Technical Note

Full-Scale Testing of Open-Ended Steel Pipe Piles inThick Varved Clayey Silt Deposits along the

Delaware River in New JerseyYong Tan, Ph.D., A.M.ASCE1; and Guoming Lin, Ph.D., P.E., F.ASCE2

Abstract: Although many studies have been done to investigate the axial behaviors of open-ended piles in sands, few studies have beenreported for weak clayey silts. To develop reliable models for the design of open-ended steel-pipe piles driven into 29-m-thick varved clayeysilt deposits, a series of full-scale field load tests including large-strain dynamic tests and static cyclic axial-compression-load tests was con-ducted on two groups of instrumented piles. Through analysis of the test data, soil parameters were back-calculated for estimation of pile capaci-ties using the static-bearing-capacity formulas and cone-resistance-basedmethods. The comparisons between the calculated results and the fieldload test data demonstrated that the following considerations can be adopted in the design of static compression capacities of an open-ended pipepile penetrating through thick varved clayey silts to end-bearing in dense cohesionless soils: (1) a fully plugged condition can be assumed, (2)cone resistance with an upper limit of 4,788 kPa (100 ksf) can be used for unit base resistance on the soil plug, and (3) exterior unit shaft re-sistance can be estimated using two-thirds of the total unit shaft resistance.DOI: 10.1061/(ASCE)GT.1943-5606.0000777. © 2013 AmericanSociety of Civil Engineers.

CE Database subject headings: Pipe piles; Steel piles; Clays; Silts; Compression; Load tests; Design; New Jersey; Full-scale tests.

Author keywords: Open-ended pipe pile; Varved clayey silt; Static cyclic axial-compression-load test; Soil parameters; Designconsiderations.

Introduction

Both open-ended and close-ended steel-pipe piles are widely used inoffshore and harbor projects. Owing to the possible formation ofa “soil plug” inside an open-ended pipe pile during pile driving, thebehavior of open-ended piles is more complicated than that of close-ended piles. Via experiments, many researchers have investi-gated the load capacities of open-ended pipe piles in sands (e.g.,Paikowsky and Whitman 1990; O’Neill and Raines 1991; Paik andSalgado 2003). Raines et al. (1992) observed that a pipe pile drivenin the fully plugged mode behaved like a close-ended pile.Paikowsky and Whitman (1990) pointed out that plugging is pos-sible once the ratioRpd of pile penetration to pile diameter exceeds 20in dense sand. McVay et al. (2004) suggested that an open-endedpipe pile in sands might still behave like a plugged pile even if thepile were unplugged during pile driving. In contrast, only a fewstudies (e.g., Bozozuk et al. 1979) were devoted to open-ended pilesin clayey silt deposits. Considering that open-ended pipe piles arewidely used in practice, it is worthwhile to study the behavior ofopen-ended pipe piles driven into thick compressible clayey siltformations.

This study investigated the axial behavior of open-ended steel-pipe piles driven into 29-m-thick varved clayey silt deposits alongthe Delaware River in New Jersey. It was achieved through a seriesof full-scale field load tests including large-strain dynamic tests andstatic cyclic axial-compression-load tests on two groups of test piles.Based on the load tests and back-analysis, considerations for bothpile tip and shaft resistances were recommended, and soil parame-ters were derived for site-specific design of the axial pile capacitiesusing static bearing-capacity formulas and cone-resistance-basedmethods.

Site and Subsurface Description

The project involved planning of three liquefied natural gas (LNG)storage tanks (74.4 m in inner radius) with a capacity of 150,000 m3

on the southeast bank of the Delaware River in Gloucester County,New Jersey. Prior to construction, the subsurface conditions at thesite were explored by a series of field and laboratory tests, includingstandard penetration test (SPT) borings, piezocone penetration test(CPTu) soundings, field vane shear tests, and unconsolidated un-drained (UU) triaxial tests. As shown in Fig. 1, the subsurface soils atthe site can be stratified into six distinct soil formations. The top 1.5-m-thick soils (Layer 1) were fills dredged from the Delaware Riverand placed on site in the 1960s. The soils underlying the fill werepredominantly loose to medium-dense sands and gravels (Layer 2)to a depth of 3 m below the ground surface (BGS). Directly belowLayer 2, a thick layer of very soft to firm varved clayey silts (Layer 3)extended to a depth of 32 m BGS. The varved clayey silts consistingof visible laminations were recent deposits of fine-grained soilsalong the shores of the Delaware River as a result of warmertemperatures and the rise of ocean levels after the glacial period.Below Layer 3, medium-dense to very dense sands and gravels

