ARTICLE

Axial compressive response of large-capacity helical and drivensteel piles in cohesive soilMohamed Elkasabgy and M. Hesham El Naggar

Abstract: The axial compression performance of large-capacity helical piles is of significant interest because they can offer anefficient alternative to conventional piling systems in many applications such as in oil processing facilities, transmission towers,and industrial buildings. This paper presents the results of seven full-scale axial compression load tests conducted on 6.0 and9.0 m large-capacity helical piles and a 6.0 m driven steel pile. The results are considered essential to qualify and quantify theperformance characteristics of large-capacity helical piles in cohesive soils. The test piles were close-ended steel shafts with anouter diameter of 324 mm. The test helical piles were either single or double helix, with a helix diameter of 610 mm andinterhelix spacing that varied between 1.5 and 4.5 times the helix diameter. The subsurface soil properties at the test site weredetermined using field and laboratory testing methods. The 6.0 m piles were tested 2 weeks after installation, while the 9.0 mpiles were tested 9 months after installation. The load–settlement curves were presented to better understand the behaviour oftest piles. An ultimate capacity criterion was proposed to estimate the ultimate load of large-capacity helical piles. The testhelical piles developed ultimate resistances up to 1.2–1.8 times that of the driven pile. The load-transfer mechanisms of large-capacity helical piles were studied, and it was found that soil disturbance during pile installation had a significant effect on thepile failure mechanism regardless the value of the interhelix spacing to helix diameter ratio. The mobilized soil strengthparameters were back-calculated and compared with the estimated intact soil strength parameters.

Key words: helical piles, driven piles, full-scale testing, compression, cohesive soils, load transfer, cone penetration test.

Résumé : La performance en compression axiale de pieux hélicoïdaux a grande capacité comporte un intérêt significatif puisqueceux-ci peuvent offrir une alternative efficace aux systèmes de pieux conventionnels dans plusieurs applications, telles que lesinstallations de traitement du bitume, les tours de transmission et les édifices industriels. Cet article présente les résultats desept essais de compression axiale a l’échelle réelle réalisés sur des pieux hélicoïdaux a grande capacité de 6,0 et 9,0 m et sur unpieu foncé en acier de 6,0 m. Les résultats sont considérés comme essentiels pour qualifier et quantifier les caractéristiques dela performance des pieux hélicoïdaux a grande capacité placés dans des sols cohésifs. Les pieux d’essai avaient des arbres d’acierfermés aux extrémités avec un diamètre de 324 mm. Les pieux d’essai hélicoïdaux avaient une hélice simple ou double, avec undiamètre d’hélice de 610 mm et un espacement inter-hélice qui variaient entre 1,5 et 4,5 fois le diamètre de l’hélice. Les propriétésdu sol au site d’essai ont été déterminées a l’aide de méthodes expérimentales en laboratoire et sur le terrain. Les pieux de6,0 m ont été testés deux semaines après leur installation, tandis que les pieux de 9,0 m ont été testés neuf mois aprèsl’installation. Les courbes de charge–tassement sont présentées pour offrir une meilleure compréhension du comportement despieux d’essai. Un critère de capacité ultime est proposé pour estimer la charge ultime de pieux hélicoïdaux a grande capacité. Lespieux hélicoïdaux testés ont développé des résistances ultimes jusqu’a 1,2 a 1,8 fois celle du pieu foncé. Les mécanismes detransfert de charge des pieux hélicoïdaux a grande capacité ont été étudiés, et il a été démontré que le remaniement du soldurant l’installation du pieu a un effet significatif sur le mécanisme de rupture du pieu peu importe la valeur du ratio del’espacement inter-hélice sur le diamètre de l’hélice. Les paramètres de résistance mobilisée du sol sont rétro-calculés etcomparés aux paramètres de résistance estimés pour un sol intact. [Traduit par la Rédaction]

Mots-clés : pieux hélicoïdaux, pieux foncés, essai a l’échelle réelle, compression, sols cohésifs, transfert de charge, essai depénétration du cône.

IntroductionA helical pile consists of one or more helical-shaped circular

plate(s), or helix, affixed to a steel central shaft as illustrated inFig. 1. Based on geometrical and design considerations, two helicalpile categories can be practically identified, namely small-capacityhelical piles and large-capacity helical piles. A large-capacity heli-cal pile is formed of a steel pipe of diameter >178 mm (7 in.) witha hollow circular cross section. Helical piles can be installed withshaft diameters up to 914 mm (36 in.) and with axial compressioncapacities of up to 4500 kN, depending on pile geometry andsubsurface conditions.

Helical piles are recognized as a viable foundation option inseveral engineering applications because of their ease of installa-tion, especially in limited access areas with low levels of noise andvibration, and cost effectiveness. In addition, helical piles allowimmediate loading upon installation and can be installed throughgroundwater without casing, unaffected by caving soils (Bobbittand Clemence 1987; Perko 2009).

Analytical methods have been developed to calculate thehelical pile capacity under compression or uplift loading (Mitschand Clemence 1985; Mooney et al. 1985; Narasimha Rao et al. 1991,1993). Several design methods were established based on the

Received 6 September 2012. Accepted 3 July 2014.

M. Elkasabgy. AMEC Environment and Infrastructure, 140 Quarry Park Blvd. SE, Calgary, AB T2C 3G3, Canada.M.H. El Naggar. Geotechnical Research Centre, Faculty of Engineering, The University of Western Ontario, London, ON N6A 5B8, Canada.Corresponding Author: M. Hesham El Naggar (e-mail: helnaggar@eng.uwo.ca).

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Can. Geotech. J. 52: 224–243 (2015) dx.doi.org/10.1139/cgj-2012-0331 Published at www.nrcresearchpress.com/cgj on 17 July 2014.

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individual plate anchor theories (e.g., Adams and Hayes 1967;Meyerhof and Adams 1968), while some methods are similar tothose proposed for underreamed cast-in-place piles (NarasimhaRao et al. 1993). The helical plates increase the helical pile capacitysignificantly as do the underreams in the underreamed piles.Experimental studies demonstrated that the behaviour of helicalpiles under compression and tension loads is similar (NarasimhaRao et al. 1991; Zhang 1999). Thus, the design methods are consid-ered to be interchangeable.

Tappenden (2007) investigated the accuracy of the establisheddesign methods with which the axial capacity of helical piles maybe predicted using the results of 29 full-scale axial load tests intension and compression for piles installed in western Canada.Lutenegger (2009) indicated that there was no distinct transitionfrom cylindrical shear to individual bearing behaviour based onthe results of a field study that was conducted to investigate thebehaviour of multi-helix anchors in clay. Sakr (2009) investigatedthe axial capacity of large-diameter helical piles installed in oilsands.

Several numerical studies were conducted to investigate theaxial compressive and uplift characteristics of helical piles andanchors. Merifield et al. (2006) evaluated the effect of the shape ofthe anchor plate (i.e., square, circular, or rectangular) on the axialcapacity of horizontal anchors in undrained clay through three-dimensional (3D) finite element modelling. Livneh and El Naggar(2008) investigated the load-transfer mechanism of the square-shaft slender helical piles using 3D finite element models. Kurianand Shah (2009) developed 3D and axisymmetric finite elementmodels to study the axial and lateral capacities of helical piles indrained cohesive soils. Elsherbiny and El Naggar (2013) investi-gated the axial compressive behaviour of smaller helical piles in

sand and in clay and proposed a method for their design based onthe results of the calibrated numerical model.

The installation of helical piles in cohesive soils can cause re-moulding of the adjacent cohesive soil within the zone affected bythe pile helices. This remoulding can introduce fissures within thepile cylindrical failure surface and surface heaving (Mooney et al.1985; Bradka 1997; Zhang 1999). Thus, the shear strength of cohe-sive soil decreases along the affected zone. In cohesionless soils,screwing the pile induces stress changes in the soil due to theaccompanied soil disturbance. The disturbed zone may becomeweak, which causes a cylindrical failure surface around the pile(Vesic 1971). Weech (2002) studied the soil disturbance caused bythe installation of full-scale small-diameter helical piles in sensi-tive fine-grained soils and the mobilization of axial pile capacity atdifferent times after installation. Vyazmensky (2005) developed anumerical formulation that incorporated a simplified techniquethat allowed reasonable prediction of stresses and pore pressurevariation during and after helical pile installation based on theresults of the field testing performed by Weech (2002).

Clemence and Pepe (1984) and Mitsch and Clemence (1985) per-formed field and laboratory tests and reported that during helicalanchor installation, the sand in contact with the helices was dis-placed laterally, which in turn densified the surrounding sandand increased the lateral stress. However, depending on the den-sity of sand, stress level, and shear strain level, dilation may ormay not take place.

Based on observed failure modes of helical piles installed incohesive and cohesionless soils, there are two common methodsto account for the contribution of the helices to the pile capacity,namely the cylindrical shear method and the individual platebearing method. In the cylindrical shear method (Mooney et al.

Fig. 1. Failure modes: (a) cylindrical shear method; (b) individual plate bearing method. D, helix diameter; d, pile shaft diameter; H, embedmentdepth of upper helix; Heff, pile shaft effective length; Lc, distance between top and bottom helices; S, interhelix spacing.

