelby tube sampler, a hydraulic piston sampler, and 2 different types of fixed piston samplers. Schematics of these samplers are given in Figures 3.3 and 3.4. The Shelby tube (Figure 3.3a) is the standard sampling method (ASTM D 1587) and consists of a thin-walled tube with knurled cutting edges that cut the soil at a smaller diameter than the inside diameter of the tube. This limits friction along the side walls and allows the sample to “flow” into the tube. Samplers can be described by the area ratio which is defined as the cross sectional area of the sample tube

Farmer’s Market Site

Fox Point Site

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divided by the inside cross sectional area of the tube. The smaller the area ratio the less soil that is displaced during sampler penetration. The Shelby tube has an area ratio of 9. Piston samplers are designed to further reduce soil disturbance during sampling. The piston, which remains stationary at the soil surface (at the bottom of the bore hole), creates a suction force inside the sample tube and acts to prevent soil from being dragged down while the tube is pushed into the soil. The first type of piston sampler that was used in this study was the Gregory Undisturbed Sampler (GUS) manufactured by Acker Drill Company (Figure 3.3b). It consists of a Shelby tube, piston, and a hydraulic actuator assembly that is attached to the end of the drill rod. The piston which is positioned at the cutting end of the Shelby tube is lowered to the bottom of the borehole where it is fixed. Drilling fluid is used to hydraulically push the Shelby tube into the soil while the piston is kept stationary. The second type of piston sampler used is called the fixed piston sampler (Figure 3.4a) and is manufactured by Acker Drilling Company. Like the GUS sampler, it has a Shelby tube and piston, only the piston is fixed at the bottom of the borehole using thin rods that run up the center of the drill pipe. The Shelby tube is attached to the end of the drill rod which is used to push the tube in to the soil. The third type of piston sampler (Figure 3.4b) is manufactured by the Swedish Geotechnical Institute (SGI). The sampler consists of a steel pipe with a plastic liner on the inside and a steel cutting shoe (nose cone). Relative to the Shelby tube, the sampler has thicker walls with an area ratio of 28. The piston is most similar to the fixed piston sampler only that the piston is fixed at the bottom of the borehole using a chain that runs up the center of the drill rod. The major feature that differentiates it from the other samplers is that the inside diameter of the cutting shoe is the same as the inside diameter of the sample tube. This prevents the sample from expanding within the tube during penetration and therefore limits disturbance. Like the other piston samplers, the piston creates a suction which counteracts the friction inside the sample tube.

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(a) (b)

Figure 3.3. Schematics of the (a) Shelby tube sampler and (b) GUS sampler (Acker 2003).

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(a) (b)

Figure 3.4. Schematics of the (a) fixed piston sampler and (b) SGI sampler (Acker 2003; SGI 2003).

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3.1.3 Results and Conclusions

A summary of the samples recovered from the two study sites is shown in Tables 3.1 and 3.2. Samples of organic silt were obtained in 10 of the 11 attempts with recovery lengths ranging from 10 to 61 cm. In contrast, recovery of the inorganic silt was only possible in 2 of the 15 attempts despite efforts to vary set-up time before recovery. The non-plastic and loose nature of the inorganic silts caused the soil to essentially flow out of the bottom of the sample tube during recovery. It is anticipated that the two inorganic silt samples that were recovered are in a highly disturbed condition. However, considering the plastic nature of the organic silts which helps to resist disturbance, a laboratory program was used to evaluate sample quality in this soil as described in the next section.

Table 3.1. Summary of sampling results from the Fox Point site (organic silts)

(data from Page 2004).

Boring Sampler Type Sample Depth (m) Recovery

Length (cm)

Setup Time (min)

URI-1 GUS P-1 6.7 to 7.3 38 25 URI-1 GUS P-2 8.8 to 9.5 10 25 URI-1 Shelby P-3 9.5 to 10.1 61 25 URI-1 Shelby P-4 10.2 to 10.8 53 35 URI-1 GUS P-5 11.0 to 11.6 50 20 URI-2 Shelby P-1 5.8 to 6.4 38 30 URI-2 Shelby P-2 7.3 to 7.9 0 25 URI-2 GUS P-3 8.8 to 9.5 51 40 URI-2 GUS P-4 10.1 to 10.7 15 40 URI-2 Shelby P-5 11.0 to 11.6 25 30 URI-2 Shelby P-6 11.6 to 12.2 46 20

Table 3.2. Summary of sampling results from the Farmer’s Market site (inorganic silts)

(data from Page 2004).

Boring Sampler Type Sample Depth (m) Recovery Length (cm)

Setup Time (min)

URI-3 Fixed Piston P-1 7.9 to 8.5 0 40 URI-3 SGI P-2 8.8 to 9.5 0 40 URI-3 SGI P-3 10.4 to 11.0 51 40 URI-3 Shelby P-4 11.0 to 11.6 51 45 URI-3 SGI P-5 12.8 to 13.4 0 45 URI-3 Shelby P-6 14.0 to 14.6 0 45 URI-4 Fixed Piston P-1 7.9 to 8.5 0 44 URI-4 SGI P-2 8.5 to 9.1 Bag Sample 35 URI-4 Shelby P-3 9.1 to 9.8 0 35 URI-4 SGI P-4 9.8 to 10.4 Bag Sample 40 URI-4 Fixed Piston P-5 11.0 to 11.6 0 37 URI-4 SGI P-6 12.2 to 12.8 0 40 URI-4 Shelby P-7 13.4 to 14.0 0 40 URI-4 SGI P-8 14.0 to 14.6 Bag Sample 44 URI-4 Shelby P-9 14.6 to 15.2 0 40

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3.2 Laboratory Program This section presents the results of the laboratory study to evaluate sample quality of the organic silts recovered from the Fox Point site. Testing involved either one-dimensional consolidation tests or triaxial compression tests to evaluate the level of sample disturbance. Though the approaches presented were developed from tests performed on clayey soils, it is anticipated that they are applicable to the highly plastic organic silts. The sample disturbance evaluation methods, and testing results and conclusions are given below. 3.2.1 Sample Quality Evaluation Methods Used

Methods that use 1-D consolidation test data include those proposed by Silva (1974),

Andresen and Kolstad (1979), Terzaghi et al. (1996), and Lunne et al. (1997). Silva’s (1974) method is based on the simplified construction of a corrected field compression curve that was suggested by Schmertmann (1955), shown in Figure 3.5. Silva defined a disturbance index,

od eeI ∆∆= , where e∆ = the initial sample void ratio at the minus the void ratio at the preconsolidation stress, and oe∆ = the initial sample void ratio of the sample minus the void ratio of a remolded base line curve at the preconsolidation stress. His criterion for sample quality, shown in Table 3.3, is based on tests performed on deep sea clays.

Figure 3.5. Method for determining degree of disturbance based upon

Schmertmann’s graphical procedure (Silva 1974). Andresen and Kolstad (1979) presented a method of assessing sample disturbance in

terms of the volumetric strain, voε , defined as the ratio of the change in sample volume to the initial sample volume. Their method is based on reconsolidating samples in the laboratory back to the in-situ stresses. The amount of volume change that occurred during reconsolidation was

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an indication of the quality of the sample disturbance. Their criterion for evaluating sample disturbance is presented in Table 2.4.