1Associate Professor, Dept. of Geotechnical Engineering, TongjiUniv., 1239 Siping Rd., Shanghai 200092, P.R. China; formerly, ProjectEngineer, WPC, Inc., Savannah, GA 31404 (corresponding author).E-mail: tanyong21th@tongji.edu.cn

2Senior Principal, Terracon Consultants, Inc., 2201 Rowland Ave.,Savannah, GA 31404. E-mail: glin@terracon.com

Note. This manuscript was submitted on August 19, 2011; approved onMay 14, 2012; published online on May 17, 2012. Discussion period openuntil August 1, 2013; separate discussions must be submitted for individualpapers. This technical note is part of the Journal of Geotechnical andGeoenvironmental Engineering, Vol. 139, No. 3, March 1, 2013. ©ASCE,ISSN 1090-0241/2013/3-518–524/$25.00.

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(Layer 4) were encountered to a depth of 37.8 m BGS, followedby very dense residual clayey sands (Layer 5) to a depth of 48 mBGS. Below Layer 5, there was metamorphic rock (Layer 6). Thegroundwater level was at depths of 0.6 to 1.8 m BGS, which fluc-tuated with the tides of the Atlantic Ocean.

Pile Load Test Program

Owing to the presence of the thick compressible varved clayey siltdeposits, the proposed LNG tanks were designed to be supported byopen-ended steel-pipe piles 508 mm in diameter and 12.7 mm in wallthickness. Prior to the mass production of the piles, a pile load testprogram including large-strain dynamic tests and static cyclic axial-compression-load tests was undertaken to obtain data for foundationdesign and construction planning. The pile test program incorporateda total of 10 piles driven in two groups at the proposed locations ofTanks 1 and 3 (see Fig. 2). Each test-pile group consisted of five piles,which included one primary test pile at the center with four reactionpiles surrounding it. The twoprimary test pileswere instrumentedwithvibrating-wire strain gauges (VWSGs), which were welded at six dif-ferent depth intervalsonpileDandfivedifferent depth intervals onpile Ito provide measurements from each soil stratum. Unfortunately, aftercompletion of the pile installation, only three of 12 VWSGswere func-tional on pile D, whereas eight of 10 VWSGswere functional on pile I.

To evaluate pile capacities, large-strain dynamic tests using theGRL Pile Driving Analyzer (PDA) were conducted at the end ofinitial pile driving (EOD). Restrike tests were performed 4 to 12 dayslater to investigate the setup effects (i.e., time-dependent pile ca-pacity gains). Two static cyclic axial-compression-load tests wereconducted on piles D and I 63 and 44 days after EOD, respectively.The first loading–unloading cycle was carried out using a modifiedquick-load test procedure in accordance with ASTM D1143:Standard Test Method for Pile Under Static Axial CompressiveLoad. The loadwas applied in increments of 445 kN until a total loadof 2,670 kN was reached. Thereafter, the increments were changedto 222.5 kN until pile failure. Each load increment was held for5 min. When the pile head displacement exceeded 63.5 mm, the testpile was unloaded in decrements of 25% of the maximum load.The first quick-load test was followed by three cycles of static cyclic-load tests recommended by Paikowsky et al. (1999), of which theloading rate was approximately 111.25 kN/min for loading and222.5 kN/min for unloading.

Pile Load Test Results

Fig. 3(a) plots the relationship between the measured soil plugginglength Lplug and the pile penetration depthDp, and Fig. 3(b) presentsthe development of the incremental filling ratio IFR with Dp duringpile driving. It was noticed that Lplug was smaller than Dp, and Lplugincreased almost linearly with Dp. The measured IFR was between0.9 and 1.0 whenDp was less than 35 m, which suggested a partiallyplugged condition but close to a fully coring mode. However, withthe advance of pile penetration into the underlying medium-denseto very dense sand and gravel layers, IFR decreased rapidly to about0.65 atDp 5 40 mBGS. This should be attributed mainly to the factthat the underlying sands and gravels provided much larger re-sistance than the upper soft clayey silts. Therefore, it can be expectedthat the test-pipe pileswere to be fully plugged once theywere drivenfurther into the underlying dense cohesionless soils.