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1985), a cylindrical failure surface is assumed between the topand bottom helices (Fig. 1a). The ultimate compressive capacityof the pile is the sum of the bearing resistance of the bottom helix,the frictional resistance along the cylindrical failure surface, andthe frictional resistance along the pile shaft. In the individualplate bearing method (Meyerhof and Adams 1968), the pile com-pressive capacity is calculated assuming an individual bearingfailure occurs below each helix plate (Fig. 1b). Thus, the ultimatecompressive capacity of the pile is the sum of the individual bear-ing resistance of all helices and the frictional resistance along thepile shaft.

Objectives and scope of workThe main objective of this study is to investigate the perfor-

mance characteristics of single large-capacity helical and drivenpiles installed in cohesive soils and subjected to axial compressionloading. The specific objectives of the testing program were to(i) determine the load-transfer mechanisms of helical piles,(ii) define the appropriate failure criterion for large-capacity heli-cal piles, and (iii) evaluate the axial compression capacities oftested piles. The testing program included seven full-scale axialcompression load tests carried out on six large-capacity helicalpiles and one driven steel pile.

Failure modes of helical piles

Cylindrical shear methodThe ultimate compression capacity, Q ult, of helical piles in co-

hesive soil may be given by

(1) Q ult � (Ah � At)CuNc � (�DLc)Cu � (�dHeff)�Cu

where Ah is the projected area of the helix, At is the area of the piletip, Cu is the undrained soil cohesion, Nc is the bearing capacityfactor, d is the pile shaft diameter, D is the helix diameter, Lc is thedistance between the top and bottom helices, Heff is the effectivelength of the pile shaft (equal to the embedment depth of the

upper helix, H, minus the helix diameter), and � is the adhesionfactor. Zhang (1999) stated that the method provided better capac-ity prediction for multi-helix piles with interhelix spacing (S) todiameter (D) ratio of S/D ≤ 3.0.

For cohesionless soils, the formulation proposed by Mitsch andClemence (1985) can be used to estimate the compression capacity:

(2) Q ult � (Ah � At)�′HnNq � ��

2DLc

2�Ks tan� ′ ��2

dHeff2 � ′Ks tan� ′

where Hn is the depth to the bottom helix, �= is the effective soilunit weight, Nq is the bearing capacity factor, Ks is the coefficientof lateral earth pressure, and � ′ is the effective angle of internalfriction. Zhang (1999) reported that the cylindrical shear methodprovides a better prediction of the compression capacity of pilesin sand with S/D ≤ 2.0.

Individual plate bearing methodFor cohesive soils, the individual plate bearing method can be

adopted for piles with S/D > 3.0 (Zhang 1999). The ultimate capac-ity of helical piles in cohesive soils is calculated as follows:

(3) Q ult � AtCuNc � � (AhCuNc) � (�dHeff)�Cu

For cohesionless soils, the individual plate bearing method can beused to calculate the ultimate capacity for the case of S/D > 2.0, i.e.,

(4) Q ult � At�′HnNq � �

i�1�n

Ah� ′HiNq ��2

dHeff2 � ′Ks tan� ′

where Nq and Hi are the bearing capacity factor and embedmentdepth of the upper helices, respectively.

Pile configurations and distributionSix helical piles and one driven pile were manufactured and

installed. Four of the helical piles were 6.0 m long and two were

Fig. 2. Plan view of piles, seismic cone penetration tests, and borehole. SPT, standard penetration test.

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9.0 m long. Two helical piles were fabricated as single-helix piles,and the other four piles were double-helix piles with differentinterhelix spacing ratios, S/D = 1.5–4.5. The driven pile was 6.0 mlong and had the same shaft diameter and wall thickness as thetest helical piles. Figure 2 and Table 1 provide the distribution andbasic geometrical characteristics of the test and reaction pilesemployed for compression load tests. The piles consisted of acentral steel shaft of outer diameter 324 mm and wall thickness9.5 mm. The helix plate was 610 mm in diameter, with helix pitchof 152 mm (6 in.), and thickness of 19 mm. All test piles wereclose-ended with a flush closure steel plate. The pile shafts con-form to the material and dimensional requirements of ASTM stan-dard A252-98 (ASTM 2007a), grade 2 steel. The tensile and yieldstrengths of the piles’ material were 414 and 240 MPa, respec-tively, and the elastic modulus of the pile material was 210 GPa.

The test layout complied with the specifications of the ASTMstandard D1143-07 (ASTM 2007b). The axial compression testingprogram included seven full-scale load tests, which were con-ducted in two phases: phase I and phase II (Table 1). Phase I wascarried out on the 6.0 m helical piles (SS11, SD11, SD21, and SD31)and the 6.0 m driven pile (DS), 2 weeks after pile installation,when the soil around the piles’ shaft was disturbed because of thepile installation process. Phase II of the piles testing program wascarried out on the 9.0 m helical piles (LS12 and LD12), 9 monthsafter pile installation, which was deemed to be enough time forthe disturbed soil to regain significant amount of its original stiff-ness and strength.

Subsurface conditionsThe test site is located north of the town of Ponoka, in the

central Alberta region, Canada. The surficial deposits in the testsite zone consist of Pleistocene Stagnation Moraine glacial tilldepositions (Shetsen 1990). Both in situ and laboratory tests wereperformed to characterize the index, stress history parameters,and shear strength properties of the site soils. Laboratory testswere conducted on disturbed and partially undisturbed samplescollected from one borehole at an interval of 1.5 m. Standardpenetration tests (SPTs) were performed in the borehole to deter-mine the N value at different depths. Three seismic cone penetra-tion tests (SCPTs) were conducted within the site (see Fig. 2) to adepth of 15 m. The SCPT results were used to determine the soillayering as shown in Figs. 3 and 4.

The subsurface conditions at the test site (Figs. 3 and 4) consistof the following: layer 1, a 1.0–1.5 m thick brownish and mediumdense sandy silt crust with some organic materials; layer 2, a2.8–3.1 m thick layer of brownish and moderately overconsoli-dated stiff clay to silty clay and clayey silt of medium plasticityinterbedded with seams of silt; layer 3, 0.6–1.3 m of interbeddedlayers of grey and overconsolidated very stiff silty clay, clayey silt,and clay of medium plasticity; layer 4, a 0.3–1.5 m thick yellowish

brown and dense to very dense silty sand; layer 5, a 7.2 m thicklayer of grey and overconsolidated very stiff to hard clay till ofmedium to high plasticity with interbedded seams of layers ofsandy silt, silty sand, and silt at a depth of 8.0 m; layer 6, interbed-ded layers of moderately to heavily overconsolidated sandy silt,silt, clayey silt, and silty clay. The pore pressure measurementsand dissipation tests obtained from the cone tests indicatedgroundwater level at 1.2 m below the ground surface.

A summary of the soil index properties, stress history parame-ters, and strength properties obtained from laboratory tests andfrom correlations with SPT and cone penetration test (CPT) resultsare presented in Table 2.

Pile installationFigure 5 shows a schematic diagram of the test piles. The helical

piles were installed by applying a clockwise turning moment(torque) to the pile shaft, while a rate of penetration of one helixplate pitch (152.4 mm) per revolution was considered to reducesoil disturbance. The driven pile was installed using a mechanicaldrop hammer, weighing 19 kN, and dropped through a distance ofabout 0.9 m onto the pile cushion. The piles protruded 0.6 mabove the ground surface (i.e., unsupported length).

Generally, installation of helical piles causes disturbance to co-hesive soils. This disturbance and associated reduction in strengthand stiffness are even more pronounced for structured cohesivesoils. This type of soils exhibits higher degradation in strengthand stiffness when the natural microstructure is disturbed.

The disturbance can be attributed to the advancement of thepile shaft into the ground and the rotation of the helical plates. Inthe current study, the level of disturbance due to shaft penetra-tion is expected to be higher, as the piles were installed close-ended. Therefore, it can be assumed that the test helical pilesacted as displacement piles during the installation process in away that the site's structured cohesive soils were displaced andremoulded around the pile tip. Furthermore, slight ground sur-face heave was monitored during the installation of the test piles.

In addition, the soil displacement and shearing phenomenawould cause the soil structure to collapse and lead to a buildup ofexcess pore-water pressure that causes a reduction in effectivestress. This temporary buildup of excess pore-water pressurewould cause the cohesive soil to experience significant reductionin its shear strength.

The contribution of the helical plates to the disturbance of the sitesoil can be ascribed to different phenomena. The helix disturbs soilthrough a spirally shaped cut with intervals equal to the pitch dis-tance, in which the soil traversed by the helix is sheared and dis-placed laterally and vertically. However, this level of disturbancemay increase for layered subsurface conditions, where crowding isemployed to advance the pile into harder parts of the soil profile.Furthermore, a sudden change in the applied installation torque due

Table 1. Geometrical properties of test and reaction piles.