Terzaghi et al. (1996) adapted the Andresen and Kolstad (1979) criteria. The volumetric strain criteria were used, however the designations of sample quality were changed to A (excellent) - E (very poor), as shown in Table 3.4.

Lunne et al. (1997) modified the Andresen and Kolstad (1979) approach by evaluating sample disturbance by the ratio poVV∆ , where ∆V = the change in sample volume and Vpo = the initial pore volume. Their criterion is presented in Table 3.5.

Table 3.3. Criteria for evaluating sample disturbance by Silva (1974). ID Degree of Disturbance

< 0.15 Very Little 0.15 to 0.30 Small Amount 0.30 to 0.50 Moderate 0.50 to 0.70 Much

> 0.70 Extreme

Table 3.4. Criteria for evaluating sample disturbance by Andresen and Kolstad (1979) and Terzaghi et al. (1996).

Sample Quality voε (%)

Andresen and Kolstad (1979) Terzaghi et al. (1996) < 1 Very good to excellent A

1 to 2 Good B 2 to 4 Fair C

4 to 10 Poor D > 10 Very Poor E

Table 3.5. Criteria for evaluating sample disturbance by Lunne et al. (1997) for normally consolidated soils.

poVV∆ Sample Quality

<0.04 Very Good to Excellent 0.04 to 0.07 Good to Fair 0.07 to 0.14 Poor

>0.14 Very Poor Though no formal methods have been developed to evaluate sample disturbance from

triaxial test data, the strain to failure can provide a qualitative assessment of sample disturbance (Lunne et al. 1997; Tanaka and Tanaka 1999; Santagata and Germaine 2002). Sample disturbance typically causes an increase in the strain to failure and a decrease in the undrained shear strength.

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3.2.2 Laboratory Testing Methods

A total of seven consolidation and seven triaxial tests were performed on the organic silt. A complete description of testing procedures and results can be found in Page (2004). Testing was performed on both undisturbed and reconstituted samples. Reconstituted samples were made using a batch slurry consolidometer. The consolidation tests were performed on 5-cm diameter samples using automated incremental loading equipment manufactured by the Geocomp Corp. The triaxial tests were also performed using Geocomp equipment on both 3.6 and 7.1-cm diameter samples to investigate the effect of sample size. The triaxial samples were consolidated under Ko conditions and sheared undrained. Samples were prepared for testing by cutting the sample tube into sections, carefully extruding the soil from the tube, and trimming the soil to the specified diameter. 3.2.3 Results and Conclusions

The consolidation test results are shown in Figure 3.6 where the vertical strain is plotted against the log of the vertical stress. Casagrande’s construction method was used to estimate the preconsolidation stress from the measured test data. The strain to failure was then calculated and used to evaluate sample disturbance using the methods described earlier.

The analysis of sample disturbance is presented in Table 3.6. Silva’s (1974) method suggests that the disturbance of the intact samples ranges from “small” to “moderate” amounts of disturbance. Note that the reconstituted samples which have no in situ fabric are also designated as having a “small” amount of disturbance. All other methods including Anderson and Kolstad (1979), Terzaghi et al. (1996) and Lunne et al. (1997), indicate that the samples are highly disturbed and are of poor quality. The results also suggest that the level disturbance was independent of the sampler type used.

Typical stress-strain curves from the triaxial tests are shown in Figure 3.7. Relevant parameters obtained from the test results are summarized in Table 3.7. The strain at failure shown in the table was selected using the maximum deviator stress. Strength was compared by normalizing the shear strength measured in each test by the initial overburden stress ( vouS σ ′ ). As shown by the results in Table 3.7, the sampler type had no clear effect on either the strain to failure or the shear strength. The strengths of the undisturbed samples were higher than the strength of the reconstituted sample. However, the strain to failure was within the same range suggesting that the recovered samples are likely disturbed.

The effect of sample size was investigated by comparing adjacent test samples trimmed to different diameters. These include the sample pairs of 15 and 16, and 13 and 14. As shown by these data, the strain to failure was less in the 3.6-cm samples than in the 7.1-cm samples suggesting that the smaller samples are less disturbed. However, the undrained strength was higher for the smaller sample in one case and lower in the other. The strengths measured in both sample sizes were within 15%. Therefore, considering the poor quality of the samples and the inconsistent trends in strength, the size of the sample has no apparent effect.

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1 10 100 1000 10000

Vertical Effective Stress (kPa)Ve

rtica

l Stra

in (%

)0

10

20

30

IL-04-03 Shelby URI-1 P-3IL-04-03 Shelby URI-1 P-3IL-04-07 Shelby URI-1 P-4IL-04-09 Shelby URI-1 P-4IL2-04-02 GUS URI-1 P-1IL-04-05 Reconstituted IL-04-08 Reconstituted

Figure 3.6. Consolidation test results on undisturbed and reconstituted samples (Page 2004).

Table 3.6. Evaluation of sample disturbance from consolidation test results. Samples were recovered from Boring URI-1 (data from Page 2004).

Silva (1974)

Andresen & Kolstad (1979); Terzaghi et al.

(1996)

Lunne et al. (1997)

Test Sample Sampler Type

Id Disturbance εvo (%) Sample Quality poV

V∆ Sample Quality

2 P-1 GUS 0.25 Small 7.6 Poor; D 0.13 Poor 3 P-3 Shelby 0.33 Moderate 8.7 Poor; D 0.14 Poor 6 P-3 Shelby 0.24 Small 8.8 Poor; D 0.13 Poor 7 P-4 Shelby 0.33 Moderate 8.3 Poor; D 0.13 Poor 9 P-4 Shelby 0.31 Moderate 6.6 Poor; D 0.10 Poor 5 NA Reconstituted 0.23 Small 6.8 Poor; D 0.14 Poor 8 NA Reconstituted 0.21 Small 4.5 Poor; D 0.08 Poor

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Axial Strain (%)

0 2 4 6 8 10

Nor

mal

ized

Dev

iato

r Stre

ss (k

Pa)

0.5

0.6

0.7

0.8

0.9

1.0

MGL-04-16 2.8-inchMGL-04-13 1.4-inchMGL-04-14 2.8-inchMGL-04-15 1.4-inch

Figure 3.7. Typical stress versus strain curves obtained from the triaxial tests (Page 2004).

Table 3.7. Summary of relevant triaxial test parameters.