Case Pile Wave Analysis Program (CAPWAP) analysis indi-cated that the ultimate compression capacities of the piles at theirfinal penetration depths ranged between 1,308 and 1,922 kN at EODand between 2,915 and 3,444 kN at restrike 4–12 days later. For thetime period between EOD and restrike, the pile capacity increasesranged from 77–80% at piles A and C and from 60–137% at piles Fand H. Substantial setup effects were observed at each test pile. Thecomparisons of the estimated pile capacities at EOD and restrikeindicated that the majority of setup effects occurred within a fewdays of completion of the initial pile driving.

Fig. 4(a) shows the results of the static cyclic axial-compression-load tests on piles D and I, in which the elastic compression line andtwo commonly accepted pile failure criteria (i.e., Davisson 1972;Kyfor et al. 1991)were also included for comparison.During thefirstquick-load test cycle, pile D experienced a plunging type of failureat 3,400 kN. During the subsequent static cyclic-load test, the load-displacement curves behaved similarly to that of Cycle 1. Differentfrom pile D, pile I did not experience distinct plunging failure.According to Paikowsky et al. (1999), the load at the intersectionpoints of static cyclic-load tests corresponds to the ultimatecapacities for piles that do not experience plunging failure. Theultimate capacities of pile I estimated by Kyfor et al. (1991) andPaikowsky et al. (1999) were around 3,500 and 3,400 kN, whichcompared favorably with that of pile D.

Fig. 4(b) presents the derived annulus resistance Qann and theunit-shaft resistance per linear meter of the pile Fp from the strain-gauge data at pile I.Fp is the sum of themobilized side frictions from

Fig. 1. Typical soil profiles based on the field exploration tests

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both the exterior and interior of the pile. Qann gradually increasedwith pile displacement in the next three loading cycles. This increasewas likely attributable to the increased soil density/stiffness below40.2 m BGS, which resulted from the advance of the pile tip. Itseemed thatFp of the clayey silt was fullymobilized. However,Fp ofthe underlying dense to very dense sandy soils did not reach its peak.The derived Fp and Qann for design are summarized in Fig. 4(b).

Design Considerations

Based on findings from the literature (e.g., Paikowsky and Whitman1990; O’Neill and Raines 1991; McVay et al. 2004) and the fact that

the test piles exhibited a tendency toward plugging as they advancedinto the underlying sand and gravel layers, a fully plugged conditionwas considered for the design of the static compression capacities ofopen-ended pipe piles at this site. For a plugged open-ended pipe pile,its toe resistanceQb can be separated into annulus resistanceQann andsoil-plug resistanceQplug. According to recommendations for unit toeresistances of open-ended pipe piles tipped in cohesionless soils byAmericanPetroleumInstitute (API) (1993) andTomlinson (1994), theupper limit of unit soil-plug resistance was assumed to be 4,788 kPa(100 ksf) for the static pile capacity analysis in this case.

For a plugged open-ended pipe pile, the interior shaft resistance ismobilized to counteract the soil-plug resistance. Therefore, only theexterior shaft resistance should be considered in the estimation of its

Fig. 2. Schematic of the instrumented piles and layout of test-pile groups

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Fig. 3. Measured soil-plugging effects during pile driving: (a) relationship between Lplug and Dp; (b) relationship between IFR and Dp

Fig. 4. Static cyclic axial-compression-load test results: (a) load versus displacement at piles D and I; (b) derived annulus resistance and unit shaftresistance at pile I

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axial capacity. Hannigan et al. (1997) reported that the value of theinterior unit shaft resistance in an open-ended pipe pile wastypically on the order of one-third to one-half the exterior shaftresistance, which was influenced by soil types, soil parameters,and pile shoe configurations. However, Paik et al. (2003) observedthat the average interior shaft resistance was 36% higher than theexterior shaft resistance. Because of the contradictory findings inthe literature, some assumptions had to bemade regarding the ratiobetween the interior and exterior unit shaft resistances to calculateunit exterior shaft resistance. The assumptions should satisfy theequilibrium of forces and yield reasonable exterior shaft re-sistance. Based on the aforementioned design considerations forthe pile toe resistance, the back-analysis of the load-test result onpile I showed that when the exterior unit shaft resistance was abouttwo-thirds the total unit shaft resistance for Layers 1 to 4 and two-fifths for Layer 5, the calculated pile capacity matched well withthe measured value.