Pile No. Type Length (m)No. ofhelices

Interhelixspacing, S (m) H/D S/D

Installationtorque, T (kN·m) KT (m−1)

Phase I (piles tested 2 weeks after installation)SS11 Test (helical) 6.0 1 — 8.9 — 83.0 10.2SD11 Test (helical) 6.0 2 0.9 7.5 1.5 115.2 11.6SD21 Test (helical) 6.0 2 1.8 6.0 3.0 112.9 10.0SD31 Test (helical) 6.0 2 2.7 4.5 4.5 92.2 11.3DS Test (driven) 6.0 — — — — — —

Phase II (piles tested 9 months after installation)LS12 Test (helical) 9.0 1 — 13.9 — 94.9 20.5LD12 Test (helical) 9.0 2 0.9 12.4 1.5 93.4 26.5R Reaction (helical) 6.0 2 2.1 7.9 3.0 — —

Note: S, interhelix spacing; H/D, embedment ratio defined as the ratio of the embedment depth of the upper helix, H, to the helixdiameter, D; T, pile installation torque; KT, empirical capacity-to-torque factor.

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to the change in strength and stiffness parameters between the soillayers would cause alteration in the rate of the pile penetration,which can cause more soil disturbance especially for multi-helixpiles.

The final measured average torques at the end of helical pileinstallation and total embedment depths are given in Table 1. Itshould be noted that the distribution of the torque values alongthe full embedment depths of test piles and the driving record ofthe driven pile were not available by the manufacturing company.It can be noticed from Table 1 that torque values for piles SS11,SD11, SD21, and SD31, tested in phase I, varied between 83 and115.2 kN·m. As explained later, the torque values were affected bythe presence of large-size particles, which could be the possiblecause of having large torque values for the double-helix piles SD11and SD21, as some difficulties were experienced in installing thosepiles to the designated depth. The presence of the large-size par-

ticles also prevented advancing the cone soundings SCPT-1 andSCPT-2, which were relocated to the new locations shown in Fig. 2.

The torque values required to install the double-helix pilesSD11, SD21, and SD31 were �11%–39% higher than the torque valuerequired for the single-helix pile SS11. It can be seen that thetorque value required for the double-helix pile SD31 is smallerthan those values recorded for the double-helix piles SD11 andSD21 by �19%–20% as a result of the interhelix spacing and soillayering. In addition, the torque values required to install the9.0 m single-helix and double-helix piles LS12 and LD12, tested inphase II, were 94.9 and 93.4 kN·m, respectively.

Pile instrumentationThe helical piles were instrumented with 7–12 levels of strain

gauges, depending on the pile length and number and spacing of

Fig. 3. Measured CPT profile at SCPT-2 (1 bar = 100 kPa). W.L., groundwater level.

Fig. 4. Schematic diagram for soil stratigraphy cross sections along (a) 6.0 m and (b) 9.0 m helical piles. �, bulk unit weight.

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helices (see Fig. 5). Half-bridge strain gauge circuits were affixedon the inner wall of the piles at specified locations. For the 6.0 mpiles, the first level of gauges were installed on the outer surfaceof the piles, while the first, second, and third levels of gauges wereinstalled on the outer surface for the 9.0 m piles. Each level ofgauges encompassed four half-bridges allocated equidistantlyalong the shaft circumference. The pile was cut into short seg-ments, the gauges were installed at an approximate distance of250 mm from the edge of each segment, and then gauges andwires were covered with a fiber cloth and epoxy. The pile seg-ments were then welded together to form the test pile while carewas exercised to ensure ground flush and grind welds. The straingauges and wires were protected by fiberglass insulation from theheat during the welding process.

Test setup and proceduresSchematic diagrams and photos of the load test setups are illus-

trated in Fig. 6 for helical piles and driven pile. The followingsections describe the setup loading system, the test reaction sys-tem, the displacement measurement devices, and test procedures.

Loading systemA calibrated hydraulic jack, actuated with a hydraulic pump,

with capacity of 1800 kN was used to apply the compression load,while a calibrated load cell of capacity 4500 kN was employed tomeasure the load. For the double-helix pile LD12, two hydraulicjacks each of capacity 1800 kN were necessary to apply the com-pression load. A hydraulic pressure transducer, with capacity of69 MPa, was attached to the hydraulic jack to measure the pres-sure applied at the pile head. A main beam and two secondarybeams were utilized as the reaction frame for testing helical piles.For the driven pile, the reaction frame comprised only one mainbeam.

Reaction systemThe reaction system for testing helical piles was composed of

four reaction helical piles, each of shaft diameter 324 mm, length6.0 m, helix diameter 685 mm, and interhelix spacing of threetimes the helix diameter. The reaction piles had at least 4.5 timesthe capacity of the 6.0 m test piles and at least 1.6 times thecapacity of the 9.0 m test piles. However, the reaction system forthe test driven pile was composed of two helical piles, with acapacity of at least 2.4 times the capacity of the driven pile. Thedistances between the test and reaction piles are provided inFig. 2. The test load was transferred to the reaction piles usinghigh-strength all-threaded steel bars 38 mm in diameter and1.22 m in length. The main reaction beam was placed on andconnected to the top of the two secondary beams and centeredover the test helical pile. For the driven pile, the main reaction

beam was placed on the two reaction piles and centered over thetest pile.

A steel cap plate of nominal thickness of 40 mm was centeredand welded on the head of each test pile. Two plates with thick-nesses of 25 and 14 mm were placed on top of the cap plate. Thehydraulic jack was placed on top of the steel cap, and the load cellwas situated between the jack and the main reaction beam usingsteel plates of nominal thickness of 25 mm in between.

Displacement measurementsThe vertical pile movement was measured using four calibrated

linear displacement transducers (LDTs) with accuracy of 0.01 mmand maximum mechanical travel of 100 mm. The transducerswere mounted on two independently supported 102 mm box-section steel reference beams. The lateral movement of the testpile was monitored using two dial gauges of accuracy 0.25 mm(0.01 in.).

Test proceduresThe axial compression load test was carried out following the

quick maintained load test method recommended by ASTM stan-dard D1143-07 (ASTM 2007b). Each pile was loaded in increments of10%–15% of the anticipated ultimate load. Each increment wasmaintained for a period of 5 minutes, with displacement andstrain readings recorded by the data acquisition system. The loadincrements were applied until continuous jacking was required tomaintain the test load, which was considered as failure of the testpile. The pile was then unloaded by removing the test load in fourequal decrements, while keeping the load constant for a timeinterval between 2.5 and 5 minutes. When reaching a zero appliedload, enough time was allowed to monitor the pile head displace-ment readings to assess the rebound behaviour. Most tests werecarried out up to a pile head settlement of �60 mm, which corre-sponded to a settlement of 10% of the helix diameter.

Load test results

Load–settlement curvesThis section presents the results of the axial compression

load tests conducted on piles tested in phases I and II. The load–settlement curves of phase I are given in Figs. 7a and 7b for thehelical piles (SS11, SD11, SD21, and SD31) and the driven pile (DS),respectively. Figure 7c presents the load–settlement curves ofphase II for helical piles LS12 and LD12.

Figure 7 shows that the maximum applied load ranged from994 to 1745 kN for the 6.0 m helical piles tested in phase I and from2057 to 2552 kN for the 9.0 m helical piles tested in phase II. Themaximum settlement for helical piles varied from about 43 to61 mm. The maximum applied load was 811 kN for the driven pile

Table 2. Summary of soil properties.

Property Layer 1 Layer 2 Layer 3 Layer 4 Layer 5 Layer 6

Soil type Silt crust Clay to silty clayand clayey silt

Silty clay, clayey silt,and clay

Silty sand Clay till Sandy silt, clayey silt,silt, and silty clay

Subsurface condition Medium dense Stiff Very stiff Dense to very dense Very stiff to hard Very stiffThickness (m) 1–1.5 2.8–3.1 0.6–1.3 0.3–1.5 7.2 1.7� (kN/m3) 18.2 17.8 17.8 18.8 17.5 18.0Wc (%) 16 17 17 18 17–28 —WLL (%) — 30–39 30–39 — 32–47 —IP (%) — 14–17 14–17 — 12–29 — 0.35 0.42 0.46 0.35 0.48 0.49OCR 10 10 14 6 15 7Cu (kN/m2) — 85 137 — 177 150� ′ (°) 31 — — 42 — —

Note: �, bulk unit weight; Wc, natural moisture content; WLL, liquid limit; IP, plasticity index; , Poisson’s ratio (drained value for layers 1 and 4 and undrained valuefor layers 2, 3, 5, and 6); OCR, overconsolidation ratio; Cu, undrained soil cohesion; � ′, angle of internal friction. �, , OCR, Cu, and � ′ were estimated from the resultsof the SCPT and SPT tests.

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Fig. 5. Schematic diagram of test piles: (a) SS11; (b) SD11; (c) SD21; (d) SD31; (e) LS12; (f) LD12.

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tested in phase I, while the maximum settlement reached wasabout 41 mm. It can be noted that the driven pile showed initialstiffer response than piles SS11, SD11, and SD21 up to the ultimatecapacity.

The installation of helical and driven piles is expected to causelarge deformations within the annular zone surrounding the pile.This would result in some reduction in the soil shear strengthwithin this annular zone. The actual degree of strength reductionand extent of annular zone are not well defined. It is expected,however, that the soil shear strength immediately after installa-tion would be close to the remoulded strength. As the time passesafter pile installation, it is expected that the soil regains some ofits shear strength.