Test Boring Sample Sampler Type Sample

Diameter (cm)

Strain to Failure

(%) Su/σ’vo

15 URI-1 P-1 GUS 3.6 2.0 0.41 16 URI-1 P-1 GUS 7.1 4.1 0.46 8 URI-1 P-3 Shelby 3.6 3.5 0.51 9 URI-1 P-4 Shelby 3.6 5.3 0.42 13 URI-1 P-4 Shelby 3.6 1.9 0.38 14 URI-1 P-4 Shelby 7.1 3.4 0.38 18 NA NA Reconstituted 3.6 3.8 0.36

3.3 Conclusions

The purpose of this section was to investigate the quality of the samples obtained from the field for use in the liquefaction testing program. Samples of inorganic and organic silt were recovered from two study sites using four different samplers including a: 1) Shelby tube, 2) GUS sampler, 3) fixed piston sampler, and 4) SGI sampler. The recovery of organic silts was significantly better than the recovery of inorganic silts. In most cases, the inorganic silts did not have sufficient stiffness and/or plasticity to remain intact within the sample tube during recovery.

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By field observation alone, it was concluded that undisturbed samples of inorganic silts cannot be obtained using conventional sampling equipment.

The organic silts, which had much better sample recovery, were evaluated in the laboratory to assess their level of disturbance. One-dimensional compression and triaxial tests were performed to provide a quantitative measure of sample quality. In most evaluation methods employed, the sample quality was determined to be poor. There was no clear trend indicting that one sampling method was better than any other. The results of the triaxial tests on the recovered samples were similar to the results obtained on reconstituted samples, suggesting that they are disturbed. Effects due to sample size were not apparent. Based on these findings, the samples obtained from the field in this study are not suitable for laboratory liquefaction testing.

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4.0 Background on Liquefaction Potential of Silts

This section presents a review of existing literature on the subject of liquefaction potential of silts. The purpose of the literature review is to answer two fundamental questions: 1) Do silts similar to the ones found in Rhode Island have the potential to liquefy in an earthquake?, and 2) How would liquefaction potential of these soils be evaluated according to the current standard of practice? These questions are addressed in subsequent sections. 4.1 Liquefaction Potential of Silts Evidence of liquefaction can be observed in the field during actual earthquakes or evaluated directly in the laboratory using cyclic triaxial or cyclic simple shear tests. These two aspects are discussed in detail. 4.1.1 Field Observations

Field evidence of liquefaction suggests that silts are susceptible to liquefaction in large earthquakes. In the Chile earthquake of 1965, vertical cracking was reported on the crest of Cerro Blanco dam that was constructed of mine tailings slimes (Dobry and Alvarez 1967). In the Isu-Ohima, Japan earthquake of 1978 (Magnitude, M = 7.0), a retaining dike failure was attributed to the liquefaction of tailings slimes contained behind the dike (Ishihara et al. 1980; Okusa et al. 1980). The 1989 Loma Prieta earthquake (M = 7.0) in California, liquefaction of silty clays and clayey silts was believed to be responsible for the lateral spreading and building damage that was observed following ground shaking (Boulanger et al. 1997). Most recently, observations of ground failure and building damage from the 1999 Kocaeli, Turkey earthquake (M = 7.4) were attributed to liquefaction of the underlying low plasticity silts (Martin et al. 2004; Bray et al. 2004). 4.1.2 Laboratory Investigations

Investigations by various researchers also indicate that silt soils can liquefy under cyclic loading in the laboratory. The three primary soil parameters that affect the liquefaction resistance are the void ratio (or relative density), plasticity characteristics, and the soil fabric. Void Ratio. Triaxial tests by Ishihara et al (1980), El Hosri et al. (1984) and Polito and Martin (2001) showed that liquefaction resistance increases as the void ratio decreases, and is consistent with the behavior of sands. Plasticity. Plasticity has the effect of increasing the liquefaction resistance by resisting collapse of the soil fabric during shearing. Tests by Ishihara et al (1980) and Prakash and Sandoval (1992) indicated that liquefaction resistance increases with increasing plasticity index (PI). However, there is contradictory evidence as to what values of PI silts become non liquefiable. El Hosri et al. (1984) showed that silts with a PI of less than 10% are susceptible to liquefaction. Liquefaction was not observed in samples with a PI greater than 15% (Puri et al. 1996). Yet, Ishihara et al. (1980) observed liquefaction in tailings slimes up to a PI of 28%. Prakash et al.

16

(1998) and Guo and Prakash (1999) suggested that liquefaction resistance is lowest at a critical PI value of about 4-5%. Fabric. Fabric refers to the arrangement of soil particles and can affect the cyclic strength. Cyclic triaxial tests by Singh (1994) and Puri et al. (1996) and cyclic shear tests by Clukey et al. (1985) showed that liquefaction resistance is higher in undisturbed silt samples than in samples reconstituted in the laboratory. Singh (1994) also indicated that test results are sensitive to sample preparation. These findings suggest that in-situ fabric or aging effects may play a significant role in the liquefaction resistance of silts.

4.2 Standard of Practice for the Evaluation of Liquefaction Potential of Silts In engineering practice it is often necessary to evaluate the liquefaction potential of soil at a given site. The standard of practice for evaluating liquefaction potential has been use of the “simplified procedure” developed primarily by H.B. Seed and his colleagues since the 1970’s. A conference held by the National Center for Earthquake Engineering Research (NCEER) in 2001 (Youd and Idriss 2001) provided an update on these procedures and is considered to be the current standard of practice. This section will review these methods and how they would be used to evaluate the liquefaction potential of Rhode Island silts which consists of 100% fines (see Section 2). Both the cyclic resistance and the seismic demand are addressed. 4.2.1 Cyclic Resistance of Soil

Cyclic resistance refers to the ability for a soil to resist liquefaction and is evaluated in the simplified procedure using in situ test data. Since the standard penetration test (SPT) and the cone penetration test (CPT) are most commonly used in Rhode Island, only SPT and CPT-based approaches are presented. 4.2.1.1 SPT-Based Approach

Liquefaction resistance can be evaluated from blow counts measured in the standard penetration test. The liquefaction resistance curve that would be used in such an analysis is presented in Figure 4.1. This curve was established by observing cases where liquefaction occurred (or did not) in actual earthquakes. A boring was performed at the site after the earthquake and a blow count was selected that represented the average of the weakest strata in the soil profile. The blow counts were corrected for overburden stress for energy delivered to the SPT hammer to obtain the corrected blow count, (N1)60.

At the same sites, the cyclic stress ratio was estimated from the peak ground acceleration (amax) using a simplified method given by the following equation (Seed and Idriss 1971).

dvo

vo rg

aCSR

'65.0 max

σσ

= (Eq. 4.1)

where g = acceleration due to gravity, σvo = total vertical stress, σ’vo = vertical effective stress and rd = stress reduction factor. The stress reduction factor has a value of 1.0 at the ground surface and decreases with depth.

17

0

0.1

0.2

0.3

0.4

0.5

0.6

0 10 20 30 40 50(N1)60

CR

RClean Sand

Fines Content ≥35

Liquefaction

No Liquefaction

M=7.5

Figure 4.1. SPT-based liquefaction curves (after Youd and Idriss 2001).