Comparison Between Load Test Results andCalculations

Calculation by Static Bearing-Capacity Formulas

Based on the aforementioned design considerations for both pile toeresistance and shaft resistance, the axial pile capacities Q of the testpiles were calculated by the static bearing-capacity formulas usingthe derived soil annulus resistance and shaft resistance [Eq. (1)]:

Q ¼ Qb þ Qs ð1Þ

where toe resistance Qb and shaft resistance Qs are calculated byEqs. (2) and (3), respectively:

Qb ¼ Qann þ Qplug ¼ qann � Sann þ qplug � Splug ð2Þ

Qs ¼ Pfs � c � Dh ¼ Sb × fp � c � Dh ¼ Sb ×Fp � Dh

ð3Þ

whereQann 5 annulus resistance, equal to 800 kN; Qplug 5 soil-plugresistance; qann 5 unit annulus resistance on the pile rim; Sann 5 areaof pile rim; qplug 5 unit base resistance on the soil plug, equal to thecone penetration test (CPT) cone resistance qc, with an upper limit of4,788 kPa (100 ksf); Splug 5 area of the soil plug; fs 5 unit exteriorshaft resistance, equal to b × fp; b 5 two-thirds for Layers 1 to 4 andtwo-fifths for Layer 5; fp 5 total unit shaft resistance for each soillayer, equal to Fp=c; Fp 5 derived unit shaft resistance per linearmeter of the pile shown in Fig. 4(b); c 5 pile perimeter; and Dh 5depth interval for each soil layer. Fig. 5 compares the measuredpile capacities with those calculated by the static bearing-capacityformulas. It was noticed that the calculated pile capacities were inagreement with the measured pile capacities at piles D and I from thestatic axial-load tests and the estimated pile capacities at piles A, C, F,and H by CAPWAP analysis at restrike.

Calculation by CPT-Based Methods

Toderive pile capacities usingCPTdata, correlations betweenCPTdataand unit soil resistance needed to be established. Among the parametersobtained from CPT tests, tip resistance is considered to be the mostreliable. Consequently, almost all CPT-based methods for axial pilecapacity analysis used the correlation coefficients between CPT tip re-sistance and soil resistance. Both the uncorrected and corrected CPT tipresistances had been used to develop such correlations, which includedthe measured cone resistance qc (e.g., De Kuiter and Beringen 1979;Bustamante and Gianeselli 1982—Laboratoire Central des Ponts etChausees (LCPC) method), the net cone resistance qn (Almeida et al.1996), and the effective cone resistance qe (Eslami and Fellenius 1997).Although it has beenwidely recognized that the corrected tip resistancesqe and qn are more reliable than the uncorrected tip resistance qc, thecorrelations between the uncorrected tip resistance qc and the correctedtip resistances qe and qn and shaft resistance were established in thisstudy for references. For each soil layer, the correlation coefficient for theshaft resistancea is defined as the ratio of the average unit exterior shaftresistance obtained from the load-test result at pile I to the average CPTtip resistance. The derived a values for each soil layer in the case of qc,qe, and qn are summarized in Table 1. The table shows that a was notsensitive to qe or qn. As expected, qc yielded relatively smallera values.

Fig. 5. Comparison of pile load testing results with those calculated

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Toverify thederiveda values, the calculated pile capacities usingthe CPT-based methods were compared with the measurements atthe test piles. Since the derived a values for qc, qe, and qn were allback-calculated from the load-test result at pile I, only the pilecapacities derived from qe are presented here. The pile capacities Qwere still calculated by Eqs. (1) to (3). Unlike the previous calcu-lation, here fs is calculated by