Pile ultimate capacityConceptually, the ultimate (or failure) load of a single pile is

reached when rapid displacements occur under sustained or

slightly increased loads. However, pile load tests do not alwaysexhibit a well-defined peak load due to practical limitations onloading equipment and test setup. The load–settlement curvesobtained from the axial compression tests can normally be sim-plified into three distinct regions (Fig. 8): initial linear–elasticregion with steep slope; transition nonlinear region where themovements are largely disproportional to the load increments;and final linear region that shows small stiffness behaviour.

Point (L1) in Fig. 8 corresponds to the end of the initial linearregion, and point (L2) represents the onset of the final linear re-gion. Most of the ultimate load interpretation methods define theultimate capacity of piles to fall within the transition nonlinearregion or rarely within the final linear region. Hirany andKulhawy (2002) stated that creep displacements become signif-icant at loads that lie after the transition nonlinear region,which can cause difficulty in maintaining a constant load on

Fig. 6. Axial compression load test setups: (a) helical pile, schematic view; (b) driven pile, schematic view; (c) helical pile, four reaction piles;(d) driven pile, two reaction piles. LDT, linear displacement transducer.

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pile. Consequently, the displacements measured during thisloading zone would lead to ultimate load that is subjected toerror due to the fluctuating applied load. Thus, it is desirable toobtain the ultimate load of the pile from the upper limit of thetransition nonlinear zone and before the onset of the finallinear region.

Several definitions for the ultimate load of piles have been pro-posed to interpret the pile load carrying capacity from its load testdata based on mathematical rules for different types of piles(Fellenius 1980). Terzaghi (1942) defined the ultimate load as theload at which the pile settlement exceeds 10% of its diameter.This criterion is expected to overestimate the capacity of large-

Fig. 7. Axial compression load at pile head versus settlement: (a) helical piles (phase I); (b) driven pile (phase I); (c) helical piles (phase II).

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diameter piles and some pile types. In this study, the ultimate loadcapacities of the test piles were established using four commonlyused interpretation methods: Davisson’s method (Davisson 1972);tangent intersection method (Mansur and Kaufman 1956);O’Neill and Reese (1999) method; and Livneh and El Naggar (2008)method. These interpretation methods aim at locating the pileultimate capacity within the transition nonlinear region to reachan allowable load (design load) that lies in the initial linear region,thus avoiding excessive settlement under operating loads. Thegeneral procedures of these methods are briefly summarized inthe following text. It should be noted that for any pile that showsplunging failure behaviour, the ultimate load is taken as equal tothe maximum test load.

Davisson method (offset limit method)The Davisson method is primarily intended for driven piles

tested according to the quick maintained loading methods. Thefailure load is defined as the load corresponding to the pile headmovement that exceeds the pile elastic compression by a value of4 mm plus a factor equal to the pile diameter divided by 120, i.e.,

(5) Sp �QLEpA

� (4.0 � 8d)10�3

where Sp is the pile head settlement at ultimate load; Q is theapplied load; L is the pile length; Ep and A are the pile Young’smodulus and cross-sectional area, respectively; and d is the tipdiameter, equal to the helix diameter for helical piles and theshaft diameter for driven piles.

Tangent intersection methodThe ultimate load capacity is taken as the load associated with

the intersection of two tangents to the load–settlement curve, onerepresents the elastic range and the other represents the progres-sive settlement.

O’Neill and Reese (1999) methodThe method defines the ultimate axial compression capacity of

drilled shafts as the load that produces a pile head settlement of5% of the pile tip diameter. The tip diameter is taken equal to thehelix diameter for helical piles and the shaft diameter for drivenpiles.

Livneh and El Naggar (2008) methodThe interpretation method is intended for slender helical piles

with small diameter helices. The ultimate load is defined as theload corresponding to pile head movement that exceeds the elas-tic compression of the pile by 8% of the pile largest helix diameter,

D (i.e., eq. (6)). This method was adopted in the current study forinterpreting the ultimate capacity of helical piles only. This defi-nition was considered suitable for slender helical piles, as thehelix bearing capacity represents a considerable component ofthe pile capacity.

(6) Sp �QLEpA

� 0.08D

Inspecting the load–settlement curves in Fig. 7, it is noted thatthe load–settlement curves of the test piles showed the threedistinctive regions (initial linear, transition nonlinear, and finallinear) within the induced settlements. Generally, the near-linearfailure region was achieved at a net settlement of >3.5% of the pilehelix diameter for helical piles (i.e., 0.035D) and 3.5% of the pileshaft diameter for the driven pile (i.e., 0.035d). Thus, the ultimateload is defined as the load applied at the pile head, resulting intotal settlement at the pile head equal to 0.035D plus the elasticdeflection of the pile, i.e.,

(7) Sp �QLEpA

� 0.035D

where, for the driven pile, the helix diameter is replaced with theshaft diameter, d.

The application of the developed ultimate capacity criterion isillustrated on the load–settlement curves as shown, for example,in Fig. 9 for helical pile SS11 (phase I), driven pile DS (phase I), andhelical pile LS12 (phase 2). The 6.0 m helical piles showed a notice-able nonlinear transition zone followed by a linear zone withmoderate slope (i.e., no sign of plunging failure), reflecting theeffect of the dense to very dense silty sand layer and the clay tillsoils where the pile tips rested in. The 9.0 m helical pile LS12,however, displayed almost plunging failure because of the softersoil at its tip. In contrast, pile LD12 displayed the initial linear andthe transition nonlinear regions but the final near-failure linear re-gion was not reached due to the limitation of the load test setupcapacity. However, it experienced a higher curvilinear tendency,which reflected the increased bearing resistance from the helix. Thedriven pile, in general, exhibited a more defined ultimate load.

The ultimate load of the test piles were interpreted using the afore-mentioned methods and the proposed criterion in this study (eq. (7)),and the results are presented in Table 3. It can be concluded that aconsiderable scatter exists for the load capacities evaluated usingdifferent interpretation criteria. The table summarizes the load ca-pacity ratio for the piles, which is defined as the ratio of the differ-ence between the load capacity obtained from each of the differentultimate load criteria and the present study’s criterion to that estab-lished from the present study criterion.

The current study failure criterion estimated ultimate capaci-ties varying between 721 and 1331 kN for the 6.0 m piles (phase I),and 1952 and 2477 kN for piles LS12 and LD12 (phase II). The cor-responding pile settlements were between 13.2 and 25.6 mm forthe 6.0 m piles, and between 30 and 33 mm for piles LS12 andLD12. The ultimate capacities estimated by Davisson criterion var-ied between 629 and 884 kN for the 6.0 m piles and between 1847and 1922 kN for piles LS12 and LD12, with corresponding settle-ments that varied between about 10 and 18 mm. However, thecapacities estimated using the tangent intersection criterion werebetween 645 and 900 kN for the 6.0 m piles and between 1808 and1757 kN for piles LS12 and LD12, with corresponding settlementsthat varied between about 8 and 18 mm. It can be noted thatDavisson and tangent intersection criteria significantly underes-timated the ultimate capacities of helical piles by about 18%–34%and 14%–43%, respectively. However, the difference decreased for

Fig. 8. Typical load–settlement curve for drilled foundation.

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the cases of the 9.0 m helical pile, LS12, and the driven pile DS;possibly because these piles have a more defined ultimate load.

The ultimate capacities estimated using O’Neill and Reese cri-terion varied between 751 and 1478 kN for the 6.0 m piles andabout 1952 and 2477 kN for piles LS12 and LD12, with correspond-ing settlements of about 31 mm for helical piles and 16 mm for thedriven pile. It can be seen that the O’Neill and Reese criterionoverestimated the ultimate capacities by about 6%–12% for the6.0 m helical piles, with lesser difference for the driven pile, DS1.The ultimate capacities estimated by the Livneh and El Naggarcriterion, proposed for slender helical piles, were about 994–1623 kN for the 6.0 m piles, and about 2057 kN for pile LS12, withcorresponding settlements between about 52 and 59 mm. Thecriterion highly overestimated the ultimate capacities of thetested helical piles by about 18%–44%, but the difference was5% for pile LS12. Furthermore, it exceeded the maximum loadreached at the end of the load test of piles SD11 and LD12. Conse-quently, it can be concluded that the Davisson, tangent intersec-tion, Livneh and El Naggar criteria are not suitable to estimate theultimate compression capacity of large-diameter helical piles, es-pecially those with helices resting in very stiff to hard soils.

As presented in Table 3, the ultimate capacities for the 6.0 mhelical piles varied between 845 and 1331 kN, while the capacity ofthe 6.0 m driven pile was 721 kN. The ultimate capacity of thesingle-helix pile, SS11, was about 18% higher than that of thedriven pile, DS. However, the ultimate capacity of the double-helix piles (SD11, SD21, and SD31) was about 45%–85% higher thanthat of the driven pile. The ultimate capacity of the helical pilesoccurred at pile head displacement values that were �85% higherthan the displacement at the ultimate capacity of the driven pile.This can be attributed to the beneficial contribution of the largediameter helices on the mobilized end bearing capacity. Gener-ally, it demonstrates the advantage of helical piles over drivenpiles in ground conditions similar to what was encountered in thetest site. However, the range of ratios of axial compression capac-ity of the helical piles to the driven pile reported herein (i.e.,1.18–1.85) is smaller than the range reported by Sakr (2011) as3.0–5.0.