The CSR estimated at each site was plotted against (N1)60. A line was then drawn through

the data points separating the sites that showed evidence of liquefaction from the ones that did not. Since the resulting curve describes the resistance of the soil, the CSR is re-labeled as the cyclic resistance ratio (CRR) in Figure 4.1. The lower curve in Figure 4.1 is referred to as the “clean sand” curve and was established from the data in clean sands. It was observed in the field cases that an increase in the fines content tended to shift the curve to the left. To account for the effect of soil fines content, a correction can be applied to the measured blow counts to obtain an equivalent clean sand blow counts, (N1)60cs. For soils having greater than 35% fines the following equation can be used.

601601 )(2.15)( NN cs += (Eq. 4.2)

Alternatively, Equation (1) can be used to shift the clean-sand base curve to use directly

with the SPT measured in soils with >35% fines. This curve, which is applicable to the Rhode Island silts, is also shown in Figure 3.1 and is positioned to the left of the clean sand base curve. The liquefaction resistance curves presented in Figure 4.1 are for a magnitude (M) 7.5 earthquake. To use these curves for seismic events other than 7.5, a magnitude scaling factor (MSF) is applied to the resistance as described later in Section 4.2.1.2. Corrections for a sloping ground surface and overburden stresses can also be applied.

18

4.2.1.2 CPT-Based Approach

Liquefaction resistance can also be estimated from cone penetration test resistance. The CPT-based liquefaction resistance curve that is the standard of practice was developed by Robertson and Wride (1998) was obtained in a similar fashion as the SPT-based curve. The clean sand curve in Figure 4.2 forms the basis of the liquefaction resistance evaluation.

0

0.1

0.2

0.3

0.4

0.5

0.6

0 50 100 150 200 250(qc1N)cs

CR

R

Clean Sand

Liquefaction

No Liquefaction

M=7.5

Figure 4.2. CPT-based liquefaction resistance curves (after Robertson and Wride 1998).

To use the clean sand curve, the measured cone tip resistance must first be corrected for overburden stress and normalized to a reference pressure to obtain the corrected tip resistance (qc1N):

a

c

n

vo

aNc P

qPq

=

'1 σ (Eq. 4.3)

where Pa = Reference pressure (100kPa), σ’vo = vertical effective stress, qc = measured cone tip resistance, and n = exponent depending on soil type. Values of n range from 0.5 for sands to 1.0 for clays. A value between 0.5 and 1.0 would be appropriate for silts and sandy silts. As a first step, the cone data must be used to differentiate sands and silts from clays. Clays are known to be resistant to liquefaction and the analysis will be complete if these soils are encountered. This is accomplished by calculating a soil behavior type index (Ic) using an n =1.0:

19

( ) ( )[ ] 5.022 log22.1log47.3 FQIc ++−= (Eq. 4.4)

where,

a

voc

n

vo

a

PqP

Q)(

'σ

σ−

= (Eq. 4.5)

%100)(×

−=

voc

s

qf

Fσ

(Eq. 4.6)

and σvo = total vertical stress, and fs = CPT sleeve resistance. If the Ic calculated with n = 1 is greater than 2.6, the soil is clayey and is resistant to liquefaction. However, it is recommended that the samples of the soils be recovered and evaluated using other criterion. The Chinese criterion was recommended in Youd and Idriss (2001), however, recent criteria proposed by Bray et al. (2004b) are considered to be more reasonable. If the calculated Ic < 2.6, the soil is likely more granular in nature. The Ic should be recalculated using Equation (4) using a n value of 0.5. If the Ic < 2.6 the soil is nonplastic and granular and an n value of 0.5 should also be used in the liquefaction analysis. If the Ic > 2.6 the soil is likely very silty and plastic and a n value of 0.7 should be adopted for the liquefaction analysis. It was recommended that any soil having a Ic > 2.6 be evaluated using other liquefaction criteria (i.e. Bray et al. 2004b). To use the curve in Figure 4.2, the corrected tip resistance should be converted to an equivalent clean sand value (qc1N)cs using the following expression

NcccsNc qKq 11 )( = (Eq. 4.7)

where Kc is a correction factor that is a function of the soil index given by the following equations.

For 64.1≤cI , 0.1=cK (Eq. 4.8)

For 64.1>cI , 88.1775.3363.21581.5403.0 234 −+−+−= ccccc IIIIK (Eq. 4.9)

Like the SPT-based liquefaction resistance curve, the CPT-based curve in Figure 4.2 is only applicable to a magnitude 7.5 earthquake. Magnitude scaling factors should be applied to adjust these curves for other magnitude events. 4.2.1.2 Magnitude Scaling Factors

The SPT and CPT-based liquefaction resistance curves presented above are only applicable to an earthquake having a moment magnitude of 7.5. To use these curves for earthquakes having a magnitude other than 7.5, a magnitude scaling factor (MSF) is used where

20

MSFCRRCRR M ⋅= = 5.7 (Eq. 4.10)

The MSF increases the liquefaction resistance for smaller events and decreases the resistance for larger events. Numerous magnitude scaling factors have been proposed. However, based on the NCEER workshop participants (Youd and Idriss 2001), the range of factors reproduced in Figure 4.3 is recommended for use in engineering practice.

0

0.5

1

1.5

2

2.5

3

5 5.5 6 6.5 7 7.5 8Earthquake Magnitude, M

Mag

nitu

de S

calin

g Fa

ctor

, MS

F

Figure 4.3. Magnitude scaling factors recommended by NCEER for use

in engineering practice (after Youd and Idriss 2001). 4.2.2 Seismic Demand

Seismic demand in the simplified approach is characterized by an earthquake magnitude and a shear stress level imposed within the soil deposit. The earthquake magnitude is already accounted for by scaling the resistance curves using the magnitude scaling factors discussed previously. Therefore, the remaining parameter that must be evaluated is the shear stress caused by an earthquake of interest. Shear stresses can be evaluated in one of two ways.

First, the cyclic stress ratio (CSR) can be evaluated from an anticipated peak horizontal ground acceleration using Equation (4.1). Alternatively, a site-specific response analysis can be performed using computer programs such as SHAKE that compute more accurate estimates of shear stress within the soil profile. In a site response analysis, a soil profile is input along with

21

stiffness and damping properties of the soil. An earthquake time history is applied to the base of the soil profile and the maximum cyclic stress ratio (CSRpeak) can be computed directly at depth. The CSR used in the simplified approach would be calculated as 0.65 · CSRpeak.

22

5.0 Laboratory Study of Liquefaction Resistance This chapter presents the results of the laboratory investigation to evaluate the liquefaction resistance of Rhode Island silts. Since high quality undisturbed samples could not be obtained in the field, laboratory test were preformed on reconstituted samples. The test method, sample preparation procedures, and results are presented. Some preliminary estimates of earthquake demand for a site in Providence, RI are also made to provide a relative comparison to the measured liquefaction resistance. 5.1 Soils Tested The soils tested in the triaxial program consisted of organic silt and inorganic silt. The organic silt was obtained offshore Quonset Point as part of a project on beneficial uses of dredge material (Baxter et al. 2004). All recovered samples of organic silt were combined and rinsed to remove salt. The plasticity index (PI) at the time of testing was 11 and the grain size curve is shown in Figure 5.1. The inorganic silts were obtained from a construction site in Cranston, RI located just south of the Wellington Avenue railroad bridge. Silt was recovered from auger spoils that had been removed during installation of drilled shafts at the site. The silt in non-plastic and the grain size curve is also shown in Figure 5.1.