fs ¼ a × qe ð4Þ

where a 5 derived correlation coefficient based on qe; Dh 5 depthinterval for CPT soundings; qann 5 qe over an influence zone de-fined by Eslami and Fellenius (1997); and qplug 5 qe over the in-fluence zone with an upper limit of 4,788 kPa (100 ksf). As shownin Fig. 5, the predicted capacities using the qe-based procedurematched well with those from the load tests and the CAPWAPanalysis at restrike. Additionally, the estimated pile capacities bythe qc-based LCPC method (Bustamante and Gianeselli 1982), theqe-based method (EF1997; Eslami and Fellenius 1997), and theqn-based method (ADL1996; Almeida et al. 1996) were also includedin Fig. 5 for comparison. Apparently, both the LCPC method andthe EF1997 and ADL1996 methods overestimated the pile capaci-ties. However, compared with the LCPC method, both the EF1997and ADL1996 methods demonstrated better match with thecapacities determined by the field axial-load tests and the CAPWAPanalysis at restrike. This difference should be attributedmainly to thefollowing two factors: (1) the LCPC method disregards the porepressures acting on the cone shoulder, and (2) the LCPC methodemploys total vertical stress, whereas the long-term pile behavior isgoverned by effective stresses.

The axial pile compression capacities were also estimated on thebasis of SPT blow counts. The calculation used two computer pro-grams, i.e., (1) Florida Bridge (FB)-Deep (Hoit et al. 2002) and (2)DRIVEN1.2 (Mathias andCribbs 1998). In the prediction procedures,a blow count of 60 was set as an upper limit for soil strength; i.e.,blow count over 60 was automatically truncated to be 60. In ad-dition to a generalized soil profile and average SPT blow countswithin each soil layer required for the FB-Deep program, the inputsoil parameters for DRIVEN 1.2 included the averaged undrainedshear strength Su. Based on the field and laboratory test results,20 and 50 kPa were adopted as the lower and upper limits for Su.Comparedwith themeasured pile capacities by the static axial-loadtests and estimated by CAPWAP analysis at restrike, FB-Deepslightly underestimated the pile capacities, but DRIVEN 1.2 sig-nificantly overestimated the pile capacities in the underlying sandand gravel layer.

Conclusions

Although the open-ended steel pipe piles experienced a partiallyplugged condition yet close to a fully coring mode during the

process of penetrating through the upper thick varved clayey siltdeposits, a fully plugged condition can be considered in the design oftheir static compression capacities if they were end-bearing in densecohesionless soils. In such a case, unit base resistance on the soilplug equal to the CPT cone resistance with an upper limit of 4,788kPa (100 ksf), recommended by API (1993) and Tomlinson (1994),and exterior unit shaft resistance equal to two-thirds the total unitshaft resistance in the varved clayey silts can be adopted in designusing the well-known static bearing-capacity formulas or theqe-based method.

Since both pile tip and shaft resistances are significantly affectedby soil properties, the derived soil parameters Qann, Fp, b, and a inthis study are limited only to the varved clayey silt deposits along theDelaware River. They might not be applicable to silty soils fromdifferent sites, which needs to be verified by further studies.

Acknowledgments

Many organizations and people contributed to the success of this re-search project, and special thanks are due to Mr. Junius Allen of BPand the former Geotechnical Engineers of WPC, Inc., i.e., Mr. WuYang, Mr. Joshua Adams, Mr. Jian Fang, Mr. Donovan Ledford,and Dr. Ed Hajduk. The financial support offered by NSFC GrantNo. 51109125, the Fundamental Research Funds for the CentralUnivs. No. 0230219134, Innovation Program of Shanghai Munici-pal Education Commission No.13ZZ027, and the project sponsoredby SRF for ROCS, SEM are gratefully acknowledged. Dr. Ye Lufrom Shanghai Univ. is appreciated for her contribution to improv-ing the quality of this paper. Finally, the insightful comments andsuggestions from the three anonymous reviewers, the AssociateEditor, and the Editor are sincerely appreciated.

References

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Table 1. Comparison of Correlation Coefficients Between qc, qe, and qn

Soil type

Derived correlation coefficienta in this study

qc qe qn

Hydraulic fill 0.017 0.018 0.018Clayey silt 0.019 0.024 0.021Sand and gravel 0.0028 0.003 0.003Residual soil 0.0049 0.005 0.005

Note: qc 5 measured cone resistance; qe 5 effective cone resistance;qn 5 net cone resistance.

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