Figure 10 depicts the applied load normalized by the ultimatecapacity (Q /Q ult) and the settlement, St, normalized by the helixdiameter, D, for helical piles and by the shaft diameter, d, for thedriven pile. It can be noted from Fig. 10a and capacity values ofthe 6.0 m helical piles that the silty sand layer that existed at theelevation of the pile tip increased the pile capacity and eliminatedthe plunging failure trend.

The site investigation program showed that the silty sand layervaries in thickness along the test site and in some locations maynot exist in the whole interhelix zone for all the 6.0 m helicalpiles. For some piles, the interhelix zone was located within thesilty clay and silty sand layers. In addition, the strength and stiff-ness of the clay till layer varied along the test site and with depth.A higher degree of soil disturbance is expected to occur due toinstallation under the upper helices compared to the bearing soilunder the lower helices, which can significantly compromisetheir end bearing resistance. These factors affected the generaltrend of the ultimate capacity of the 6.0 m helical piles.

The pile design capacity is usually evaluated as one-half of itsultimate capacity as established from the load test. The load–settlement curves show that the pile settlement at the designcapacity is quite small. The measured settlements at the designcapacities for all piles tested in this study were <1% of the helixdiameter for the helical piles and <1% of the pile shaft diameterfor the driven pile. The settlement at the pile design capacity wasslightly less for the driven pile compared to the helical piles.However, the settlement of the driven pile at the same load valueis greater than that of the helical pile. This demonstrates that theperformance of helical piles is superior to that of the driven pile atthe same load level.

Fig. 9. Developed ultimate capacity criterion: (a) helical pile SS11(phase I); (b) driven pile DS (phase I); (c) helical pile LS12 (phase II).

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Distribution of axial load along helical pilesThe capacity of helical piles is developed from three load-

transfer mechanisms, namely resistance along the shaft above thehelices, resistance along the interhelix zone, and resistance at thebottom helix. To establish the contributions of these load-transfermechanisms, the distribution of the axial compression load alonghelical piles during the load test was evaluated from the analysisof the strain gauge readings at different elevations along the pile.Since the driven pile, DS, had no strain gauges, the load distribu-tion for this pile was not included in this section.

The load distribution curves at various load levels were estab-lished and are shown in Figs. 11 and 12. For single-helix piles, theultimate capacity is made of shaft friction and end bearing resis-tance from the bottom helix – pile tip. For double-helix piles, theultimate capacity is composed of four components: shaft friction,bearing resistance on the upper helix, shaft friction along theinterhelix zone, and bearing resistance from the bottom helix –pile tip. It can be seen from the load distribution curves that thedouble-helix piles experienced the individual bearing behaviouras the upper helix mobilized a considerable value of bearing re-sistance.

By examining Figs. 11b and 12b, it is interesting to note that pilesSD11 and LD12 with interhelix spacing, S/D = 1.5, showed mainlyindividual plate bearing failure mechanism instead of the cylin-drical shear failure. Previous studies by Narasimha Rao et al.(1989), Narasimha Rao et al. (1991), Narasimha Rao et al. (1993), andZhang (1999) examined the effect of S/D on the failure mechanismof helical piles of small shaft diameter under compression anduplift loads. These studies showed that the cylindrical shear fail-ure was predicted for helical piles in clay soil with S/D ≤ 1.5–2.0,while for S/D > 2.0 the individual plate bearing failure mechanismwould dominate.

The observed behaviour in the current study reflects the distur-bance of structured glacial clayey and silty deposits within theinterhelix zone due to pile installation. The behaviour of pilesSD11 and LD12 suggests that the upper helix transferred load tothe disturbed soil underneath it, resulting in substantial deforma-tions before there was sufficient displacement of the interhelixsoil relative to the intact soil outside the helices zone to mobilizecylindrical shear resistance. This behaviour happened to a lesserdegree in the case of pile LD12, as the pile was tested 9 monthsafter installation, allowing some regain of soil strength.

For the double-helix piles SD21 and SD31 with widely spacedhelices (S/D = 3.0 and 4.5, respectively), a plastic wedge of soil isexpected to develop underneath each helix. To mobilize pile fail-ure, a soil mass located outside the plastic wedge, and outside thedisturbed zone, would be displaced and deformed. This soil masswould not be as large as in the case of undisturbed soil beneath

and around the helices. Thus, the bearing resistance of the upperhelix may depend to some extent on the shear strength of the soillocated outside the disturbed soil in the interhelix zone.

The contribution of each load-transfer mechanism is calculatedfrom the load-transfer curves, and the results are illustrated inFig. 13. These figures show the mobilized skin friction and endbearing resistances at the ultimate capacity estimated by thecriterion proposed in the present study. The dotted line, whichindicates the ultimate capacity, was drawn at a total pile headsettlement equal to the elastic pile shortening plus a net settle-ment of 0.035D. The figure demonstrates that the initial shaftresponse was very stiff, and its capacity was mobilized at a settle-ment of about 1.0% of the shaft diameter and 0.55% of the helixdiameter. In addition, the end bearing underneath the helices –pile tip was very small initially, and was mobilized fully at a muchlarger settlement. It is important to note that shaft resistance wasmobilized fully, and the ultimate load defined by the proposedcriterion generally is defined at the end of the curvilinear regionof the end bearing resistance curve of the bottom helix – pile tip.This confirms the adequacy of the proposed criterion to estimatethe ultimate capacity of large-capacity helical piles.

The percentage contributions of the different components atthe ultimate capacity, Q ult, and at the design load (0.5Q ult) aresummarized in Table 4. Careful examination of the load distribu-tion curves and the analysis of load-transfer mechanism resultspresented in Table 4 yield the following observations:

• The percentage of shaft capacity and interhelix capacity de-creased as the applied load increased because most of the fric-tional capacity was mobilized within the first 3–4 mm of pilesettlement. In contrast, the percentage of helices bearing ca-pacity increased as the applied load increased because highersettlement is required to mobilize the end bearing component.It is also noted that the end bearing resistance continued toincrease with the increase of the applied load even after theestimated ultimate capacity, especially for piles that showed asignificant curvilinear tendency in the load–settlement curves.

• The percentage of shaft resistance decreased from 11% to 2%, forthe 6.0 m helical piles as the embedment ratio, H/D, decreasedfrom 8.9 to 4.5. However, the adhesion in the interhelix sectionincreased from 1.3% to 3.1% as the interhelix spacing ratio in-creased from 1.5 to 4.5. A slightly higher shaft resistance (12.2%)was mobilized for the 9.0 m helical piles.

• All double-helix 6.0 m piles experienced individual plate bear-ing failure mechanism with the upper helix contribution of6%–10% of the ultimate capacity. This is attributed to the distur-bance of the structured interhelix glacial clayey and silty depos-its due to pile installation.

Table 3. Ultimate pile capacities, settlements, and load capacity ratio established from different criteria.

Pile No.

Davisson Tangent intersection O’Neill and Reese Livneh and El Naggar Present study criterion

Load (kN)(Ratio %)* Settl. (mm)

Load (kN)(Ratio %)* Settl. (mm)

Load (kN)(Ratio %)* Settl. (mm)

Load (kN)(Ratio %)* Settl. (mm) Load (kN) Settl. (mm)

Phase I (piles tested 2 weeks after installation)SS11 629 (−26) 11 731 (−13) 18.2 894 (6) 30.5 994 (18) 51.8 845 24.4SD11 884 (−34) 11.5 840 (−37) 10.4 1478 (11) 30.5 —** —** 1331 25.0SD21 778 (−31) 11.8 645 (−43) 7.5 1259 (12) 30.5 1623 (44) 54.0 1126 24.8SD31 860 (−18) 11.9 900 (−14) 13.7 1116 (6) 30.5 1281 (22) 53.0 1049 24.5DS 691 (−4) 10 677 (−6) 9.3 751 (4) 16.0 —** —** 721 13.5

Phase II (piles tested 9 months after installation)LS12 1847 (−5) 17.7 1808 (−7) 16.8 1952 (0) 30.5 2057 (5) 58.6 1952 30.3LD12 1922 (−23) 17.9 1757 (−29) 15.2 2477 (0) 30.5 —** —** 2477 32.7

Note: Settl., settlement.*Load capacity ratio (values in parentheses) defined as the ratio (%) of the difference between load capacity obtained from each of the different ultimate load criteria

and the present study's criterion to that established from the present study criterion.**Ultimate capacity could not be obtained by the Livneh and El Naggar criterion.

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• The double-helix 9.0 m pile LD12 experienced individual platebearing failure with the upper helix contribution of 22.5% ofthe ultimate capacity.

• The bearing resistance contribution of the bottom helix – pile tipwas about 85%–90% of the ultimate capacity of the 6.0 m piles(phase I) and about 67% and 87% for the 9.0 m piles (phase II).