0

20

40

60

80

100

0.0001 0.001 0.01 0.1 1Particle Diameter (mm)

Per

cent

Fin

er

Inorganic SiltOrganic Silt

SandSiltClay

Figure 5.1. Grain size distribution of organic and inorganic silts used in this study.

23

5.2 Test Method

Cyclic triaxial tests were performed on 7.2-cm diameter samples using Geocomp Loadtrac/II equipment, shown in Figure 5.2. To simulate level ground conditions samples were isotropically consolidated to 100 kPa. To ensure saturation a minimum backpressure of 500 kPa and a B-parameter of at least 0.95 was required. A sinusoidal cyclic shear stress was applied at a frequency of 1 Hz and whose maximum amplitude was equal in both compression and tension.

Figure 5.2. Photograph of the cyclic triaxial apparatus used in this study.

5.3 Sample Preparation Methods

Sample preparation methods involving placement from a slurry are preferred in laboratory liquefaction studies because they limit particle segregation and produce a fabric that is most representative of alluvial soils (Kuerbis and Vaid 1998). In our study very loose samples of silt could be prepared from a slurry, however, vibration was ineffective in uniformly densifying the silts. Therefore, to investigate the behavior of silts at higher densities (lower void ratios) the moist tamping method was also used and both procedures are described below.

24

5.3.1 Slurry Consolidation To prepare the samples for testing a slurry was prepared by mixing dry soil and water to a

water content of about 45%. The slurry was poured into a two-piece sample mold that was attached to the bottom platen of the triaxial cell with a membrane stretched to the sides of the mold. A longer mold was constructed to accommodate the additional bulking of the slurry. The slurry was tamped to remove air bubbles and consolidated one dimensionally inside the mold under a vertical stress of 40 kPa. After consolidation the membrane was placed on the top cap and a 20 kPa suction was applied to the sample to maintain the effective stress so that the mold could be removed.

The slurry method was very effective in producing loose uniform samples. The sample uniformity is illustrated in Figure 5.5(a) which shows a profile of dry density of the inorganic silt measured along the length of the triaxial sample at 1 mm increments. The high resolution of the density was obtained using a gamma-density logger which is commonly used in offshore exploration to obtain non-destructive properties of sediment core samples.

1.4 1.5 1.6

0

2

4

6

8

10

12

14

Dis

tanc

e A

long

Sam

ple

(cm

)

1.4 1.5 1.6 1.4 1.5 1.6 1.4 1.5 1.6

(a) (b) (c) (d) Figure 5.3. Density profile (in g/cc) of samples prepared using a slurry method (a) and moist

tamping method with 0, 1, and 2 percent undercompaction (b-d).

5.3.2 Moist Tamping

Moist tamping involves compacting moist soil in multiple layers. Though the method

was originally intended to represent compacted soils, it was needed to prepare higher density specimens. To provide easier sample preparation and to quantify the applied tamping effort, a

25

modification to the undercompaction method (Ladd 1978) was used. In Ladd’s method bottom layers are initially undercompacted such that the compaction of subsequent layers results in a uniform density sample. In our method, rather than tamping each layer to a specific height, each layer was compacted to specified tamping effort using a drop-hammer, steel rod, and acrylic piston. A sample is prepared within a standard split mold using 8 layers, each layer tamped 25 times with the drop hammer. The basis for establishing the energy required to compact each layer is shown in Figure 5.4. This figure which shows average dry density verses the logarithm of drop height (tamping effort) was obtained by compacting samples using the same drop height for all layers and then measuring the final density of the sample. Each curve on the figure represents a sample of constant initial molding water content namely 5% and at the optimum water content of about 18% determined in a Standard Proctor Test.

1.20

1.30

1.40

1.50

1.60

1 10 100Drop Height (cm)

Dry

Den

sity

(g/c

c)

w m = 5%w m = 18%

Figure 5.4. Dry density verses compaction energy for inorganic silts at varying molding water

contents. As shown by Figure 5.4, preparing the sample at a water content of 18% requires less tamping effort to achieve a given sample density. A log-linear function provided a good fit to the measured data in the form

21 CHC +=ρ (Eq. 5.1)

26

where ρ = average dry density, and H = hammer drop height, and C1, C2 = constants particular to the soil type. Using Equation (1) it is assumed that the initial density of each n-layer (ρn) is a function of the hammer drop height for each n-layer (Hn) by the following

−=

1

2expC

CH nn

ρ (Eq. 5.2)

Given a target density ( tρ ) the initial density of each n-layer (ρn) is increased from the bottom of the sample to its top in a linear fashion by the following

12)1()1(−

−+−=N

n ttn

µρµρρ (Eq. 5.3)

where µ = percent undercompaction, and N = total number of layers. The total tamping energy (Et) required to prepare the sample is obtained from the potential energy.

∑⋅= nbht HNWE (Eq. 5.4)

where Wh = weight of hammer, Nb = number of blows per layer (= 25). Typical profiles of dry density obtained with the gamma-density logger for the undercompacted samples are shown in Figure 5.3. The samples shown in the figure have been prepared to initial undercompaction levels of 0, 1%, and 2%. Lines are also shown on the density profiles to show the average trend. It is clear from Figure 4.3 that the density is variable within each layer having a maximum of about ±0.05 g/cc relative to the local average. Though the trends in average density tends are subtle, the sample with zero undercompaction (Figure 5.3b) shows a higher density at the bottom of the sample than the top, and the sample with 2% undercompaction (Figure 5.3d) has a lower density at the bottom that the top. The sample compacted at 1% (Figure 5.3c), however, has a uniform average density and showed uniform bulging at failure in the triaxial tests. 5.4 Liquefaction Resistance This section presents the results of the cyclic triaxial tests performed on samples of reconstituted organic and inorganic silts. A preliminary estimate of the demand from a typical design level earthquake in Providence, RI is also presented for a relative comparison to the cyclic resistance obtained in the laboratory. 5.4.1 Organic silts

Typical results obtained in the cyclic triaxial tests are shown in Figure 5.5. The top of the figure indicates the applied cyclic stress ratio verses the number of cycles. Also shown are the axial strain and the pore pressure ratio (Ru), defined as the ratio of the excess pore pressure to the initial effective confining stress. Liquefaction by definition occurs when Ru = 1 or when the effective stress is zero. However, as shown in Figure 5.5, significant axial strain was observed even though Ru never reached unity. This behavior is likely attributed to non-uniform pore

27

pressure distribution within the sample that has been observed in tests on silt samples (Konrad and Wagg 1991; Zhou et al. 1995). To avoid this problem, failure was defined in this study at 1% and 5% double amplitude (DA) axial strain.

-0.5

0

0.5

CS

R

-20

0

20

Axi

al S

train

(%)

0 2 4 6 8 10 12 14 16 18 20-0.5

0

0.5

1

Number of Cycles

Ru

Figure 5.5. Typical cyclic triaxial test results of samples prepared by slurry consolidation.