Bearing pressure of helices compared with CPT resistanceThe soil resistance below the helices – pile tip and the soil resis-

tance to the advancement of cone penetrometer are functions of thesoil shear strength. In addition, the loading mechanism is similar inboth cases. Thus, the ultimate bearing pressure below the helices –pile tip may be correlated to the cone tip resistance.

Figure 14 plots the profiles of corrected cone tip resistances, qt,measured on site at locations SCPT-1, SCPT-2, and SCPT-3 alongwith the bearing pressures below the helices and pile tip. It isnoted that the bearing pressures below the upper helix of thedouble-helix 6.0 m piles tested 2 weeks after installation (phase I)were significantly less than the range of cone resistance measure-ments. This is likely due to the soil disturbance caused by pileinstallation. Furthermore, the bearing pressures below the bot-tom helix – pile tip of the same 6.0 m piles were significantly lessthan the CPT resistance measured in the silty sand layer, probablydue to the variability of the thickness of silty sand layer. By com-paring profiles of SCPT-1, SCPT-2, and SCPT-3, it can be seen thatthe silty sand layer was shallower towards the west side of the testsite (see Figs. 2 and 4).

The bearing pressure below the bottom helix – pile tip of the9.0 m piles tested 9 months after installation (phase II) was higherthan the CPT resistance, but the bearing pressure below the upperhelix of pile LD12 was approximately similar to some extent to thelower value of the cone resistance profiles. This is ascribed to thethixotropic behaviour of the soil, as it regained most of its shearstrength with time after pile installation and the existence ofinterbedded layers of silty sand within the clay till layer. Accord-ingly, the cone tip resistance along with an adequate soil investi-gation program can provide a reasonable estimate for the bearingresistance of the bottom helix – pile tip.

Inferred soil resistance mobilized by helical pilesThe average values of unit shaft adhesion along the upper 4.0 m

of pile shaft were calculated for all test helical piles at the ultimateloads, and the results are presented in Table 5. It is observed thatthe mobilized shaft adhesion value in the stiff clay to silty claylayer (layer 2) for pile SS11 was 37% higher than the correspondingvalues calculated for the other 6.0 m piles. This may be ascribed tothe possible higher disturbance induced by the multi-helix pilesand the potential effect of layered soil medium on increasingdisturbance as discussed earlier. In addition, a significant differ-ence of about 70%–135% exists between the mobilized unit shaftadhesion values along the stiff clay to silty clay layer (layer 2) forpiles SD11, SD21, and SD31. This is attributed to the minimal ad-hesion developed over the shaft length immediately above theupper helix. In the current study, the shaft adhesion was ne-glected along a length equal to one helix diameter, D, above theupper helix, and the corrected unit shaft adhesion values werecalculated for piles SD11, SD21, and SD31 and provided in Table 5.The revised unit shaft adhesion values along the stiff clay to siltyclay layer (layer 2) are in agreement at different depths along theupper 4 m of pile shaft. It should be noted that no correction wasrequired for the unit shaft adhesion values for piles SS11, LS12, andLD12 along the upper 4 m, as the helix was deep enough belowstrain gauges of level SG4 (see Table 5). Similar results were re-ported by Zhang (1999). The researchers suggested that the shaftadhesion could not be mobilized along a length of one helix di-ameter, D, above the upper helix because of the shadowing effect.It was proposed to use an effective shaft length for calculating theshaft adhesion, given as Heff = H − D.

Fig. 10. Normalized load – settlement curves: (a) helical piles(phase I); (b) driven pile (phase I); (c) helical piles (phase II).

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The unit shaft adhesion values along the silt crust (layer 1) andin the stiff clay to silty clay (layer 2) for piles LS12 and LD12 tested9 months after installation are significantly higher than the cor-responding values for the 6.0 m helical piles tested 2 weeks afterinstallation. The unit shaft adhesion values increased in averageby about 16%–50% for the silt crust and between about 54% and130% for the stiff clay to silty clay layer.

The undrained and drained soil cohesion, Cu and C=, respec-tively, and the effective angle of internal friction, � ′, mobilized bythe different sections of the helical piles, were back-calculatedusing the design equations given in the section “Individual platebearing method” and the obtained load-transfer curves at ulti-mate loads, employing a trial-and-error technique.

The formulations and correlations provided by Bishop (1966),Kulhawy and Mayne (1990), Mitchell (1993), Stark and Eid (1994),and Terzaghi et al. (1996) showed that the angle of internal fric-tion, � ′, of the glacial clay till was found to vary between approx-imately 28° and 35°. Thus, it was deemed reasonable to use anaverage value of � ′ = 31° for the glacial clay till layer.

Cemented and heavily overconsolidated clays mobilize effec-tive cohesion, C=, under loading (Kulhawy and Mayne 1990). Mesriand Abdel-Ghaffar (1993) related the short-term value of C= to thepreconsolidation pressure, p

′ , and the current stress state, i.e.,C= = 0.024p

′ . This relationship gives an effective cohesion for theglacial clay till layer that varies between 18 and 26 kPa. Thus, a C=value of 20 kPa is considered reasonable in the analysis.

Fig. 11. Load distribution curves for piles tested in phase I: (a) SS11; (b) SD11; (c) SD21; (d) SD31.

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The average undrained cohesion mobilized along the pile shaft inthe different silty clay and clayey silt layers was calculated using theshaft diameter, d = 0.324 m, and an adhesion factor, �, that variedbetween 0.4 and 1.0 as per the recommended values by the CanadianFoundation Engineering Manual (CFEM 2006), such that

(8) Cu � Q shaft/��dHshaft

where Hshaft is the length of the shaft section corresponding to acertain soil layer, and Q shaft is the ultimate pile shaft capacity. Themobilized drained shear strength parameters of the silty crustand clay till were obtained as follows:

(9) Q shaft � �C ′�dHshaft ��2

dHshaft2 � ′Ks tan�

where Ks and � are the coefficient of lateral earth pressure andpile–soil friction angle, respectively, obtained from Kulhawy(1984) and the CFEM (2006).

The average undrained cohesion mobilized by the upper helixfor all 6.0 m double-helix piles was back-calculated consideringthe individual plate bearing method, and using the bearing capac-ity factor, Nc = 9.0, as follows:

(10) Q upper-helix � NcCuAh

where Q upper-helix is the ultimate bearing capacity of the upperhelix, and Ah is its projected area. The average drained shearstrength parameters mobilized by the upper helix for the 9.0 mdouble-helix pile, LD12, were back-calculated using Nc = 9 and Nq,considering the Meyerhof (1976) approach but using a factor ofsafety of 2.0, as follows:

(11) Q upper-helix � Ah�C ′Nc � vo′ Nq�

where vo′ is the effective overburden pressure.

Table 6 provides the back-calculated mobilized shear strengthparameters for the 6.0 m helical piles tested 2 weeks after instal-lation. The results show that the soil disturbance had a significanteffect on the mobilized undrained and drained shear strengthparameters along the shaft and upper helix sections. The back-calculated � ′ for the silt crust (layer 1) was about 75%–88% of theintact value obtained from the site investigation program (� ′ =31o). The mobilized undrained cohesion, Cu, of the stiff clay – siltyclay and clayey silt (layer 2) was about 20%–45% of the shearstrength of the intact layer. In contrast, the reduction in the un-drained cohesion of the very stiff silty clay – clayey silt – clay (layer3) was about 60%.

Table 7 provides the results for the 9.0 m helical piles, LS12 andLD12, tested 9 months after installation. For the silt crust (layer 1),the mobilized � ′ was about 17%–42% higher than the values cal-culated for piles tested in phase I. The mobilized Cu values in-creased by about 45%–250% (in layer 2) and by about 53%–67% (inlayer 3) relative to the values mobilized for piles tested in phase I.The mobilized drained shear strength parameters, C= and � ′, ofthe clay till layer were in average of about 70% and 74% of theinterpreted intact values, respectively. However, it can be con-cluded that the soil disturbance occurred in the soil during instal-lation persisted long after installation, which was illustrated incomparing the mobilized shear strength parameters 9 monthsafter pile installation with the intact ones due to the structurednature of the soils encountered in the test site.

It should be noted that the induced effective stresses in the sitesoils during the penetration of the pile shaft and helices exceededtheir yield stresses, which resulted in significant vertical and ra-dial strains that caused the interparticle bonds to break and ex-cess pore-water pressure to generate due to the collapse of the soilstructure. Therefore, the soil adjacent to the pile shaft and helicesexperienced a significant drop in its shear strength that was man-ifested in the lesser mobilized shaft unit friction resistances andstrength parameters after installation. As such, Roy and Lemieux(1986), through in situ tests, indicated that the soil memory of claysoil adjacent to driven piles would be erased due to soil remould-ing.

Fig. 12. Load distribution curves for piles tested in phase II: (a) LS12; (b) LD12.

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Fig. 13. Load transferred to soil for piles: (a) SS11; (b) SD11; (c) SD21; (d) SD31; (e) LS12; (f) LD12.