The results of multiple tests on identical samples were combined to obtain a liquefaction

resistance curve. This curve relates the cyclic resistance ratio to the number of cycles to cause liquefaction. The curve in Figure 5.6 is based on 5 identical samples of organic silt prepared from a slurry and sheared at stress levels. As shown on the figure, the CRR decreases with an increase in the number of cycles. The choice of the failure criterion (1% DA or 5% DA strain) had little effect on the resulting curve. Figure 5.6 represents the liquefaction resistance of a young, loose alluvial organic silt deposit. Densification, prior stress history, and aging have the effect of increasing the liquefaction resistance, and therefore, the curve is anticipated to be a lower bound of cyclic strength.

28

0.0

0.1

0.2

0.3

0.4

0.5

1 10 100Number of Cycles

CR

R1% DA Strain5% DA Strain

Organic Silt- Quonsett Pointe = 0.72 to 0.74

Figure 5.6. Liquefaction resistance of organic silts prepared by slurry consolidation.

5.4.2 Inorganic Silts

The cyclic resistance of the inorganic silts prepared from a slurry at constant void ratio are given in Figure 5.7. Relative to the organic silts, the cyclic strength of the inorganic silts was significantly lower. Like the organic silts, this curve represents the liquefaction resistance of a young, loose alluvial silt deposit and is also considered to be a lower bound strength. To investigate the effect of soil density, samples were prepared using the moist tamping method. It was identified in the literature that for sands, moulding water content has an effect on the cyclic resistance of sands (Ladd 1977). Therefore, to quantify this effect, two cyclic resistance curves were generated at about the same void ratio; one curve was generated from samples prepared at a molding water content of 5% and a second was prepared near the optimum water content of 17%. As expected, a lower tamping energy was required to prepare the samples at their optimum water content because the water acts as a “lubricant” between soil grains. The results, shown in Figure 5.8, indicate that the liquefaction resistance of the sample prepared at the optimum water content is also lower by about 15%. As discussed earlier, it is anticipated that the slurry method provides a more representative fabric to the in situ condition. Therefore, the moist tamping results were compared to the slurry results at about the same void ratio, as shown in Figure 5.9. Figure 5.9 suggests that if the samples are prepared at their optimum water content, then the cyclic strength is comparable to the samples prepared from a slurry.

29

0.0

0.1

0.2

0.3

0.4

0.5

1 10 100Number of Cycles

CR

R

1% DA Strain

5% DA Strain

Inorganic Silt- Wellington Ave.e = 0.84 to 0.88

Figure 5.7. Liquefaction resistance of inorganic silts prepared by slurry consolidation.

0.0

0.1

0.2

0.3

0.4

0.5

1 10 100Number of Cycles

CR

R w m =5%; E t =193 N-m; e=0.87-0.90

w m =18%; E t =109 N-m; e=0.88-0.90

Inorganic Silt- Wellington Ave.1% DA Strain

Figure 5.8. Liquefaction resistance of inorganic silts prepared by moist tamping at

approximately constant void ratio at two different molding water contents and tamping efforts.

30

0.0

0.1

0.2

0.3

0.4

0.5

1 10 100Number of Cycles

CR

RMoist TampingSlurry Consolidation

w m = 5%

w m = 18%

Inorganic Silt- Wellington Ave.1% DA Strain

Figure 5.9. Comparison of liquefaction resistance of samples prepared by slurry consolidation

and moist tamping methods at approximately constant void ratio.

0.0

0.1

0.2

0.3

0.4

0.5

1 10 100Number of Cycles

CR

R

e = 0.88-0.90

e = 0.82

Inorganic Silt- Wellington Ave.1% DA Strain

Figure 5.10. Liquefaction resistance of inorganic silts at two different densities prepared by

moist tamping at the optimum water content.

31

Higher density samples were prepared by tamping at the optimum water content using greater tamping effort. The effect of density is shown in Figure 5.10, which shows the cyclic resistance obtained for the denser sample is higher than for the loose samples. The increase in liquefaction resistance with decreasing void ratio is therefore consistent with the findings of Ishihara et al. (1980), and Polito and Martin (2001). 5.5 Preliminary Evaluation of Earthquake Demand

Given the liquefaction resistance of the silts, it is difficult to evaluate the significance of the laboratory results without context to the seismic demand. Seismic demand of a given earthquake is quantified by the amplitude and the duration of ground shaking. The moment magnitude describes the duration of shaking and is analogous to the number of cycles obtained in a triaxial test. The cyclic stress ratio is proportional to the amplitude of shaking.

A seismic study recently performed by Weston Geophysical for the Washington Bridge project was used to estimate a range of anticipated seismic demand. The seismic demand is greatly dependent upon the soil types and the characteristics of a particular site, thus these data are used only for illustrative purposes. The seismic parameters that were extracted from the seismic study (Weston 2000) for two design level seismic events are summarized in Table 5.1. The smaller event represents an earthquake with a return interval of about 500 years and the larger event has a return interval of about 2,500 years. In Rhode Island the 500-year event is typically used in the design of railroad bridge and trestle structures, bridge approach structures, retaining walls, and abutments (Weston 2000). The seismic parameters include the anticipated moment magnitude (M), peak ground acceleration (PGA) in bedrock, and a range of average cyclic shear stress ratio (CSRavg) estimated from 1-D site response analysis using synthetic bedrock ground motions. Profiles of average CSR computed at depth from the 1-D response analyses are summarized for both design level events in Figure 5.11.

To compare the design earthquake data to the triaxial test results the earthquake magnitude must first be converted to an equivalent number of uniform stress cycles. This was accomplished using Seed et al. (1975) which estimates an average of 4 cycles for the M = 5 event and 4½ cycles for the M = 6 event. Also, to obtain the field strength from the triaxial data, the test data were corrected for multi-directional shaking and shear stress orientation using the following equation (Kramer 1996):

TriaxrField CSRCCRR ⋅= 9.0 (Eq. 5.5)

where Cr = correction factor that was assumed to be 0.69 for normally consolidated soils (Castro 1975).

Table 5.1. Summary of seismic parameters obtained from the site response analysis performed

for the Washington Bridge by Weston Geophysical (Weston 2000). Average CSREQ estimated at depth Return

Interval (yr) Magnitude PGA (g) Lower Bound Upper Bound 500 5.0±1/4 0.11 0.03 0.20

2,500 6.0±1/4 0.19 0.07 0.35

32

0

5

10

15

20

25

30

35

40

45

0.0 0.1 0.2CSRavg

Dep

th (m

)

500-yr EarthquakePGA=0.06g

0

5

10

15

20

25

30

35

40

45

0.0 0.1 0.2 0.3 0.4CSRavg

Dep

th (m

)2,500-yr EarthquakePGA=0.19g

(a) (b)

Figure 5.11. Profiles of average CSR with depth estimated from 1-D site response analysis using four different soil profiles for a (a) 500-year seismic event and a (b) 2,500-year event

(after Weston 2000).