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The estimated overconsolidation ratios of the test site soils in-dicate moderately and highly overconsolidated soil layers. It isbelieved that such soils were subjected to aging and hardeningduring the geological process of the natural development of thesoils’ microstructure. However, they experienced a decrease intheir strength and stiffness parameters when those geologic ef-fects were removed due to the destruction of the soil microstruc-ture. Perko (2009) observed decreased pile capacities for helicalpiles installed and tested in highly overconsolidated soils. Fur-thermore, Tappenden (2007), after conducting several load testson helical piles installed in typical Alberta soils, concluded thatdisturbance to the native soils during the pile installation couldseverely reduce the expected capacity for piles installed in highlyoverconsolidated or structured soils. The reduction in the soilstrength parameters as inferred from the measured shaft resis-tances can also be ascribed to testing the phase I piles during thewet season and probable formation of very small gaps between

the pile shaft and the surrounding soil due to undetected minorimperfections during the welding of the pile shaft segments.

After pile installation, the effective stress of soil adjacent to thepile increases due to the dissipation of the excess pore pressureaccompanied by a reduction in the void ratio. The reconsolidationof soil causes the yield stress to increase in step with the increasein the effective stress, leading to an increase in the strength pa-rameters of soil layers. Further increase in the strength and stiff-ness of the reconsolidated soils could be achieved by the processesof aging and thixotropic hardening. The aging process employs adecrease in volume of soil and an increase in the vertical yieldstress under a constant effective vertical stress. An additional in-crease in the vertical yield stress and hence the soil strength mayoccur due to the thixotropic hardening phenomenon in which thethixotropic bonding between the soil particles takes place, whenthe rate of change in void ratio is slow enough. The thixotropiccohesive soil would have the tendency to regain the interparticleforce balance and reorganize the adsorbed water–cation structureto a lower energy state. It should be noted that Mitchell (1993)indicated that thixotropic hardening may account for low to me-dium sensitivity of the cohesive soils.

The mobilized undrained cohesion values underneath the up-per helix embedded in layer 2 were about 70%–74% of the undis-turbed values and about 50% for the case of the upper helixembedded in layer 3. These high percentages are probably due tothe propagation of the failure surface outside the disturbed zoneinto the undisturbed soil beyond the helix diameter, or propaga-tion of failure surface into the underlying very stiff silty clay –clayey silt (layer 3). For the upper helix section in pile LD12, themobilized C= and � ′ values of clay till were about 90% and 93.5% ofthe intact parameters, respectively. These high percentages areattributed to the propagation of the failure surface outside thedisturbed zone.

The ultimate end bearing capacities mobilized by the bottomhelix – pile tip were calculated using the intact shear strengthparameters and compared with the measured values, as it wasbelieved that pile installation process would introduce minimaldisturbance below the bottom helix – pile tip. Due to the variabil-ity in the thickness of the silty sand layer located near the 6.0 mpile tip, the shear strength parameters of the clay till layer and thesilty sand layer were used to calculate the end bearing resistancesof the bottom helix – pile tip for all 6.0 m piles (SS11, SD11, SD21,and SD31). The ultimate bearing capacity of the bottom helix – piletip was calculated as follows:

(12) Q Bottom � (Ah � At)�C ′Nc � vo′ Nq�

where At is the area of the pile tip, and the bearing capacity factorsNc and Nq were obtained from Meyerhof (1976) using a safety factorof 2.0 (assuming behaviour at helical pile tip is similar to boredpiles). The average calculated ultimate end bearing capacity forthe 6.0 m piles considering clay till was 872 kN (i.e., equivalent tothe end bearing pressure of about 3000 kPa). The measured valuesvaried between 752 and 1131 kN. Thus, the percentage differencebetween the predicted capacity and the measured capacity variedbetween 15.9% and −22.9%. Therefore, the ultimate capacity of thebottom helix – pile tip can be estimated using the intact shearstrength parameters. Furthermore, Fig. 14 illustrates that themeasured end bearing pressures below the bottom helix – pile tipof the 6.0 m piles are in close agreement with the cone tipresistance. In contrast, the average calculated ultimate bearingresistance for the 6.0 m piles using the drained shear strengthparameter of the silty sand layer (� ′ = 42°) was 2565 kN (i.e.,equivalent to the end bearing pressure of 8780 kPa), consideringthe Meyerhof (1976) limiting point resistance of q l = 0.5Nq tan(� ′),where q l is in t/ft2 (1 ft = 0.3048 m). This significantly overesti-mated the capacity by 127%–241%. Moreover, Fig. 14 shows that the

Table 4. Percentage of load transferred by different sections of helicalpiles.

Pile

Load Pile section SS11 SD11 SD21 SD31 LS12 LD12

0.5Q ult Shaft 20.3 7.2 6.4 4.9 23.6 17.00.5Q ult Upper helix — 11.1 4.7 10.9 — 19.40.5Q ult Interhelix — 2.6 5.0 5.2 — 1.60.5Q ult End bearing* 79.7 79.1 83.9 79.0 76.4 62.0Q ult Shaft 11.0 4.1 3.3 2.0 12.7 9.4Q ult Upper helix — 9.7 10.0 5.9 — 22.5Q ult Interhelix — 1.3 1.6 3.1 — 1.2Q ult End bearing* 89.0 84.9 85.1 89.0 87.3 66.9

*End bearing resistance from bottom helix and pile tip.

Fig. 14. Bearing pressures at failure generated by helices and piletip compared with cone tip resistances (qt).

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measured end bearing pressures were much less than the cone tipresistance in the silty sand layer. Hence, it can be concluded thatthe end bearing resistances mobilized below the bottom helix –pile tip of the 6.0 m piles were developed mainly from the clay till

layer, as the thickness of the silty sand layer below the piles wassmall.

The measured values of the ultimate bearing capacity of the9.0 m piles LS12 and LD12 were 1704 and 1658 kN, respectively.

Table 5. Unit shaft adhesion for helical piles along upper 4.0 m of pile shaft.

PileStrain gauge depth (m)(stain gauge level) Soil layer

Uncorrected unitshaft adhesion (kPa/m)

Corrected unitshaft adhesion (kPa/m)

SS11 0–1.35 (SG1–SG2) Layer 1, silty to sandy silt 6.2 6.21.35–2.85 (SG2–SG3) Layer 2, stiff clay to silty clay 24.0 24.02.85–4.35 (SG3–SG4) Layer 2, stiff clay to silty clay 24.2 24.2

SD11 0–1.60 (SG1–SG2) Layer 1, silty to sandy silt 7.7 7.71.60–3.10 (SG2–SG3) Layer 2, stiff clay to silty clay 18.0 18.03.10–4.50 (SG3–SG4) Layer 2, stiff clay to silty clay 10.6 17.4

SD21 0–1.30 (SG1–SG2) Layer 1, silty to sandy silt 6.6 6.61.30–2.45 (SG2–SG3) Layer 2, stiff clay to silty clay 17.1 17.12.45–3.0 (SG3–SG4) Layer 2, stiff clay to silty clay 7.3 15.2

SD31 0–1.60 (SG1–SG2) Layer 1, silty to sandy silt 5.9 5.91.60–2.73 (SG2–SG3) Layer 2, stiff clay to silty clay 9.5 17.3

LS12 0–1.5 (SG1–SG2) Layer 1, silty to sandy silt 8.9 8.91.5–3.0 (SG2–SG3) Layer 2, stiff clay to silty clay 35.5 35.53.0–4.35 (SG3–SG4) Layer 2, stiff clay to silty clay 36.1 36.1

LD12 0–1.5 (SG1–SG2) Layer 1, silty to sandy silt 9.4 9.41.5–3.0 (SG2–SG3) Layer 2, stiff clay to silty clay 36.7 36.73.0–4.61 (SG3–SG4) Layer 2, stiff clay to silty clay 37.2 37.2

Note: Shaded values indicate revised unit shaft adhesion values.

Table 6. Back-calculated shear strength parameters mobilized by piles in phase I.

Mobilized shear strength Ratio of mobilized to intact

Shaft Upper helix Shaft Upper helix

Layer Cu (kPa) � ′ (o) Cu (kPa) Cu (%) � ′ (%) Cu (%)

Pile SS11Layer 1, silt crust — 25.0 — — 80.6 —Layer 2, stiff clay to silty clay and clayey silt 38.5 — — 45.3 — —Layer 3, very stiff silty clay, clayey silt, and clay 54.0 — — 39.4 — —

Pile SD11Layer 1, silt crust — 27.0 — — 87.1 —Layer 2, stiff clay to silty clay and clayey silt 17.6 — — 20.7 — —Layer 3, very stiff silty clay, clayey silt, and clay — — 68.1 — — 50.0

Pile SD21Layer 1, silt crust — 27.5 — — 88.7 —Layer 2, stiff clay to silty clay and clayey silt 16.5 — 59.5 19.4 — 70.0

Pile SD31Layer 1, silt crust — 23.5 — — 75.8 —Layer 2, stiff clay to silty clay and clayey silt 17.3 — 33.0 20.4 — 38.8

Table 7. Back-calculated shear strength parameters mobilized by piles in phase II.