A comparison of earthquake demand and soil cyclic resistance is given in Figure 5.12. The liquefaction resistance curves shown for both the organic and inorganic silts represent a wost-case scenario since these samples are in their loosest condition. It is important to note that the soil’s liquefaction resistance will increase with increasing relative density and aging effects. As shown in Figure 5.11, both the 500-year and 2,500-year events have stress levels in some cases that exceed the strength of the soil, suggesting the potential for liquefaction.

33

0.0

0.1

0.2

0.3

0.4

0.5

1 10 100Number of Cycles

CR

R Fi

eld (C

SR

EQ)

CSREQ (M=6)

CSREQ (M=5)Loose Inorganic Silt

Loose Organic Silt

Figure 5.12. Range of average CSRs estimated for the Washington Bridge (Weston 2000)

compared to the CRRs of loose organic and inorganic silt determined in the laboratory. Two design levels are shown including a 500-year event (M = 5) and a 2,500-year event (M = 6).

5.6 Summary and Conclusions This section presented a laboratory study of liquefaction resistance of Rhode Island silts. Reconstituted samples of inorganic and organic silts were tested in the cyclic triaxial apparatus. Sample preparation using slurry consolidation was preferred because it was most representative of young alluvial soils. However, vibration was ineffective in densifying the slurried samples and only loose specimens could be produced. To test samples at higher densities the moist tamping method was employed. A comparison of the sample preparation methods indicated that if the samples were tamped at their optimum water content as determined in a standard Proctor test, than the cyclic strength was comparable to the strength measured in the slurried samples at the same void ratio. Liquefaction resistance curves were obtained for both the organic and inorganic silts. These curves were compared to seismic demand parameters obtained from a study performed at the Washington Bridge by Weston Geophysical. For a 500-year seismic event, which is the design level for most bridge structures in Rhode Island, the loose inorganic silt appeared to be the most susceptible to liquefaction. However, increasing the density of the silt increased its liquefaction resistance.

34

6.0 Conclusions This report presented a laboratory study to evaluate the liquefaction resistance of silts commonly encountered in Rhode Island. In Section 2, a field study was presented that evaluated the sample disturbance of soils recovered from two different sites. This study indicated that it was very difficult to recover loose samples of inorganic silts which tended to flow out of the sample tube during recovery. Samples of organic silts were successfully recovered from the field, but after detailed evaluation of sample disturbance using consolidation tests, the organic silts were found to be of poor quality and not suitable for liquefaction testing. In Section 3, a review of the geotechnical literature confirms that both non-plastic and plastic silts are susceptible to liquefaction from cyclic loading. This was supported by evidence of ground failure and building damage in the field during actual earthquakes as well as from laboratory studies conducted on silty soils. A review of the standard of practice for evaluation of liquefaction potential was also presented. These “simplified” methods which are based on in situ tests such as the SPT and CPT are largely based on data obtained in sands. The uncertainty in these methods stresses the need to find more reliable methods of evaluating liquefaction potential in silts. Finally, in Section 4, a laboratory investigation was performed to quantify the liquefaction resistance of both organic and inorganic silts. The test specimens were reconstituted in a manner to best represent the in situ condition of a young alluvial soil. Because of this, the liquefaction resistance curves obtained in the laboratory are anticipated to be lower bound estimates of strength, as other factors such as prior stress history and aging are known to increase liquefaction resistance. To gain perspective on the significance of the results, a range of seismic demand was compared to the cyclic resistance of the silts. The seismic demand parameters were taken from a seismic design study that was performed in 2000 for the Washington Bridge in Providence, RI. Though the demand will depend on the particular site characteristics, it provides a preliminary evaluation of liquefaction potential. This comparison suggested that the 500-year seismic event, which is typically used in the design of bridge structures, has sufficient stresses to liquefy the loose inorganic silt. However, the liquefaction resistance increases as the density of the soil increases. Therefore, future work should focus on using in-situ test methods that capture both the in situ fabric and density such that the evaluation is representative of actual site conditions. 7.0 Acknowledgements This study was made possible by a grant from the University of Rhode Island Transportation Center and the Rhode Island Department of Transportation. Special thanks to Mathew Page for the evaluation of sample quality. Drilling and sampling for this study was performed by Guild Drilling and Geologic Drilling. Their help is greatly appreciated.

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8.0 References Acker (2003). Drilling Equipment Catalog. Acker Drill Company, Scranton, PA. Andresen, A. A., and Kolstad, P. "The NGI 54-mm Samplers for Undisturbed Sampling of Clays

and Representative Sampling of Coarser Materials." International Symposium on Soil Sampling, Singapore, 13-21.

Baxter, C.D.P., King, J.W., Silva, A.J., Page, M.J., Calabretta, V.V. (2004) “Site

Characterization of Dredged Sediments and Evaluation of Beneficial Uses”. Geotechnical Special Publication No. 127, ASCE, 150-161.

Boulanger, R.W., Mejia, L.H., and Idriss, I.M. (1997). “Liquefaction at Moss Landing During

Loma Prieta Earthquake.” Journal of Geotechnical and Geoenvironmental Engineering, 123(5), 453-467.

Boulanger, R.W., Meyers, M.W., Mejia, L.H., Idriss, I.M. (1998). Behavior of a Fine-grained

Soil During the Loma Prieta Earthquake.” Canadian Geotechnical Journal, v 35, n 1, p 146-158.

Bray, J.D., Sancio, R.B., Durgounoglu, T., Onalp, A., Youd, T.L., Stewart, J.P., Seeed, R.B.,

Cetin, O.K., Bol, E., Baturay, M.B., Christensen, C., and Karadayilar, T. (2004). “Subsurface Characterization at Ground Failure Sites in Adapazari, Turkey”. Journal of Geotechnical and Geoenvironmental Engineering, 130(7), 673-685.

Bray, J.D., Sancio, R.B., Riemer, M.F., and Durgunoglu, T. (2004b) “Liquefaction Susceptibility

of Fine-Grained Soils”, Proceedings of the 11th International Conference on Soil Dynamics and Earthquake Engineering and 3rd International Conference on Earthquake Engineering, V. 1, 655-662.

Castro, G. (1975) “Liquefaction and Cyclic Mobility of Sands”, Journal of the Geotechnical

Engineering Division, ASCE, 101(GT6), 551-569. Clukey E.C., Kulhawy, F.H., and Winters, W.J. (1985). “Wave Induced Liquefaction and

Scouring in Yukon Prodelta Silts.” Civil Engineering in the Arctic Offshore, Proceedings of the Conference Arctic ’85, ASCE, 838-850.

D'Andrea, R. (1966). In-situ Shear Strength Characteristics of a Saturated Silt, M.S. Thesis,

Department of Civil Engineering, University of Rhode Island. DMJM Harris. (2002). Geotechnical Design Summary Report for the Approaches to the

Wellington Avenue Railroad Bridge. Prepared for Rhode Island Department of Transportation, Warwick-Cranston Rhode Island

Dobry, R. and Alvarez, L. (1967). “ Seismic Failures of Chilean Tailings Dams.” Journal of the

Soil Mechanics and Foundations Division, ASCE, 93(SM6), 237-260.