Mobilized shear strength Ratio of mobilized to intact

Shaft Upper helix Shaft Upper helix

Layer Cu (kPa) C= (kPa) � ′ (o) C= (kPa) � ′ (o) Cu (%) C= (%) � ′ (%) C= (%) � ′ (%)

Pile LS12Layer 1, silt crust — — 32.0 — — — — 103.0 — —Layer 2, stiff clay to silty clay and clayey silt 55.0 — — — — 65.0 — — — —Layer 3, very stiff silty clay, clayey silt, and clay 82.5 — — — — 60.2 — — — —Layer 5, clay till — 14.0 23.0 — — — 70.0 74.0 — —

Pile LD12Layer 1, silt crust — — 33.5 — — — — 108.1 — —Layer 2, stiff clay to silty clay and clayey silt 58.0 — — — — 68.2 — — — —Layer 3, very stiff silty clay, clayey silt, and clay 87.0 — — — — 63.5 — — — —Layer 5, clay till — 14.0 23.0 18.0 29.0 — 70.0 74.0 90.0 93.5

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However, the average calculated ultimate bearing capacity usingthe intact clay till strength parameters was 1090 kN, representinga difference of –36.0% and −34.2% for piles LS12 and LD12, respec-tively. In contrast, the Meyerhof theory overestimates the bearingcapacity of the bottom helix – pile tip (2565 kN) by about 50%–55%.The measured ultimate end bearing resistance below the bottomhelix – pile tip of the 9.0 m helical piles lies between the ultimatecapacity calculated using the clay till drained strength parametersand that calculated using the silty sand drained strength param-eters. This can be ascribed to the interbedding of silty sand andclay till layers.

Installation torque to helical pile capacityThe relationship between helical pile ultimate capacity and

installation torque has been used as a general rule of thumbin industry since the 1960s (Perko 2009). The empirical torque–capacity relationship was first developed by Hoyt and Clemence(1989) for small diameter helical piles or anchors resisting upliftloads. The relationship is also reported in the CFEM (2006). Theempirical relationship can be presented as follows:

(13) Q ult � KTT

where KT is the empirical capacity-to-torque factor and has unitsof m−1 (ft−1), and T is the average installation torque. A number oftheoretical capacity-to-torque correlations were developed by sev-eral researchers, including Perko (2001), Perko (2009), and Tsuhaand Aoki (2010).

Based on the results of the full-scale compression load testsconducted on large-diameter helical piles in the current studyand the measured torque values, the back-calculated capacity-to-torque factors for the tested piles are presented in Table 1. Thecapacity-to-torque factors, KT, for piles tested in phase I were10.2 m−1 (3.1 ft−1) for the single-helix pile SS11 and between 10.0 m−1

(3.0 ft−1) and 11.6 m−1 (3.5 ft−1) for the double-helix piles SD11, SD21,and SD31. Moreover, the KT values for the single-helix and double-helix piles, LS12 and LD12, tested in phase II were 20.5 m−1 (6.2 ft−1)and 26.5 (8.0 ft−1), respectively, which are higher than those ob-tained in phase I of testing by about 101% for the single pile and byabout 128%–165% for the double-helix piles. It should be noted thatthe obtained torque values for piles LS12 and LD12 are consideredlow for the estimated consistency of the till deposits around thepile tip, which resulted in the high KT numbers calculated forthose piles. The low torque values may be attributed to the layer-ing nature of the site soil, which affects the bearing strata underthe helices at different depths or a possible source of error inmeasuring the torque. Thus, the torque and KT values for pilesLS12 and LD12 should be cautiously considered.

The CFEM (2006) recommended the use of a KT value of10.0 m−1 (3.0 ft−1) for helical piles with shaft diameter approaching200 mm. Tappenden (2007) reported a KT value of 9.19 m−1 for pilesof shaft diameter between 140 and 400 mm installed in westernCanadian soils and loaded in compression or tension. For piles ofshaft diameter 267 mm, installed in oil sands, and loaded in com-pression, Sakr (2009) reported a KT of 23.9 m−1. It can be seen thatthe calculated KT values for the piles tested in phase I are in goodagreement with that recommended by the CFEM (2006) andslightly higher than that reported by Tappenden (2007).

It should be mentioned that the installation torque is depen-dent on several factors such as soil strength, existence of large-size coarse particles (i.e., such as cobbles and boulders), pile shaftdiameter and shape (i.e., circular or square shaft), helix diameter,pitch size, number of helices, applied compression load on pileduring installation, etc. In addition, the recommended KT valuesreported in the literature were mainly for piles under tensionloading or did not differentiate between piles under compressionor tension loads. Thus, such differences between KT obtained for

compression loading and tension loading of piles can also be at-tributed to the fact that the bottom helix rests on relatively un-disturbed soil in compression loading, but in tension loading theupper helix bears on the soil affected by the installation of thehelices (i.e., disturbed soil). Hence, it is recommended by the au-thors to use the torque–capacity empirical relationship as a qual-ity assurance measure during pile installation.

Summary and conclusionsIn total, seven axial compression tests were conducted, and the

measured pile axial response and strain values were used toevaluate the piles’ ultimate capacities and the load-transfer mech-anisms. The mobilized shear strength parameters were back-calculated for helical piles tested 2 weeks and 9 months after pileinstallation. Based on the experimental results and their analysis,the following conclusions can be drawn:

1. An ultimate capacity criterion was proposed to estimate theultimate load of large-capacity helical piles. The ultimate ca-pacity is defined as the load corresponding to the pile headsettlement that equals the pile elastic deformation plus 3.5%of the helix diameter. The proposed failure criterion is appro-priate for large-diameter helical piles, as it identifies thefailure load within the nonlinear region of the pile load–settlement curve.

2. The Davisson (1972) criterion and tangent intersectionmethod significantly underestimated the ultimate capacitiesof the test large-capacity helical piles, while the criterion de-veloped by Livneh and El Naggar (2008) for slender helicalpiles significantly overestimated their ultimate capacities.

3. The ultimate axial capacity of the 6.0 m double-helix pileswere about 24%–57% higher than that of the single-helix pileof similar pile length and diameter.

4. The ultimate axial compression capacity of the 9.0 m double-helix pile was about 27% higher than that of the single-helixpile of similar pile length and diameter.

5. The ultimate capacity of the 6.0 m single-helix pile washigher by about 18% than the capacity of the 6.0 m driven pile,while the ultimate capacities of the 6.0 m double-helix pileswere higher by about 45%–85% than the capacity of the 6.0 mdriven pile, considering the site subsurface conditions. Theseresults confirm the advantage of helical piles compared to theconventional driven piles of similar pile embedded lengthand shaft diameter.

6. The measured settlements at ultimate compression capacityof helical piles were 23.8–25.6 mm for the 6.0 m helical pilestested 2 weeks after pile installation, and 30–33 mm for the9.0 m helical piles tested 9 months after pile installation. Thisdifference in settlement is due to the nature of soil under-neath the bottom helices. In the case of the 6.0 m helicalpiles, the helices were bearing on layers 2 and 3 (stiff siltyclay, clayey silt, and clay), while for the 9.0 m piles, the heliceswere bearing on layer 5 (clay till). The till withstood largersettlement before failure occurred. The settlement at failurein the case of the 6.0 m driven pile, tested 2 weeks after pileinstallation, was 13.6 mm. This is because the driven pile hasderived most of its capacity through shaft resistance, whilehelical piles have derived the majority of their capacitiesthrough end bearing (see Table 4). It is well known that endbearing resistance requires large settlement to mobilize.

7. The measured settlements at the design capacities (i.e., byemploying a factor of safety of 2.0) for all piles tested in thisstudy were <1% of the helix diameter for the helical pilesand <1% of the pile shaft diameter for the driven pile.

8. The tested double-helix piles showed mainly individual platebearing failure mechanism, even for piles with interhelixspacing of 1.5 times the helix diameter, contrary to the com-mon cylindrical shear failure mechanism expected for such

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interhelix spacing ratio. This is attributed to the high degreeof soil disturbance and softening of the structured site soils.

9. The end bearing capacity of the bottom helix – pile tip couldbe reasonably predicted using the cone tip resistance profilesmeasured during the cone penetration testing in the intactsoil and also from using the Meyerhof bearing capacity the-ory for deep foundations.

10. The mobilized shaft shear resistance of the helical piles de-creased due to soil disturbance associated with pile installa-tion. The mobilized strength parameters of stiff to very stiffsilty clay and clayey silt soils were about 20%–40% of theintact values. However, the mobilized shaft resistance in-creased significantly with time (during the 9 month period)to about 65% of the intact values for stiff to very stiff clayeysilt and silty clay soil.

11. The mobilized soil strength parameters underneath the up-per helix were higher than those mobilized along the pileshaft for the same soil layers. This was attributed to the ex-tension of the failure surface underneath the upper helixbeyond the edge of the helices into intact soil that did notexperience any disturbance during pile installation.

It should be noted that the experimental results are site specificand must be used with caution when applied elsewhere.

AcknowledgementsThe research described herein received direct support from the

Natural Sciences and Engineering Research Council of Canada(NSERC), The University of Western Ontario, and Almita PilingInc., Ponoka, Alberta. The authors would like to thank Erden Ertor(The University of Western Ontario) and David Sparrow and PaulVan Der Raadt (AMEC E&I, Inc., Calgary, Alberta) for their support.

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