36

El Hosri, M.S., Biarez, J., Hicher, P.Y. (1984) “Liquefaction Characteristics of Silty Clay”

Proceedings of the 8th World Conference on Earthquake Engineering, 277-284. Garga, V.K. and McKay, L.D. (1984). “Cyclic Triaxial Strength of Mine Tailings”. Journal of

Geotechnical Engineering, ASCE, 110(8), 1091-1105. Guo, T. and Prakash, S. (1999). “Liquefaction of Silts and Silt-Clay Mixtures”. Journal of

Geotechnical and Geoenvironmental Engineering, ASCE, 125(8), 706-710. GZA (2002). Final Geotechnical Interpretative Report Improvements to Route I-195 Providence

River Bridge East Abutment, South Main Street Bridge, and Associated Retaining Walls. Providence, RI

GZA (1998). Geotechnical Report Improved Access from Route 95. Report 31945, Providence,

RI Ishihara, K., Troncoso, J. Kawase, Y., and Takahashi, Y. (1980). “Cyclic Strength

Characteristics of Tailings Materials”, Soils and Foundations, Japanese Society of Soil Mechanics and Foundation Engineering, 15(1), 127-142.

Konrad. J.-M. (1985). “Undrained Cyclic Behavior of Beaufort Sea Silt.” Civil Engineering in

the Arctic Offshore, Proceedings of the Conference Arctic '85, ASCE, 830-837. Konrad, J.-M., and Wagg, B.T. (1991). “Undrained Cyclic Loading of Anisotropically

Consolidated Clayey Silts.” Journal of Geotechnical Engineering, ASCE, 119(5), 929-947.

Kuerbs, R. and Vaid, Y.P. (1988). “Sand sample preparation- the slurry deposition method”. Soils and Foundations, JSSMFE, 28(4), 107-118.

Ladd, R.S. (1977). “Specimen preparation and cyclic stability of sands”. Journal of the

Geotechnical Engineering Division, ASCE, 103(GT6), 535-547. Ladd, R.S. (1978). “Preparing test specimens using undercompaction”, ASTM Geotechnical

Testing Journal, 1(1), 16-23.

Lunne, T., Berre, T., and Strandvik, S. (1997). "Sample Disturbance Effects in Soft Low

Plasticity Norwegian Clay." Proceedings of Conference on Recent Developments in Soil and Pavement Techniques, Rio de Janeiro, 81-102.

Martin, J.R., Olgun, C.G., Mitchell, J.K., and Durgunoglu, H.T. (2004). “High Modulus

Columns for Liquefaction Mitigation.” Journal of Geotechnical and Geoenvironmental Engineering, ASCE, 130(6), 561-571.

Nacci, D. (1980). Geotechnical Properties of Organic Silt, M.S. Thesis, Department of Civil

Engineering, University of Rhode Island.

37

Okusa, S., Anma, S., Maikuma, H. (1980). “Liquefaction of Mine Tailings in the 1978 Izu-Oshima-Kinkai Earthquake, Central Japan.” Proceedings of the 7th World Conference on Earthquake Engineering, v 3, 89-96.

Page, M.J. (2004) Sample Disturbance and Strength of Silts. M.S. Thesis, Department of Civil

Engineering, University of Rhode Island. Polito, C.P. and Martin, J.R. (2001). “Effects of Nonplastic Fines on the Liquefaction Resistance

of Sands”. Journal of Geotechnical and Geoenvironmental Engineering, ASCE, 127(5), 408-415.

Prakash, S. and Sandoval, J.A. (1992). “Liquefaction of Low Plasticity Silts.” Soil Dynamics and

Earthquake Engineering, vol. 11, 373-379. Prakash, S., Guo, T., and Kumar, S. (1998). “Liquefaction of Silts and Silt-Clay Mixtures.”

ASCE Geotechnical Special Publication No. 75, Geotechnical Earthquake Engineering and Soil Dynamics III, 337-348.

Prakash, S. and Puri, V.K. (1982). “Liquefaction of Loess Soil”, Third International Earthquake

Microzonation Conference, Seattle, WA, Vol. II, 1101-1107. Puri, V.K., Das, B.M., and Prakash, S. (1996) “Liquefaction of Silty Soils”. International

Journal of Offshore and Polar Engineering, ISOPE, 6(4), 308-312. Robertson, P.K., and Wride C.E. (1998) “Evaluating Cyclic Liquefaction Potential Using the

Cone Penetration Tests”, Canadian Geotechnical Journal, 35, 442-459. Santagata, M. C., and Germaine, J. T. (2002). "Sample Disturbance Effects in Normally

Consolidated Clays." Journal of Geotechnical and Geoenvironmental Engineering, ASCE, 128(12), 997-1006.

Schmertmann, J. H. (1955). "The Undisturbed Consolidation Behavior of Clay." Transactions

ASCE, 1201. Seed H.B. and Idriss, T.L. (1971). “Simplified Procedure for Evaluating Soil Liquefaction

Potential.” Journal of Geotechnical and Geoenvironmental Engineering, ASCE, 97(9), 1249-1273.

Seed, H.B., Idriss, I.M., Makdisi, F.I., and Banerjee, N. (1975). Representation of Irregular

Stress Time Histories by Equivalent Uniform Stress Series in Liquefaction Analyses, EERC 75-29, Earthquake Engineering Research Center, University of California, Berkeley.

Silva, A. J. (1974). “Marine Geomechanics: Overviews and Projections.” Deep Sea Sediments,

A. L. Inderbitzen, ed., 45-76.

38

Silva, A. J., Baxter, C. D. P., and Calabretta, V. (2003). Beneficial Uses of Dredge Material from QDP Intermodal Port Terminal. Report Prepared for the URI Transportation Center, Kingston, RI.

Singh, S. (1994). “Liquefaction Characteristics of Silts”. ASCE Geotechnical Special Publication

No. 44, Ground Failures Under Seismic Conditions, 105-116. Tanaka, H., and Tanaka, M. (1999). "Key Factors Governing Sample Quality." Japan. 25 Terzaghi, K., Peck, R. B., and Mesri, G. (1996). Soil Mechanics in Engineering Practice, John

Wiley & Sons, INC, New York Weston (2000). Seismic Design Recommendation for the Washington Bridge No. 200,

Providence, RI, Report by Weston Geophysical Corp., dated August 14, 2000. Youd, T.L., and Idriss, I.M. (2001). “Liquefaction Resistance of Soils: Summary Report from the

1996 NCEER and 1998 NCEER/NSF Workshops on Evaluation of Liquefaction Resistance of Soils”. Journal of Geotechnical and Geoenvironmental Engineering, 127(4) 297-313.

Zhou, J., Lee, W., and Zhou, K. (1995). “Dynamic Properties and Liquefaction Potential of

Silts”. Proceedings of 1995 International Conference on Earthquake Geotechnical Engineering, v 2, 833.